69 (2023) 7-8 http://www.sv-jme.eu Since 1955 Contents 289 Matej Razpotnik, Thomas Bischof, Miha Boltežar: The Dynamics of Tapered-roller Bearings – A Bottom-up Validation Study 299 Anh-Tuan Dang, Dang-Viet Nguyen, Dinh-Ngoc Nguyen: Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts 308 Fikret Sönmez: Machining of Hard-to-cut AISI 4462 Duplex Stainless Steel with an Environmentally Friendly Approach with Vortex Tube 317 Ivan Marc, Tomaž Berlec: Inventory Risk Decision-Making Techniques Using Customer Behaviour Analysis 326 Kaan Emre Engin: Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting ž of Stainless-Steel Sheet 339 Benchaabane Chaouki, Kirad Abdelkader, Aissani Mouloud: Recent Advancement via Experimental Investigation of the Mechanical Characteristics of Sisal and Juncus Fibre-Reinforced Bio-Composites 352 Vishwa Priya Vellingiri, Udhayakumar Sadasivam: Effect of Vibrator Parameters and Physical Characteristics of Parts on Conveying Velocity Journal of Mechanical Engineering - Strojniški vestnik Papers 7-8 year 2023 volume 69 no. 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Strojniški vestnik -Journal of Mechanical Engineering Aškerčeva 6, 1000 Ljubljana, Slovenia, e-mail: info@sv-jme.eu Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8 Contents Contents Strojniški vestnik - Journal of Mechanical Engineering volume 69, (2023), number 7- 8 L jubljana, July- August 2023 ISSN 0039-2480 Published every t w o months Papers Matej azpotnik, Thomas Bischof, Miha Boltežar: The Dynamics of Tapered-roller Bearings A Bottom-up Validation Study Anh-Tuan Dang, Dang-Viet guyen, Dinh- goc guyen: Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts Fikret S nmez: Machining of Hard-to-cut AISI 4462 Duplex Stainless Steel ith an Environmentally Friendly Approach ith Vortex Tube Ivan Marc, Tomaž Berlec: Inventory isk Decision-Making Techni ues Using ustomer Behaviour Analysis Kaan Emre Engin: Finite Element Analysis of otch Depth and Angle in otch Shear utting of Stainless-Steel Sheet Benchaabane haouki, Kirad Abdelkader, Aissani Mouloud: ecent Advancement via Experimental Investigation of the Mechanical haracteristics of Sisal and Juncus Fibre- einforced Bio- omposites Vish a Priya Vellingiri, Udhayakumar Sadasivam: Effect of Vibrator Parameters and Physical haracteristics of Parts on onveying Velocity 28 2 308 317 326 33 352 Strojniški of of Mechanical Engineering 63(2017)3, XXX-4 289-298 Strojniškivestnik vestnik- Journal - Journal Mechanical Engineering 69(2023)7-8, © of Mechanical All rights reserved. © 2017 2023Journal The Authors. CC BY Engineering. 4.0 Int. Licensee: SV-JME Review Paper — DOI: 10.5545/sv-jme.2017.4027 DOI:10.5545/sv-jme.2023.592 Original Scientific Paper Received review: 2016-11-04 Received for for review: 2023-03-30 Received revised form: 2017-01-14 Received revised form: 2023-05-03 Accepted for publication: 2017-02-12 Accepted for publication: 2023-05-31 The Dynamics of Tapered-roller Bearings – A Bottom-up Validation Study Matej Razpotnik1 - Thomas Bischof2 - Miha Boltežar1,* 1 University of Ljubljana, Faculty of Mechanical Engineering, Slovenia 2 ZF Friedrichshafen AG, Germany Rolling-element bearings are one of the most important elements when predicting the noise of rotating machinery. As a major connecting point between the rotating and non-rotating parts, their dynamic properties have to be accurately known. In this investigation we present a bottom-up approach to characterising the dynamics of the rolling-element bearing. A special test device was designed and built to assess the quality of the well-established analytical modelling approach of Lim and Singh. Two types of bearings were tested, i.e., the ball and tapered-roller types. The dynamic properties were observed by investigating the frequency-response functions. In addition, non-rotating as well as rotating test scenarios were checked. It was shown that the ball bearing model adequately predicts the system’s response, whereas the tapered-roller bearing model requires modifications. These results were further confirmed with a quasi-static load-displacement numerical evaluation, where a full finite-element model serves as the reference. Keywords: dynamic bearing model, tapered-roller bearing, bearing stiffness matrix, vibration transmission Highlights • • • • Test device for bottom-up investigation of bearing’s dynamics is built. Ball and tapered-roller bearings are tested at different speeds and axial preloads. The system is evaluated numerically and experimentally. Ball bearing model is validated whereas tapered-roller one needs improvements. 0 INTRODUCTION Every rotating machinery contains bearings. They represent the connecting points between the rotating and non-rotating parts and so are very important elements in the chain of vibration transmission. The dynamic properties of rolling-element bearings have been studied for many decades; however, due to their complex contact-related phenomena the topic remains important in ongoing research. A first general theory for elastically constrained ball and roller bearings was developed by Jones [1]. Later this theory was further extended by Harris [2]. The theory was very general and focused more on static and fatigue-life calculations than on the vibration transmission through the bearings. Simplified bearing models were introduced by other researchers, with the bearings being modelled as an ideal boundary condition for the shaft, as presented by Rao [3]. Meanwhile, the idea of interpreting the bearings with a simple one- or two-degrees-of-freedom (DOFs) model with linear springs was introduced by While [4] and Garigiulo [5]. Later, more accurate dynamic bearing models were derived. A major improvement in predicting the vibration transmission through rolling-element bearings was made by Lim and Singh [6] and [7] and in parallel by de Mul [8]. They derived a model that provides a comprehensive bearing-stiffness matrix. The model is capable of describing the nonlinear relation between the load and the deflection, taking into account all six DOFs and their interplay. These authors also presented system studies [9] and [10] for model-validation purposes. A good agreement between the measurements and the analytical model was shown for the ball and the cylindrical roller bearings as the two most distinct examples of different contact types. The six-DOFs model is the basis for the widely used industrial standard ISO/TS 16281 [11] as well as for many subsequent studies. Recently, a thorough review of mechanical model development of rolling-element bearing was presented by Cao et al. [12]. The authors classify modelling approaches into five different techniques and comprehensively discuss the current progress of development as well as identify future trends for research. Despite great computational power available these days, modelling of the bearings primary remains on the analytical level. Contact related phenomena and non-linearities lead to huge and often unstable finite-element method (FEM) models. However, connecting analytical models with numerical ones is crucial in predicting the proper behaviour of a modern system. Guo and Parker [13] presented a stiffness-matrix calculation *Corr. Author’s Address: Name of institution, Address, City, Country, xxx.yyy@xxxxxx.yyy *Corr. Author’s Address: University of Ljubljana, Faculty of Mechanical Engineering, Aškerčeva 6, 1000 Ljubljana, Slovenia, miha.boltezar@fs.uni-lj.si 1 289 “output” — 2023/6/22 — 8:23 — page 2 — #2 Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 using a finite-element/contact-mechanics model. On the other hand analytical approaches demand advanced methods to solve the system of nonlinear equations. Fang et al. [14] recently presented a comprehensive study of the speed-varying stiffness of ball bearings under different load conditions. They proposed a novel mathematical method for solving an implicit set of nonlinear equations based on a new assumption of the initial conditions. To mitigate the numerical difficulties of time integration procedure, induced by rolling-elements coming and leaving the contact, Razpotnik et al. [15] extended the model from Lim and Singh [6]. They implemented modular smoothing of the load-displacement characteristics in the region of contact-state transition. To improve the calculation accuracy of non-hertzian contact pressure the high-precision half space theory was adopted by Kabus et al. [16] and [17]. Tapered-roller bearings (TRBs) are widely used in rotor dynamics. They are usually treated as a special case of a cylindrical roller bearing with a non-zero contact angle. The difference, in fact, is much more significant because TRBs have two different contact angles (the inner ring-roller and outer ring-roller contacts) and also because of their additional roller-flange contact. Cretu et al. [18] and [19] analysed the dynamics of TRBs under fully flooded conditions. The assumption of an elastohdydrodynamic (EHD) lubrication regime is common to the majority of TRB studies. In this way the friction forces can be either calculated or, even more commonly, neglected. Tong and Hong [20] analytically studied the characteristics of TRBs subjected to combined radial and moment loads. The same authors [21] investigated the influence of the roller profile and the speed on the stiffness of a TRB. Zhao et al. [22] studied the effect of gyroscopic moment on the damage of a tapered-roller bearing, which are found to occur under high-speed and high-load conditions, such as high-speed trains. Roda-Casanova and Sanchez-Marin [23] presented an illustrative study of the contribution of the deflection of the TRB to the misalignment of the pinion in a pinion-rack transmission. They stressed the importance of having accurate knowledge of the elasticity of the bearings. Houpert [24] studied the torque generated by the friction forces in a TRB, where he emphasised an important fact, i.e., TRBs are subjected to a high roller-flange torque. The roller-flange contact, which largely affects the power loss, was investigated also by Ai et al. [25]. In addition, Tong et al. [26] made numerical evaluation of the effect 2 290 of misalignment on the generated friction forces and consequently evaluated the power loss of a TRB. It was shown that already a small misalignment can have a significant influence on the generated torque. Experimental investigations of a TRB’s dynamics are rare in the literature. Zhou and Hoeprich [27] measured the torque generated at different contacts in a TRB; however, they focused on the losses and not on the dynamics. Gradu [28] also analysed the TRB losses and compared them with equivalent ball bearing. Wrzochal et al. [29] presented a new device for measuring the friction torque in rolling-element bearings of different types, where the main goal was to establish a reliable device for quality control measurement. Discrepancy between theoretical and measured friction torque was presented and discussed. A comparative study, as presented by Zhang et al. [30] for angular-contact ball bearings, would also be beneficial for TRBs. Further, since TRBs are often used in applications that do not require high speed, the influence of friction on the dynamic properties would be generally welcomed. In this paper a numerical and experimental characterisation of a TRB’s dynamics is presented. First, a general bearing modelling technique is introduced, where the analytical model of Lim and Singh [6, 7] is embedded into a FEM model. Afterwards, a special test device is presented. There follows a description of a workflow for a bottom-up validation study. A TRB is mainly investigated, whereas ball bearing is also tested. The results in the form of frequency-response functions (FRFs) are compared for the measurements and the simulations. The non-rotating as well as low-speed-rotating scenarios are presented. Finally, a quasi-static load-displacement numerical analysis was performed to additionally verify the results. 