Since 1955 „ > Strojniški vestnik Journal of Mechanical Engineering no. 7-8 year 2020 volume 66 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 Cover: Top left to right: SEM images of polyamide 6 membranes with average fiber diameter of 325 nm and 122 nm. Center: Demonstration of the polymer fibers prepared by Nanospider methodology Bottom left to right: SEM images of polyamide 6 membranes with "spider nets" and beads. Image courtesy: Education and Research Institute of Nanostructures and Biosystems, Saratov State University, Russian Federation 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)7-8 Contents Contents Strojniški vestnik - Journal of Mechanical Engineering volume 66, (2020), number 7 8 Ljubljana, July August 2020 ISSN 0039-2480 Published monthly Papers Alexandra Aulova, Marko Bek, Leonid Kossovich, Igor Einri: Needleless Electrospinning of PA6 Fibers: The Effect of Solution Concentration and Electrospinning Voltage on Fiber Diameter i Dalibor Petrovič, Maijan Dodič, Nenad Kapor: A New Design Solution for Aircraft Wheels that Reduces Overpressure in the Tire while Retaining its Absorption Power and its Dimensions 431 Renata Mola, Tomasz Bucki: Characterization of the Bonding Zone in AZ9 AlSil Bimetals Fabricated by Liquid-Solid Compound Casting Using Unmodified and Thermally Modified A1SÍ12 Alloy 439 Himmat Singh, M. S. Niranjan, Reeta Wattal: A Study for the Nanofinisliing of an EN-31 Workpiece with Pulse DC Power Supply Using Ball-End Magnetorheological Finishing 449 Zifeng Zhang, Hongxun Fu, Xuemeng Liang, Xiaoxia Chen, Di Tan: Comparative Analysis of Static and Dynamic Performance of Nonpneumatic Tire with Flexible Spoke Structure 458 Anastasios Tzotzis, César García-Hernández, José-Luis Huertas-Talón, Panagiotis Kyratsis: 3D FE Modelling of Machining Forces during AISIH ard Turning 47 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8,421-430 © 2020 Journal of Mechanical Engineering. All rights reserved. D0l:10.5545/sv-jme.2020.6713 Original Scientific Paper Received for review: 2020-04-09 Received revised form: 2020-05-27 Accepted for publication: 2020-06-01 Needleless Electrospinning of PA6 Fibers: The Effect of Solution Concentration and Electrospinning Voltage on Fiber Diameter Alexandra AulovaJ* -M arko Bck1 -L eonid Kossovich2 4 gorEmri13 1 University of Ljubljana, Faculty of Mechanical Engineering, Slovenia 2 Saratov State University, Russian Federation 3 Footwear Research Center, Tomas Bata University, Czech Republic Needleless electrospinning is the process of forming thin material fibers from the open surface of its solution or melt in a strong electrostatic field. Electrospun non-woven materials are used in various applications that require specific fiber diameters and pore size distributions. Fiber diameter depends on the properties of the polymer solution and manufacturing conditions. A needleless electrospinning process using the Nanospider setup was investigated using the commonly used polyamide 6 (PA6) solution in a mixture of acetic and formic acids. Polymer solutions with different polymer concentrations were characterized by viscosity surface tension and electrical conductivity. An increase in polymer content in the solution resulted in the exponential increase of the solution viscosity, polynomial increase of electrical conductivity and had almost no effect on surface tension. The effect of the polymer concentration in the solution, as well as electrospinning voltage on fiber diameter and diameter distribution, was investigated using scanning electron microscopy images. The average fiber diameter linearly increases with the increased polymer concentration and also demonstrates an increase with increased electrospinning voltage, although less pronounced. Therefore, a change in the PAß solution concentration should be used for the robust adjustment of fiber diameter, while changes in electrospinning voltage are more appropriate for fine tuning the fiber diameter during the process of needleless electrospinning. Keywords: needleless electrospinning, nanospider, PA6, fiber diameter, solution concentration, voltage Highlights • Needleless electrospinning is used for manufacturing non-woven materials. • Properties of non-woven materials, e.g. fiber diameter and pore size distribution, depend on the material's properties and the electrospinning conditions. • By increasing the polymer concentration, the viscosity of the solution increases exponentially, the electrical conductivity increases polynomially and the surface tension does not change. • The average fiber diameter linearly increases with the increase in polymer solution concentration, as well as electrospinning voltage; however, it is less pronounced compared to solution concentration. • For the robust adjustment of fiber diameter, it is practical to change the PA6 solution concentration, while for fine tuning, voltage changes can be applied. 0 INTRODUCTION Electrospinning is the process used to produce polymer fibers and membranes from a polymer solution or melt under high electrostatic forces. Electrospun materials are used in various applications, including textiles [1], insulation materials, tissue scaffolds [2], patches, filters [3] and separation membranes [4], The functionality of membranes for different applications is determined by the material, physicochemical surface properties of the fibers, fiber diameter and pore size distribution. Within this study, we investigated how the fiber diameter is affected by solution concentration and electrospinning voltage. Together with membrane thickness, the fiber diameter determines the pore size distribution of the membrane. Fiber diameter determines the free surface of the membrane and pore size distribution determines the barrier and permeability properties of the material. Fiber diameter and its distribution are strongly dependent on the polymer solution/melt properties and electrospinning conditions [5] and [6], Controlling the fiber diameter during electrospinning allows for the formation of membranes with selected pore size distribution and is a crucial part of the electrospinning process. The formation of electrospun fibers happens from the polymeric melt or solution between the two electrostatically charged electrodes. This process starts with the localization of the free charges in the polymer solution close to the surface in the electrostatic field [7], As soon as the concentration of the charges is high enough the surface of the solution forms the so-called Taylor cone [8] and [9]. The charge carriers continue to concentrate on the top of the cone, and this finally results in a jet of polymer solution being pulled by the electrostatic field towards the receiving electrode [7], While the jet is travelling towards the receiving electrode, the solvent in the solution evaporates and *Corr. Author's Address: University of Ljubljana, Faculty of Mechanical Engineering, Aškerčeva 6,1000 Ljubljana, Slovenia, alexandra.aulova@fs.unMj.si 421 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 421-430 the dry fiber is deposited. The process of solvent removal is governed by the properties of the solvent, the distance between electrodes and the enviromnental conditions during the electro spinning (temperature, airflow, humidity) [5] and [7], There are three main properties of the polymer solution that are important for the electrospinning process: viscosity, surface tension and electrical conductivity. Their combination determines the behavior of the solution during electrospinning and directly affects the fiber diameter. The viscosity of the polymer solution plays a crucial role in the electrospinning process as it affects the formation of the Taylor cone as well as the stabilization of the solution jet while travelling to the receiving electrode. During the first stage of fiber formation, high viscosity results in energy losses, but it also provides a more efficient process since more material is deposited. Additionally, a high enough viscosity of the polymer solution prevents the jet from breaking up into droplets and therefore stabilizes the jet [5] and [10], Solutions with too high viscosity will not provide any fibers and solutions with too low viscosity will result in multiple droplets instead of a fiber (electrospraying). Usually, viscosity is the parameter that significantly affects the electrospinning process and, consequently, the thickness of the electrospun fibers. Typically higher viscosity of the solution corresponds to the larger fiber diameters [5], [11] and [12], The viscosity of the solution depends on the polymer concentration, its molecular weight and weight distribution, as well as molecular topology. It also depends on the loading conditions (frequency, shear rate) and enviromnental conditions (temperature, humidity), which are related to the electrospinning parameters. Surface tension forces tend to reduce the surface of the liquid per mass unit. These forces are a result of interactions between solvent and polymer molecules (both topology and electromagnetic potentials). Surface tension forces (coupled with viscosity) are the main counterforces to the electrostatic forces that are pulling the fibers. Surface tension also determines the behavior of the fibers while forming and moving to the collecting electrode [5] and [12], Electrical conductivity measures the ability of the material to carry an electrical charge and represents an intrinsic property of the solution. Conductivity mostly depends on the solvent's electrical conductivity [5], [12] and [13], It is also affected by the polymer's molecular weight, polymer's concentration and temperature. While surface tension and electrical conductivity can be changed using additives (surfactants or ionic salt), they also affect other properties and change the balance between surface tension, electrical conductivity and viscosity [11] and [14], The empirical ranges of viscosity, surface tension and electrical conductivity required for successful electrospinning are defined, for example, in [5]; however, the range of values is relatively broad and can be used only in the needle electrospinning method. The needle or capillary electrospinning method is a process where one of the electrodes uses a needle through which the polymer solution/melt is pushed into the electrostatic field under pressure. Needleless electrospinning, as its name suggests, does not use a needle or nozzle to introduce the polymer solution into the electrostatic field but instead uses a free surface or droplets of the solution on the electrode in the form of a rotating cylinder or string. Such a system has been introduced to the market under the commercial name Nanospider. The formation of fibers in this system occurs from the bottom rotating electrode to the top receiving electrode [15] and [16], In this case, the process is more complex compared to conventional needle electrospinning, as the formation of Taylor cones happens not only on the surface of one polymer droplet but in multiple regions on the electrode. One critical advantage of this technology compared to nozzle technology is increased production capacity. Multiple fibers form at the same time without clogging the capillaries. Needleless technology also provides a more homogeneous electrostatic field and the possibility to produce largeer dimensions of electrospun material [15], [17] and [18], The physics of needleless electrospinning process is different compared to the capillary method [5], [7] and [9]; as the Taylor cones form on a free surface subjected to electrostatic field, the shape of this surface is important. During capillary electrospinning, a Taylor cone forms on the droplet from the capillary, whos size is controlled by the pressure in the capillary. In the case of Nanospider technology, multiple Taylor cones appear on the surface of the thin film covering the cylinder electrode or droplets on the string electrode. The thickness of the polymer solution film and the size and density of droplets depend on the speed of electrode rotation and the solution's properties. In both cases, however, the solution's properties play the most crucial role in the formation of the electrospinning surface, as well as fiber dimensions and quality. Due to these differences 422 Aul ova, A. - Bek, M. - Kossovich, L. - Emri, I. Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 421-430 in physics of the electro spinning process, a well-known theory for capillary electrospinning cannot be directly applied to needleless fiber production technology. Differences in the shape of the solution's surface (a drop at the capillary or a drop on the string) and the interplay of multiple polymer solution jets in the electrostatic field during the material formation requires different settings and polymer concentrations for successful needleless electrospinning. Within this work, the needleless electrospinning process of polyamide 6 (PA6) dissolved in a mixture of acetic and formic acids is investigated. PA6 is a material widely used in a variety of applications including the automotive and textile industries, medicine and engineering. Its high mechanical strength in a wet and dry state [19], hydrophilicity and chemical resistance [20] make it an appropriate material for desalination, textiles [21], various filtration [22] and medical applications [23] water purification and gas separation. Additionally, PA6 can be modified with antibacterial/antifungal agents [24] and is often used in electrospun form for filtration, medical applications and textiles. PA6 has been well studied using nozzle electrospinning [11], [14], [22] and [25] whereas, to the best of our knowledge, only limited research has been done using the Nanospider needleless methodology [21] and [24], Researchers focused on the antibacterial additives to PA6 and morphological properties of the fiber's surface depending on the distance between electrodes [21] and [24], There is a lack of information on the processing conditions of needleless electrospinning, which is essential for fiber properties and is the focus of this paper. Within the paper, we conducted systematic research of the effect of polymer solution concentration on the critical properties of the polymer solution: its viscosity, surface tension and electrical conductivity. We investigated the effect of these parameters and the electrospinning voltage on the formation of PA6 fibers and their diameter. An electrospun material composed of fibers with the chosen diameter is the first step towards the production of materials with a controlled pore size distribution, as are required in different applications. 1 EXPERIMENTAL METHODS 1.1 Materials Light-stabilized PA6 under the trade name Ultramid B24 N03 was provided through an in-kind donation by BASF (Germany). The material lias a number average molecular weight (Mn) of 21,100 g/mol and a weight average molecular weight (Mw) of 55,600 g/mol. The material producer describes it as a material grade for the production of textile fibers that is especially suitable for high-rate spinning. PA6 is dissolvable in the formic acid; therefore, for this study, a mixture of 2:1 by weight of acetic (AA, 100 % glacial, EMSURE, Sigma-Aldrich, USA) and formic (FA, 99 %, Carlo Erba, France) acids were used to prepare the PA6 solutions. Both acids are weak carboxylic acids, which only partially dissociate in the solution. However, formic acid is a polar solvent with low polarity, which, in combination with PA6, results in the polyelectrolytic behavior of the solution [13], This increases its electrical conductivity to values high enough for electrospinning [26], In the mixture of acids, the formic acid acts as the aggressive PA6 solvent, while the presence of acetic acid increases the boiling point of the solvent (118 °C for acetic acid and 108 °C for formic). This means that the jet of the solution during electrospinning will dry closer to the receiving electrode at the stage when the jet trajectory has already rotated in the normal direction to the receiving electrode [5], This results in a higher product yield for Nanospider electrospinning technology. 1.2 Preparation of Solutions The solutions for electrospinning and characterization were prepared following the same procedure. The components of the solution were mixed in the following sequence: polymer granulate, formic acid and acetic acid. Afterwards, the solution was mixed for approximately 8 hours at 200 rpm using a magnetic stirrer. For mixing solutions with a larger content of the polymer, the magnetic stirrer was heated up to 50 °C. Measurements were started 48 hours after the solutions were prepared. The solutions were subjected to electrospinning in a laboratory setup three times. For each of the tests, a new batch of the solution was prepared following the same protocol. This was done in order to avoid the ageing of the solution and introducing the error of solution preparation into the whole process. Due to the smaller amount of the solution prepared for characterization, the PA6 granulate 423 Needleless Electrospinning of PA6 Fibers: The Effect of Solution Concentration and Electrospinning Voltage on Fiber Diameter 413 Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 421-430 was dried in order to avoid drastic errors in solution composition due to the polyamide water uptake. Drying was performed using a humidity analyzer (KE-DBS60-3, KERN, Germany) for 15 minutes at 80 °C using % rams of polymer granulate at once. Five solution concentrations were prepared and their compositions are presented in Table 1 Table 1. Solution compositions Concentration [%] 11 12 13 14 15 PA6 mass [g] 66 72 78 84 90 FA volume [ml] 178 176 174 172 170 AA volume [ml] 356 352 348 344 340 1.3 Characterization of Solutions 1.3.1 Viscosity Shear viscosity was measured using a rotational rheometer (MCR 302, Anton Paar, Austria) using a concentric cylinder in a cup. A cylinder with a 27-îmn diameter and 1-mm gap between the cylinder and cup was used for measuring. To prevent solvent evaporation we used a solvent trap. Measurements were performed at ambient conditions by linearly increasing the shear rate from 1 1/s to 150 1/s in 180 s. Three repetitions per solution were made. 1.3.2 Electrical Conductivity The electrical conductivity coefficient of the prepared solutions was measured using a Metrohm Conductometer 912 (Switzerland) equipped with a PtlOOO electrode (measuring ranges from 0.015 mS/cm up to 250 mS/cm), which determines the conductivity of liquid samples between the two platinum electrodes. Measurements were taken at ambient conditions six times for each of the solutions, after ft stabilization time. 1.3.3 Surface Tension The surface tension coefficient was determined using a force tensiometer (K20, Kruss, Germany) equipped with the Wilhelmy plate, at ambient conditions. Measurements were performed 6 times for each of the solutions. The surface tension was determined as average over ® s of measurement. 1.4 Preparation of Electrospun Membranes The electrospinning process was performed using a Nanospider LAB (Elmarco, Czech Republic) laboratory setup and a string electrode. As a substrate, a polypropylene spunbonded material was used. The distance between electrodes was fixed to 140 mm and the speed of the substrate was set to 0.08 m/min. All five solutions with different PA6 concentrations were electrospun at three voltage levels: 70 kV, 75 kV and 80 kV, respectively. The choice of this voltage range was made based on past experience using the Nanospider electrospinning setup as well as available research works [21], [24] and [25], In total, 45 membrane samples were produced: 5 different electrospun solutions at 3 different voltages and 3 repetitions at each voltage level. 