1 BEARING MODELLING TECHNIQUE Rolling-element bearings can be modelled as a part of a wider system in several different ways. Most often the system is studied by utilising a FEM model. The bearings are, due to their complex contact-related phenomena, represented by a special element that embeds the analytically calculated bearing-stiffness matrix Kb . This technique introduces the so-called spider elements (commercially known as RBE3 element), as shown in Fig. 1. A spider element connects a raceway of a ring to one, central node. The motion of that central node depends on the weighted average of the motions at a set of connected grid points [31]. Two spider elements are needed, Razpotnik, M, – Bischof, T. – Boltežar, M. Razpotnik, M. Bischof, T. Boltežar, M. “output” — 2023/6/22 — 8:23 — page 3 — #3 Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 2 TEST DEVICE Kb A simple test device was designed and built for validation purposes. It consists of the housing, a long shaft and a special nut, as shown in Fig. 2. It is important to note the shape of the nut, which is only in contact at both ends. This design ensures that the load dependency of the thread contact is negligible. A bottom-up validation approach was (a) nut housing shaft Fig. 2. A simplified technical drawing of the test device, where all three possible configurations are shown Kb (b) Fig. 1. A cross-section view of a bearing in a FEM model with zoomed-in areas; a) ball bearing, and b) TRB each for one bearing ring. Central nodes, connected by Kb , are located at exactly the same position. Fig. 1 separates them just for illustrative purposes. Bearing rings can be coarsely meshed since the mesh is only used to represent the shape of the ring and to serve as a connecting body for the spider element. Rolling elements are not modelled, since all contact-related phenomena are covered by the stiffness matrix. The source of the accuracy of the presented technique is not in the spider element itself, but in the bearing-stiffness matrix. In this study we implement the bearing-stiffness matrix from Lim and Singh [6] and [7]; however, any relevant bearing theory yielding the stiffness matrix can be implemented. There are two types of FEM analyses used in the presented study, i.e., frequency-response modal analysis and quasi-static load-displacement analysis. Both assume that rolling elements are not rotating. However, the former is used to obtain dynamic response of the system when excited and subjected under different axial loads, and the latter is used to obtain load-displacement characteristics of a bearing. utilised. Therefore, solid rings are inserted at first in order to prove the linearity of the system. Afterwards, two types of rolling-element bearings were tested, i.e., ball bearing and TRB, with the properties given in Table 1. In order to eliminate the influence of the surroundings, the device was tested with free-free boundary conditions (BCs). These conditions were achieved by hanging the test device via housing by thin ropes. The FRFs were measured between different Table 1. Bearings used in the test device. type designation d [mm] D [mm] B [mm] ball tapered-roller 6006 32006-X 30 30 55 55 13 17 parts of the system. An excitation was applied with a modal hammer, whereas the acceleration was measured by the accelerometer. The transfer path from the shaft to the housing is of special interest, since the bearing’s dynamics are the most clearly seen there. The test device makes it possible to apply different axial preloads to the bearings and consequently to the entire system by turning the nut with respect to the shaft. The applied axial force is measured with the strain gauges located at the housing ribs. Additionally, the system can be investigated while the shaft is either stationary or rotating (up to 6000 rpm). For this purpose a special motor can be mounted to the system. In doing so, the free-free BCs are maintained, as shown in Fig. 3. Dynamics Tapered-rollerBearings Bearings ––AA Bottom-up Study TheThe Dynamics of ofTapered-roller Bottom-upValidation Validation Study 2913 “output” — 2023/6/22 — 8:23 — page 4 — #4 Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 measurements/ simulations 100 N Preload 300 N 600 N 1000 N FRFs Comparison between measured and calcluated FRFs Fig. 3. Experimental setup with attached motor Fig. 5. Workflow of the validation procedure for the non-rotating test device 3 WORKFLOW The main goal was to evaluate the quality of the analytically calculated dynamic bearing model. The assessment was made for a non-rotating scenario by comparing the dynamic properties of the system in the form of FRFs. We focused on a representative FRF, namely Accelerance, where the excitation was performed on the shaft and acceleration was obtained at the housing, as shown in Fig. 4. All of the possible setups (solid rings, ball bearings, TRBs) undergo the same testing procedure. Four different axial preloads were inserted into the system. Thus, four different FRFs were obtained. Fig. 5 shows Fig. 4. Excitation point and accelerometer position the corresponding workflow. The same workflow is followed for measurements and simulations. Finally, the FRFs are compared. The rotating version of the test device was investigated only experimentally. A run-up investigation was utilised. By that we can see the potential change of the system’s dynamics due to rotor-dynamic effects appearing in the rolling bearings, e.g., centrifugal forces and gyroscopic effects. Some researchers have pointed out these effects during high-speed applications [1], [14], [21], [22], and [32], where the effects start to become noticeable in the region between 5000 rpm and 10000 rpm, depending on the bearing type and load case. However, it is important to note that our testing did not exceed 6000 rpm. When the system rotates, there is no additional external excitation. The system is mainly excited by the white noise coming from the bearings. The resulting acceleration is measured and shown in a Campbell diagram. 4 292 4 RESULTS 4.1 Solid Rings With solid rings inserted it is possible to verify whether the system without bearings is linear or not. The linearity implies the load-independent dynamic properties of the system. Ideally, the presented system should be load-independent; however, due to contact issues, especially the thread contact, the load independence has to be experimentally proven. Fig. 6 shows the results. The measured FRFs, given in the form of Accelerance (A), are shown with a gray and black colour. The red curve corresponds to simulation. The preload was considered in simulations as well, but its effect is completely unnoticeable, thus only one line sufficiently represents the simulation results. It is clear that all the measured curves correlate well with each other regarding the eigenfrequency position. Damping, however, decreases with an increased preload. Also, the calculated FRF predicts the measured behaviour correctly. However, the peak around 1.45 kHz is more damped in the measured results. Proving the linearity of the system with solid rings is an important step. All the non-linearities in the succeeding investigations (when real bearings are inserted) can now be associated with the bearing’s behaviour. 4.2 Ball Bearings The ball bearings, as given in Table 1, are inserted into the system. Fig. 7 shows the amplitude comparison between the measured and calculated results. It is clear that some peaks move their position with the increased preload, while the others do not. Those involving the modes of the shaft are affected, while the others are not. Fig. 8 shows the modes of the marked regions from Fig. 7. The stiffness of the bearing plays a crucial role there. All of them are pure modes of the shaft, only the mode at 2490 Hz is a combination Razpotnik,M. M, – Bischof, Bischof, T.T.– Boltežar, Razpotnik, Boltežar,M.M. “output” — 2023/6/22 — 8:23 — page 5 — #5 104 103 102 101 100 10−1 10−2 10−3 π 0 −π 1 meas, 100N meas, 300N meas, 600N meas, 1000N sim γ2 [ / ] φ [rad] A [1/kg] Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 0 0 500 1000 1500 2000 2500 Frequency [Hz] 3000 3500 4000 Fig. 6. Load-dependency of the FRF (from top to bottom: magnitude, phase, coherence) for the test device with the inserted solid rings of shaft and the local movement of the housing ribs. Additionally, it is clear that the higher the preload, the higher is the bearing stiffness. This also causes the eigenfrequencies to increase. Comparing the spectra it can be seen that the frequency span of each affected eigenfrequency is around 100 Hz for the measured as well as the calculated FRFs. Some peak positions, however, differ slightly, but the general behaviour is well predicted. It is important to note that no FEM model updating was performed. Doing so would obviously align the calculated results completely with the measured ones. So far, we have presented the non-rotating version of the test device. Since the bearings are meant to rotate it is crucial to determine whether the analytical bearing-stiffness matrix is an adequate representation of the bearing’s dynamics also under operating conditions. A run-up test was performed. An extension to the motor was additionally mounted to the test device while maintaining its free-free BCs (see Fig. 3). The run-up test sequentially increases the motor speed from 500 rpm to 6000 rpm with a step of 100 rpm. At each step the acceleration on the housing (the same position as for the FRF investigation) was measured. The resulting Campbell diagram is shown in Fig. 9 for the preloads of 300 N and 1000 N. The preload of 100 N was found to be too loose for the run-up investigation. Already a slight torque induced by the motor caused a disturbance that changed the axial preload. At 300 N this effect is not noticeable any more. It is clear that the eigenfrequencies governed by the bearing stiffness do change with a higher preload in a similar manner to the non-rotating version. They are marked with red arrows. On the other hand, their position does not change while increasing the RPM in the investigated RPM region. Another important conclusion is that the locations of the eigenfrequencies remain at practically the same positions as in the non-rotating investigation. The distinct change of eigenfrequencies dominated by the bearing stiffness is evident. Comparing the results for the non-rotating and rotating versions we notice that one dominating peak is missing in the rotating version, i.e., the one at 2490 Hz. The eigenmode of this peak is a combination of housing and shaft movements and is apparently changed due to the extension mounted to the housing. The comparison between the measurements and the simulations shows good agreement for the non-rotating as well as for the rotating setup. As such it can be concluded that the analytical bearing-stiffness matrix seems to be an adequate representation of the actual bearing’s dynamics for the ball type in the observed speed range. 4.3 Tapered-roller Bearings The TRBs, as given in Table 1, are inserted into the system. Fig. 10 shows the amplitude comparison between the measured and calculated results. Both spectra have marked regions where the eigenfrequencies shift with respect to the inserted preload. Comparing the spectra it is clear that the positions of the regions differ tremendously. Investigating the eigenmodes gives us an insight into the problem. Fig. 11 shows all the calculated eigenmodes of the test device within the marked Dynamics Tapered-rollerBearings Bearings –– AA Bottom-up Study TheThe Dynamics of ofTapered-roller Bottom-upValidation Validation Study 2935 “output” — 2023/6/22 — 8:23 — page 6 — #6 A [1/kg] A [1/kg] Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 104 103 102 101 100 10−1 10−2 10−34 10 103 102 101 100 10−1 10−2 10−3 meas, 100N meas, 300N sim, 100N 0 500 1000 1500 meas, 600N sim, 300N 2000 2500 Frequency [Hz] sim, 600N 3000 meas, 1000N sim, 1000N 3500 4000 Fig. 7. Comparison between the measured (upper part) and calculated (lower part) load-dependency of the chosen FRF for the test device with inserted ball bearings (a) (c) Since the non-rotating scenario has a huge gap between the measurements and the calculations we did not continue to the rotating scenario. Instead we tried to shed some light on possible causes for the observed differences with the help of a detailed FEM bearing model, as discussed in detail in the next section. (b) (d) 5 NUMERICAL INSIGHT Fig. 8. Selected calculated eigenmodes of the test device with ball bearings; a) Eigenmode at 590 Hz, b) Eigenmode at 1650 Hz, c) Eigenmode at 2490 Hz, and d) Eigenmode at 3085 Hz frequency band from Fig. 10. All the presented modes are governed by the bearing stiffness and do actually correspond to the first, second and third modes of the shaft. To show the measured modes we need to do a complete experimental modal analysis (EMA). For this purpose we used an approach with a high-speed camera. The measured eigenmodes are shown in Fig. 12. The first eigenmode obviously represents the first mode of the shaft, whereas the second mode represents the combination of the dominant local housing movement and the second mode of the shaft. The arrows in Fig. 12 indicate the direction of motion, where a different colour stands for a different phase. From the presented results it can be concluded that the analytically calculated bearing-stiffness matrix for the TRBs exhibits behaviour that is much too weak. As such the current modelling approach is not an adequate representation of the TRB’s dynamics. 6 294 The analytical bearing-stiffness model seems to be a good representation of reality for the ball bearing, but considerably worse for the TRB. To find the origin of the problem we built a complete, detailed, full FEM bearing model for both bearing types (see Table 1) as depicted in Fig. 13. The goal is to compare the load-displacement characteristics, where the load is exerted incrementally in the axial direction. A quasi-static load-displacement analysis was performed on the full FEM bearing model. Further, the slope of the load-displcement curve is extracted, representing the total axial stiffness in the loaded direction, which is also compared. Due to its completeness, the FEM model represents the reference. Besides the full FEM model and the analytical model from Lim and Singh [6] and [7], the results from the widely used standard ISO/TS 16281 [11] are also included. The results in the form of load-displacement characteristics and the corresponding stiffness for the ball bearing are shown in Fig. 14. All three approaches result in similar characteristics. There is a minor gap between both analytical approaches, whereas the FEM yields a slightly higher stiffness at a high preload. Razpotnik,M. M, – Bischof, Bischof, T.T.– Boltežar, Razpotnik, Boltežar,M.M. “output” — 2023/6/22 — 8:23 — page 7 — #7 Strojniški vestnik - Journal of Mechanical 289-298 Strojniški vestnik - Journal of MechanicalEngineering Engineering69(2023)7-8, 63(2017)3, XXX-4 A [1/kg] A [1/kg] Fig. 9. Campbell diagrams for the test device with inserted ball bearings loaded under 300 N and 1000 N 104 103 102 101 100 10−1 10−2 10−34 10 103 102 101 100 10−1 10−2 10−3 meas, 100N meas, 300N sim, 100N 0 500 1000 1500 meas, 600N sim, 300N 2000 2500 Frequency [Hz] sim, 600N 3000 meas, 1000N sim, 1000N 3500 4000 Fig. 10. Comparison between the measured (upper part) and calculated (lower part) load dependency of the chosen FRF for the test device with inserted TRBs (a) (b) (c) Fig. 11. Selected calculated eigenmodes of the test device with TRB; a) Eigenmode at 820 Hz, b) Eigenmode at 1960 Hz, c) Eigenmode at 3300 Hz However, it can be concluded that all the models are sufficiently well correlated. In the same manner we incrementally load the TRB in the axial direction. The results are shown in Fig. 15. The curves quantitatively differ significantly; however, they reflect the same tendency. Concerning the load-displacement characteristics, it is interesting that already both analytical approaches differ to a great extent. The same two approaches exhibit a similar level in the stiffness characteristic. The FEM model, on the other hand, yields a factor of two higher stiffness. DynamicsofofTapered-roller Tapered-roller Bearings Bearings ––AABottom-up Study TheThe Dynamics Bottom-upValidation Validation Study 7 295 “output” — 2023/6/22 — 8:23 — page 8 — #8 Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 0.12 δz [mm] 0.10 0.08 0.06 FEM Analytical Standard 0.04 0.02 (a) 0.00 0 100 200 300 400 500 600 Fz [N] (a) 700 800 900 1000 300 400 500 600 Fz [N] (b) 700 800 900 1000 ×104 FEM Analytical Standard kz [N/mm] 5 (b) 4 3 2 1 Fig. 12. Measured eigenmodes of the test device with inserted TRB (using high-speed camera); a) Eigenmode at 1250 Hz, b) Eigenmode at 2490 Hz 0 0 100 200 Fig. 14. Ball bearing being incrementally loaded in the axial direction; a) load-displacement characteristic, b) corresponding total stiffness (a) (b) Fig. 13. Full FEM model of a bearing; a) ball bearing, b) TRB 6 DISCUSSION Modelling the dynamics of ball bearings with the presented approach seems appropriate. The discrepancy between the measurements and the simulations is negligible. On the other hand, the TRBs have a significant mismatch between the measured and calculated results. This is true for the dynamic testing (comparing FRFs) as well as for a simple quasi-static load-displacement investigation. The reason for such a disagreement is the inadequate bearing-stiffness model. The shortcomings can be outlined as: 1. The TRB has different contact angles in between the inner ring-roller and the outer ring-roller. The theory assumes that all the contacts are 8 296 happening at the nominal contact angle, which is defined for the axis going through the centre of a roller. Consequently, no axial force is generated that pushes the rollers out of the initial contact and no flange is needed to prevent the rollers from escaping the contact. The roller-flange contact is thus neglected. This contact carries only minimal load; however, it should not be neglected, especially if the TRB is loaded in the axial direction only. It is important to note that the literature provides some theories that take into account different contact angles and flange contacts [8] and [21]. However, those theories provide a negligibly different stiffness matrix compared to the theory, which does not include the mentioned contacts. Having looked at the results, the stiffness should be of factorial difference. 2. Friction effects are neglected in all models, i.e., analytical, standard and FEM model. When the bearing operates in the hydrodynamic regime, the friction is expected to be very small and as such it is justified to neglect it in the stiffness calculation. However, when the bearing operates in the boundary or mixed-lubrication regime, the friction coefficients are expected to have a significant influence on the bearing’s stiffness. Further, the dynamic load from the eigenmodes is expected to be small enough to not cause the transition from the stick to the slip state in the contact along the roller’s line of action. That Razpotnik,M. M, – Bischof, Bischof, T.T.– Boltežar, Razpotnik, Boltežar,M.M. “output” — 2023/6/22 — 8:23 — page 9 — #9 Strojniški vestnik - Journal of Mechanical Engineering 289-298 Strojniški vestnik - Journal of Mechanical Engineering69(2023)7-8, 63(2017)3, XXX-4 δz [mm] 0.008 99 REFERENCES REFERENCES FEM Analytical Standard 0.006 0.004 0.002 0.000 0 3.5 100 200 300 400 ×105 500 600 Fz [N] (a) 700 800 900 1000 kz [N/mm] 3.0 2.5 2.0 1.5 1.0 0.5 FEM Analytical Standard 0.0 0 100 200 300 400 500 600 Fz [N] (b) 700 800 900 1000 Fig. 15. TRB being incrementally loaded in the axial direction; a) load-displacement characteristic, b) corresponding total stiffness being said, the stiffness of the TRB would have been significantly increased when the friction phenomena were also taken into account. 7 CONCLUSION A bottom-up approach to characterise a bearing’s dynamics is presented. A special test device was designed and built to assess the quality of the well-established modelling approach. The dynamic properties of the system were measured in the form of FRFs, where load-dependent nonlinearities, resulting from the bearings were observed. It was shown that the ball bearing model yields appropriate results, whereas the TRB model requires modifications. These outcomes were confirmed with a quasi-static, load-displacement numerical insight, where a full FEM model serves as a reference. In future work it will be of great interest to see how a TRB reacts when loaded in other than the pure axial direction. In addition, the influence of a different lubrication regime would shed some light on questions relating to the friction. The presented study represents a good starting point for a possible new TRB model derivation. 8 ACKNOWLEDGEMENT The authors would like to thank the ZF Friedrichshafen AG for supporting this research. [1] Jones, A.B. (1960). A general theory for elastically constrained ball and radial roller bearings under arbitrary load and speed [1]conditions. Jones, A. B., of Basic A general theoryvol.for82,elastically Journal Engineering, no. 2, p. constrained ball and radial roller bearings under DOI:10.1115/1.3662587. 309-320, arbitrary load and speed conditions, of Basic [2] Harris, T.A. (1984). Rolling Bearing Analysis,Journal John Wiley, New Engineering, ASME 82 (1960) 309. York. [2] Harris, T. A., Rolling bearing analysis, John Wiley, New [3] Rao, J.S. (1983). Rotor Dynamics, John Wiley, New York. york, 1984. [4][3]While, M.F. (1979). 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Licensee: SV-JME DOI:10.5545/sv-jme.2023.539 Original Scientific Paper Received for review: 2023-01-29 Received revised form: 2023-05-15 Accepted for publication: 2023-06-06 Applyi ng Parametric Analys is in Enhancing Performance for D ouble-L aye r Scissor L ifts Dang, A.T. guyen, D.-V. guyen, D.- . Anh-Tuan Dang Dang-Viet guyen Dinh- goc guyen* Thai guyen University of Technology, Vietnam This study presents a calculation procedure of applying the statistical method in determining design parameters for double-layer scissor lifts to improve their operation (e.g., lifting height, loading, stability). By parameterizing the mounting orientation of cylinders, the working of the mechanism is summarized as functions and evaluated. To verify the accuracy of obtained equations, a 2D model of the mechanism was constructed and simulated using Working Model software. The results obtained from the simulation indicate that by adjusting the mounting positions of cylinders, the platform’s position and reactions on the device can be determined, enabling improvement of the system’s performance. The results obtained from the study show the practical significance of applying parametric methods to the calculating process and optimizing the structure of multiple-layer scissor lifts. Keywords: double-layer scissor lift, cylinder’s orientation, kinematic analysis, parametric method Highlights Applying parametric analysis for a distinctive model of a double-layer scissor lift to obtain the reactions of each joint in the whole mechanism. The proposed method can assist the calculation process without the need to construct 3D models for complex simulations. Based on the acquired reactions and position information, the structure of the system can be further optimized, reducing the production cost and the calculation time. From the initial design parameters, hydraulic cylinders can be easily selected to meet the designer’s requirements. 0 INTRODUCTION Scissor lifts are popular aerial platform lifts that use a scissor-shaped mechanism to lift objects or people over short distances in height. Although the first mechanism as used in S eden in the early 1 00s, it as not until the 1 70s that scissor lifts gained popularity as commonly manufactured lifting structures. To select the appropriate lift for the job, lifting height, and loading are factors to be considered hen choosing the e uipment. Based on the above re uirements, several studies have been conducted on devices to test their functional ability. Solmazyi it et al. [1] proposed the construction of and successfully manufactured a scissor lift device that enables lifting loads of up to 25 tons. Pan et al. [2] conducted a falling arrest test on a multi-layer scissor lift to check the structural stability, thereby ensuring tip-over safety for the system during the dropping process. Another study by Dong et al. [3] on an actual model also sho s that the tip-over potential of the scissor lift system depends on the cylinder s speed. The study also indicates that the system s stability reduces hen connecting joints are severely orn, or the structure is damaged. In addition to manufacturing and testing on real devices, other studies also focus on theoretical models to evaluate the performance of lifting systems to select the optimized layouts. Spackman [4] applied mathematical techni ues to analyse the mechanism of n-layer scissor lifts. The study not only analysed the reactions in the scissor members but also presented some design issues related to actuator placement, member strength, and rigidity. Kosucki et al. [5] suggested using a volumetric controller to regulate the speed of the cylinders. ith significant advantages such as a simple structure and lo cost, the paper presents the possibility of extending the use of volumetric control to operate lo -po er driving systems. Karag lle et al. [6] employed finite element analysis (FEA) on a 3D model created in Solid orks to ascertain the internal loads acting on each component of a single-layer system. Dang et al. [7] proposed using the parametric method to model single-layer scissor lifts and construct mathematical e uations to determine cylinder loads and reactions at the joints. Using the same approach, Todorović et al. [8] applied Harris Ha ks optimization (HH ) to reduce mass in the mechanism frames. By solving static e uations for a single-layer structure, uchor et al. [9] suggested an accurate method for selecting a hydraulic cylinder capable of lifting loads up to 3.5 tons. These studies demonstrate the significant influence of the cylinder on the operation of the lifts. It can be observed that for 3D models *Corr. Author’s Address: Faculty of International Training, Thai Nguyen University of Technology, Vietnam, ngocnd@tnut.edu.vn 299 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 constructed using graphic