1.5 Membrane Characterization All samples were investigated using a scanning electron microscope (SEM) with an autoemission cathode (MIRA II LMU, Tescan). For the SEM analysis, three samples were taken from each of the concentrations and voltage levels. Three different randomly selected places were selected at each sample, resulting in 9 testing areas per material. They were covered with 5 mn gold coating and analyzed using a SEM under 643 kx magnification. For each location, 10 fibers were selected on the upper surface of the electrospun mat, and their diameters were measured, altogether resulting in 90 diameter measurements per each of the 15 samples types obtained via electrospinning. The mean value of the fiber distribution was calculated, and the width of distribution was assessed using the uniformity coefficient [16] and [27], which is calculated similarly to the polydispersity index as the ratio of weight average, Aw and number average, A„, analogues: r - FU A,' where A, = Yunid< I>, A„ = (i) (2) where df is fiber diameter, and », denotes the number of fibers with this diameter. A uniformity coefficient equal to 1 represents a perfectly monodisperse distribution of fiber diameter, while a higher number means broader fiber diameter distribution. 424 Aul ova, A. - Bek, M. - Kossovich, L. - Emri, I. Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 421-430 2 RESULTS AND DISCUSSION 2.1 Polymer Solution Properties Polymer solutions with different PA6 content levels were characterized in order to determine how polymer concentration affects viscosity, surface tension and electrical conductivity, see Fig. 1 A combination of these three properties determines if the solution will form fibers in an electrostatic field and the quality of fibers. Error bars were calculated as maximal errors for viscosity measurements and standard deviation for electrical conductivity and surface tension. 2.1.1 Viscosity In the measured range of shear rates, up to 150 1/s all materials exhibited Newtonian behavior. Within the measured shear rates, we did not observe shear thinning behavior, which is in line with findings in [28], For the analysis, the viscosity was calculated as the average viscosity value in the measured shear rate range. The results of average viscosity measurements as a function of PA6 concentration are shown in Fig. la. The viscosity increases with PA6 concentration. Comparing the viscosity of PA6 concentrations of 11 % and 15 %, there is an increase of 300 %. The solutions are classified as concentrated [29] and demonstrate an exponential increase in viscosity with concentration (also demonstrated in [11], [14] and [25]). This increase can be evidence of an entangled regime of polymer chains [30], Chain entanglements are the overlaps of the molecular chains in the solution, and a sufficient number of them is needed for the solution to form fibers without defects [31], 2.1.2 Surface Tension the solution with the smallest concentration is about 1.3 %. The error of surface tension measurement increases with polymer concentration, starting at 4 %. This might be attributed to the fast changes of solution properties due to evaporation, which is more pronounced in higher concentrations. t/3 o u % bO R CO 1-1 I- 40 mm) for the same contact force of tire and runway, the deflection of the set of tires is much smaller than the deflection of a standard tire. If we consider the intensity of the force when the deflection of a standard tire is equal to h = 0 mm, it can be observed that the deflection of the set of tires is around 15 % smaller than in the case of a standard tire. In addition, it can be noticed that the force required to reach the deflection h = Q mm is 61 % greater in the case of a set of tires compared to a standard tire. The diagram in Fig. 10 shows the dependence of the tire track width from the deflection on part A-B. It can be observed that for the same tire deflection (up to h < 40 mm), the tire track width is larger for a standard tire than for the set of tires. This is the result of lateral expansion caused by deflection and overpressure that is two times higher in a standard tire compared to the overpressure in the set of tires. 20 40 60 80 Tire track width [mm] Diagram of change of tire track width on section A-B with respect to tire deflection From the moment of contact of the outer and inner tire (h > 40 mm), as the deflection increases, the track width of the set of tires is larger than in the track width of a standard tire. Thus, at the final deflection h = 70 mm, the track width of a set of tires is 8 % larger compared to the case of a standard tire. This is because the inner tire does not allow the outer tire to be drawn inward, and consequently, the outer tire is extending laterally (Fig. 7). 40 60 80 100 120 140 Tire track width [mm] Fig. 11. Diagram of change of tire track width on C-D section in respect to the landing force 436 Petrovič, D. - Docf/c, M. - Kapor, N. Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 431-438 The diagram in Fig. 11 presents the changes of the tire track width on section C-D in respect to the landing force. It can be observed that for the deflection of h < 40 111111, the tire track width is almost the same for standard tire and the set of tires. At the deflection of h = 0 mm, it can be noticed that for the same tires track width in section C-D, the absorption of the set of tires is around 67 % larger compared to a standard tire. The diagram of Fig. 12 shows the results of the stress analysis in the carcass for deflection of h = 0 mm on section A-B below the tread for both standard tire and the set of tires. It can be observed that the stress in the carcass is around 37 % lower in the set of tires compared to the standard tire. In addition, the stress in the carcass sections A-E and B-F below the tread (which is in contact with the runway) are the same for both standard tire and the set of tires. -t0 -20 0 20 40 Tire track width [mm] Fig. 12. Diagram of the distribution of von Mises stresses in the carcass below the tread for h = 70 mm, in section A-B -40 0 40 Tire track width [mm] Fig. 13. Diagram of the distribution of von Mises stresses in the carcass below the tread for h = 70 mm, in the section C-D The diagram of Fig. 13 gives the results of the stress analysis in the carcass section C-D below the tread for both standard tire and the set of tires, for the deflection of/7 = 70 111111. It can be seen that the stress in the carcass is 42 % lower in the case of the set of tires than in the standard tire. Furthermore, it can be observed that the stresses in the carcass sections C-G and D-H below the tread (which is in contact with the base) are the same for both standard tire and the set of tires. In contrast, the load-capacity of the set of tires is approximately 56 % higher compared to the load-capacity of a standard tire (for deflection h = 70 111111). 6 CONCLUSION This study presents a comparative analysis of stressstrain behaviour of a standard tire and the new constructive solution of the set of tires. The analysis was conducted for different aircraft landing forces by using finite elements. It was shown that the problem of analysing aircraft tires for different aircraft inclinations requires highly complex modelling procedure for both model and material. The analysis conducted in this study aimed to investigate the possibility of reducing the pressure in the tires while preserving its dimensions and the kinetic energy absorption. The change in tire geometry during landing at the time of direct contact of the tires with the runway at maximum tire deflection up to h = 70 111111 was analysed by comparing the results for a standard (outer) tires and a set of tires consisting of an outer (standard) and a smaller (inner) tire. The set of tires was exposed to the overpressure that was two times lower compared to the overpressure in the standard tire. Based on the results of comparative analyses, it can be concluded that the new technical solution leads to the increase of the tire tread width by 8 %, and therefore, it ensures a larger contact area between the tire and the runway. It results in the fact that the tire stress is approximately 40 % lower for the tire set compared to a standard tire, which further leads to higher durability of the tire in a new solution, as well as the ability to absorb around 56 % more kinetic energy than in the case of a standard tire. 7 NOMENCLATURE W jyiso W vol strain energy potential function, [MPa] isochoric deformation energy function, [MPa] volume deformation energy function, [MPa] Àh /-2- h principal stretch ratios, Ih /2 principal deviatory strain invariants, d material incompressibility parameter, [1 MPa] v Poisson ratio A New Design Solution for Aircraft Wheels that Reduces Overpressure in the Tire while Retaining its Absorption Power and its Dimensions 437 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 431-438 Cm , Cm materials constants, [MPa] J volume ratio, J — Xi Xn A3 A distance between the inner and outer tire, [mm] /i friction coefficient h deflection of a tire, [mm] p0 initial pressure, [MPa] E Young's modulus, [daN/mm2] G Shear modulus, [daN/mm2] 8 REFERENCES [1] Aircraft Tire Care and Maintenance Manual, from www goodyearaviation.com, accessed on 2019-06-11, Goodyear, Akron. 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D0l:10.5545/sv-jme.2020.6703 Original Scientific Paper Received for review: 2020-04-02 Received revised form: 2020-06-13 Accepted for publication: 2020-06-17 Characterization of the Bonding Zone in AZ91 AlSil B imetals Fabricated bjL iquid-Solid Compound Casting Using Unmodified and ThermalljM odified AlSil A lloy Renata Mola - Tomasz Bucki* Kielce University of Technology, Faculty of Mechatronics and Mechanical Engineering, Poland Liquid-solid compound casting was used to produce two types of AZ91/AISH2 joints. The magnesium alloy was the cast material poured onto a solid aluminium alloy insert with an unmodified or modified structure. The bonding zone obtained for the unmodified insert was not uniform in thickness. There was a eutectic region (Mg17A!12 + a solid solution of A! in Mg) in the area closest to the AZ91. The region adjacent to the /US/12 had a non-uniform structure with partly reacted Si particles surrounded by the Mg2Si phase and agglomerates of Mg2Si particles unevenly distributed in the Mg-AI intermetallic phases matrix. Cracks were detected in this region. In the AZ91/AISH2 joint produced with a thermally modified /US/12 insert, the bonding zone was uniform in thickness. The region closest to the AZ91 alloy also had a eutectic structure. However, significant microstructural changes were reported in the region adjacent to the modified /US/12 alloy. The microstructure of the region was uniform with no cracks; the fine Mg2Si particles were evenly distributed over the Mg-AI intermetallic phase matrix. The study revealed that in both cases the microhardness of the bonding zone was several times higher than those of the individual alloys; however, during indenter loading, the bonding zone fabricated from modified /US/12 alloy was less prone to cracking. Keywords: compound casting, magnesium alloy, aluminium alloy, grain refinement, microstructure, microhardness Highlights • The AZ91/AISH2 joints were fabricated through compound casting by pouring liquid AZ91 onto solid /US/12 inserts. • An insert with an unmodified structure and a thermally modified insert were used. • The thermal modification resulted in the refinement of the /US/12 alloy microstructure. • The study aimed to determine how the thermal modification of the aluminium alloy insert affected the microstructure, composition, and microhardness of the bonding zone. 0 INTRODUCTION Compound casting is an inexpensive and efficient method for joining two dissimilar metals or alloys to produce bimetallic parts, often complex in shape, with no geometric or dimensional restraints. The literature on the topic shows that compound casting lias been used for years to connect various dissimilar alloys, e.g., steel and cast iron [1] to [4], steel and alloyed steel [5] and [6], steel and A1 [7] and [8], A1 and Cu [9], dissimilar A1 alloys [10] and [11], dissimilar Mg alloys [12] or A1 and Mg alloys [13] to [27], There are two variants of the method: liquid-solid compound casting and liquid-liquid compound casting, with the former being more common. Liquid-solid compound casting consists of pouring one liquid alloy directly onto a solid insert made of another alloy, placed in a mould cavity. Liquid-liquid compound casting [5], [11] and [22], in contrast, involves pouring two molten alloys together directly into a mould. To prevent their mixing, a thin plate made of still another alloy is placed in the mould to act as an interlayer [22], An alternative solution is to rapidly cool one of the liquid alloys at the joint interface so that it starts to solidify when in contact with the other alloy poured [11]. Currently, magnesium and aluminium alloys are widely used in the automotive and aviation industries [28] to [30] as the application of lightweight materials aids in reducing the vehicle mass and, consequently, fuel consumption. A bimetal based on the two alloys retains the favourable properties of both and, as such, is an extremely attractive engineering material. Mg/Al bimetals can be obtained using various methods [31], As stated above, research on the joining of Mg alloys to A1 alloys using compound casting has been conducted extensively in recent years. Some studies have investigated joining pure Mg to pure A1 [14], [17] and [24], In this case, three zones can be differentiated at the interface: a eutectic (Mg17Al2 + a solid solution of A1 in Mg) adjacent to the Mg, the Mg7 Alj intermetallic phase in the middle and the Al3Mg2 intermetallic phase close to the Al. Some works dealing with Al/Mg bimetals produced by compound casting reveal that good results are obtained when Mg is poured onto an Al insert [14] and [17], The developments in Mg/Al compound casting indicate that a reaction can occur between liquid Mg and the oxide film on the surface of an Al insert. As the reaction leads to the removal of the passive film, there is direct contact between the metals, which allows the *Corr. Author's Address: Kielce University of Technology, Faculty of Mechatronics and Mechanical Engineering, 25-314 Kielce, Poland, tbucki@tu.kielce.pl 439 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 joint to be formed. The use of molten A1 and solid Mg [14], however, results in the occurrence of gaps at the interface; this is related to the presence of unreactive oxides on the surface of an Mg insert. One study on the subject [24] suggests that it is possible to form a uniform bonding zone for Al/Mg rods produced by pouring A1 around an Mg insert. The major problem encountered during the fabrication of Mg/Al bimetal joints is the occurrence of Mg-Al intennetallic phases at the interface. It is commonly known that Mg-Al intennetallic phases are hard but brittle. Thus, when a continuous zone of these phases forms close to the Al-based insert, the joint strength is low (20.2 MPa to 39.9 MPa) [14], For this reason, structure modification to improve the strength properties of a bimetallic joint seems a promising approach. Previous research has included studies of compound casting applied to join Mg alloys to A1 alloys, e.g., Mg/AlMgl [13], Mg/ A413 [15], AM50/A319 [16], AZ91/A356 [18] and [22], AZ9A AlSiT [19] and [20], ZE4 AlSil [21], AZ31/AlZn (AlZn6, AlZnl5, AlZn30) [23], and AZ31/ AW-6060 [25], The experimental data presented in the above-mentioned works show that the microstructure of the bonding zone undergoes modification because of the presence of other alloying elements in the alloys joined. As pointed out in [15] and [19] to [21], the use of AISi alloy in the form of a solid insert leads to the formation of the Mg-Al intennetallic phases as well as the Mg2Si phase in the bonding zone close to the A1 alloy. Applying ZE4 magnesium alloy [21] also results in some modification of the bonding zone structure in the area close to the Mg alloy. The Mg-Zn-RE and Al-RE intennetallic phases are detected there. The use of A1 alloys containing a high amount of Zn [23] causes a change in the morphology and microstructure of the interface, where Zn-rich intennetallic phases are predominant. As noted in [13], [22], [26] and [27], significant modification of the bonding zone structure can also be achieved by applying an interlayer made of a different metal: Mn [13], Zn [22] and [26] and Ni [27] between Mg and A1 in compound casting. The research described in [13] shows that an Mn interlayer prevents Mg-Al intennetallic phases from fonning at the Mg/Al interface. The results presented in [22] and [26] indicate that the bonding zone of the AZ31/AW-6060 joint fabricated with a Zn interlayer is mainly composed of Mg-Al-Zn phases and it is characterized by high shear strength, higher than that obtained without the Zn interlayer [26], Li et al. [27] reported that the application of a Ni interlayer could also limit the fonnation of brittle Mg-Al intennetallic phases and improve the strength of the joint. This article focuses on the fabrication of a light bimetallic material, composed of AZ9 magnesium alloy and AlSil2 aluminium alloy, using the liquidsolid compound casting method. The present study aimed to investigate how the thennal modification process of the aluminium alloy insert would affect the microstructure, composition, and microhardness of the bonding zone fonned between the alloys joined. 1 EXPERIMENTAL DETAILS AZ91 magnesium alloy, containing 9.14 wt% Al, 0.64 wt% Zn and 0.23 wt% Mn, was used as the cast material. Two types of AlSil2 aluminium alloy (10.97 wt% Si, 0.43 wt% Mn, 0.26 wt% Fe, 0.15 wt% Mg and 0.15 wt% Cu) were considered: unmodified (as-received) and thennally modified. AlSil2 was in the fonn of cylindrical inserts 30 mm in diameter and 10 mm in thickness. The unmodified inserts were cut from an AlSil ingot. The thennally modified inserts were prepared by pouring AlSil2 alloy onto a thick steel plate. The microstructures of both insert types are shown in Fig. 1. The thennal modification caused significant refinement of the AlSil alloy Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 microstructure, as depicted in Fig. ft . Before compound casting, the surfaces of the A1 alloy inserts were ground using SiC papers up to 800 grit, polished and cleaned with ethanol. The liquid-solid compound casting process proceeded as follows. First, an insert was placed in the cavity of a steel mould. Then, both were preheated to 300 °C. No protective atmosphere was used because of the ability of the molten Mg alloy to react with the oxide film on the surface of the A1 alloy. Next, the AZ91 alloy (100 g) was melted in a furnace under an argon atmosphere at a temperature of 660 °C. Finally, the Mg alloy was poured under normal atmospheric conditions into the mould with the aluminium alloy insert inside. The microstructural analysis of the AZ9A AlSil joints was performed on cross-sections prepared using standard metallographic techniques, with the final polishing being carried out with aluminium oxide slurry (0.3 jim). A Nikon ECLIPSE MA200 optical microscope (OM) and a JEOL JSM-5400 scanning electron microscope (SEM), equipped with an energy-dispersive X-ray spectrometer (EDS), were employed for the examinations. The microhardness of the bonding zone was measured with a MATSUZAWA MMT Vickers tester at a load of 2 RESULTS AND DISCUSSION Fig. 2 shows low and high magnification OM images of the microstructure of the AZ9 AlSil bimetal joint produced with an unmodified AlSil2 insert. It is visible that the bonding zone is continuous but nonuniform in thickness (120 |im to 200 |im) or structure. The area of the bonding zone close to the AZ9 alloy is thinner and darker, while that close to the AlSil2 is thicker with fine and coarse particles unevenly distributed in the light matrix. The contour of the interface with the AlSil2 alloy is highly irregular in shape (Fig. 2b). Fig. 3 shows details of the microstructure of the bonding zone observed through scanning electron microscopy. The results of the EDS quantitative analysis conducted at points 1 to 11 as marked in this figure, are provided in Table 1 The high magnification image in Fig. 3a reveals that the darker region adjacent to the AZ91 has a structure with two phases: dark and light (points 1 and 2, respectively). The Al-Mg equilibrium phase diagram [32] and the EDS results suggest that the chemical composition of the dark phase corresponds to a solid solution of A1 in Mg, whereas the light phase is the Mg7 Al2 intermetallic phase. Below the eutectic, there are light dendrites (point 3) of the Mg7 Al2 phase. Fig. 3b depicts the region adjacent to the AlSil2 alloy. The chemical composition of the light matrix is not uniform. The Mg7 Alj intermetallic phase was detected in the matrix close to the eutectic (point 4), and the Al3Mg2 phase (point 5) was observed close to the AlSil2. As can be seen from Fig. £ , the microstructure of the unmodified AlSil alloy comprised needle-like eutectic Si and Chinese script-like phases distributed regularly in the A1 matrix. The Chinese script-like Al-Si-Mn-Fe phases (point 6), as well as the needle-like and irregularly shaped Si particles (points 7 and 8, respectively), were also present in the bonding zone (Fig. 3b). The Al-Si-Mn-Fe phases were stable; they did not undergo transformation through reactions at the interface, so their Chinese script-like shape remained unchanged. 20 pm Fig. 2. 0M images of the bonding zone microstructure for the bimetallic joint with an unmodified /US/12 insert; a) low magnification; b) high magnification 441 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 AISH2 Fig. 3. SEM images of the bonding zone for the bimetallic joint with the unmodified /US/12; a) dose to the AZ91; b) close to the /US/12; cj details of the zone close to the /US/12 and the corresponding EDS line scan results Table 1. Results of the EDS quantitative analysis at points marked in Fig 3 Point Mg Al Zn Si Mn Fe [at.%] 1 84.15 15.85 - - - 2 64.36 35.42 0.22 - - 3 63.08 36.92 - - - 4 62.76 37.24 - - - 5 41.76 58.24 - - - 6 1.64 69.13 13.28 8.17 7.78 7 - 0.56 99.44 - - 8 - 0.44 99.56 - - 9 62.86 4.77 32.37 - - 10 39.49 60.51 - - - 11 55.75 44.25 - - - It was found that the Si particles observed in the bonding zone were surrounded by a darker phase. At a certain distance from the Si particles, the darker phase was fragmented and dispersed in the light-coloured matrix. This phenomenon is clearly visible in Fig. 3c. The EDS linear analysis along the marked line reveals that the darker phase area is rich in Mg and Si. According to the Mg-Si binary phase diagram [33], the solubility of Si in Mg is extremely low, and Si atoms react with Mg atoms and form the Mg2Si phase. The results of the EDS quantitative analysis in this area (point 9) confirm the presence of the Mg2Si phase. Some of the Si particles are too large to be fully consumed in the reaction with Mg during compound casting. Locally in the bonding zone, there are single Mg2Si particles and their agglomerates, where smaller Si particles were fully consumed. The light matrix area close to the Mg2Si "halo" is slightly lighter in colour (point 10), and its chemical composition resembles that of the Al3Mg2 intennetallic phase. In the slightly darker area (point 11) surrounding the Al3Mg2 particles, the atomic percentage of Mg is higher and more similar to that of the Mg7 Al2 intennetallic phase. The decrease in the Mg content in the light matrix close to the dark halo resulted from the formation of the Mg2Si phase. Figs. 2 , b and Figs. 442 Mo/a, R. - Buck/, I Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 Fig. 4. OM ¡mage of the microstructure of the bonding zone between AZ91 and modified /US/12; a) lower magnification; b) higher magnification 3 a and b reveal some pores and cracks in the bonding zone. The defects were caused by volume changes associated with the phase transformation at the reactive interface during liquid-solid compound casting. Long parallel cracks can be seen propagating across the light matrix between areas where the transformation of large Si particles into the Mg2Si phase occurred. The volume changes accompanying the formation of new phases resulted in transformation stresses, which occurred locally and initiated the formation of cracks in the matrix composed of brittle Mg-Al intermetallic phases. It is evident that the presence of numerous defects, such as cracks and pores, as well as residual stress negatively affects the properties of the joint. The microstructural observations also showed that the contour of the interface between the bonding zone and the AlSil2 alloy was highly irregular. The long needles of Si present in the AlSil alloy disturbed the diffusion reactions between the AZ91 and the AlSil2, which resulted in a non-uniform thickness of the bonding zone. The OM images in Fig. 4 show the microstructure of the bonding zone obtained when a thermally modified AlSil2 insert was used. From the low magnification view in Fig. 4a, it is evident that the refinement of the insert microstructure led to substantial changes in the joint microstructure. In this case, the bonding zone has a uniform thickness of about 250 |im. The interface between the AZ91 alloy and the bonding zone and that between the bonding zone and the AlSil alloy are flat. The region of the bonding zone adjacent to the AZ91 alloy is thinner and darker than the region close to the AlSil2, as was the case with the joint fabricated with an umnodifted AlSil2 insert. The thinner and darker region, however, can be further divided into two subregions (marked A and B). The thicker and lighter region closest to the AlSil2 (marked C) has a uniform structure with fine darker particles distributed evenly in the light matrix. Fig. 4b shows a higher magnification image where the microstructure of the three regions can be seen. Details of the bonding zone were examined using SEM/EDS. Fig. 5a shows a low magnification SEM image of the microstructure of the bonding zone and the linear distribution of elements along the marked line. As can be seen from the linear analysis, the bonding zone is mainly composed of Mg, Al, and Si. The concentration of Mg close to the AZ9 is slightly higher than that in the area adjacent to the AlSil2. In contrast, the concentration of Al increases slightly with increasing distance from the area close to the AZ91. Si was detected only in regions B and C of the bonding zone. The high magnification images in Figs. 5b to d reveal details of the microstructure of the three regions (A, B and C). The EDS quantitative analysis was carried out at points 1 to 11 marked in these figures. The results are provided in Table 2 A lamellar structure is visible in the AZ91 alloy close to the bonding zone (Fig. 5b). At point 1, a high concentration of Al was detected, which suggests that during compound casting, enrichment of AZ9 with Al occurs locally close to the bonding zone and plate-like precipitates of the Mg7 Al2 phase form in the solid solution of Al in Mg [34], This phenomenon was also observed inthe AZ91/AlSil2 joint fabricated using an umnodified AlSil nsert. Changes in the AZ9 alloy microstructure caused by the diffusion of Al were visible at a distance of about 200 |im from the bonding zone. Fig. 5b reveals that region A of the bonding zone close to the AZ9 alloy has a eutectic structure. The eutectic is composed of an Mg7 Alj intermetallic phase (light phase - point 443 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 Fig. 5. SEM images of the bonding zone between AZ91 and modified /US/12; a) low magnification image with EDS line scan results; b) high magnification image of zone A; c) high magnification image of zone B; dj high magnification image of zone C with EDS line scan results 2) and a solid solution of A1 in Mg (dark phase - point 3). In region B (Fig. 5c), the light dendrites of the Mg7 Alj intennetallic phase (point 4) are surrounded by a eutectic (Mg7 Alj (point 5) + a solid solution of A1 in Mg (point 6)). The region also contains fine grey particles. The EDS result obtained for point 7 suggests that these particles are the Mg2Si phase. Fig. 5d shows details of the inicrostructure of region C adjacent to the AlSil2 alloy and a distribution of elements along the marked line. In this region, there are Si-rich grey particles regularly distributed in the 444 light matrix rich in Mg and Al. The EDS quantitative analysis was carried out for the light matrix at points 8 to 10. The analysis at point 8 indicates the Mg7 Alj intennetallic phase. The data obtained for the region closer to the AlSil2 alloy (points 9 and 10) suggest the Al3Mg2 intennetallic phase. The quantitative analysis of the gray particles (point 11) indicates the Mg2Si phase. From the linear distribution of elements, it is clear that the fine, white particles found in this area are rich in Al, Si, Mn and Fe. In the bonding zone fonned between the AZ91 and the thennally modified Mo/a, R. - Buck/, I Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 Table 2. Results of the EDS quantitative analysis at points marked in Fig. 5 Point Mg Al Zn Si [at.%] 1 85.11 14.77 0.12 - 2 65.42 34.18 0.40 - 3 79.76 20.24 - - 4 61.68 38.32 - - 5 63.23 36.77 - - 6 74.76 25.24 - - 7 61.18 5.51 - 33.31 8 62.10 37.90 - - 9 40.67 59.33 - - 10 40.18 59.82 - - 11 60.27 5.82 - 33.91 Fig. 6. An OM image of the microstructure of the AZ91/AISH2 joint with some local porosity in the bonding zone AlSil2, there were no cracks that could reduce the mechanical properties of the joint. Fig. 6 shows that, in this case, only single pores occurred locally in the bonding zone. It can be assumed that the formation of the bonding zone between AZ91 and AlSil2, described in this study, was as follows. The compound casting process involved pouring liquid AZ9 alloy heated to 660 °C onto a solid AlSil2 alloy insert preheated to 300 °C, which led to the interdiffusion of the elements at the reactive interface as well as partial melting of the AlSil alloy surface. A thin layer of the AlSil insert Fig. 7. Microhardness of the bonding zone oftheAZ91/AISH2 joint fabricated using unmodified /US/12; a) low magnification; b) high magnification of the area close to theAZ91; c) high magnification of the region close to theAISH2 445 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 (100 |im to 150 |.nn) and the AZ91 alloy solidifying at the interface were mutually enriched with the elements present in the two alloys as a result of diffusion and fusion processes. When the temperature of the AZ91/ AlSil bimetal fell after casting, the solidified liquid formed a continuous transition zone containing Mg-Al intennetallic phases. The microstructure of the zone was not homogeneous because of the concentration gradient in the reactive interface. Compound casting with the use of thermally modified AlSil2 aluminium alloy inserts led to the formation of a bonding zone that was more uniform in thickness and contained no cracks. In this case, the microstructure of the bonding zone close to the AlSil2 was relatively uniform. Fine Mg2Si particles were distributed evenly in the Mg-Al intennetallic phases matrix. There were no Si particles that were not fully consumed in the reaction with Mg or the coarse Chinese script-like Al-Si-Mn-Fe phases. Figs. 7 and 8 show impressions left by the Vickers indenter in the AZ91/AlSil2 joints fabricated using an unmodified and a thermally modified AlSil insert, respectively. Whichever the case, the microhardness of the AZ91 alloy was in the range of 56.2 HVto 65.5 HV. The unmodified and thermally modified AlSil alloy inserts differed in microhardness, which was 43 HVto 48.5 HV (Fig. 7) and 51.7 HVto 53.8 HV (Fig. 8), respectively. As is evident from these figures, the impressions in both bonding zones were much smaller than those in the alloys joined, which indicates the high hardness of the interface. The microhardness values were reported to change across the bonding zones because of its non-uniform microstructure. For the AZ91/AlSil2 joint produced with an umnodified AlSil2 insert (Fig. 7), the microhardness of the eutectic region close to the AZ91 alloy was ranged from 156 to 170.4 HV Slightly higher values (171.6 HV to 180.5 HV) were reported in the transition area between the eutectic and the Mg7 Al2 phase region, where dendrites of the Mg7 Al2 phase were found. Higher microhardness (198.6 HV to 252.2 HV) was observed in the region close to the AlSil aluminium alloy. The values, however, varied significantly. The highest microhardness was reported in the areas where Mg2Si phase agglomerates or not fully consumed Si AZ91 m SOPfl gfff c Im^wmm AIS.i12 AlSil 2 100 urn Fig. 8. Microhardness of the bonding zone of the AZ91/AISH2 joint fabricated using modified /US/12; a) low magnification; b) high magnification of the area close to theAZ91; c) high magnification of the region close to theAISH2 446 Mo/a, R. - Buck/, I Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,439-448 particles surrounded by the Mg2Si phase occurred in the Mg-Al intennetallic phases matrix. As can be seen from the higher magnification images in Figs. 0 and c, no cracks were observed near the impressions left in the eutectic region; they developed, however, at the corners and near the edges of the impressions in the bonding zone close to the AlSil2 alloy, which suggests that the Mg-Al intennetallic phases constituting the matrix of this region were brittle. In the bonding zone of the AZ9 AlSil bimetal fabricated from modified AlSil2 (Fig. 8), the microliardness of the eutectic region ranged between 158.5 HV and 171.3 HV The microliardness of the region close to the AlSil did not differ much either (224.7 HV to 233.8 HV). This was a consequence of a more uniform structure of this region, which contained fine Mg2Si phase particles distributed regularly in the Mg-Al intennetallic phases matrix. As can be seen from Fig. 8b, no cracks are propagating from the impression in the eutectic region. Cracks propagating from the impression corners (Fig. 8c) were detected only in the region adjacent to AlSil2, where Mg2Si particles were embedded in the Al3Mg2 intennetallic phase matrix, which indicates that this intennetallic phase had higher brittleness than the Mg7 Al12 phase constituting the bonding zone matrix close to the eutectic region. 3 CONCLUSIONS The AZ91/AlSil2 bimetallic joints under analysis were fabricated through liquid-solid compound casting. The process involved pouring liquid AZ9 magnesium alloy heated to 660 °C onto a solid AlSil2 aluminium alloy insert placed in a steel mould, both preheated to 300 °C. Two types of AlSil2 inserts were used in the experiment: unmodified and thennally modified. It was found that the thennal modification resulted in the refinement of the AlSil alloy microstructure and was consequently responsible for the microstructural changes in the bonding zone. When an unmodified insert was used, the bonding zone was non-unifonn in thickness. The region closest to the AZ$ alloy had a eutectic structure (Mg7 Al2 + a solid solution of A1 in Mg). The region adjacent to the AlSil2 alloy was non-unifonn. The Mg-Al intennetallic phases matrix contained partly reacted Si particles sunounded by the Mg2Si phase, single fine Mg2Si particles and their agglomerates as well as irregularly distributed coarse Al-Si-Mn-Fe phase particles. Long parallel cracks were observed in this area of the bonding zone. In the case of the AZ91/AlSil2 joint fabricated using a thennally modified AlSil insert, the bonding zone was unifonn in thickness, and no cracks were detected. The region of the bonding zone close to the AZ9 alloy also had a eutectic structure. Significant microstructural changes were observed in the region close to the AlSil alloy. The microstructure of the bonding zone adjacent to the AlSil2 was unifonn. There were fine Mg2Si particles evenly distributed in the Mg-Al intennetallic phases matrix. 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Archives of Foundry Engineering, vol. 10, no. 1, p. 23-26. 448 Mola, R. - Bucki, T. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8,449-457 © 2020 Journal of Mechanical Engineering. All rights reserved. D0l:10.5545/sv-jme.2020.6681 Original Scientific Paper Received for review: 2020-03-22 Received revised form: 2020-06-12 Accepted for publication: 2020-07-01 A Studjf or the Nanoflnishing of an EN-3®^ kpiece with Pulse DC Power SuppljU sing Ball-End Magnetorheological Finishing Himmat Singh - M. S. Niranjan* - Reeta Wattal Delhi Technological University, Mechanical Engineering Department, India Ball-end magnetorheological finishing (BEMRF) is a high-level nanoflnishing process used in finishing different kinds of surfaces including flat, 2- and 3-dimenslonal, and curved surfaces In the present study, a pulse DC power supply is used to energize the electromagnet of magnetorheological (MR) finishing tool. The experiments have been conducted on EN-31 flat workpiece surface with and without pulse DC power supply using a magnetizing current (MC) 2.5 A, a working gap (WGJ of 1.5 mm and a rotational speed of the tool (RST) of 500 rpm with a feed rate of workpiece of 50 mm/min. The study has been carried out to analyse the effect of the duty cycle on the response percentage reduction in surface roughness. It has been observed that an improved response percentage reduction in surface roughness has been found with pulsating DC power supply as compared to the response percentage reduction in surface roughness obtained with DC power supply without pulse at the same process parameters. After conducting the preliminary experiments, the statistical analysis was done to analyse the effect of various process parameters on the response percentage reduction In surface roughness using response surface methodology (RSM) at 0.16 duty cycle. Keywords: BEMRF, pulse DC power supply, duty cycle, response percentage reduction in surface roughness Highlights • Ball-end magnetorheological finishing (BEMRF) is a high-level nanoflnishing process. • In the present study, the pulse DC power supply is used to energize the electromagnet of magnetorheological (MR) finishing tool. • It is noted that the better percentage reduction In surface roughness has been achieved with a pulse DC power supply than without pulse DC power supply. • It is a/so observed that the surface texture achieved by pulse DC power supply is more even as compared to that obtained without pulse DC power supply. 