soft are, the calculation process is only performed for limited positions, thus making it difficult to evaluate the overall impacts of the cylinder s operation. In contrast, employing mathematical e uations as a solving approach can greatly reduce the calculation and testing time, providing a more efficient method. In this study, a specific structure of a double-layer scissor lift consisting of t o cylinders is selected for analysis (Fig. 1). By determining the mathematical relationship bet een the movements of the cylinders and the platform, designers can choose the suitable dimensional arrangement for the cylinders that is beneficial for the re uired lifting. designers may have difficulty choosing the correct layout to represent the movement of the platform. a) b) Fig. 2. Schematic diagram of the mechanism in: a) lowest position, and b) highest position Assign parameters for assembling dimensions F = β1· a, PF = β2· a, and P λ·a ( ith 0 β1, β2 1, λmax > λ > λm i n 0 is the length ratio of the cylinder compared to one frame); the system is parameterized as sho n in Fig. 3. Fig. 1. Structure of the studied system 1 KINETIC ANALYSIS From the 3D model of the system, a 2D model is represented, as sho n in Fig. 2. The design problem can be described as follo s: for a double-layer scissor mechanism ith a fixed frame length a, designers must arrange t o given hydraulic cylinders to utilize the maximum potential of the structure. Assuming the cylinders operate bet een the zero-stroke (l Cyl .min) and full-stroke positions (l Cyl .max) ith a maximum thrust force F Cyl , and the platform has a raising re uirement of Dh. The problem then becomes one of determining the mounting position of joints P and in the t o connected frames (AF and F ) hile ensuring the stability of the lift and limiting the loading on the cylinders. Analysis of the lift construction sho s that hen the cylinder operates (the distance bet een P and changes), platform 5 moves to the corresponding height h. Although the length of the frames (a) and the arrangement dimensions (F , FP) of the cylinder are fixed, the stability of the mechanism is altered (angles γ at joints bet een frames E and F, and the distance l bet een supports A and B). Since there are too many parameter dimensions involved in the operation, 300 Fig. 3. Parameterizing the device in Fig. 2 hen cylinder extends, platform height h can be calculated using the follo ing e uation: h  2a sin  1  cos   2a . 2 2 (1) The angle γ bet een frames is determined based on the relationship of the triangle MP : Dang, A.T. – Nguyen, D.-V. – Nguyen, D.-N. cos   12   22   2 . 2 1  2 (2) Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 Substitute E . (2) into E . (1); the height of the platform can be determined by the layout dimensions of the structure: ha  2   1   2  2 1  2 support distance l changes during the operation of cylinders, the magnitude of loads acting on joints A and B also changes, influencing the balance of the device and reactions on joints. (3) . The lifting efficiency of the system can be evaluated by the lifting ratio, hich is determined by the lifting height of the platform hen the cylinder extends its length from zero-stroke to full-stroke: kh  hmax  hmin  f  max , min , 1 ,  2  . a (4) It should be noted that hen the platform is raised, γ bet een the frames increases and the loading distance bet een supports l decreases, hich is also a factor affecting the stability of lift. This distance can be calculated using the e uation:  l  a cos  a 2  1   2  2 4 1  2  2 . (5) 2 FORCE ANALYSIS As described in the modelling step, due to the symmetry of the 3D structure, it is possible to investigate the reactions on the lift by analysing force on one side of the system. If the loading is 2P , the reactions at the bearings and cylinders in Fig. 2 can be correspondingly determined in terms of W p ) 2 ( here W p represents the eight of P G (P the platform, and is the centroid of the total load P G ). In practice, the load on the system, specifically the lifting load P G on the platform, may continuously change during the operation of the device. This occurs hen orkers alk on the platform to perform repair and maintenance tasks hile the lift is still in motion. This situation not only exerts forces on joints but also causes deformation of the component frames, including bending, compression, or tensile stresses. Since these frames can be designed ith rigid and sustainable materials, the impact of forces on them can be ignored, allo ing the problem to be focused on revolution joints here bearings are assembled. In the scope of this study, this paper is only concerned ith analysing the static influence of forces on the structure (it considers that the load P G is fixed on the platform during the movement of the cylinder, and the frame eight W of each frame is at its centre). This means that the distance bet een point and support A is constant hen lifting persons or objects from height h t o h’ (see Fig. 4). Ho ever, since the Fig. 4. Location of load PG when the platform raises from the lowest to the highest position Assuming that the platform is raised slo ly enough that the effect of acceleration can be ignored. iven the load on the system and the cylinder s eights are much smaller than the eight of the frames (W ), the calculation of reactions on the structure can be outlined in the follo ing steps: First, the scissor structure is separated from moving platform 5 and ground platform 1 (see Fig. 5). Based on this figure, the reactions at supports A and B can be determined using moment e uilibrium e uations: PB  PB  PG lG , l PA  PA  PG  PB  PG  (6) PG lG . l Similarly, the reactions at supports be calculated as: PC  PC  PG lG  4W l PD  PD  PG  Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts (7) and D can l 2  PG lG  2W . l (8) PG lG  2W. l ( ) 301 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 Fig. 5. Separating component frames to determine reactions in the mechanism Fig. 6. Release the connection at joints M and N Second, release the connection at joints M and , and separate these reactions into directional c om pone tn s F x and F y. (see Fig. 6): Balancing the moment in DEB frame: M E M  0,  a   PD  PB  a cos   FMx  FNx  sin 2 2 2 a   WNy  2W  FMy  cos  0. 2 2 302 Third, continuing to release connections at E, F, and apply the moment e uilibrium e uation at point E for BE frame (see Fig.7); E   FNy  W  PB a cos (10) Dra F Dang, A.T. – Nguyen, D.-V. – Nguyen, D.-N. Nx and F M x  a2 cos 2   a  FNx sin  0. 2 2 2 from E . (10): (11) Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 Fig. 7. Release the connection at joints E and F FNx  FMx  W tan Dra F y  2 . (12) from E . (12): Pl  FNy  2 PB  2W  2  G G  W  .  l  (13) Apply the moment e uilibrium e uation for AF frame:  a   M F  PA a cos 2   FNy  W  2 cos 2 a   FNx sin  FQ 1a sin   0. 2 2  PG  2W    2 1  2 sin 2   PG  2W   2 1  2  2   1   2       FFx  FQ cos      FNx , 2   (16)   FFy   FQ  sin      FMy  PA  W. 2  (17) Subtitle β1, β2 and λ into these e uations:   PG  2W    2  1    1   2    2 FFx   W  , (18) 2 2 1  2    2   1   2  2 (14) Substitute E s. (7), (12) and (13) into E . (14) and dra the cylinder s thrust force F : FQ  present in the reaction at other joints, their magnitudes ill vary and influence the stability of the system. By applying the e uilibrium e uation for frame AF, directional reactions at joint F can be computed as: . (15) It can be observed that the coefficient l G in E . (15) has been omitted, implying that the position of loading P G does not impact the magnitude of thrust force in cylinders. Ho ever, as this parameter remains FFy  PG   3W  PG lG a 4 1  2  1   2    2  PG  2W   1   2  2 2 1  2 . (1 ) The total reactions in the remaining joints can be determined by combining their directional components: FE  FE2x  FE2y , Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts (20) 303 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 FF  FF2x  FF2y , (21) FM  FN  FN2x  FN2y . (22) A Reaction at point N 3 RESULTS To validate the accuracy of the calculation method, five 2D physical models ere created using orking G PG B N Reaction at point E Thrush force on cylinder Cylinder Reaction at point F E F Reaction at point M M C D Fig. 8. Export reactions at joints using Working Model Table 1. Compare result of forces between calculation method and 2D model simulation a = 1 m; W β1 = 0.20 β2 = 0.65 l = 0.46 β1 = 0.20 β2 = 0.65 l = 0.74 β1 = 0.20 β2 = 0.80 l = 0.62 β1 = 0.10 β2 = 0.30 l = 0.22 β1 = 0.25 β2 = 0.50 l = 0.30 = 75 N; P G = 5000 N FCyl FM FE FF FCyl FM FE FF FCyl FM FE FF FCyl FM FE FF FCyl FM FE FF lG F 1 [N] 68,865.00 807.30 67,138.51 1,288.87 17,991.60 938.33 12,454.58 1,874.53 50,998.50 693.59 49,127.06 1,225.13 71,199.00 652.18 68,427.42 1,215.76 26,339.80 665.36 23,158.51 1,217.95 =1m F 2 [N] 68,876.76 807.18 67,150.27 1,287.93 17,992.64 938.66 12,455.55 1,875.04 51,102.72 695.64 49,227.89 1,228.80 71,372.29 653.41 68,545.39 1,218.03 26,351.61 666.31 23,169.49 1,219.00 D lG [%] 0.02 (0.01 0.02 0.07 0.01 0.04 0.01 0.03 0.20 0.30 0.20 0.30 0.24 0.19 0.17 0.19 0.04 0.14 0.05 0.09 F 1 [N] 68,866.50 2,657.36 66,835.78 5,224.77 17,990.90 4,385.43 9,637.26 8,770.51 51,006.00 2,647.89 48,696.98 5,257.09 71,226.80 2,678.20 67,549.09 5,334.71 26,339.80 2,664.44 22,373.90 5,301.81 with F 1 is reaction measured from Working Model; F and D 2 is reaction obtained by calculation;= 304 F 1−F F 2 2 ×100 [ ] is the difference between the two methods. Dang, A.T. – Nguyen, D.-V. – Nguyen, D.-N. =5m F 2 [N] 68,876.76 2,657.27 6,846.14 5,224.64 17,992.64 4,386.01 9,638.13 8,771.54 51,102.72 2,650.97 48,839.48 5,261.63 71,372.29 2,681.35 67,706.62 5,341.76 26,351.61 2,664.92 22,385.29 5,302.57 D [%] 0.01 0.00 0.02 0.00 0.01 0.01 0.01 0.01 0.19 0.12 0.29 0.09 0.20 0.12 0.23 0.13 0.04 0.02 0.05 0.01 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 Model soft are to measure the reactions at the joints (refer to Fig. 8). The obtained results from this step are then compared ith the results calculated using e uations, as presented in Table 1. The first t o models represent the results of the same structure ith cylinder lengths of 0.46 m and 0.74 m, respectively. The remaining models depict structures ith random cylinder orientations. The differences of less than 0.3 , demonstrate the accuracy of the calculation model in determining the reactions in the mechanism. This indicates that e can accurately determine the design parameters ithout the need to construct or simulate the system in 3D virtual models, hich can be time-consuming. For detailed analyses, constraints such as the range of frame angular movement are applied (γm i n 10 and γmax 120 ), the operational length of the cylinder (50 mm for zero-stroke and 80 mm for full-stroke, corresponding to λm i n 0.5 and λmax 0.8 for the given model ith a 1 m), and the maximum thrust force F Cyl 16 k . Based on this information, t o graphs of lifting ratio k h and the maximum loading for the cylinder at zero-stroke (lo est position) are constructed, as presented in Fig. . The graph also indicates that to achieve the highest efficiency for the cylinder (maximum lifting ratio), the cylinder orientation must be selected as β1 0.31 and β2 0.65; Additionally, the loads in the other joints of this system can also be calculated, as sho n in Fig. 10. E . (3) sho s that the position of the platform is determined by the displacement of the cylinder l Cyl and the arrangement coefficients β1, β2. This implies that Fig. 9. Graphs constructed from the given data (β1, β2); a) lifting ratio kh, and b) maximum thrust force FCyl Fig. 10. Structure of the new system: a) highest and lowest positions of the lift; and b) reactions on joints corresponding the movement of platform Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts 305 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 The results from Fig. 11 sho that hen the cylinder extend or raises the platform at a slo er rate, there is a smaller difference in velocity during the movement. onversely, faster movement of the cylinder leads to system instability, as varying velocity not only causes vibration but also creates acceleration and inertia forces. Another application of the parametric method is selecting the appropriate cylinder in case the position of the cylinder has been given. For example, if the dimensions for the cylinders arrangement are β1 0.3 and β2 0.6, the structure of the cylinder (length, stroke and force) can be calculated and summarized as sho n in Fig. 12. According to the re uirement for structural stability, the angle γ bet een frames should be adjusted bet een 10 degrees and 120 degrees, meaning the cylinder length should be selected bet een 0.31 meters and 0.8 meters, according to E . (2). Suppose the platform is re uired to be raised 0.5 meters. In that case, e can either select the cylinder ith a 32 cm initial length and 12 cm stroke the lifting velocity of the platform can be determined by taking the derivative of E . (3): h  A 2 1  2  2   1   2     (23) , here λ is the extend rate of the cylinder. The e uation also indicates that by controlling the oil pump system, designers have the ability to modify the system s operation, particularly the movement of the platform. This feature can be achieved by using a time-control unit for the pumping system, allo ing for complex platform movements. such as ac uiring the complex movement of the platform by using a timecontrol unit for the pumping system. Fig. 11 illustrates the platform movement for a specific system selected from Fig. 10, in hich the cylinder moves at different constant speeds of 1 m s, 2 m s, and 5 m s. This corresponds to operation times for cylinder extension from zero stroke to full stroke length of 60 s, 150 s, and 300 s, respectively. Platform's displacement h [m]1.00 Platform's displacement Platform's velocity 2 Platform's velocity 1 Platform's velocity 3 Platform's velocity [mm/s] . λ1 = 5 [m/s] 0.75 15 10 0.50 . λ2 = 2 [m/s] 0.25 - 5 . λ3 = 1 [m/s] 0.5 0.55 0.6 0.65 0.7 0.75 0.8 Cylinder's length [m] 0 Fig. 11. Movement of platform for selected structure with cylinder moving extends with different speeds Platform's height h [m] 2 1.75 m 1.5 Maximum thrush F60 cyl [kN] 45 34.88 kN 1 γ<10° 1.25 m 0.76 m γ>120° 30 13.91 kN 0.5 0 15 0.26 m λ2 = 0.61÷0.80 λ1=0.32÷0.44 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 Cylinder’s length [m] Fig. 12. Data of cylinder corresponding to the raising height and loading requirement (a = 1 m; β1 = 0,3; β2 = 0,6; P 306 Dang, A.T. – Nguyen, D.