0 INTRODUCTION In non-traditional finishing processes, the application of magnetic fields has become very important for nano-level finishing on a variety of surfaces. Nano finishing of critical shapes is always difficult to control despite the high demand for nano-finishing. The traditional finishing processes, such as honing, grinding, etc., produce finished surfaces, but these finishing processes produce some thermal and residual stresses on the workpiece surface. It, therefore, becomes a challenge to finish these types of components [1]. Due to advancement in new materials and complex shapes of workpiece geometry, some newly developed advanced finishing processes are employed to resolve such types of problems. One important parameter encountered is surface roughness, which plays a crucial role in high-quality products. These processes are helpful for polishing of any type of materials [2] and [3], The magnetorheological finishing process, which has more flexibility for process control and a high level of finishing with close dimensional tolerances, can be achieved without leaving any defect on surfaces or subsurfaces. A finishing spot is formed at the tip of finishing tool in magnetorheological finishing, which acts as a semisolid finishing tool that moves over the workpiece surface during finishing of any kind of workpiece surface [4]. Magnetorheological (MR) fluid has more flexibility during the finishing operation; it changes from liquid to semisolid in a very short duration due the effect of the magnetic field. MR fluid is prepared with ferro-magnetic particles and various types of abrasive particles mixed with some base fluid, such as paraffin oil. In advancement of MR fluid for better finishing, bidisperse MR polishing fluid samples have been prepared with different percentages of carbonyl iron powder (CIPs) of standard CS and HS grade (BASF, Germany), and the response has been compared with monodisperse MR polishing fluid. The magnetorheological characterization of bidisperse and monodisperse MR polishing fluid samples have been studied at different magnetic fields with the use of a magnetorheometer [5], In MR finishing, different process parameters, such as current, working gap, feed rate and central core rotation have been analysed regarding the response surface roughness. The composition of MR polishing fluid has also been studied; it was observed that the CIPs concentration is the most influential parameter *Corr. Author's Address: Delhi Technological University, Mechanical Engineering Department, India, himmat.diwakar48@gmail.com 449 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,449-457 on the response surface roughness in comparison to other process parameters for the finishing of hardened AISI 52100 steel [6], Many finishing techniques based on magnetic fields have been developed. In these techniques, MR fluid behaves like a semisolid finishing spot in the presence of a magnetic field, which is used for finishing action on the workpiece These magnetic finishing techniques include magnetic abrasive finishing (MAF) [7], magnetorheological jet finishing (MRJF) [8], magnetic float polishing (MFP) [9], magnetorheological finishing (MRF) [4], rotational magnetorheological abrasive flow finishing (R-MRAFF) [10], and ball-end magnetorheological finishing (BEMRF) [11], MAF is a process that has been used to finish flat surfaces, two dimensional (2D) surfaces, and any complex shape of workpiece with high dimensional accuracy. A better-quality product is achieved by the MAF process with a high level of surface finish without any defect on the surface or sub-surface [12], In the MAF process, direct current (DC) power supply has been used to energize the electromagnet of finishing. Due to the higher demand for the improvement in surface roughness, the MAF process has been used with Pulse DC power supply instead of DC power supply without pulse. Experiments have been conducted with the use of DC power supply and pulsating DC power supply, which showed that there is better surface finish with DC power supply compared to DC power supply without pulse using the same process parameters. It is also observed during continuous DC power supply that the abrasive particles are not so effective for finishing the workpiece surface after a certain time because the cutting edge of abrasive particles loses its finishing capability in finishing zone. While using pulsating DC power supply, the orientation of abrasive particles may be changed or new effective grains of abrasive come in contact with the workpiece surface during finishing, which promotes better surface finish [13], Many existing magnetic finishing processes have been discussed in the literature, which are not suitable for finishing three dimensional (3D) complexly shaped workpieces, such as narrows cut in workpiece. For the advancement of the MR finishing process with some modification, a BEMRF process has been developed for finishing of any kind of workpiece surface (flat, curved surface, 2D, 3D and stepped surfaces, etc.). The electromagnetic coil of a finishing tool is energized, and MR fluid is used at tip of the tool formed a semisolid ball. The semisolid ball-shaped MR fluid is responsible for the finishing action on the workpiece surface irrespective of any kind of workpiece materials. The BEMRF process can finish any kind of workpiece surface as achieved by finishing in computer navigated controlled (CNC) milling machine for three-dimensional surfaces [14], This process has major applications in the optics industry, aerospace, and automotive component, among others. In the present study, an attempt has been undertaken to energize the electromagnetic coil of the BEMRF tool with the help of pulse DC power supply for finishing of a flat EN-31 workpiece which may give improved surface roughness and may increase the efficiency of finishing action of BEMRF process. 1 METHODS Various components of the BEMRF setup are shown in Fig. 1 The setup consists of an electromagnetic coil which is energized as soon as the DC supply is switched ON, which results in the formation of a semisolid hemisphere or ball-shaped finishing spot at the extremity end of the tool as shown in Fig. lb. Fig. 1. a) Pulse DC power supply, b) BEMRF tool with MR fluid on workpiece, c) chiller for cooling the electromagnetic coil, dj BEMRF setup, e) EN-31 finished workpiece, fj schematic of die with workpiece dimension in mm This semisolid ball of MR fluid formed at the tip of tool approximately 15 mm diameter is used for finishing of EN-31 workpiece. A chiller (Fig. lc) is 450 Singh, H. - Niranjan, M.S. - Wattal, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 449-457 installed to control the constantly rising temperature of the electromagnetic coil due to the continuous supply of current. 1.1 Mechanism of Material Removal In the BEMRF process, the DC power supply is used to energize the electromagnetic coil due to which a semisolid hemisphere or ball-shaped finishing spot of MR polishing fluid is established at the extremity end of the tooltip. The semisolid finishing spot is used for finishing of surfaces such as flat, curved surface, D , 3D, and stepped surfaces irrespective of the movement of finishing tool. The semisolid ball-shaped finishing spot lias more pliability to move over any kind of workpiece surface for finishing. In conventional MR finishing process, the abrasive particles in contact with the workpiece do not change their orientation in the entire finishing time due to fixed magnetic flux density. As a result, the cutting edges of the abrasive particles become dull, resulting in poor surface finish. In order to further enhance the finishing efficiency, the fresh active abrasive particles are required again and again to finish the workpiece surface, which is not possible with conventional MR finishing process. To overcome this limitation, an improvement has been made in the conventional MR finishing process by providing fluctuating DC power supply. Fig. 2. Mechanism of material removal in BEMRF process; a) ball is formed at finishing tool with pulse DC power supply (TON condition), b) ball is formed at finishing tool with pulse DC power supply (TOFF condition) Pulse DC power supply is given to the BEMRF finishing tool, which produces a fluctuating magnetic field on the tip of MR finishing tool. Therefore, fresh active abrasive grains may come outward or the orientation of abrasive particles may be changed and the finishing action of abrasive particles over the surface of workpiece in direct contact enhances the finishing efficiency of BEMRF process. During the process, the viscosity of MR polishing fluid reduces as soon as the DC power supply is switched off for very short duration, which results in changing the semisolid ball towards liquid a state. As soon as the power supply is ON, the orientation of the abrasive particles is changed due to the fluctuating magnetic field or some new active abrasive grains may move outward, as shown in Fig. 2a. As soon as the power supply is OFF, the orientation of the abrasive particles is changed due to the fluctuating magnetic field, as shown in Fig. 2b. The frequent ON and OFF of DC power supply results in orientation change of abrasive particles, hence fresh abrasive particles come in contact with the workpiece surface. 2 EXPERIMENTAL The experimental setup has an electromagnetic coil, pulse DC power supply, thermocouple, and a chiller for cooling the electromagnetic coil is shown in Fig. 1 A semisolid hemisphere ball of MR polishing fluid and pulsed BEMRF setup is given in Fig. ft and d, respectively. MR polishing fluid changes from liquid to semisolid state and behaves like a semisolid hemispherical ball due to the magnetic flux density produced by the electromagnetic coil. A homogenous mixture of MR polishing fluid was synthesized with silicon carbide abrasive particles of 800 mesh size (25 vol%) with density d 3.33 gm/ cm3, ferro magnetic iron particles (CIP CS grade, 0 vol%) with density d 7.8 gm/cm3 and 55 vol% of base fluid. When a magnetic field is present at the tip of the tool, ferro-magnetic iron particles (CIPs) engaged to each other form a chain-like columnar structure in which abrasive particles are positioned between magnetic particles (CIPs). A semisolid MR polishing fluid ball is formed at the tip of tool, which moves over the workpiece surface and hence material is sheared off the workpiece surface in the form of very small chips. The workpiece position and its die are given in Fig. If. The total height of the workpiece is 10 mm, the depth of the slot is 8 mm and, during the finishing operation, the workpiece is kept slightly above 2 mm from slot depth. The preparatory experiments were conducted to attain the intended purpose of the A Study for the Nanofinishing of an EN-31 Workpiece with Pulse DC Power Supply Using Ball-End Magnetorheological Finishing 451 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,449-457 finishing process and to develop the ranges of the duty cycle parameters. Duty cycle r is given as: r=TONl (Ton+Toff). (1) Here TON and TOFF denotes on-time & off-time of pulse DC power supply respectively. The preliminary experiments have been carried out on an EN-31 flat workpiece surface with and without pulse DC power supply given to electromagnetic coil. The approximate values for parameters are selected on the basis of the literature review done in this work [15] to [17], These values contributed to the notable results observed in the studies highlighted by the above references. These experiments have been conducted using process parameters made up of; working gap 1.5 mm, tool rotational speed 500 rpm, current 2.5 A, and finishing time of 30 min. These parameters were employed with DC power supply at different duty cycles with and without pulses. The response parameter percentage reduction in surface roughness (%ARa) was calculated using; n/ . „ initial roughness - final roughness =-----x 100 %. initial roughness A Talyor Hobson surface analyser was used with cut-off length 0.8 mm and data length of 4 mm to measure surface roughness. 3 RESULT AND DISCUSSION The roughness profile of EN-31 workpiece surface with DC power supply without pulse is shown in Fig. 3. The percentage reduction in surface roughness was calculated and given in Table ti nd 2 The roughness profile of the finished EN-31 workpiece surface was drawn after conducting experiments with pulse DC power supply at 0.16 duty cycles, as shown in Fig. 4b and %ARa has been calculated, which is given in Table 1. It is observed that the %ARa lias been found better by conducting the experiments with pulse DC power supply at 0.16 duty cycle as compared to %ARa obtained by DC power supply without pulse. The surface roughness profile of finished EN-31 workpiece surface has been drawn after conducting experiments with pulse DC power supply at 0.27 duty cycle is shown in Fig. 5b. The percentage reduction in surface roughness was calculated for preliminary experiments at various duty cycles and given in Table 1 A total of eighteen experiments were performed with three repletion for each sample (total number of samples = 6). It has been observed from the experimental study that the best %ARa was found at 0.16 duty cycle. It is observed from Fig. 6 that the highest %ARa is found to be 37.09 % at 0.16 duty cycle and is lowest i D 02 0.4 D 6 d8 i i j r* i.® ie x it *- ° x o a ¿>.t M J.O mn , , a) b) Fig. 3. aj Surface roughness profile of EN-31 before finishing (Ra = 0.187 ym, Rq = 0.244 ym, Rz = 1.56 m); b) surface roughness profile after finishing using DC power supply without pulse (Ra = 0.161 ji/m, Rq = 0.203 ji/m, Rz = 1.13 /jmJ Table 1. Preliminary experimentation details with pulse supply Exp. no. Duty cycle on-time [ms] off-time [ms] Pulse time Initial ña [/um] Final ña [/jm] %ARa 1 0.45 4 5 9 0.197 0.172 12.69 2 0.36 4 7 11 0.194 0.161 17.01 3 0.16 4 21 25 0.186 0.117 37.09 4 0.27 4 11 15 0.197 0.149 24.36 5 0.61 4 3 7 0.197 0.185 6.09 6 0.67 4 2 6 0.185 0.176 4.86 Table 2. Preliminary Experimentation without pulse supply Exp no._MC [A]_RST [rpm] WG [mm] Time [min] Initial Ra pum] Final ifa pum]_%ARa 1 2.5 500 1.5 30 0.187 0.161 13.44 452 Singh, H. - Niranjan, M.S. - Wattal, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 449-457 Fig. 4. a) Roughness profile before finishing ofEN-31 workpiece surface (Ra = 0.187 ji/m, Rq = 0.235 ji/m, Rz = 1.42 /jmJ; b) Surface roughness profile after finishing with pulse DC power supply at 0.16 duty cycle (Ra = 0.117 ji/m, Rq = 0.151 ji/m, Rz = 0.95 /jmJ; 02 0 < 06 06 I !I U 1.« II 2 22 21 2« 26 3 12 M It Hmn bj Fig. 5. a) Roughness profile before finishing ofEN-31 workpiece surface (Ra = 0.197 ¡jm, Rq = 0.255 ¡jm, Rz = 1.4 /jmJ; b) surface roughness profile after finishing with pulse DC power supply at 0.27 duty cycle (Ra = 0.149 ji/m, Rq = 0.198 ji/m, Rz = 1.16 ji/mJ 37.05 24.36 Duty Cycle 0.3 0.5 0.7 Fig. 6. Relationship between percentage reductions in surface roughness with the duty cycle at 0.67 duty cycle. The %ARa is found to be 13.44 % on conducting the experiments with DC power supply without pulse. It is due to the adopted technique that the fresh active abrasive grains change their orientation and come in contact with workpiece surface directly during finishing action, which enhances the finishing efficiency of BEMRF process. During the process, the viscosity of MR fluid reduces as soon as the DC power supply is switched off for a very short duration. As soon as the power supply is ON, the orientation of the abrasive particles get changed due to the fluctuating magnetic field, or some new active abrasive grains may come outward. The frequent ON and OFF of DC power supply results in orientation change of abrasive particles; thus, fresh abrasive grains may come in contact with the workpiece surface. The optical microscopic views of finished EN-31 workpiece surface with and without pulse DC power supply are shown in Fig. 7. The best texture of surface finished using the pulse DC power supply was found at duty cycle 0.16, as shown in Fig. 7c, as compared Fig. 7. Optical microscopic views of finished surface; a) with DC power supply without pulse, b) with pulse DC power supply at 0.27 duty cycle, c) with pulse DC power supply at 0.16 duty cycle A Study for the Nanofinishing of an EN-31 Workpiece with Pulse DC Power Supply Using Ball-End Magnetorheological Finishing 453 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,449-457 Fig. 8. SEM images of workpiece surface: a) before finish (lays are present on the initial grinded surface), b) finished workpiece surface (Improved surface texture) at 0.27 duty cycle, c) finished workpiece surface (better improved surface texture) at 0.16 duty cycle, d) finished workpiece surface (less improved surface) with DC power supply without pulse to the surface finished with DC power supply without pulse, as shown in Fig. 7a. From optical microscopic views of the finished surface, it was found that the surface texture is observed improved as the duty cycle decrease from 0 2 o Scanning electron micrograph (SEM) of the EN-31 workpiece surface is shown at 100 |im resolution and 600 x magnification as outlined in Fig. 8. The lays are clearly visible in the initial grinded surface, as shown in Fig. 8a. It is observed from Fig. 8c that a more uniform finished surface is obtained with pulse DC power supply at 0.16 duty cycle as compared to the finished surface obtained with DC power supply without pulse, as shown in Fig. 8d. From Fig. 8b and c, it is shown that as the duty cycle decreases from (E to (ft. and a better surface texture is achieved at (ft dut y cycle. 4 DESIGN OF EXPERIMENT The present work utilizes central composite design (CCD) under response surface methodology to design the experiments based on the preliminary study. In this regression analysis, three parameters are used (II'G. MC and R3> ) to determine the significance of these parameters on output responses %ARa. The most appropriate value of duty cycle was found to be 0.16, for which the highest %ARa observed. Therefore, the 0.16 duty cycle was taken for all experiments. The run order and results of output responses for the finishing of EN-31 using BEMRF process are shown in Table 3. 5 RESULTS 5.1 Analysis of Surface Roughness Insignificant terms having /»-value greater than 0.05 are eliminated by using backward elimination and the pooled ANOVA results for surface roughness are 454 presented in Table 4. Table 3 shows the %ARa for the EN-31 workpiece before and after finishing through the BEMRF process. Table 3. Design and result of output response in surface roughness (%ARa) Std order Run order MC [A] RS [rpm] WG [mm] %ARa 16 1 2.5 700 1.5 34.49 12 2 3 400 1 39.17 4 3 2.5 300 1.5 28.25 1 4 3.5 500 1.5 40.69 10 5 2 400 1 26.6 14 6 3 600 2 31.12 15 7 3 400 2 29.12 13 8 2.5 500 1.5 33.48 17 9 2 400 2 24.38 20 10 2.5 500 2.5 23.22 6 11 1.5 500 1.5 20.68 7 12 2 600 2 26.99 11 13 2.5 500 1.5 32.06 5 14 2 600 1 27.9 9 15 2.5 500 0.5 37.03 2 16 2.5 500 1.5 32.48 3 17 3 600 1 40.37 19 18 2.5 500 1.5 32.48 18 19 2.5 500 1.5 32.1 8 20 2.5 500 1.5 32.23 In the tool Design-Expert, the quadratic model is selected on the basis of the lack of fit tests, since the cubic model is aliased. Table 4 shows the significant terms after analysis of variance (ANOVA). The /»-value less than 0.05 shows that the model parameters are significant. In this reduced model of ANOVA, the following terms A, B, C, AC, A2 and C2 are found to be significant as shown in Table 4./-value is 108.61, which indicates that the model is significant. There is only (W/<> possibility of this to occur due to noise. Singh, H. - Niranjan, M.S. - Wattal, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 449-457 Table 4. ANOVA for %ARa Source Sum of Squares D0F MSE /-value />-value Model 565.92 6 94.32 108.61 <0.0001 sign A-A/C 341.6 1 341.6 393.36 <0.0001 £ B-RS1 23.99 1 23.99 27.62 0.0002 ra rr C - WG 156.56 1 156.5 180.28 <0.0001 £ AC 32.68 1 32.68 37.64 0.0376 A2 4.65 1 4.65 5.36 0.0087 C2 8.27 1 8.27 9.53 Residual 11.29 13 0.868 Lack of fit 9.91 8 1.24 4.48 0.0577 not sig Pure error 1.38 5 0.2766 Cor Total 577.21 19 The equation for %ARa in term of actual factor as is given below: =-33.6820 + 29.77 xA+0.0122 xB + 20.679 x C -8.08 x Ax C-1.68 x A2-2.24 x C2 6 DISCUSSION The results obtained after finishing of EN-31 using BEMRF process and the effect of parameters on %ARa are discussed in this section. 6.1 Effect of Rotational Speed of Tool (RST) Fig. 9 outlines the effect of RST on %ARa at various WG and at constant MC 2.5 A & 0.16 duty cycle. The %ARa slightly increasing with the increase in R3> at all WG. The effect of R3> reveals that it is the least contributing parameter on %ARa with a 4.17 % (from model analysis). 6.2 Effect of Magnetizing Current (MC) Fig. 10 describes the effect of MC on %ARa at various R3> and at a constant 0.16 duty cycle & working gap 1.5 lmn. The %ARa increased with the increasing in MC at all RT . The MC is the highest contributing parameter on %ARa with 59.18 % (from model analysis). It is also seen from the perturbation or 3D surface diagram, as shown in Fig. 12a and b. As the magnetic flux increases, the magnetic force acting on SiC abrasive particle increases due to which high %ARa is achieved at (ft dut y cycle. 