-V. – Nguyen, D.-N. 0 G = 5 kN; W = 75 N) Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 299-307 (corresponding to hm i n 0.26 m and hmax 0.76 m) ith a maximum capacity of P max 34.88 k or select another one ith a 61 cm initial length and 1 cm stroke (hm i n 1.25 m, and hmax 1.75 m) ith a maximum capacity P max 13. 1 k (having a longer stroke but a smaller thrust force re uirement). 4 CONCLUSIONS This study analyses the applicability of the parametric method in designing double-layer scissor lifts. The conclusions dra n from the study are presented as follo s: By using simple dimensional parameters, it is possible to determine the important information of double-layer scissor lifts, such as the platform s height, cylinder s thrush force, and loading in revolution joints. This enables easy component selection ithout the need for complex 3D modelling or experimental tests. The parametric method allo s for efficient selection of the appropriate cylinder based on design parameters and the given re uirements. By analysing the reactions and position of the lift, it is possible to optimize the structure of the system, leading to cost reduction and shorter calculation time. 5 ACKNOWLEDGEMENTS The authors ish to thank Thai guyen University of Technology for supporting this ork. 6 REFERENCES [1] Solmazyiğit, İ., Başkurt, R.C., Ovalı, İ., Tan, E. (2022). Design and prototype production of scissor lift platform 25 tons capacity. The European Journal of Research and Development, vol. 2, no. 4, p. 326-337, DOI:10.56038/ejrnd.v2i4.177. [2] Pan, C.S., Powers, J.R., Hartsell, J.J., Harris, J.R., Wimer, B.M., Dong, R.G., Wu, J.Z. (2012). Assessment of fall-arrest systems for scissor lift operators: computer modeling and manikin drop testing. Human Factors, vol. 54, no. 3, p. 358-372, DOI:10.1177/0018720811425024. [3] Dong, R.G., Pan, C.S., Hartsell, J.J., Welcome, D.E., Lutz, T., Brumfield, A., Harris, J.R., Wu, J.Z., Wimer, B., Mucino, V., Means, K. (2012). An investigation on the dynamic stability of scissor lift. Open Journal of Safety Science and Technology, vol. 2, no. 1, p. 8-15, DOI:10.4236/ojsst.2012.21002. [4] Spackman, H. M. (1989). Mathematical Analysis of Scissor Lifts. Naval Ocean Systems Center San Diego Ca. [5] Kosucki, A., Stawiński, Ł., Morawiec, A., Goszczak, J.. (2021). Electro-hydraulic drive of the variable ratio lifting device under active load. Strojniški vestnik - Journal of Mechanical Engineering, vol. 67, no. 11, p. 599-610, DOI:10.5545/svjme.2021.7320. [6] Karagülle, H., Akdağ, M., Bülbül, İ. (2022). Design automation of a two scissors lift. The European Journal of Research and Development, vol. 2, no. 4, p. 178-191, DOI:10.56038/ejrnd. v2i4.192. [7] Dang, A.T., Nguyen, D.N., Nguyen, D.H. (2021). A study of scissor lifts using parameter design. Sattler, KU., Nguyen, D.C., Vu, N.P., Long, B.T., Puta, H. (Eds) Advances in Engineering Research and Application. Lecture Notes in Networks and Systems, vol. 178, Springer, Cham, p. 75-85, DOI:10.1007/978-3-030-64719-3_10. [8] Todorović, M., Zdravković, N. B., Savković, M., Marković, G., Pavlović, G. (2021). Optimization of scissor mechanism lifting platform members using HHO method. The 8th International Conference, Transport And Logistics, p. 91-96. [9] Čuchor, M., Kučera, Ľ., Dzimko, M. (2021). Engineering design of lifting device weighing up to 3.5 tons. Transportation Research Procedia, vol. 55, p. 621-628, DOI:10.1016/j. trpro.2021.07.095. Applying Parametric Analysis in Enhancing Performance for Double-Layer Scissor Lifts 307 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 © 2023 The Authors. CC BY 4.0 Int. Licensee: SV-JME DOI:10.5545/sv-jme.2023.578 Original Scientific Paper Received for review: 2023-03-17 Received revised form: 2023-04-27 Accepted for publication: 2023-05-15 Machining of H ard-to-cut AISI 4462 D uplex S tainless Steel w ith an Environmentally F riendly A pproach w ith Vortex T ube S nmez, F. Fikret S nmez* Manisa elal Bayar University, Hasan Ferdi Turgutlu Faculty of Technology, Turkey Machining is a manufacturing process that can be used to produce precision machine parts and has many advantages. The first is the ability to achieve superior surface quality. Tool wear is an inevitable phenomenon that occurs during machining. It is affected by many machining conditions; therefore, this process needs to be monitored and controlled. In this study, tool wear and surface roughness tests were carried out on AISI 4462 duplex stainless-steel materials, known to be a hard-to-cut material. For this purpose, tool wear and surface roughness analyses were implemented by using the environmentally friendly vortex tube cooling system in addition to wet turning conditions for the first time. For both methods, experiments were conducted at a 1 mm depth of cut, 120 m/min cutting speed, and 0.1 mm/rev feed with a 90 mm cutting length for each pass. Both tool wear and surface roughness were examined at the end of each pass. The analysis showed that wet turning gave better results in terms of tool life (19.8 minutes of tool life) compared to 11.1 minutes of tool life in vortex turning. In contrast, the surface roughness values differed up to two times in some experiments, and the vortex tube experiments gave better surface roughness values in all passes. In addition, the vortex tube experiments showed less built-up-edge (BUE) formation than the wet-turning experiments. Keywords: vortex tube, machining, duplex stainless steel, tool life, surface roughness Highlights Tool wear is one of the most important outputs of the machining process. Estimating tool wear and modifying machining conditions based on it are important. Tool wear is directly dependent on cutting conditions, and unsuitable conditions dramatically increase wear. It is possible to reduce the surface roughness values by using a vortex tube. The surface roughness worsens due to the increase in tool wear. 0 INTRODUCTION Machining has many advantages over many other manufacturing methods. It is possible to achieve lo surface roughness ith machining, produce machine parts ith high dimensional accuracy, and ensure the repeatability of the manufacturing process [1]. It is also kno n that machining processes are influenced by many factors, especially cutting speed, feed, and depth of cut [2] and [3]. During machining, friction occurs bet een the tool and the orkpiece [4]. Due to this friction, the cutting tool ears out over time [5]. Tool ear is inevitable, although it is possible to reduce tool ear by choosing the right machining conditions. Machine parts can be machined ith dry and et machining methods [1]. In addition, the machining process can also be carried out using an environmentally friendly vortex tube. It is possible to transmit air cooled do n to 50 to the contact area bet een the cutting tool and the orkpiece by using a vortex tube. The vortex tube, also kno n as the an ue-Hilsch vortex tube, is a piece of e uipment that separates compressed air into t o s irling streams, one hot and the other cold [6]. In addition, different cooling temperatures can be obtained by using various pressures and gases [7] and [8]. Thus, 308 it may be possible to improve tool life and surface roughness. Machining can be used to produce parts for a ide variety of needs from many different materials. Ho ever, the machinability of some materials is more difficult compared to other materials. Stainless steel is one such material. Stainless steel materials are hardto-cut due to their lo thermal conductivity and high hardening tendency [9]. Duplex (ferritic-austenitic) stainless steels are one of the toughest stainless steels [1] and [10]. Stainless steel is one of the essential materials used by many sectors, despite being a hardto-cut material. In addition, the development of these materials by various methods may increase the use of s t a i nl e s s s t e e l [11] and [12]. Therefore, numerous researchers have attempted to investigate this material group in detail. Szczotkarz et al. [13] analysed the values of flank ear and crater ear on AISI 316 stainless steel under different machining conditions. The researchers use dry, minimum uantity lubrication (M L) and minimum uantity cooling lubrication (M L) ith the addition of extreme pressure and anti- ear (EP A ) cooling methods. The researchers emphasized that tool ear is accelerated under inappropriate cooling conditions. They also highlighted that ear is intense and tool life is completed uickly, *Corr. Author’s Address: Hasan Ferdi Turgutlu Faculty of Technology, Department of Mechanical Engineering, Manisa Celal Bayar University, Manisa 45140, Turkey, sonmezfikret@gmail.com Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 especially in dry machining conditions. hen et al. [14] investigated the machinability performance of stainless-steel materials under different cutting parameters, specifically, cutting speed, feed, and depth of cut under dry cutting conditions. In their study, they inspected tool ear and surface roughness values. In the 18-minute tool life experiments, the researchers found that the tool ear first increased rapidly, then progressed at a constant speed, and then increased rapidly again. They also stated that increased tool ear had a negative effect on surface roughness. Derani et al. [15] performed tool ear and surface roughness tests on AISI 316 (austenitic) stainless steel under dry-cutting conditions. The researchers examined both nose ear and flank ear in the experiments and emphasized that nose ear accelerates faster than flank ear initially. They found that these t o types of ear gradually increased, and finally, nose ear increased significantly. Ho ever, they also emphasized that despite an increase in tool ear, the surface roughness remained largely constant. Singh Bedi et al. [16] investigated tool ear on AISI 304 stainless steel under dry cutting conditions. They emphasized the increase in tool ear values over time. The researchers found flank ear, crater ear, builtup-edge (BUE), abrasion ear, and adhesion ear on the cutting tool in their investigation. The researchers also emphasized that the coating of the cutting tool is lost over time due to ear. Pek en and Kalyon [17] carried out a tool ear analysis on AISI 430 stainless steel. Three different feeds and cutting speeds ere used in these analyses. As a result of the analysis, the researchers suggested that the cutting speed and feed directly affect the tool life. The researchers also determined notch ear, crater ear, and flank ear on the cutting tool. a mur [18] carried out analyses of tool ear and surface roughness using dry machining, M L, and vortex tube cooling methods. The analyses ere carried out using different cutting speeds and feeds. It has been observed that the results are close to each other in the experiments ith the vortex tube and the experiments ith the M L. Therefore, in this study, it as understood that the vortex tube could be an alternative option to et machining for machining. In their experiments ith the vortex tube, ukor et al. [6] pointed out that in some machining experiments, vortex machining may offer better tool life than et machining. The researchers also discovered that overall production costs could be minimized using this environmentally friendly method. Pinar et al. [19] emphasized that similar results can be obtained ith these t o cooling methods in their experiments in hich they compared vortex tube and et machining processes. Thus, the researchers stated that it is possible to perform et machining processes using the vortex tube method. Valic et al. [3] used different cutting speeds, feeds, and depth-of-cut parameters in their studies on 20 r13 martensitic stainless steels. The researchers combined the vortex tube and M L system and proposed an alternative method for machining stainless steel. In this proposed method, the air-oil mixture (vortex tube M L) resulted in an additional temperature reduction of -38.1 at the tool- ork interface compared to the use of M L alone. The number of studies in the field of sustainable machining has been increasing. Korkmaz et al. [20] adopted dry, M L, and nano-M L methods for machining hard-to-cut nickel-based superalloys. As a result of the experiments, the researchers concluded that tool ear could be reduced by up to 60 using sustainable methods. Airao et al. [21] used ultrasonicassisted turning under dry, et, M L, and L 2 methods on hard-to-cut Ti-6Al-4V material. As a result of the experiments, the researchers proposed that L 2 and ultrasonic vibration could considerably reduce the specific cutting energy ithout sacrificing surface roughness or tool life. Shah et al. [22] used environmentally friendly lubrication techni ues in their study. Electrostatic minimum uantity lubrication (EM L), hybrid nanoparticles immersed in EM L (H EM L), and electrostatic lubrication (EL) techni ues ere used in experiments ith precipitation-hardened stainless steel. As a result, it is comprehended that the H EM L method provides up to 10 lo er po er consumption compared to other methods. In contrast, the EL techni ue has the lo est tool ear. In addition, the researchers proposed that it is possible to reduce the R a value by approximately 30 ith the EL techni ue. In the examination of the existing literature, it is seen that hile there are numerous studies on stainless steel, the number of studies on duplex stainless steel is limited. In addition, the number of studies using the vortex tube cooling method used in this study is very limited. Also, no study as found in hich tool ear and surface roughness values ere analysed over time. In this study, tool ear analysis as carried out using AISI 4462 duplex stainless steel (a hard-tocut material). Time-dependent analyses of tool ear and surface roughness values ere also conducted. Furthermore, the effect of tool ear on the change in surface roughness of the machined part as thoroughly investigated. Machining of Hard-to-cut AISI 4462 Duplex Stainless Steel with an Environmentally Friendly Approach with Vortex Tube 309 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 1 MATERIALS AND METHODS 1.1 Materials In this study, AISI 4462 ( 2 r iMo 2253) duplex stainless steel materials ith a diameter of 75 mm and a length of 110 mm ere used in the experiments. The chemical and mechanical properties of this material are sho n in Table 1. Mechanical properties Values C 0.015 Cr 22.69 Yield strength [MPa] 488 Mo 3.11 Ni 4.84 Tensile strength [MPa] 692 1.2 Methods In the experiments, in addition to the conventional cooling approach of et turning (5 concentration), an environmentally friendly vortex cooling approach as also used. A vortex tube ( an ue-Hilsch vortex tube) as used to achieve a temperature of 23 at 5 bar pressure and measured using a K-type thermocouple and a Fluke 17 multimeter (Fig. 1). Table 1. The chemical and mechanical properties of AISI 4462 Element wt.% ear more clearly, the cutting speed as preferred as 120 m min. IS 3685 [23], the primary standard for tool ear testing, as used to determine the cutting parameters other than the cutting speed. For this purpose, the parameters recommended by IS 3685 ere used, i.e., a depth of cut of 1 mm, a feed of 0.1 mm rev, and a cutting-edge radius of 0.4 mm. Mn 1.40 Hardness [HBW] 218 In the machining experiments, T M 160404M3 ( rade TP2500) cutting inserts supplied by SE , hich ere developed for P-class materials (steels), ere used to clearly observe the tool ear. The utilized insert as suitable for average cutting conditions (designated as the P25 class). In addition, this insert as coated ith Ti ( , ) Al2 3 layers by VD (chemical vapour deposition) method. Moreover, a PT 2020K16 tool holder ith a 0 approach angle is preferred. It is crucial to determine the cutting speed, hich has a prominent effect on tool ear. Depending on the material used (AISI 4462 duplex stainless steel), the cutting speed recommended by the tool manufacturer (SE ) is used as a reference. Ho ever, to observe the effect of Fig. 1. Vortex tube Turning experiments ere carried out ith a cutting length of 0 mm in each pass. After each Fig. 2. Experimental setup 310 Sönmez, F. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 Fig. 3. Measurement setup; a) tool wear measurement, and b) surface roughness measurement turning pass, the surface roughness values of the orkpieces ere measured according to the IS 21 20-3:2021 (formerly IS 4288) standard [24]. The tool ear as measured by removing the cutting tool from the tool holder. Both flank ear values and notch ear values ere observed in tool ear examinations. In addition, the cutting tools ere examined by scanning electron microscope (SEM) and SEM energy dispersive -ray (SEM&ED ) analysis at the end of all the experiments (Fig. 2). The surface roughness values ere measured ith a Mahr M300 terminal and a Mahr Marsurf D18 measuring device via Bluetooth connection. The tool ear measurements ere performed using a Mahr MM200 toolmaker s microscope e uipped ith an M-shot MD30 camera via digital image analysis soft are (Fig. 3). 2 RESULTS AND DISCUSSION 2.1 Tool wear Tool ear is one of the principal factors to monitor in machining. During the experiments, the cutting tool as removed from the tool holder after each pass, and the tool ear as measured, as sho n in Fig. 3. The tool ear inspections performed at each pass are sho n in Fig. 4. In Fig. 4, time-dependent tool ear (flank and notch) can be clearly seen. Time-dependent tool Fig. 4. The tool wear (flank wear) inspections at each pass Machining of Hard-to-cut AISI 4462 Duplex Stainless Steel with an Environmentally Friendly Approach with Vortex Tube 311 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 ear analysis according to tool ear inspections and measurements sho n in Fig. 4 is sho n in Fig. 5. Single-point cutting tools are typically used in lathes. The tool life control of these cutting tools is carried out according to IS 3685 [23], hich specifies certain limits for tool ear. If the tool ear is regular, the flank ear limit value (V b ) is 300 m, and if the ear is irregular, the notch ear limit value (V bm ax ) is 600 m. Therefore, both flank and notch ear values ere measured in the tool ear measurements. The experiments ere terminated hen the flank ear or notch ear value reached the limit value. Both flank ear and notch ear increased rapidly in both et turning and vortex cooling conditions (Fig. 5). In both cooling conditions, notch ear increased faster than flank ear; this is a kno n situation. Derani et al. [15] found that notch ear is more than t ice as high as flank ear in the first phase of tool ear. The flank ear, in contrast, increased rapidly and then continued to do so at a constant rate almost to the end of the experiments. Acceleration of flank ear to ards the end of tool life is a kno n fact [15] and [25]. Ho ever, in this study, an acceleration zone of the flank ear value as not detected. otch ear is directly influenced by cooling conditions. The ear rate of the notch ear decreased ith the et turning. In the experiments performed ith the vortex tube, it as determined that the value of notch ear increased dramatically at the end of the 6t h pass. In the experiments, flank ear and notch ear, hich are the tool life criteria, ere reached almost simultaneously. As a result of 1 .8 minutes of tool life under the et turning conditions (14t h pass), the tool life as completed according to both flank ear and Fig. 5. Tool wear Fig. 6. The SEM images of the cutting tools SEM images of cutting tools; a) wet turning after 19.8 minutes of cutting time, and b) vortex turning after 11.8 minutes of cutting time 312 Sönmez, F. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 notch ear criteria. In the experiments ith the vortex tube, the tool life as completed after 11.1 minutes (7t h pass). hen both cooling methods are evaluated together, it is comprehended that the experiments ith the vortex tube have an efficiency of 56 compared to the et turning. SEM analyses ere carried out to compare the cutting edges concerning each other. Fig. 6 sho s both the nose and top surfaces of the cutting tools employed ith both cooling conditions. Similar types of tool ear are observed on the tools in both cooling conditions ( et and vortex tube). The BUE formation is commonly observed in the machining of stainless steels [4]. Ho ever, in the et turning conditions, significant BUE formation as detected on the top surface, hereas in the vortex tube experiments, BUE formation as limited (Fig. 6). The BUE formation observed in the et turning cutting tool as also investigated using SEM-ED analysis (Fig. 7). Fig. 7a presents three selected areas for inspecting the chemical composition of the cutting insert via SEM-ED . As can be clearly seen in Fig. 7d, the chemical composition includes significant amounts of r (21 ), i (4 ), and Mo (2 ) as determined by the examination made in the BUE observed area (red rectangle). The composition is consistent ith the chemical composition of the AISI 4462 material used in the experiments and directly confirms the formation of BUE. In addition to the high content seen in Fig. 7b, the presence of many other elements indicates both the loss of the insert coating and the apparent BUE formation. A significant Al content as detected in Fig. 7c. It is comprehended that it is an un orn surface since the cutting tool coating had Ti ( , ) Al2 3 layers. 2.2 Surface Roughness Measurements Results ne of the most significant features desired in the mechanical parts manufactured ith machining is the achievement of the expected surface roughness values. In this study, the surface roughness measurements ere carried out on the computer numerical control ( ) turning centre ithout removing the orkpiece. Five different surface roughness measurements ere taken on the orkpiece, and the average of these values as calculated. The average surface roughness values obtained are illustrated in Fig. 8. Fig. 7. The SEM-EDX analysis of the wet turning cutting tool after 19.8 minutes of cutting time; a) top surface of the cutting tool, b) EDX analysis of area 3, c) EDX analysis of area 2, and d) EDX analysis of area 1 Machining of Hard-to-cut AISI 4462 Duplex Stainless Steel with an Environmentally Friendly Approach with Vortex Tube 313 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 It is kno n that surface roughness can vary depending on tool ear. In many studies, an increase in tool ear causes the deterioration of surface roughness values [26] and [27]. Selecting the appropriate experiment conditions has an outstanding influence on the surface roughness. Maruda et al. [28] used sustainable M L methods in turning steel materials in their study. In addition, M L M L&EP A as also investigated. As a result of the study, it as encountered that the surface roughness could be improved by up to 2 ith M L and up to 70 ith M L&EP A . Therefore, a detailed study of the effect of cooling conditions on surface roughness as conducted. In the et turning condition, it is observed that the surface uality deteriorates over time due to tool ear. Ho ever, it is comprehended that the surface uality has improved after 15.1 minutes of tool life. This situation is also seen in some literature studies [15]. utting tools used in turning can provide better surface uality as a result of ear to ards the end of their life [15] and [29]. The most effective parameter in evaluating surface roughness is the average surface roughness (R a) parameter. Ho ever, the R a parameter can only provide general information about the surface; therefore, the maximum height of the profile (R z ) parameter is also analysed (Fig. ). Fig. 8. The average surface roughness (R a) values Fig. 9. Maximum height of profile (R z ) values Interesting results ere obtained in this study. hile it ould be expected that the surface roughness values ould be better under et turning conditions, better results ere ac uired in the vortex tube experiments. The feed and cutting speed values used in the experiments and the chip breaker form of the cutting tool directly affect the chip formation. In the examinations, it as comprehended that the cutting tool could not partially perform the chip-breaking function under the current experimental conditions, and the chip could not be removed from the cutting area at some moments. Although the tool life is completed early in the experiments ith the vortex tube, it is understood that much better surface uality can be obtained because it helps to remove the chips from the orkpiece surface. As a result, much better surface uality as achieved in the experiments ith the vortex tube. In addition, this surface uality obtained as not directly affected by tool ear, and superior results ere obtained throughout all passes. This is thought to be due to the prominent BUE formation seen in cutting processes using et turning in tool ear analyses (Fig. 7a). Similar to the R a parameter, hich gives principal information about surface roughness, vortex cooling creates better surface uality in the R z parameter (Fig. ). Under et machining conditions, the R z parameter is 5.45 times of R a parameter on average, under the vortex turning conditions R z parameter is 5.75 times of R a parameter on average. Tool life, orkpiece surface roughness, and sustainability are the main topics of discussion in industrial machining processes. ithin the scope of this study, it as found that the use of the vortex tube, hich may contribute to sustainability, could be a satisfactory alternative in industrial practice, especially since it offers better surface uality. 