6.3 Effect of Working Gap (WG) Fig. 11 explains the effect of WG on %ARa at various MC and at constant duty cycle & R3> 500 rpm. 200 400 600 Rotationalspeedoftool [rpm] Fig. 9. Effect of RST on %ARa at 2.5 A Magnetizing current [A] Fig. 10. Effect ofMC on %&Ra at WG 1.5 mm 60 50 40 30 20 10 MC-1.5 --»-MC-2 - A- MC-2.5 - * MC-3 - ■ MC-3.5 * J ^ V \ ^---■ ♦*.......* 0 0.5 1 1.5 2 2.5 3 Working gap [mm] Fig. 11. Effect of WG on %&Ra at 500 rpm The %ARa decreases with the increase in WG at all MC. The effect of WG reveals that it is second most contributing parameter on %ARa with 27.12 % (from model analysis). It is also seen from the perturbation or 3D surface diagram, as shown in Fig. 12a and c. A Study for the Nanofinishing of an EN-31 Workpiece with Pulse DC Power Supply Using Ball-End Magnetorheological Finishing 455 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8,449-457 Fig. 12. a) Perturbation diagram for %ARa (A - current, B - working gap, C- current), b) and c) 3D surface for 96ARa 7 CONCLUSIONS The following conclusions are drawn on the study with ball-end magnetorheological finishing process with and without pulse DC power supply. 1 It is noted that the best percentage reduction in surface roughness (%ARa) was achieved with pulse DC power supply than without pulse DC power supply. 2 The %ARa is found as 37.09 % at 0.16 duty cycle and 24.36 % at 0.27 duty cycle while 13.44 % with DC power supply without pulse. 3. It was also observed that the surface texture achieved by pulse DC power supply is more monotonous as compared to without pulse DC power supply. 4 The effect of different rotational speeds of the tool, R3> is found to be the least contributing parameter on %ARa with a reduction of 4.17 % at (ft dut y cycle. 5. The %ARa increased with the increase in magnetizing current, MC at all R3> . It is the highest contributing parameter on %ARa with a reduction of 59.18 % at 0.16 duty cycle. 6. The %ARa decreases with the increasing in I I'd at all MC. It is second most contributing parameter on %ARa with a reduction of 27.12 %. 8 ACKNOWLEDGEMENTS I would like to show our gratitude to the Delhi technological university for providing experimental setup and for sharing their pearls of wisdom with me during the course of this research. 9 NOMENCLATURES d density, [g/cm3] MC magnetizing current, [A] WG working gap, [mm] Ra surface roughness, ||im| R3> rotational speed of tool, [rpm] %ARa response parameter percentage reduction in surface roughness, [%] r duty cycle, [-] Ton on-time of pulse DC power [ms] Toff off-time of pulse DC power [ms] 8 REFERENCES [1] Kordonskl, W, Gollnl, D. (1998). Magnetorheological suspension-based high precision finishing technology (MRF). Journal of Intelligent Material Systems and Structures, vol. 9, no 8, p 650 654, DOI:10.1177/1045389X9800900811 [2] Jain, V.K., Jayswal, S.C, Dixit, P.M. (2007). Modeling and simulation of surface roughness In magnetic abrasive finishing using non-uniform surface profiles. Journal of Materials and Manufacturing Processes, vol. 22, no. 2, p. 256-270, D0l:10.1080/10426910601134096 456 Singh, H. - Niranjan, M.S. - Wattal, R. Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 449-457 [3] Zhong, Z.W. (2008). Recent advances in polishing of advanced materials. Journal of Materials and Manufacturing Processes, vol. 23, no. 5, p. 449-456, D0l:10.1080/10426910802103486. [4] Kordonski, W.I., Jacobs, S.D. (1996). Magnetorheological finishing. International Journal of Modern Physics B, vol. 10, no. 23-24, p. 2837-2848, D0I:10.1142/S0217979296001288. [5] Niranjan, M.S., Jha, S. (2014). Flow behaviour of bidisperse MR polishing fluid and ball end MR finishing. Procedia Materials Science, vol. 6, p. 798-804, D0I:10.1016/j. mspro.2014.07.096. [6] Maan, S., Singh, A.K. (2018). Nano-surface finishing of hardened AISI 52100 steel using magnetorheological solid core rotating tool. The International Journal of Advanced Manufacturing Technology, vol. 95, p. 513-526, D0I:10.1007/ s00170-017-1209-x. [7] Shinmura, T., Takazawa, K., Hatano, E., Aizawa, T. (1985). Study on magnetic abrasive process-process principles and finishing possibility. Japan Society of Precision Engineering, vol. 19, no. 1, p. 54-55. [8] Kordonski, W.I., Shorey, A.B., Tricard, M. (2006). Magnetorheological jet finishing technology. Journal of Fluids Engineering, vol. 128, no. 1, p. 20-26, D0I:10.1115/1.2140802. [9] Komanduri, R. (1996). On material removal mechanisms in finishing of advanced ceramics and glasses. CIRP Annals, vol. 45, no. 1, p. 509-514, D0I:10.1016/S0007-8506(07)63113-8. [10] Das, M., Jain, V.K. Ghoshdastidar, P.S. (2012). Nanofinishing of flat workpieces using rotational-magnetorheological abrasive flow finishing (R-MRAFF) process. The International Journal of Advanced Manufacturing Technology, vol. 62, p. 405-420, D0I:10.1007/s00170-011-3808-2. [11] Singh, A.K., Jha, S., Pandey, P. (2012). Magnetorheological ball end finishing process. Materials and Manufacturing Processes, vol. 27, no. 4, p. 389-394, D0I:10.1080/1042691 4.2011.551911. [12] Shinmura, T., Takajava, K., Hatano, E. (1985). Study on magnetic abrasive process - Application to plane finishing. Bulletin of the Japan Society of Precision Engineering, vol. 19, no. 4, p. 289-291. [13] Jain, V.K., Singh, D.K., Raghuram, V. (2008). Analysis of performance of pulsating flexible magnetic abrasive brush (P-FMAB). Machining Science and Technology, vol. 12, no. 1, p. 53-76, D0I:10.1080/10910340701883538. [14] Singh, A.K., Jha, S., Pandey, P.M. (2011). Design and development of nanofinishing process for 3D surfaces using ball end MR finishing tool. International Journal of Machine Tools & Manufacture, vol. 51, no. 2, p. 142-151, D0I:10.1016/j.ijmachtools.2010.10.002. [15] Singh, A.K., Jha, S., Pandey, P.M. (2015). Performance analysis of ball end magnetorheological finishing process with MR polishing fluid. Materials and Manufacturing Processes, vol. 30, no. 12, p. 1482-1489, D0I:10.1080/10426914.2015 .1019098. [16] Sidpara, A., Jain, V.K. (2012). Theoretical analysis of forces in magnetorheological fluid based finishing process. International Journal of Mechanical Sciences, vol. 56, no. 1, p. 50-59, D0I:10.1016/j.ijmecsci.2012.01.001. [17] Singh, A.K., Jha, S., Pandey, P.M. (2013). Magnetorheological finishing process, mechanism of material removal in ball end magnetorheological finishing process. Wear, vol. 302, no. 1-2, p. 1180-1191, D0I:10.1016/j.wear.2012.11.082. A Study for the Nanofinishing of an EN-31 Workpiece with Pulse DC Power Supply Using Ball-End Magnetorheological Finishing 457 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8,458-466 © 2020 Journal of Mechanical Engineering. All rights reserved. D0l:10.5545/sv-jme.2020.6676 Original Scientific Paper Received for review: 2020-03-19 Received revised form: 2020-06-01 Accepted for publication: 2020-07-01 Comparative Anal^ is of Static and Dji amic Performance of Nonpneumatic Tire with Flek ble Spoke Structure Zifeng Zhang ongxun Fu* ^X uemeng Liang ^X iaoxia Chen -D i Tan Shandong University of Technology, School of Transportation and Vehicle Engineering, China Based on ABAQUS software, the three-dimensional finite element models of195/50R16 radial tire and three kinds of nonpneumatic tire, i.e spoke plate, honeycomb and grid type, are established, and the static and dynamic performance of the four finite element models of tire are analysed. The results show that in the static condition analysis, the ground pressure of the nonpneumatic tire is distributed on both sides of the tread; the stress concentration of the nonpneumatic tire mainly occurs in the grounding area near the support spoke; the deformation area of nonpneumatic tire mainly appears in the area of grounding and adjacent to the grounding spoke; In the analysis of the tire's steady rolling condition, the ground imprint of a nonpneumatic tire is similar to that of a bar rectangle, and the ground pressure distribution is relatively uniform. The static and dynamic characteristics of honeycomb nonpneumatic tire are close to that of a radial tire. The research work of this paper will provide some reference for the structural design and parameter optimization of nonpneumatic tires. Keywords: nonpneumatic tire, ground pressure distribution, load-carrying characteristics, dynamic characteristics Highlights • Three numerical models of nonpneumatic and pneumatic tires are introduced. • The radial stiffness, the grounding performance and the stress and deformation of the spokes in the static characteristics of three nonpneumatic and pneumatic tires are analysed. • The regulation of axial displacement and the change of ground pressure in the dynamic characteristics of three kinds of nonpneumatic tire and pneumatic tire are compared and analysed. • The static and dynamic characteristics of honeycomb nonpneumatic tire are similar to those of pneumatic tire. 1 INTRODUCTION The nonpneumatic tire does not use compressed air, which is an essential part of the pneumatic tire; instead it relies on the wheel's integrated structure. Compared with the pneumatic tire, the nonpneumatic tire has the advantages of being maintenance-free, explosion-proof, puncture-proof and lias a strong bearing capacity. According to China's transportation department, the proportion of traffic accidents caused by pneumatic tires is about 50 % [1], Many scholars in the tire industry analyse and study the structure and bearing capacity of flexible structure nonpneumatic tire [2] to [4], Such tires are safer than pneumatic tires. The most representative is the design of honeycomb structure; the structure of which has been studied for more than 30 years. Honeycomb materials are mainly used in military equipment to absorb high energy impact [5] to [7]; they are also used in for structural designs of high strength and low density [8], In 2008, a honeycomb tire without internal pressure was developed by the Madison Polymer Research Center, Wisconsin, USA [9], In order to improve the strength and vibration performance of the wheel, the honeycomb hexagon structure is mutually supported based on the bionics principle; the optimal design of the structure can improve the strength of the wheel, avoid the problem of tire burst and damage, and the wheel structure also has the advantages of low noise and low friction heat generation. A design schematic of a mechanical elastic safety wheel with a non-inflatable structure was put forward in [10] and [11] Gawwad et al. [12] studied the interaction between tire and ground, and analysed the influence of tire camber on tire performance. Tonuk and Unlusoy [13], Goldstein [14], and Oislii et al. [15] explored the magnitude of the force and the moment in the tire imprint under rolling, steering and braking conditions, as well as the steering stiffness, braking stiffness and return stiffness of the whole tire. Baranowski et al. [16] set up a multistage testing procedure of the tire and proved the accuracy of the test results and the finite element results. Kucewicz et al. [17] evaluated the effect of mesh on finite element results by studying the mesh sensitivity. Baranowski et al. [18] studied the mechanical properties of rubber at different speeds under quasi-static and strong dynamic conditions. Padovan et al. [19], Pldaparti [20] and Ebbott [21] established a three-dimensional model of a pneumatic tire, and explored the mechanical properties of cord-rubber composite and its influence on tire properties through the finite element software. Kim et al. [22] and [23] explored the grounding characteristics of 458 *Corr. Author's Address: Shandong University ofTechnology, Zibo, China, fuhongxun615@163.com Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 a honeycomb nonpneumatic tire under different vertical loads and different honeycomb structures, and compared them with the traditional pneumatic tire under the same working conditions based on ABAQUS software. Kim et al. [24] applied ABAQUS software to study the static grounding characteristics and steady rolling of a honeycomb nonpneumatic tire, and discussed the factors affecting its vibration characteristics. Kim et al. [25] and Ju et al. [26] conducted a parametric study and experimental design of the honeycomb nonpneumatic tire through the finite element analysis, explored the influence of three variables, namely, spoke thickness, spoke angle and shear band thickness, on the rolling resistance of the nonpneumatic tire, and optimized its geometric structure. Song et al. [27] used the finite element method to study the temperature field distribution of tire in a rolling state. Nishiyama et al. [28] developed an algorithm for transforming finite element and discrete element to improve the calculation efficiency. Based on ABAQUS software, this paper makes static and dynamic analysis of radial tire and three kinds of nonpneumatic tires with different structures, the main research contents are as follows: (1) The tire's finite element model is established, including radial pneumatic tire, spoke plate nonpneumatic tire, honeycomb nonpneumatic tire and grid type nonpneumatic tire; (2) Based on ABAQUS/Standard module, the static analysis and comparison of the tire finite element model are carried out, including load-bearing characteristics, grounding imprint, stress and deformation of spoke plate, etc.; (3) Based on the ABAQUS/Standard module, the dynamic performance of the tire finite element model is compared and analysed, including the displacement regulation of the tire centre point and the distribution of the ground pressure during the rolling process. 2 THE ESTABLISHMENT OF TIRE FINITE ELEMENT MODEL 2.1 Pneumatic Tire Reference Model Numerical analysis is an important part of tire research. By setting different material properties, different model structural parameters and different boundary constraints, various characteristics of tire can be effectively studied [29] and [30], Based on ABAQUS software, a three-dimensional finite element model of 195/50R16 radial tire is established, as shown in the left part of Fig. 1 The material properties of different components of radial pneumatic tire are different, so the pneumatic tire should be divided into different zones to give different material properties to the tread. crown belt layer, belt layer#l, belt layer#2, inside liner, sidewall, bead wire and other parts. In this paper, the method of defining the rebar element in the rubber matrix element is used to establish the model. Belt Tread Side wait Inside liner I J Bead filler ffl Bead wire Fig. 1. 3D finite element model of pneumatic tire 2.1.1 Material Characteristics of Pneumatic Tire Rubber material is a kind of hyperelastic material, which lias approximate volume incompressibility and nonlinearity. In order to better describe the mechanical properties of rubber materials in tire, the Neo-Hookean model, the most common molecular statistical constitutive model of rubber materials, is selected in this paper. Table 1 shows the parameters used in the model; the table is provided by China Triangle Tire co., Ltd. The strain energy function of Neo-Hookean model describes the compressible rubber material as follows: W = ^A(lnJ)2+^Ju(/1-3)-JulnJ. (1) The strain energy function of Neo- Hookean model describes the incompressible rubber material as follows: W = Cw{Il- 3). (2) where W is strain energy density, X elongation, J volume ratio before and after deformation ,u material stress class constant, /, the first invariant of principal elongation ratio, and Q material constants. Table 1. Rubber material parameters Rubber material Q) A Density [1e-9T/mm3] Tread 0.5 0.04 1.11 Sidewall 0.6 0.03 1.12 Inside liner 1.5 0.01 1.10 Apex 6.0 0.003 1.10 The two layers of reinforcement embedded in the tread are made of high-strength steel; the elastic Comparative Analysis of Static and Dynamic Performance of Nonpneumatic Tire with Flexible Spoke Structure 459 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 modulus E = 2.1 x 105 Mpa and Poisson's ratio |i= 0.29. The tliickness of the reinforcement layer is 0.5 mm. This steel wire material property is also applicable to the nonpneumatic tire; this material property is no longer described for the nonpneumatic tyre part. 2.2 Structural Parameters and Modeling of the Nonpneumatic Tire The static analysis and dynamic comparative analysis are carried out for the tire. Therefore, in the modelling stage, the control variables are the same for the general structural parameters of the spoke, honeycomb and grid type nonpneumatic tire. Only the number and structure of spokes as supporting parts is different. The unified structural parameters are as follows: The outer diameter is 602 mm, the inner diameter is 390 mm, the tread tliickness is 2 mm, and the spoke tliickness is 5 mm. The three-dimensional finite element models of spoke, honeycomb and grid type nonpneumatic tires are shown in Fig. 2. 2.2.1 Material Characteristics of the Nonpneumatic Tire The nonpneumatic tire mainly consists of the following parts: (1) tread; (2) flexible spokes; (3) two reinforcing layers embedded in the tread. The tread is made of rubber with a tliickness of 12 mm and is located at the outermost layer of the tire. The relationship between the two reinforcing layers and tread position is shown in Fig. 3. Fig. 3. Two reinforcing layers and tread position relationship The tread rubber is dumbbell-shaped, and the upper and lower ends are fixed on the tensile tester, the stress-strain relationship is obtained by the tensile stress-strain tester (Instrument model: WDW-100), as shown in Fig. 4. 20 30 Strain [%] Fig. 4. Material parameters of tread rubber Fig. 2. 3D finite element models; a) spoke, b) grid, and c) honeycomb type of nonpneumatic tires Polyuretliane not has only the high elasticity of rubber, but also high strength of plastic, excellent comprehensive performance, wear-resistant, flame-retardant characteristics, long service life, simple production process, green enviromnental protection and many other characteristics. It is an ideal material for manufacturing tires. The carcass of the tire is made of polyuretliane elastic material. According to the 460 Zhang, Z.F. - Fu, H.X. - Liang, X.M. - Chen, X.X. - Tan, D. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 literature, the relationship between stress and strain is shown in Fig. 5 [31], scheme of four types of tires. Both tires and rims are tied using the TIE command. 100 150 200 250 300 strain [%] Fig. 5. Stress and strain relationships of uniaxial, biaxial and planar tension in nonpneumatic tire 2.3 Finite Element Simulation Settings Hexahedral elements have high accuracy and good convergence, and are thus better for use as solid elements. Therefore, a C3D8R [32] unit is used for ground structures and nonpneumatic tire structures. Among other uses, SFM3D4 units are used for the belt and cord layers in the pneumatic tire, and an M3D4R unit is used for the reinforcement layer in the nonpneumatic tire. The same part structure of the four tires is divided by the same mesh type and mesh size. The number of elements for nonpneumatic and pneumatic tires is controlled between 100,000 and 9) depending on the specific tire structure; the number of elements for the road surface is unified at 8,000. The number of elements and nodes is no longer stated separately. Both static and dynamic analyses were performed using implicit algorithms. The contact properties were set as tangential behaviour and normal behaviour, and the friction coefficient was set as 0.7 using the friction formula of the penalty function. In order to improve the calculation efficiency, the road surface is uniformly set as a rigid body, the ground to be completely fixed, the road surface as the first contact surface, and the tread as the second contact surface. The axial positions of pneumatic and nonpneumatic tires are the loading points. In order to avoid the repetition and jumble of the picture, the finite element model of inflatable tire is only used to replace the finite element model of other three types of nonpneumatic tires in Fig. 6 to show the simulation Fig. 6. Static and dynamic tire finite element simulation scheme 3 NUMERICAL ANALYSIS OF STATIC PERFORMANCE In the static analysis, the main research is the deformation of the tire, so the ground and the hub are set as rigid bodies, and the ground is completely fixed, in order to reduce the calculation amount and shorten the calculation time. The static analysis of pneumatic tire includes three analysis steps: inflation, contact, and loading, while the analysis of nonpneumatic tire only includes two analysis steps: contact and loading. The vertical load of the pneumatic tire and the nonpneumatic tire is 3000 N in the loading analysis step, analysis and research on grounding performance of pneumatic and nonpneumatic tyres. The vertical load (500 N to 6000 N) was applied to compare the stiffness change trend of nonpneumatic tire and pneumatic tire. The large deformation switch is turned on in all analysis steps. 3.1 Comparison of Static Load-carrying Characteristics of Tire 3.1.1 Radial Stiffness Analysis When the internal pressure of the pneumatic tire is 9 kPa, the comparison diagram of the load-carrying characteristic curve of the pneumatic tire and the nonpneumatic tire with the spoke, honeycomb and grid type formats is shown in Fig. 7, and the reference point is the centre point of the tire. In order to fully reflect the authenticity of the numerical analysis. Fig. 8 shows the static load test of the pneumatic tire. The tire is placed on the loading mechanism to apply different radial loads (Instrument model: CSS-88100), laser level instruments are used to ensure excellent level of tires, applied load, get accurate test Comparative Analysis of Static and Dynamic Performance of Nonpneumatic Tire with Flexible Spoke Structure 461 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 6000 Simulated Measured 5000 z 4000 - 3000 « 2000 5 10 15 20 Displacement [mm] Fig. 8. Static ioad test of pneumatic tire Displacement [mm] Fig. 7. Comparison of tire bearing characteristics data. The material properties used for fimte element simulation are entirely consistent with the material properties obtained from the actual tire. The error between the experimental data and the numerical analysis is within 5 %, which is enough to reflect the real situation. It can be seen from the comparison diagram of tire load-bearing characteristic curve: (1) the load-carrying characteristic curve of pneumatic tire, spoke plate and honeycomb nonpneumatic tire is approximate to the first-order function, among them, the radial stiffness of honeycomb nonpneumatic tire is the most similar to that of pneumatic tire, while the radial stiffness of spoke nonpneumatic tire is weaker; therefore, the load-carrying characteristics and vehicle driving comfort of honeycomb nonpneumatic tire are similar to those of pneumatic tire; (2) among the four types of tires, when the radial load is less than 3500 N, the radial stiffness of the grid type nonpneumatic tire is the largest, so the bearing capacity is the best, and its radial stiffness is about 4 times of the radial stiffness of the pneumatic tire, when the vertical load is more than 3500 N, the radial stiffness of the grid type nonpneumatic tire is similar to that of the spoke nonpneumatic tire, but the load-carrying capacity of the grid type nonpneumatic tire is much higher than that of the spoke plate nonpneumatic tire. 3.1.2 Analysis of Grounding Performance Under the action of the rated load of 3000 N, the static deformation of the pneumatic tire, spoke plate, honeycomb and grid type nonpneumatic tire is shown in Fig. 9 It can be seen that the ground pressure distribution of pneumatic tire is similar to the ellipse. and the ground pressure concentration phenomenon appears at the groove position of the ellipse central pattern and the ground pressure distribution presents a large to small distribution from the ellipse centre to the surrounding; The distribution of the ground pressure of the nonpneumatic tire in the form of spoke, honeycomb and grid type is similar to the bar rectangle. From the 3D mapping surface, it can be seen that the two sides are not closed, and the ground pressure of the nonpneumatic tire is concentrated at both sides of the connection between the flexible spoke and the tread. The transverse length of the ground pressure distribution of the spoke, honeycomb and grid type nonpneumatic tire is larger than that of the pneumatic tire, while the longitudinal width of the ground pressure distribution is smaller than that of the pneumatic tire, which improves the axial stability of the vehicle. 3.1.3 Stress and Deformation of Spoke The stress and deformation nephogram of the pneumatic tire, spoke plate, honeycomb and grid type nonpneumatic tire under the rated load of 3000 N are shown in Fig. 10; the high stress of pneumatic tire is mainly concentrated in the ground tread and sidewall area, while the high stress of nonpneumatic tire is mainly concentrated in the spoke plate and tread belt layer. The stress peak value of the spoke nonpneumatic tire are 0.41 and 0.54 times of that of the honeycomb and grid type nonpneumatic tire, respectively, and the stress distribution of the spoke nonpneumatic tire is more uniform than that of the honeycomb and grid type nonpneumatic tire. Where the stress 462 Zhang, Z.F. - Fu, H.X. - Liang, X.M. - Chen, X.X. - Tan, D. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 Fig. 9. Static deformation and grounding mark of a) (inflatable) pneumatic tire, b) spoke plate type nonpneumatic tire, c) honeycomb nonpneumatic tire, and d) grid type nonpneumatic tire under rated load of 3000 N a) b) d) Fig. 10. Stress and deformation nephogram of a) inflatable pneumatic tire, b) spoke plate type nonpneumatic tire, c) honeycomb nonpneumatic tire and d) grid type nonpneumatic tire under rated load of 3000 N Comparative Analysis of Static and Dynamic Performance of Nonpneumatic Tire with Flexible Spoke Structure 463 c) Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 concentration is high, the strain energy density is high, which leads to a decrease of tire fatigue life. 4 DYNAMIC FINITE ELEMENT ANALYSIS OF TIRE Using ABAQUS software to establish the tire steady-state rolling finite element model, this paper mainly explores the pressure change and pressure distribution of the tire ground part and the displacement regulation of the tire centre point when the tire rolls freely under the condition of the rated load of 3000 N and the angular speed of 31.4 rad/s. The finite element model is modelled in a unified coordinate system. In the inflation analysis step, fix the central point of the pneumatic tire and apply (I MPa pressure on the inner surface of the tire to simulate the inflation condition. Grid division is the same as the type of grid division in static analysis, and it will not be described in detail. 4.1 Displacement Regulation of Tire Center Point The displacement regulation of tire centre point is of great significance to vehicle driving comfort and vehicle structural stability. Fig. 11 shows the change rule of the displacement of the central point of the pneumatic tire and the nonpneumatic tire with time. Fig. 11. Center point displacement curve of tire The displacement curve of the central point of the pneumatic tire is under the inflation condition in the period of 0 s to 1 s, the tire centre point is fixed; 1 s to 2 s is the loading step, and the tire centre point moves down; 2 s to 2.25 s is the acceleration step, and the angular velocity at the tire centre point increases from 0 rad/s to 31.4 rad/s; the period of 2.25 s to 2.5 s is steady rolling condition. The displacement curve of the centre point of the nonpneumatic tire in the period of 0 s to 1 s is the loading step, and the centre point of the tire moves down; The period of 1 s to 1.2 s is the acceleration step, the tire angular velocity increases from 0 rad/s to 31.4 rad/s, and the displacement curve fluctuates; The period of 12 s to 2 s is steady rolling stage, and the vertical displacement of tire centre point changes periodically. Under the condition of pure rolling, the angular acceleration, m of the tire is 0 and the tangent speed, v of the centre of the tread is the forward driving speed. i of the vehicle, the equation is as follows: v = reG) o =v0. (3) Fig. 12 shows the three radii of the tire. The vertical load has a direct influence on effective rolling radius and load radius. Pacejka [33] proposed a more comprehensive interpretation of the effective rolling radius of the tire. Fig. 12. The rolling radius of the tire; Re effective rolling radius, Ri radius of tire under load, Rf radius of tire under no load; a>n pure rolling angular velocity Both the pneumatic tire and nonpneumatic tire are defonnable bodies with hysteresis, under the influence of inertia force and centrifugal force in the rolling process, the radius of each particle relative to the tire axis is different, so the tire presents the curve characteristics of wave shape in the rolling process. It can be seen from the comparison of the displacement regulation of the centre point of the nonpneumatic tire and the pneumatic tire in Fig. 9 that the fluctuation amplitude of honeycomb nonpneumatic tire is the closest to that of pneumatic tire. Among them, the honeycomb nonpneumatic tire is most similar to the pneumatic tire in driving comfort and structural stability during steady rolling. Under the action of a 3000 N radial load, the radial sinkage of the spoke type nonpneumatic tire is too large, and 464 Zhang, Z.F. - Fu, H.X. - Liang, X.M. - Chen, X.X. - Tan, D. Strojniski vestnik - Journal of Mechanical Engineering 66(2020)7-8, 458-466 its driving comfort and vehicle structure stability in the steady rolling process are worse than those of the pneumatic tire, so it is not suitable for the vehicle with larger curb weight. In the steady-state driving process under the action of 3000 N radial load, the amplitude of the grid type nonpneumatic tire changes too much, but the sinking of the grid type nonpneumatic tire is the smallest compared with the pneumatic tire, the spoke nonpneumatic tire and the honeycomb nonpneumatic tire; therefore, it can be concluded that the structure of the grid type nonpneumatic tire needs to be optimized, the number of nonpneumatic tire units needs to be increased, and the amplitude of the nonpneumatic tire can be reduced. To increase the vehicle's driving comfort and tire grip by increasing the sinking of grid type nonpneumatic tire, the material stiffness must be improved. 5 CONCLUSIONS Based on ABAQUS software, 3D finite element models of 195/50R16 radial pneumatic tire, spoke plate type nonpneumatic tire, honeycomb nonpneumatic tire and grid type nonpneumatic tire are established, and their static and dynamic analysis are carried out respectively. 1. Under the action of the rated load of 3000 N, the grounding area of the pneumatic tire has axisymmetric deformation, and the deformation area is mainly distributed in the grounding part and sidewall; Because the rubber material of the sidewall of the radial pneumatic tire is thin, it is consistent with the actual situation. 2 The deformation area of the nonpneumatic tire is mainly distributed in the grounding part and the spokes near the grounding part. The grounding imprint of the pneumatic tire is similar to the bar shape rectangle, and the grounding part near the spoke plate has the phenomenon of ground pressure concentration; when the radial load is less than 3500 N, the load-carrying capacity of the grid type nonpneumatic tire is the best; when the radial load is more than 3500 N, the load capacity of honeycomb nonpneumatic tire is the best. When the polyurethane support structure of the nonpneumatic tire uses the same material model, the radial stiffness of the spoke plate nonpneumatic tire is too small, which is not suitable for vehicles with too large curb weight; however, the radial stiffness of the grid type nonpneumatic tire is too large, which is suitable for vehicles with too large curb weight; The load-carrying characteristics of the honeycomb nonpneumatic tire and radial pneumatic tire are similar. 3. When the tire is rolling, the displacement curve of the tire centre point fluctuates during acceleration, and the vertical displacement of the tire centre point changes periodically during steady rolling; the overall performance of grid type nonpneumatic tire is not ideal, so its structure needs to be further optimized. 4 In the analysis of tires' dynamic characteristics, the amplitude of their axis position not only has an inevitable relationship with the structure of tires but also has a direct correlation with the vertical load; therefore, it is necessary to reasonably choose the style and type of inflatable tire and nonpneumatic tires according to the requirements of ride comfort and vertical load conditions. 6 ACKNOWLEDGEMENTS This work was supported by the National Natural Science Foundation of China under Grant 51775320 and project ZR2018PEE008 supported by Shandong Provincial Natural Science Foundation. 7 REFERENCES [1] Zhao, Y.Q., Zang, L.G., Chen, Y.Q., LI, B„ Wang J. 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Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8,467-478 © 2020 Journal of Mechanical Engineering. All rights reserved. D0l:10.5545/sv-jme.2020.6784 Original Scientific Paper Received for review: 2020-03-22 Received revised form: 2020-06-12 Accepted for publication: 2020-07-01 3D FE Modelling of Machining Forces during AISI410 H ard Turning Anastasios TzotzisJ* -C ésar García-Hernández1 4 osé-Luis Huertas-Taln 1 -P anagiotis Kyratsis2 1 University of Zaragoza, Department of Design and Manufacturing Engineering, Spain ^University of Western Macedonia, Department of Product and Systems Design Engineering, Greece Hard turning is one of the most used machining processes in industrial applications. This paper researches critical aspects that influence the machining process of AISI 4140 to develop a prediction model for the resultant machining force-induced during AISI 4140 hard turning, based on finite element (FE) modelling. A total of 27 turning simulation runs were carried out in order to investigate the relationship between three key parameters (cutting speed, feed rate, and depth of cut) and their effect on machining force components. The acquired numerical results were compared to experimental ones for verification purposes. Additionally, a mathematical model was established according to statistical methodologies such as the response surface methodology (RSM) and the analysis of variance (ANOVA). The plurality of the simulationsyielded results in high conformity with the experimental values of the main machining force and its components. Specifically, the resultant cutting force agreement exceeded 90 % in many tests. Moreover, the verification of the adequacy of the statistical modelled to an accuracy of 8.8 %. Keywords: AISI 4140 turning, machining forces, 3-dimensional finite element modelling, response surface methodology Highlights • A3DFE model has been established for the hard turning of AISI 4140. • In addition, a statistically based prediction model for the resultant machining force has been developed. • Through RSM and ANOVA, the effect of the cutting parameters on the machining force components and their relationship were investigated. • The numerical results were in complete accordance with the experimental ones; the relative error was estimated within the range of-10 %tol2% for most of the cases. • The prediction model offered an accuracy of 8.8 %. 0 INTRODUCTION Hard turning is a cost-effective machining operation since it can reduce finish grinding of parts and, in some cases, eliminate it. The term "hard turning" refers to the turning operation of hardened steel with hardness between 58 and 62 HRC [1]. Additionally, hardened steel is an essential work material in industrial applications. Therefore, many researchers developed an increased interest in the investigation of hard turning and similar machining processes [2] to [6], One of the latest advances in machining studies is the implementation of the finite element method (FEM) with the aid of specialized software. In recent years, the B orthogonal cutting finite element (FE) model have proved to be a valuable tool for many researchers. Klocke et al. [7] simulated the high-speed orthogonal turning of AISI 1045 steel using commercial software. Yen et al. [8] developed a methodology to predict the tool wear evolution and tool life in orthogonal cutting using FEM simulations. Arrazola et al. [9] employed a 2D FE model with the use of arbitrary Langrangian Eulerian (ALE) formulation for prediction purposes of the serrated chip during AISI ® machining. Similar studies have been conducted to examine the generated chip morphology and cutting forces via D modelling for homologous materials [10] to [12], but this model has several restrictions that limit its field of application. In contrast, the constant advancement of computational resources resulted in the emergence of 3D modelling, which does not have the limitations of D models. Therefore, FE modelling in three dimensions can help researchers to study several aspects of machining at a greater extent. Tool wear analysis and temperature distribution on the cutting tip of tools has been studied extensively [13] to [15], as well as the prediction of the cutting forces and residual stresses, and the optimization of the machining conditions [16] to [18], Three-dimensional modelling is also used for investigation purposes on the turning of hardened steels; in most cases, the results are experimentally validated. Guo and Liu [19] established a geometric model and a general practical, explicit 3D FE model to analyse the hard turning of AISI 52100; the model predictions provide reasonable accuracy for several cutting results. Later, Ozel et al. [20] utilized 3D FE modelling to predict chip formation, forces, temperatures, and tool wear during hard turning of AISI 4130 with poly crystalline *Corr. Author's Address: University of Zaragoza, Campus Rio Ebro, C/ Maria de Luna 3,50018, Zaragoza, Spain, atzotzis@unizar.es 467 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 467-478 cubic boron nitride (PcBN) tools. Lian et al. [21] proposed a structural model for soft/hard composite-coated textured (SHCCT) tools and confirmed it with a three-dimensional numerical simulation. The proposed model was applied to AISI 1045 hard turning and was analysed via orthogonal experiments for different coating thickness, material, and ratios of the soft/hard coatings. Magalhaes et al. [22] aimed to provide a better understanding of the mechanical and thermal loads involved in cutting, with respect to the variation of the tool's edge discretization. To do so, they prepared numerical simulations of AISI 5115 steel hard turning using finite element analysis (FEA). The advent of more advanced inserts, such as PcBN, polycrystalline diamond (PCD), chemical vapour deposition (CVD) and ceramics, lead to more efficient machining and overall to better quality finished parts. In particular, ceramic tools are preferred when machining hardened steels; therefore, the analysis of hard turning with ceramics and the effects of various conditions is a research area that can benefit from the implementation of FEM. Hu and Huang [1] and [14] researched the influence of tool shape and cutting angles on the contact stresses, sliding speed, and temperature with