314 3 CONCLUSIONS Machining is a complex process influenced by numerous factors. Tool life and surface roughness are of immense importance in controlling this process. In this study, tool ear and surface roughness experiments ere carried out on the material AISI 4462 ( 2 r iMo 2253), hich is duplex stainless Sönmez, F. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 308-316 steel and belongs to the class of hard-to-cut materials. The experiments ere conducted at a cutting speed of 120 m min, a depth of cut of 1 mm, and a feed of 0.1 mm rev. Experiments ere conducted ith both et and vortex tube cooling methods for the first time on AISI 4462. Experiments ere carried out ith a 0-mm cutting length for each pass, and at the end of each pass, both tool ear and surface roughness ere investigated. As a result of these experiments, the follo ing results emerged: Tool life under et turning conditions as completed after 1 .8 minutes (14t h pass) based on flank ear and notch ear standards. Ho ever, in the experiments ith the vortex tube, the tool life as completed in 11.1 minutes (7t h pass). The BUE formation on cutting tools as observed in both methods. In both et and vortex cooling conditions, both notch and flank ear ere rapidly orn ithin 1.4 minutes of cutting time (first pass). Both types of ear continued to increase at a constant ear rate under et turning conditions. Ho ever, under the vortex cooling conditions, the tool as rapidly orn, especially after 6.6 minutes of cutting time, and the tool life as completed. Similar types of ear ere observed in both et and vortex cooling conditions. In particular, abrasion ear and the BUE formation ere evident. Interestingly, ho ever, the vortex tube experiments revealed less BUE formation than the et-turning experiments. The interesting situation as presumably an indication of chip control problems. In vortex cooling conditions, better chip control as achieved in all experiments. In particular, the selected lo feed value (0.1 mm rev) may have caused the insert s chip breaker geometry not to function appropriately. Since the loss of chip control is due to the lo feed, further studies may investigate increasing the feed values to analyse the effects of et and vortex cooling methods. Better surface roughness as expected under the et machining conditions; ho ever, better surface roughness values (R a and R z ) ere determined in the experiments ith the vortex tube. In general, vortex cooling resulted in t o times better surface uality, and the lo est surface roughness as obtained ith about 0.65 m for R a and about 3 m for R z . This could be explained by the fact that the BUE formation as lo er in the experiments performed ith the vortex tube. In addition, this situation sho ed that a possible conse uence of the chip control problem may be seen in the surface roughness. Under et machining conditions, the surface roughness deteriorated up to t o times due to tool ear, reaching approximately 2.2 m. This deterioration as not observed in the vortex tube experiments. Stable surface roughness values ere mostly obtained in the vortex tube experiments, hile decreasing surface roughness values ere obtained in some experiments. Ho ever, it is interesting to note that some improvement in surface roughness values as detected to ards the end of the tool life under et machining conditions. The vortex tube, a sustainable approach, could be a suitable alternative in industrial applications, especially because it provides a higher surface uality. According to the data obtained ithin the scope of this study, investigating the use of a vortex tube ith the M L system in detail in future orks may be of interest. Experiments can be performed by using M L and a vortex tube together. It ould also be interesting to use nanoparticles in M L applications. In addition, these processes can be investigated using the finite element method. 4 REFERENCES [1] Stephenson, D.A., Agapiou, J.S. (2018). Metal Cutting Theory and Practice. CRC press, Boca Raton, DOI:10.1201/9781315373119. [2] Tzotzis, A., García-Hernández, C., Huertas-Talón, J.-L., Kyratsis, P. (2020). 3D FE modelling of machining forces during AISI 4140 hard turning. Strojniški vestnik - Journal of Mechanical Engineering, vol. 66, no. 7-8, p. 467-478, DOI:10.5545/svjme.2020.6784. 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Maruda, R.W., Wojciechowski, S., Szczotkarz, N., Legutko, S., Mia, M., Gupta, M.K., Nieslony, P., Krolczyk, G.M. (2021). Metrological analysis of surface quality aspects in minimum quantity cooling lubrication. Measurement, vol. 171, art. ID 108847, DOI:10.1016/j.measurement.2020.108847. Niaki, F.A., Mears, L. (2017). A comprehensive study on the effects of tool wear on surface roughness, dimensional integrity and residual stress in turning IN718 hard-to-machine alloy. Journal of Manufacturing Processes, vol. 30, p. 268280, DOI:10.1016/j.jmapro.2017.09.016. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 © 2023 The Authors. CC BY 4.0 Int. Licensee: SV-JME DOI:10.5545/sv-jme.2023.577 Original Scientific Paper Received for review: 2023-03-17 Received revised form: 2023-05-16 Accepted for publication: 2023-05-30 Inventory R isk D ecision-Making Techniq ues U sing Customer Behaviour Analys is Marc, I. Berlec, T. Ivan Marc Tomaž Berlec* University of Ljubljana, Faculty of Mechanical Engineering, Slovenia More recent research shows the significant impact of accurate demand forecasting on the operation of supply chain system and thus on the performance of the company. Inventories in the production process could represent waste, which results in higher storage costs and consequently a higher product price, which in turn reduces company's competitiveness on the market. Nevertheless, a company must implement a lean production process and consequently carefully control storage and inventory costs. The introduction of a lean production process is closely linked to the risk of stock-outs, and knowledge of this risk in relation to customer habits is therefore a useful piece of information for the line manager's decision-making. This paper will present a mathematical model that relates customer demand for a product to the inventory level in the warehouse or between the work operations of the production process and the risk of potential penalties that arises with the introduction of a lean production process. With this model we can simulate, how to improve the production processes with still acceptable risk, with the goal of achieving a balance between stocks and the leanness of the production process. The paper demonstrates the use of a mathematical model on a concrete example from practice for risk simulation when choosing different production scenarios resulting from changed customer behaviour. Keywords: lean production, customer demand, risk simulation, inventory optimisation Highlights A mathematical model is presented that allows fast and easy simulations of inventory management based on consumer buying behaviour. The probability density function of demand was chosen in the mathematical model. The changes in the production process can be monitored in terms of inventory costs and the risk that there may be insufficient products available to meet customer needs. By simulating three different scenarios, the optimum inventory level was found to be 30,000 pieces, at which the inventory costs are minimal with a moderate risk of 17 %. 0 INTRODUCTION o adays, companies decide to implement lean production hich aims to reduce costs by eliminating all activities that do not add value [1]. In doing so, companies usually first decide to reduce inventories, that create unnecessary costs, mainly by tying up capital and occupying production areas. f course, caution should be exercised hen reducing inventory, as excessive inventory reduction can lead to business conse uences and risks. It is important to determine here the lo est limit of inventory minimization can be and hich parameters have a dominant influence. The aforementioned problem can be described ith a mathematical model that can serve as a basis for simulating various scenarios that management can predict based on hat is happening on the market. This paper presents the results of a follo -up study [2] in hich, by using a mathematical model, a significant correlation bet een a product demand E [ X , the optimal inventory level z 0 and the average cost of product (A P), as ell as the impact of excess inventory and storage costs E [ C] i n t he production process on the average cost of product as demonstrated. An excess inventory level in the production process as found to raise the average cost of product A P, eaken a company s position on the market and increase the company s inefficiency. To this end, control charts have been developed that can be used to control the inventory level in the ork operations of the production process and, conse uently, to control the average cost of product A P. The paper [2] concludes ith the measures to be implemented to correct the production flo y0 and to correct the optimal inventory level z 0 in the production process depending on the changed demand E [ X . Introducing lean principles into the production process means making changes to the existing production process, hich often entails risks such as interruptions in the production process, nonfulfilment of contractual obligations to customers, loss of good ill, loss of market, loss of revenue, noncompetitiveness, payment of penalties, etc. This paper ill therefore focus on identifying a risk β that may arise in the production process as a conse uence of inventory level optimisation. A shortage of needed products may mean that customer re uirements are not met and the conse uences may be penalties, loss *Corr. Author’s Address: University of Ljubljana, Faculty of Mechanical Engineering, Aškerčeva 6, SI 1000 Ljubljana, Slovenia, tomaz.berlec@fs.uni-lj.si 317 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 of customers, loss of market share or even loss of market. Using a mathematical model and simulation, e ill demonstrate the possibility of implementing a lean production process using an example of inventory optimisation, so that line management ill also have control over the risks associated ith inventory optimisation in the production process. The mathematical model allo s to simulate different production scenarios according to the customer s re uirements and to determine the level of inventories so that the average cost of product is minimised. The mathematical model can also be extended along the entire supply chain, since every link in the chain, especially the eakest one, is important for the performance, competitiveness and profitability of the entire supply chain. The aim of this paper is to present a mathematical model and its use in a simulation to find the best ratio bet een customer demand for a product, the optimal inventory level, including the safety inventory, the optimal storage costs, risk, lean production process due to inventory and the average cost of product A P. An important advantage of the simulation based on the developed mathematical model is to eliminate those scenarios here the A P could exceed the market price due to too high storage costs. 