the aid of 3D FEM and experimental testing; they also established a new type of tool life model for nano-ceramic tools, which includes several parameters. Moreover, they studied the effects of cutting speed on the high-speed turning of AISI 1013 with ceramics by using similar methodologies. An adequate number of solely experimental research studies can be found in the literature, related to the hard turning of steel; however, the implementation of FEM in such studies remains limited, especially during the investigation of the modern standardized turning inserts. In the present paper, the components of the turning force-induced during the hard turning of AISI ® is examined with respect to several combinations of cutting speed, feed, and depth of cut. The study lias been carried out with the aid of a commercially available finite model analysis (FEA) software (DEFORM3D™). In addition to the established FE model, a prediction model for the main machining force based on statistical methods has been developed. Furthermore, both the FE model and the statistical prediction model were validated via comparison with equivalent experimental results that are available in the literature [23], 1 METHODOLOGIES 1.1 Experimental Layout The experimental values [23] used in this study were acquired with the aid of a three-component dynamometer, (Kistler 9257B) and a standard data acquisition system which includes a charge amplifier, a data acquisition card (A/D2855A3), and the appropriate software (DynoWare 2825A1-1). The turning experiments were performed with the aid of a universal lathe type SN , are the vectors that contain the regression coefficients. Y = b0 + b2X2 +b3X3 +b4X{ +b5 X; +b6 X: + b-j X \X 2 + \XxX3 + b9X2X3. (4) Eq. (5) presents the complete statistical model for the resultant machining force based on the aforementioned formula and the data of the verified FE model (see Table 3). Fmam = -112.2 + 0.566V + 820/ + 422. lap -0.00074F2 -1860f2 +125ap2 -0.671 J +0.7081 a/? + 3124 fap, (5) where Fmain is the resultant machining force in N, V is the cutting speed in m/min,/is the feed in mm/rev and ap is the depth of cut in mm. The design of experiments (see Table 3) contains the estimation of the resultant cutting force for all 2 combinations of machining parameters, derived from the experiments [23], the simulations and the statistical model. The comparison of these results shows an increased correlation. Specifically, the highest level of agreement between the numerical values and the experimental ones is observed in the test number 26 (relative error 0.2 %), whereas the lowest in test number 19 (relative error 12.4 %). Between the values obtained from the regression model and the numerical values, test number $ yielded an agreement of almost 100 %. In contrast, the lowest level of accordance (91.2 %) was found to be in the first test. Eventually, for the comparison case between the statistical values and the experimental ones, a high correlation is highlighted with a mean absolute percentage error of approximately 4o %. Furthermore, the best level of agreement (99.6 %) was achieved in the second test, whereas the worst (86.6 %) in the first one. By observing Table 3 and the charts of Fig. 4, the following statements can be made for the FE model: The radial force is the dominant of the three components. Higher values of feed rate affect all forces, but the tangential force is affected the most due to the increase of the sheared chip region. The depth of cut has a strong impact on all machining forces as anticipated; as the tool cuts deeper in the material, the tool-workpiece contact length increases. For instance, the main cutting force increases approximately 43.6 % (from 3D FE Modelling of Machining Forces during AISI 4140 Hard Turning 473 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 467-478 Table 3. Main machining force comparison between experimental, simulated and statistical values Cutting parameters ^main Vc [m/min] /[mm/rev] ap [mm] Experiments FE model Regression model 1 80 0.08 0.10 59.9 56.9 51.9 2 115 0.08 0.10 67.5 69.0 73.6 3 150 0.08 0.10 76.0 79.5 92.0 4 80 0.11 0.10 76.3 73.9 67.2 5 115 0.11 0.10 91.5 89.7 88.3 6 150 0.11 0.10 92.4 99.6 106.0 7 80 0.14 0.10 94.2 88.4 80.8 8 115 0.14 0.10 97.5 102.2 101.1 9 150 0.14 0.10 121.3 119.5 118.1 10 80 0.08 0.20 116.1 118.2 128.5 11 115 0.08 0.20 134.3 142.8 159.6 12 150 0.08 0.20 152.2 163.1 187.4 13 80 0.11 0.20 154.2 164.1 146.3 14 115 0.11 0.20 172.2 178.4 176.8 15 150 0.11 0.20 186.5 189.2 203.8 16 80 0.14 0.20 177.3 189.7 162.4 17 115 0.14 0.20 208.6 206.1 192.1 18 150 0.14 0.20 227.3 223.4 218.5 19 80 0.08 0.30 189.8 207.7 207.6 20 115 0.08 0.20 225.7 232.1 248.1 21 150 0.08 0.30 238.0 249.6 285.3 22 80 0.11 0.30 244.6 250.2 227.9 23 115 0.11 0.30 264.1 266.2 267.7 24 150 0.11 0.30 267.7 281.2 304.2 25 80 0.14 0.30 282.3 284.5 246.4 26 115 0.14 0.30 300.9 301.5 285.5 27 150 0.14 0.30 316.9 320.7 321.3 223.4 N to 320.7 N) for the same conditions (Vc = 150 m/min,/= 0.14 mm/rev) and an increase in depth of cut from 0.20 mm to 0.30 mm. Lastly, as cutting speed rises, the turning forces decrease in most cases, so does the main cutting force. For example, at ap = 0.30 mm, f = (4 mm/rev and Vc =150 m/min 115 m/min and 80 m/min the resultant force is equal to 320.7 N, 301.5 N and 284.5 N respectively. The primary reason for this tendency is that an increase in temperature at the shear plane region, resulting in the plastic softening of the deformation zone which ultimately lowers the shear strength of the material. Fig. 5 illustrates the relative error percentage between the values of resultant machining force derived from the regression model and the experiments, as well as between the regression model and the simulations. The graph indicates that both lines follow a similar trend with the exception of two points (tests number 6 and 19) where more abrupt increase in experimental values occurred. 15 15 1 5 10 15 20 25 27 Test No experiment vs regression ^^ simulation vs regression Fig. 5. Relative error comparison between simulations and experiments Moreover, the maximum error was found to be -13.4 % and -8.8 % for the regression versus 474 Tzotz/s, A. - Garcia-Hemandez, C. - HuertasrTaldn, J.L. - Kyrats/s, P. Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 467-478 experiments and the regression versus simulations, respectively (test number 1 for both cases). In contrast, the lowest value of error was determined to be -0.4 % (test 2) for the regression versus experiments and close to zero (test 19) for the regression versus simulations. Finally, the mean absolute percentage error was estimated 2 % for the regression versus simulations case and 45 % for the regression versus experiments case. 2.3 Validation of the Statistical Model Due to the number of independent variables taken into account in current research the validity of the fit was analysed with ANOVA. A standard confidence level of 95 % was used for all intervals throughout the analysis, which revealed a successful fit of the model with an adjusted R-squared of 99.72 %. Furthermore, according to the significance level of 0.05 and to Table 4 the terms that contribute the most to the model are the ap and the fxap with a /»-value equal to CD as well as the constant with p = 0.005. Last but not least, I 'xap and / have great impact on the model with /»-values of 0.060 and 0.089 respectively, even though are higher than 0.05. The sum of squares and the degrees of freedom for the analysis are included in Table 4. With the total sum of squares which is the sum of squared deviations due to each of the nine factors and the sum of squares attributed to the error, it is possible to determine the dispersion of data points. In addition, the mean square is the ratio of the sum of squares to the degree of freedom and the /-value is the ratio of the mean square of the regression model to the mean square of residual error. Lastly, the fact that the /»-value of the regression was estimated 0.000 indicates the very high correlation of the model and eliminates the probability of yielding unusual results. With the validation of goodness of fit, a residual analysis was performed to check the accuracy of the model. The graphs that are illustrated in Fig. 6 proves that the model has very good accuracy. In particular, the normal probability plot (Fig. 6a) shows a normality in the distribution of the residuals with no serious departures from the straight line. Additionally, the residuals versus the fitted values (Fig. 6b) indicate a constant variance of the residuals since they are almost evenly scattered on both sides of the reference line. The overall normality is present in the residuals versus the order graph also (Fig. 6d). It is observed that there are no systematic faults, and the residuals are independent of one another. Eventually, the normality in the distribution of the error percentages can be displayed in the error histogram (Fig. 6c) and proved by the fit line. The analysis of the developed prediction model was carried out with the aid of 3D response surface plots for visualizing the data gathered from Fig. 4 and Table 3. That is, the combined effect of the machining conditions and the depth of cut on the generated radial, tangential and feed forces were investigated. Fig. 7 illustrates the plotted 3D surfaces for each depth of cut value based on the polynomial solutions. The cutting speed and the feed are the input parameters of the polynomial with values within the investigated range. Table 4. ANOVA results for the main machining force Source Degree of freedom Sum of squares Mean square /value /»-value Regression 9 165668 18407.6 1012.4 0.000 Residual error 17 309 18.2 Total 26 165977 R-sq (adj) = 99.72 % Term PE Coefficient SE Coefficient /value /»-value Constant -112.8 35.1 -3.19 0.005 V 0.566 0.359 1.58 0.134 f 820 455 1.80 0.089 ap 422.1 92.9 4.55 0.000 r- -0.00074 0.00142 -0.52 0.609 / -1860 1934 -0.96 0.350 ap2 125 174 0.72 0.482 Vxf -0.67 1.17 -0.57 0.574 Vxap 0.708 0.352 2.01 0.060 fxap 3124 410 7.61 0.000 3D FE Modelling of Machining Forces during AISI 4140 Hard Turning 475 Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 467-478 Fig. 6. Residual analysisgraphs: a) probability plot, b) residuals versus fitted values, c) error histogram, and d) residuals versus order hence 80 m/min to 150 m/inin for cutting speed (step of 10 m/min) and 0.08 mm/rev to 0.14 mm/rev for feed (step of 0.01 mm/rev). According to Fig. 7 it is concluded that: The depth of cut affects the resultant machining force significantly; as the depth of cut increases, so does the force. Similarly, higher values of feed have a great impact on the main machining force. Eventually, even though higher cutting speeds result in lowering the machining force, the effect is limited. Fig. 7. 3D plots of the Fmai„ for each depth of cut Conclusively, six extra simulation runs were accomplished to further validate the prediction model of Fmain by utilizing randomly selected conditions 476 from within the range of the data employed in this study, forming the following sets: I, II and III with I = (I m/min,/= (2 mm/rev and ap = 0.15 mm. 0.25 mm and 0.30 mm respectively. IV, V and VI with Vc = 130 m/min,/= (8 mm/rev and ap = 0.15 mm, 0.25 mm and 0.30 mm respectively. The results are presented in Table 5, in which it is highlighted that the relative error is low for all cases. Table 5. Confirmation of prediction model for F, Set Simulated Predicted Relative error Fmain Emain [%] 1 147.8 133.6 -9.61 II 239.3 225.3 -5.85 III 260.4 272.2 4.53 IV 140.5 130.4 -7.19 V 233.8 218.0 -6.76 VI 248.9 262.8 5.58 3 CONCLUSIONS In this study, the development of a 3D FE model, as well as a prediction model for the main machining force induced during hard turning of AISI 9 were presented. A series of 27 3D simulations were conducted under different conditions of cutting speed and feed in addition to the three different depths of cut. The obtained numerical results were validated via experimental values that are available in the literature. Tzotz/s, A. - Garcia-Hemandez, C. - HuertasrTaldn, J.L. - Kyrats/s, P. Strojnlskl vestnik - Journal of Mechanical Engineering 66(2020)7-8, 467-478 and it was observed that are in high agreement that surpasses 90 % in most of the runs. The accuracy (8.8 %) and goodness of fit of the statistical model, dictate that both the developed models (FE and statistical) can securely predict the resultant machining forces when applied within the scope of this study. In conclusion, the following remarks are pointed out: Higher values of depth of cut and feed rate significantly increase machining forces, especially the depth of cut is the factor that effects Fmam the most. Specifically, according to the simulated values of Fmain (see Table 3), an average increase of about ® % in the resultant cutting force is observed when the depth of cut changes from (8 mm to CD mm. The equivalent shift from 0.20 mm to 0.30 mm amplifies Fmain by approximately 50 %. Similarly, when feed changes from 0.08 mm/ rev to 01 1 mm/rev and from (1 1 mm/rev to Gl mm/rev, the resultant cutting force gains an increase of about Wo and 6 % respectively. 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Strojniški vestnik- Journal of Mechanical Engineering 66(2020)7-8 Vsebina Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 66, (2020), številka 7 8 Ljubljana, julij-avgust 2020 ISSN 0039-2480 Izhaja mesečno Razširjeni povzetki (extended abstracts) Alexandra Aulova, Marko Bek, Leonid Kossovich, Igor Emri: Brezigelno elektropredenje vlaken PA6: vpliv koncentracije raztopine in električne napetosti na premer vlaken SI 53 Dalibor Petrovič, Maijan Dodič, Nenad Kapor: Nova konstrukcijska rešitev letalskih koles za znižanje nadtlaka v pnevmatiki, ki ohranja absorpcijsko zmogljivost in dimenzije kolesa SI 54 Renata Mola, Tomasz Bučki: Karakterizacija območja spoja pri sestavljenih bimetalnih ulitkih AZ91/ A1SÍ12, izdelanih iz nemodificirane in toplotno modificirane zlitine A1SÍ12 SI 55 Himmat Singh, M. S. Niranjan, Reeta Wattal: Študija nanoobdelave dela iz jekla EN-31 po postopku magnetoreološke obdelave z okroglo delovno površino in impulznim enosmernim napajalnikom SI 56 Zifeng Zhang, Hongxun Fu, Xuemeng Liang, Xiaoxia Chen, Di Tan: Primeijalna analiza statičnih in dinamičnih lastnosti nepnevmatskih koles z upogljivimi naperami SI 57 Anastasios Tzotzis, César García-Hernández, José-Luis Huertas-Taln, Panagiotis Kyratsis: 3D-modeliranje sil pri trdem struženju jekla AISI4140 po metodi končnih elementov SI 58 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 53 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-04-09 Prejeto popravljeno: 2020-05-27 Odobreno za objavo: 2020-06-01 Brezigelno elektropredenje vlaken PA6: vpliv koncentracije raztopine in električne napetosti na premer vlaken Alexandra AulovaJ* -M arko Bck1 -L eonid Kossovich2 4 gorEmri13 1 Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija 2 Državna univerza Saratov, Ruska federacija 3 Univerza v Zlinu Tomas Bata, Center za raziskave obutve, Češka republika Elektropredenje je proces proizvodnje polimernih vlaken in membran iz polimerne raztopine ali taline ob prisotnosti elektrostatičnih sil. Tako narejeni materiali se uporabljajo v številnih aplikacijah: tekstil, izolacijski materiali, tkivni odri, obliži, filtri in ločilne membrane. Brezigelni postopek elektropredenja je sorazmerno nov izdelovalni postopek, pri katerem ne potrebujemo igle ali šobe za uvajanje raztopine polimera v elektrostatično polje. Namesto tega pri tem postopku uporabimo cilinder ali žico, ki se vrtita in sta delno potopljena v polimerno raztopino. Tako dobimo površino, prekrito z raztopino, ali površino, na kateri so posamezne kapljice raztopine. Tvorba vlaken poteka od spodnje vrteče se elektrode do zgornje sprejemne elektrode. V primeijavi s klasično izdelovalno metodo brezigelni postopek omogoča izdelavo večje količine materiala v enakem času. Kljub temu pa zaradi nekoliko drugačne in kompleksnejše fizike brezigelni izdelovani postopek ni tako dobro raziskan in razumljen. Funkcionalnost materialov, narejenih z elektropredenjem, je opredeljena z materialom, s fizikalno-kemijskimi lastnostmi površine vlaken, premerom vlaken in s porazdelitvijo velikosti por med vlakni. Premer vlaken določa prosto površino membrane, medtem ko porazdelitev velikosti por membrane, ki je odvisna od njega in debeline membrane, opredeljuje barierne lastnosti in prepustnost materiala. Premer in porazdelitev velikosti vlaken sta močno odvisna od lastnosti raztopine/taline polimera in pogojev elektropredenja. Nadzor premera vlaken med elektropredenjem omogoča nastanek membran z izbrano porazdelitvijo velikosti por in je zaradi tega ključen del izdelovalnega procesa. V okviru tega dela smo preiskovali brezigelni postopek elektropredenja poliamida 6 (PA6), raztopljenega v mešanici ocetne in mravljične kisline. Zaradi visoke mehanske trdnosti v mokrem in suhem stanju, hidrofilnosti in kemične odpornosti PA6 se uporablja v različnih aplikacijah, vključno z avtomobilsko industrijo, s tekstilom, z medicino in s tehniko. Poleg tega se lahko v material PA6 vključi protibakterijska ali protiglivična sredstva in se zaradi tega pogosto uporablja tudi za elektropredenje za namen filtracije, medicinskih aplikacij in tekstila. Glede na to, da gre pri brezigelni tehnologiji elektropredenja za novejši izdelovalni postopek, je na voljo sorazmerno malo informacij o procesnih pogojih, ki so bistvenega pomena za izdelavo vlaken z izbranimi lastnostmi. V sklopu te raziskave smo načrtno preiskali vpliv koncentracije polimerne raztopine na viskoznost, površinsko napetost in na električno prevodnost. Raziskali smo tudi, kako ti parametri skupaj z električno napetostjo pri postopku izdelave vlaken vplivajo na tvorbo vlaken PA6 in njihov premer. Nadzorovanje izdelovalnih parametrov in izdelava vlaken z izbranim premerom je prvi korak, ki omogoča izdelavo materialov z izbrano porazdelitvijo velikosti por v membrani. Ugotovili smo, da koncentracija polimerne raztopine pa tudi električna napetost, ki ji je raztopina izpostavljena, vplivata na premer vlaken, kar je skladno z obstoječimi raziskavami. Poleg tega pa smo ugotovili še, daje bolj grobe popravke premera vlaken lažje doseči s spremembami koncentracije polimerne raztopine, medtem ko je s spremembo napetosti elektropredenja mogoče doseči drobne prilagoditve premera vlaken. Ugotovitve tega prispevka predstavljajo osnovo za naše nadaljnje delo, ki bo vključevalo preiskavo vpliva premera in debeline vlaken na porazdelitev por membrane, ter učinkovitost filtriranje takšnih membran. Ključne besede: brezigelno elektropredenje, nanospider, PA6, premer vlaken, koncentracija raztopine, električna napetost *Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za strojništvo, Aškerčeva 6,1000 Ljubljana, Slovenija, alexandra.aulova@fs.unMj.si SI 53 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 54 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-02-07 Prejeto popravljeno: 2020-05-15 Odobreno za objavo: 2020-05-28 Nova konstrukcijska rešitev letalskih koles za znižanje nadtlaka v pnevmatiki, ki ohranja absorpcijsko zmogljivost in dimenzije kolesa Dalibor Petrovič1 - Maijan Dodič1 -N enad Kapor? 1 Univerza za obrambo, Vojaška akademija, Srbija, 2 Univerza Megatrend, Fakulteta za civilno letalstvo, Srbija Letalske pnevmatike nosijo težo letala le z nadtlakom, njihova konstrukcija pa se v zgodovini ni bistveno spreminjala. Čeprav je znano, da delovni tlak vpliva na življenjsko dobo pnevmatik, do zdaj še ni bil obravnavan problem nižanja nadtlaka v pnevmatiki za enako nosilnost in ohranitev geometrijskih lastnosti. K reševanju tega problema je težko pristopiti s standardno konstrukcijo kolesa, saj znižanje nadtlaka zahteva povečanje pnevmatike. Večje pnevmatike pa imajo več slabosti in najpomembnejša med njimi (poleg povečanja teže) je prostor, ki ga zasedajo. Glede na dejstvo, da znižanje tlaka podaljša življenjsko dobo pnevmatik, se postavlja vprašanje, kako bi bilo to mogoče izkoristiti brez spreminjanja obstoječe geometrije. Pnevmatika se med trdim pristankom letala močno deformira, tj. njeno površino potegne navznoter tako v prečnem kot v vzdolžnem prerezu. V tem primeru celotno težo letala nosi rama pnevmatike in dejanska stična površina je majhna. Raziskovalci so se v zadnjem času posvečali analizi standardne oblike pnevmatik in njenim izboljšavam. Alternativne konstrukcije koles glavnega pristajalnega mehanizma, ki bi omogočile izboljšanje lastnosti pnevmatik, nasprotno niso bile predmet raziskav. Teža je najpomembnejši omejujoči dejavnik v letalstvu in zamisel o prilagoditvi sestava letalskega kolesa za zmanjšanje pritiska in povečanje absorpcije kinetične energije ob ohranitvi dimenzij pnevmatike je na prvi pogled protislovna in neizvedljiva. To je veijetno tudi razlog za to, da v literaturi ni mogoče najti raziskav v tej smeri. V pričujočem članku je predstavljena nova konstrukcijska rešitev, kije sestavljena iz standardne (zunanje) pnevmatike in manjše (notranje) pnevmatike, ki je vstavljena v standardni pnevmatiki. Par pnevmatik tvori edinstven sistem za absorpcijo kinetične energije ob stiku letala s pristajalno stezo. Manjša (notranja) pnevmatika je vgrajena v notranjem torusu zunanje pnevmatike in pozicionirana tako, daje njen zunanji torus na določeni razdalji od vrha notranjega torusa zunanje pnevmatike. Notranja pnevmatika preprečuje, da bi standardno pnevmatiko potegnilo navznoter in zagotavlja njen stik s pristajalno stezo, ko je pnevmatika že deformirana. Ta razdalja je odvisna od deformacije zunanje pnevmatike, tj. zunanji torus manjše (notranje) pnevmatike je na razdalji, kjer začne standardno (zunanjo) pnevmatiko vleči navznoter. Položaj notranje pnevmatike tako preprečuje deformacijo zunanje pnevmatike navznoter in jo sili k ekspanziji vstran. Na ta način se poveča širina kontaktne površine zunanje pnevmatike. Manjša (notranja) pnevmatika tudi sama sprejema kinetično energijo in nadtlak v zunanji pnevmatiki je zato za 50 % manjši v primeijavi s standardno vrednostjo. Ta pristop odstopa od obstoječih študij v literaturi, ki za podaljšanje življenjske dobe predlagajo zasuk kolesa pred stikom s pristajalno stezo. Zmanjšanje nadtlaka v pnevmatiki pripomore k daljši življenjski dobi pnevmatik in površine pristajalne steze. Numerična analiza predlagane konstrukcijske rešitve je bila opravljena kot implicitna simulacija v programskem paketu ANSYS. Implicitna numerična analiza letalske pnevmatike upošteva definicijo stanj in obremenitev, definicijo geometrije in definicijo materiala. Opravljena je bila primeijalna analiza sprememb odklona, širine kontaktne površine in porazdelitve napetosti za standardno pnevmatiko in za par novih pnevmatik. Rezultati primeijalne študije so pokazali, da nova tehnična rešitev prinaša za 8 % večjo kontaktno širino pnevmatike in s tem večjo površino stika med pnevmatiko in pristajalno stezo. Napetost v pnevmatiki pri paru novih pnevmatik je zato za približno 40 % manjša kot pri standardni pnevmatiki, s tem pa je zagotovljena boljša vzdržljivost nove pnevmatike ter za 56 % večja zmožnost absorpcije kinetične energije kot pri standardni pnevmatiki. Rezultati jasno razkrivajo potencial predlagane konstrukcijske rešitve za transportna letala. Ključne besede: letalska pnevmatika, metoda končnih elementov, konstrukcija pnevmatike, kinetična energija, absorpcijski sistem, hiperelastičen material SI 54 *Naslov avtorja za dopisovanje: Megatrend univerza. Fakulteta za civilno letalstvo, Srbija, nenad.kapor@gmail.com Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 55 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-02-07 Prejeto popravljeno: 2020-05-15 Odobreno za objavo: 2020-05-28 Karakterizacija območja spoja pri sestavljenih bimetalnih ulitkih AZ91 AlSil, i zdelanih iz nemodificirane in toplotno modificirane zlitine AlSil Renata Mola - Tomasz Bučki* Tehniška univerza Kielce, Fakulteta za mehatroniko in strojništvo, Poljska Sestavljeno oz. zloženo litje je preprosta in učinkovita metoda za spajanje dveh raznorodnih kovin oz. zlitin pri izdelavi bimetalnih delov, pogosto kompleksnih oblik ter brez geometrijskih ali dimenzijskih omejitev. V zadnjih letih je bil zabeležen velik napredek na področju spajanja magnezijevih in aluminijevih zlitin. Tovrstni lahki bimetali se uporabljajo predvsem v avtomobilski industriji za zmanjšanje mase vozil in s tem posledično porabe goriva. Za spajanje magnezijevih in aluminijevih zlitin je značilno oblikovanje trdih in krhkih intennetalnih spojin. Mg/Al zlitine imajo značilno slabše mehanske lastnosti, zato se izvajajo obsežne študije za modifikacijo mikrostrukture v območju spoja bimetalov Mg/Al in izboljšanje lastnosti spojev. Namen pričujoče študije je raziskava vpliva toplotne modifikacije vložka iz aluminijeve zlitine na mikrostrukturo, sestavo in mikrotrdoto v območju bimetalnega spoja magnezijeve zlitine AZ91 in aluminijeve zlitine AlSil2, izdelanega po postopku sestavljenega litja tekoče na trdno. V eksperimentih je bila tekoča zlitina AZ91, ogreta na temperaturo 660 °C, nalita na trden vložek iz zlitine AlSil2 v jekleni formi. Vložek in forma sta bila predhodno ogreta na temperaturo 300 °C. V študiji so bili uporabljeni vložki iz nemodificirane in toplotno modificirane zlitine. Nemodificirani vložki so bili odrezani od ingota iz materiala AlSil2, modificirani vložki pa so bili pripravljeni z nalivanjem zlitine AlSil2 na debelo jekleno ploščo. Mikrostruktura spojev je bila preiskana pod optičnim mikroskopom in z vrstičnim elektronskim mikroskopom, opremljenim z energijsko disperzijskim rentgenskim spektrometrom (EDS). Izmeijena je bila tudi mikrotrdota po Vickersu. Rezultati raziskave so pokazali, da je bila s toplotno modifikacijo zlitine AlSil2 dosežena izboljšana mikrostruktura. Območje spoja med AZ91 in nemodificiranim vložkom iz AlSil2 ni bilo enakomerne debeline. Področje, najbližje zlitini AZ91, je imelo evtektično mikrostrukturo (intermetalna faza Mg7 Al2 in trdna raztopina Al v Mg). Področje v bližini AlSil2 je imelo neenakomerno strukturo z delno spojenimi delci Si, obdanimi s fazo Mg2Si, ter skupki delcev Mg2Si, neenakomerno porazdeljenimi po osnovi, ki jo sestavljata fazi Al3Mg2 in Mg7 Al2 . V tem predelu so bile ugotovljene razpoke. Območje spoja AZ91/AlSil2, izdelanega s toplotno modificiranim vložkom AlSil2, je bilo enakomerne debeline. Tudi področje, kije najbližje zlitini AZ91, je imelo evtektično mikrostrukturo, v področju ob modificirani zlitini AlSil2 pa so bile odkrite signifikantne spremembe mikrostrukture. Ta je bila enakomerna in brez razpok, fini delci Mg2Si pa so bili enakomerno porazdeljeni po osnovi z intermetalnimi fazami Mg-Al. Študija je v obeh primerih razkrila, da je mikrotrdota v območju spoja nekajkrat višja kot pri samih zlitinah. Pri spoju AZ91/AlSil2, izdelanem z modificiranim vložkom AlSil2, mikrotrdota v predelu spoja blizu zlitine AlSil2 ni signifikantno odstopala zaradi enakomerne mikrostrukture. Ta tip mikrostrukture je tudi manj občutljiv na razpoke zaradi obremenitev, ki nastanejo ob vtiskovanju. Analiza mikrostrukture in mikrotrdote opisanih spojev razkriva, da je toplotna modifikacija zlitine AlSil2 zaslužna za mikrostrukturne spremembe v predelu spoja in s tem za nove lastnosti materiala. Iz ugotovitev sledi sklep, daje tehnika sestavljenega litja tekoče na trdno primerna za izdelavo lahkih bimetalov iz magnezijevih in aluminijevih zlitin s potencialno uporabnostjo v industriji. Ključne besede: sestavljeno litje, magnezijeva zlitina, aluminijeva zlitina, udrobnjenje, mikrostruktura, mikrotrdota *Naslov avtorja za dopisovanje: Tehniška univerza Kielce, Fakulteta za mehatroniko In strojništvo, 25-314 Kielce, Poljska, tbucki@tu.klelce.pl SI 55 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 56 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-02-07 Prejeto popravljeno: 2020-05-15 Odobreno za objavo: 2020-05-28 Študija nanoobdelave dela iz jekla EN-31p o postopku magnetoreološke obdelave z okroglo delovno površino in impulznim enosmernim napajalnikom Himmat Singh - M. S. Niranjan* - Reeta Wattal Tehniška univerza v Delhiju, Oddelek za strojništvo, Indija Magnetoreološka obdelava z okroglo delovno površino (BEMRF) je visokonivojski proces nanoobdelave. V predstavljeni študiji je bil uporabljen pristop k fini obdelavi ploskega dela iz jekla EN-31 z napajanjem elektromagnetne tuljave orodja BEMRF z impulznim enosmernim napajalnikom, ki ima potencial za izboljšanje učinkovitosti procesa BEMRF. Cilj študije je analiza vpliva intennitence na odstotno zmanjšanje površinske hrapavosti (%ARa). Na površini ploskega obdelovanca iz jekla EN-31 so bili opravljeni preliminarni eksperimenti z impulznim enosmernim napajalnikom in brez njega. Magnetilni tok (\IC) pri hitrosti podajanja obdelovanca 50 mm/min je znašal 2,5 A, velikost delovne reže (IFG) 1,5 mm in vrtilna frekvenca orodja (R,T ) 500 vrt/min. Orodje BEMRF se napaja z impulznim enosmernim napajalnikom, ki na MR vrhu orodja ustvaija variabilno magnetno polje. Pogosti vklopi in izklopi enosmernega napajanja povzročijo spreminjanje orientacije abrazivnih delcev in posledično prihajajo na površino sveža aktivna abrazivna zrna ali se spreminja orientacija abrazivnih delcev. Delci, ki so v neposrednem stiku s površino obdelovanca, izboljšajo učinkovitost obdelave BEMRF. Preliminarni eksperimenti so pokazali, da je tekstura površine po obdelavi z impulznim enosmernim napajalnikom bolj monotona kot po obdelavi brez takega napajalnika. Vrednost %ARa je znašala 37,09 % pri intennitenci 0,16 ter 24,36 % pri intennitenci 0,27, pri obdelavi brez impulznega napajalnika pa je znašala 13,44 %. Najboljša površinska hrapavost je bila ugotovljena pri intennitenci 0,16. Opravljena je bila tudi statistična analiza za določitev vpliva procesnih parametrov velikosti delovne reže 1,5 mm, vrtilne frekvence orodja 500 vrt/min toka 2,5 A in časa obdelave 30 min na vrednost %ARa pri vrednosti intennitence 0,16 po metodologiji odzivne površine (RSM). Ti parametri so bili uporabljeni za obdelavo z enosmernim napajalnikom z impulzi in brez impulzov pri različnih vrednostih intennitence. Pri enakih procesnih parametrih je bilo ugotovljeno izboljšanje vrednosti %ARa z impulznim enosmernim napajalnikom v primeijavi z napajanjem brez impulzov. Med procesnimi parametri ima največji vpliv na vrednost %ARa pri vrednosti intennitence ® m agnetilni tok (A/C). V predstavljeni raziskavi je bil uporabljen impulzni enosmerni napajalnik za ustvaijanje variabilnega magnetnega polja pri obdelavi jekla EN-31 po postopku BEMFR kot izhodišče za optimizacijo MR fluidov pri impulzni magnetoreološki obdelavi z okroglo delovno površino (PBEMRF). Ključne besede: magnetoreološka obdelava z okroglo delovno površino, impulzni enosmerni napajalnik, intermitenca, odstotno zmanjšanje površinske hrapavosti, impulzna magnetoreološka obdelava z okroglo delovno površino, metodologija odzivne površine SI 56 *Naslov avtorja za dopisovanje: Tehniška univerza v Delhiju, Oddelek za strojništvo, Indija, himmat.diwakar48@gmail.com Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 57 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-02-07 Prejeto popravljeno: 2020-05-15 Odobreno za objavo: 2020-05-28 Primerjalna analiza statičnih in dinamičnih lastnosti nepnevmatskih koles z upogljivimi naperami Zifeng Zhang 41 ongxun Fu* uemeng Liang -X iaoxia Chen 4) i Tan Tehniška univerza v Shandongu, Šola za transport in avtomobilsko tehniko, Kitajska Nepnevmatska kolesa niso napolnjena s stisnjenim zrakom in so izvedena kot integrirane konstrukcije. V primeijavi s kolesi s pnevmatikami imajo več prednosti, saj ne potrebujejo vzdrževanja in ponujajo visoko nosilnost, obenem pa ni nevarnosti eksplozije oz. predrtja. Čeprav so raziskovalci v zadnjih letih zasnovali več različnih vrst nepnevmatskih koles, pa še niso bile opravljene podrobnejše primeijave mehanskih lastnosti raznih nepnevmatskih koles in koles s pnevmatikami. V članku je predstavljena primeijava treh vrst nepnevmatskih koles in tradicionalnih koles s pnevmatikami, in sicer z analizo po metodi končnih elementov ter eksperimentalna. Primeijane so različne lastnosti, prednosti in slabosti štirih koles. V programski opremi AB AQUS je bila opravljena statična in dinamična analiza kolesa z radialno pnevmatiko in treh nepnevmatskih koles različne zgradbe: (1) Pripravljeni so bili modeli koles s končnimi elementi, in sicer kolesa z radialno pnevmatiko, kolesa s ploščatimi naperami, kolesa s satovjem in kolesa z rešetkasto konstrukcijo. (2) V modulu ABAQUS/Standard sta bili opravljeni statična analiza in primeijava modelov koles s končnimi elementi, vključno z nosilnostjo, pritiskom na podlago, napetostmi in deformacijo naper idr. (3) V modulu ABAQUS/Standard je bila opravljena primeijava in analiza dinamičnih lastnosti modela kolesa s končnimi elementi, vključno z odmikom središča kolesa in porazdelitvijo pritiska na podlago med procesom kotaljenja. Iz raziskav mehanike in analize lastnosti štirih koles izhajajo naslednji zaključki: 1. Pri nazivni obremenitvi 3000 N pride do osnosimetrične deformacije območja stika pnevmatike s podlago, območje deformacije pa se pretežno razteza nad stikom s podlago in steno. Gumijast material stene radialne pnevmatike je tanek, zato se ta ugotovitev ujema z dejanskim dogajanjem. Območje deformacije nepnevmatskih koles se razteza predvsem nad stikom s podlago in bližnjimi naperami. Površina stika s podlago pri nepnevmatskih kolesih ima obliko pravokotnika, v območju stika blizu ploščatih naper pa obstaja pojav koncentracije pritiska na podlago. Pri radialnih obremenitvah, manjših od 3500 N, ima najboljšo nosilnost nepnevmatsko kolo z rešetkasto konstrukcijo, pri radialnih obremenitvah nad 3500 N pa nepnevmatsko kolo s satovjem. Če je za poliuretansko nosilno konstrukcijo nepnevmatskega modela uporabljen enak model materiala, je radialna togost nepnevmatskega kolesa s ploščatimi naperami majhna, kar ni primerno za vozila z veliko skupno maso. Radialna togost nepnevmatskega kolesa z rešetkasto konstrukcijo je velika, kar je primerno za težka vozila. Nosilnost kolesa s satovjem je primerljiva z nosilnostjo kolesa z radialno pnevmatiko. 2. Med pospeševanjem prihaja do fluktuacij krivulje odmika središča kolesa, medtem ko se vertikalni odmik središča med stacionarnim kotaljenjem periodično spreminja. Lastnosti nepnevmatskega kolesa z rešetkasto konstrukcijo niso idealne, zato bodo potrebne še dodatne optimizacije. Rezultati pričujoče študije so pomembni za konstruiranje in tovarniško proizvodnjo nepnevmatskih koles. Ključne besede: nepnevmatsko kolo, porazdelitev pritiska na podlago, nosilnost, analiza napetosti, dinamične lastnosti, analiza po metodi končnih elementov ABAQUS *Naslov avtorja za dopisovanje: Tehniška univerza v Shandongu, Šola za transport in avtomobilsko tehniko. Kitajska, fuhongxun615@163.com SI 57 Strojniški vestnik - Journal of Mechanical Engineering 66(2020)7-8, SI 58 © 2020 Strojniški vestnik. Vse pravice pridržane. Prejeto v recenzijo: 2020-02-07 Prejeto popravljeno: 2020-05-15 Odobreno za objavo: 2020-05-28 3D-modeliranje sil pri trdem struženju jekla AISI 4140 po metodi končnih elementov Anastasios TzotzisJ* -C ésar García-Hernández1 4 osé-Luis Huertas-Taln 1 -P anagiotis Kyratsis2 1 Univerza v Zaragozi, Oddelek za konstruiranje in proizvodno strojništvo, Španija 2 Univerza Zahodne Makedonije, Oddelek za produktni in sistemski inženiring, Grčija Struženje v trdo je eden najbolj razšiijenih obdelovalnih postopkov v industriji, ki lahko zmanjša potrebo po končni obdelavi z brušenjem in tako zniža proizvodne stroške. Jeklo AISI 4140 se običajno uporablja za aplikacije, ki zahtevajo visoko obstojnost. Do zdaj še ni bilo opravljenih mnogo študij obdelave jekla AISI 4140 s 3D-analizo po metodi končnih elementov (MKE) in pričujoči članek zato preučuje kritične dejavnike vpliva na proces obdelave jekla AISI 4140 s ciljem razvoja 3D-modela po MKE za napovedovanje sil pri struženju v trdo. Opravljenih je bilo skupno 27 simulacij za preučitev odvisnosti med tremi glavnimi parametri (rezalna hitrost, podajanje in globina rezanja) ter njihovega vpliva na komponente sil pri obdelavi z odrezavanjem. Za rezalne parametre so bile izbrane po tri vrednosti rezalne hitrosti (80 m/min, 115 m/min, 150 m/min), podajanja (0,08 mm/ vrt, 0,11 mm/vrt, 0,14 mm/vrt) in globine rezanja (0,10 mm, 0,20 mm, 0,30 mm). 3D-simulacije so bile opravljene s komercialno programsko opremo za analize po MKE DEFORM3D™. Predlagani model upošteva nekatere vidike, kot so začetek in napredovanje poškodb v materialu, stik med stružno ploščico in obdelovancem ter standardni robni pogoji. Modeli orodja in obdelovanca so bili pripravljeni v programskem okolju SolidWorks™. Narejene so bile tudi določene poenostavitve problema za doseganje sprejemljivega trajanja simulacij. Tako je bil namesto celotnega cilindričnega modela obdelovanca uporabljen le njegov manjši del - natančneje, pretvoijen je bil v krožni lok premera 72 mm. Za mreženje so bili uporabljeni tetraedrični elementi s štirimi oglišči, mreža paje bila zgoščena v srednjem delu površine obdelovanca. Tudi pri modelu rezalne ploščice je bila uporabljena gostejša mreža v bližini rezalnega roba. Tovrstne optimizacije običajno zagotavljajo natančnejše rezultate in skrajšanje časa računanja. Temu je sledilo modeliranje materiala. Za orodje je bila uporabljena konvencionalna keramična ploščica CNGA120408, modelirana po standardih ISO 13399. Za simulacijo napetosti tečenja obdelovanega materiala je bil uporabljen posplošeni Johnson-Cookov model, za simulacijo ločevanja materiala pa normalizirani Cockroftov in Lathamov model. Za aproksimacijo trenja med rezalnim robom orodja in obdelovancem je bil uporabljen standardni Coulombov model. Upoštevana sta bila tudi interakcija med nastalimi odrezki in neobdelano površino obdelovanca ter konvektivni prenos toplote. Rezultati numeričnih simulacij so bili za verifikacijo primeijani z rezultati eksperimentov. Postavljen je bil tudi matematični model na osnovi statističnih metod, kot sta metodologija odzivne površine (RSM) in analiza variance (ANOVA). Rezultati simulacij se dobro ujemajo z eksperimentalno določenimi vrednostmi glavne rezalne sile in njenih komponent: ujemanje rezultante rezalnih sil je v mnogih testih preseglo 90 %. Verifikacija primernosti statističnega modela je pokazala stopnjo točnosti 8,8 %. Iz tega sledijo naslednji sklepi: Višje vrednosti globine rezanja in podajanja povzročijo signifikantno povečanje rezalnih sil. Globina rezanja je dejavnik z največjim vplivom na glavno rezalno silo (/• „/). Z ozirom na simulirane vrednosti Fgl je bilo ugotovljeno približno 104-odstotno povprečno povečanje rezultante rezalnih sil, ko se je globina rezanja povečala z 0,10 na 0,20 mm. Ob prehodu z globine 0,20 mm na 0,30 mm seje sila FgJ povečala za približno 50 %. Podobno se je rezultanta rezalnih sil ob povečanju vrednosti podajanja z 0,08 mm/vrt na 0,11 mm/vrt oz. z 0,11 mm/vrt na 0,14 mm/vrt povečala za približno 24 % oz. 16 %. Rezalne sile nasprotno upadajo s povečevanjem rezalne hitrosti, pri čemer pa je vpliv rezalne hitrosti manj pomemben od vpliva globine rezanja in podajanja. Ocenjeno zmanjšanje vrednosti Fgl ob zmanjšanju rezalne hitrosti s 150 m/min na 115 m/min tako znaša približno 10 %. Ko se rezalna hitrost zmanjša s 115 m/min na 80 m/min, se vrednost Fgl v povprečju zmanjša za približno 13 %. Ključne besede: struženje AISI 4140, sile pri obdelavi, 3D-analiza po metodi končnih elementov, metodologija odzivne površine SI 58 *Naslov avtorja za dopisovanje: Univerza v Zaragozi, Campus Rlo Ebro, C/ Marla de Luna 3,50018, Zaragoza, Španija, atzotzls@unizar.es 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. 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