1 LITERATURE REVIEW In the study [2], the authors identify the impact of the optimal total inventory level in the production process on the average cost of product A P. Using a mathematical model, in hich customer demand and an intermediate inventory level in the company s production process and storage costs are linked together, a negative impact of excessive intermediate inventory level on the final price of the product can be identified. In this paper, the authors prove the assumption that, in extreme cases, storage and inventory costs can cause the average cost of product A P to rise even to a price level higher than the price recognised by the market. Seyedan and Mafakheri [3] state that demand forecasting and planning refer to forecasting the uantities and timeframes of customer re uirements. They are of the opinion that the objective of such forecasts is to achieve customer satisfaction by meeting customer needs in a timely manner. Accurate demand forecasting could improve the efficiency and robustness of production processes and related supply chains, as resources ill be matched to demands, leading to a reduction in inventories and aste. 318 Uhrin et al. [4] investigated the role of a company s external and internal resources and the impact of their variability on the degree of lean production implementation. They presented an analysis of the effects of environmental risk and past operational performance of a company on the level of lean production implementation. The paper contributes to explaining the circumstances that ultimately lead to the implementation of lean production. As a conse uence, the external and internal environment influences a company s commitment to increase lean production. Petropoulos et al. [5] assess that demand forecasting is a prere uisite for decision-making on inventories and plays a key role in supply chain management. Ho to improve the accuracy of forecasts has al ays been a focus of research in academia and business, and the study argues that forecasts are designed to help business decisions and should be assesses on the basis of their economic implications. The study compares the performance of several commonly used forecasting methods in terms of achieving inventory control objectives, taking into account a simulation approach. The authors in [6] argue that for any company, in addition to reducing lead times, cost reduction is also a necessity. Therefore, monitoring and controlling manufacturing costs over time can be an important driver for improvement. In [7], t h e authors use stochastic and hybrid models, hich they consider to be very close to reality. They explain the structure of supply chains, the decisions to be taken in a typical supply chain and the models developed for supply chain planning and optimisation. The paper further explores simulation and optimisation to solve stochastic and hybrid models, their applications in the supply chain domain and future research directions arising from the emphasis on sustainability, robustness and resilience of supply chains and opportunities. Authors in [8] highlight the accuracy of demand forecasting as it has a significant impact on the performance of the supply chain system and thus on the performance of company s business operations. An accurate forecast ill allo a company to make the best use of its resources. Synchronisation of customer orders ith production is crucial for a timely fulfilment of orders. This paper presents a system that can improve the accuracy of demand forecasting for more efficient inventory management, also in a Smart Industry 4.0 concept. In their study [9], Pusztai et al. presented a method for implementing a risk-adjusted production plan. T o examples of order allocation ith inventory costing ere presented, one ith inventory costing and one ithout inventory costing. The authors consider that Marc, I. – Berlec, T. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 such information can play a key role in the initial contract negotiations as it also includes potential risks. In their study [10], Perera et al. conclude that the success of a supply chain is highly dependent on the effectiveness of inventory management by taking into account customer behavioural patterns, hich is largely neglected by decision-makers in production. In their study [11], Mahesh ari et al. pointed out the problem of managing big data also in terms of customer behaviour. The role of big data analysis related to customer behaviour in supply chain management (S M), logistics (LM) and inventory management (IM) is of paramount importance for inventory optimisation. Soares do Amaral et al. in their study [12] highlight the importance of simulation for evaluating different hat-if system scenarios to reduce costs and risks especially in industry and service 4.0. The use of modern decision-making support analytical tools and simulation are therefore indispensable to ensure customer satisfaction and cost reduction. In the study [6] and [13], the authors present a frame ork for value flo optimisation by combining value flo cost and cost-time profiles. Value flo mapping represents a very effective tool for visualising activities ithin a production flo , focusing on the duration of activities in order to eliminate non-value-added activities. A good cost analysis system is useful and applicable in helping managers to understand the detailed costs of different short-term and long-term activities and processes, hich is hy stochastic production planning models have been proposed by various authors [14] and [15]. The paper [16] also proposes a stochastic production planning model to reduce risk, for a manufacturing company ith seasonal demand and market gro th uncertainty, to prevent excess inventory and stockout. In the paper [17] even a t o-stage stochastic Linear Programming approach is proposed to investigate a production planning problem here the nonhomogeneous characteristics of logs result in random process yields hich are modelled as scenarios ith discrete probability distributions. Paper [18] focuses on developing a stochastic approach to costing systems that considers the variability in the process cycle time of the different orkstations in the assembly line and provides a range of values for the product costs, allo ing for a better perception of the risk associated to these costs instead of providing a single value of the cost. The proposed model allo s a better analysis of the margins and optimization opportunities as ell as investment appraisal and uotation activities. eal production systems are characterised by a high degree of variability and uncertainty, hich can have a drastic impact on the price of a product. Uncertainty factors in production processes are mainly demand, cycle time and available resources. Afonso et al. [18] also used a stochastic approach to product costing in a production process. Based on their ork, it is possible to further analyse the variation in costs associated ith risks. They consider that the use of descriptive statistics makes sense because it gives the ability to understand and evaluate the behaviour of cycle times and their impact on costs. The approach allo s for a better detection of the risks associated ith product costs. In his study [19], the author concludes that non-economic theories, such as psychological or psychoanalytical theories, also allo a better understanding of other factors influencing consumer behaviour that need to be taken into account. onsumer behaviour is considered to be a holistic approach and consumer behaviour is based on perceptions to ards a product. The theories and models developed in the study can serve as a basis for determining consumer buying behaviour. In [20], adhakrishnan et al. investigate the effective inventory management in the supply chain and argue that inventory management is one of the important areas of supply chain management. They have developed a novel and efficient approach using a genetic algorithm that clearly determines the maximum possible excess inventory level and the shortage level re uired to optimise inventory in the supply chain. In the paper [21], Prasert attana et al. investigate material ordering and inventory control in supply chain systems. The effect of control policies is analysed under three different configurations of supply chain systems. The authors consider that the problem is solvable using an evolutionary optimization method kno n as differential evolution (DE). The results sho that the incentive scheme compliance policy is appropriate and outperforms other policies and can improve the efficiency of the hole system and of all members in the supply chain management frame ork. Based on the literature revie , a ne methodology for decision-making support on inventory levels is proposed in conjunction ith an analysis of the changing customer habits based on a mathematical model and simulation. e have not yet found a similar solution in the literature. 2 METHODOLOGY Kno ing the purchase behaviour of customer, a probability density function of the product demand can be determined. This information is provided by Inventory Risk Decision-Making Techniques Using Customer Behaviour Analysis 319 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 the sales department, using data on sales in previous periods and forecasts of future sales. The probability density function of product demand f X (x ) is approximated as closely as possible to the actual purchase behaviour of the customer. The chosen probability density function, (E . (1)) of the demand X for product is a good approximation of the behaviour of customer demand in practice [2]. X  f X  x   a 2 x  e  ax .  0 2 , a (2) 2 = x, a y0 = x . (3) (4) If the probability density function of customer demand f X (x ) is kno n, the probability density function of arehouse stock fluctuations f Z (z ) (E s. (5) to (7) can also be determined. Z  y0  X , Z  z0  Z , X  z0  y0  Z , Z  fZ  z   f X  x  f Z  z   a 2  z0  y0  z   e f Z  z   ae dx , dz  a  z0  y0  z  (5) , (6)  az0  2   az  e az , (7) here Z is inventory fluctuations due to variable demand X , y0 constant production flo e ual to the average demand uantity E [ X ] , X variable product demand, a parameter of probability density function of customer demand, z 0 initial inventory level, and z inventory level in arehouse or bet een orking operations. Fig. 1 sho s the inventory cost fluctuations p, q , h in a arehouse or bet een ork operations. herein in Fig. 1 C(z ) represents storage cost ⌊€ ⌋, C(z ) p storage cost of a product hen it is not in stock € P , C(z ) q safety inventory costs € P , C(z ) h storage cost of a product hen it is in stock € P , s safety inventory P , and z the number 320   az0  2  arehouse P , and P (1) If the function f X (x ) is multiplied by x and integrated on the interval 0 to , mathematical hope E [ X is obtained. The theoretical mean value of E [ X ] (E . (2)) is then e uated ith the mean value x obtained from real customer demand data (Eg. (3)) and the constant production flo y0 (E . (4)). E  X    x  f X  x  dx  of pieces in stock in the represents piece. Fig. 1. Cost fluctuations in a warehouse or between work operations Here, the key uestion is, of course, ho much inventory and ho much safety inventory can there be bet een ork operations, and hat are the costs and risks of: 1) not paying penalties for not delivering products to the customer on time and 2) having excess production in the arehouse and potentially being uncompetitive because of excessive storage costs The simplest solution is to opt for more inventory in the arehouse or bet een ork operations, hich means that excess production is created, hich in turn, from the point of vie of lean principles, represents astes such as redundant transport, arehousing, processing, poorer product uality, increased arehousing costs, inefficiency, uneconomic and conse uently uncompetitive business. The solution to the problem and the elimination of aste is seen in the introduction of a lean production process ith a focus on optimising inventory levels in the arehouse and bet een ork operations. e believe that this is feasible if e have a good kno ledge of the customer s buying behaviour and of the risk hen interruptions (shortages) may occur in the production process. In this case, e can set an objective that the total aste in the production process or logistics chain should be minimal, hich can be achieved by minimising the initial inventory z 0. Marc, I. – Berlec, T. Fig. 2. The figure shows the probability density function of inventory fluctuations Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 e no define a mathematical algorithm in E s. (8) to (10) that ill relate the parameters ith each other, namely z 0 (optimal initial inventory level), the risk β of being out of stock at a given moment in the arehouse or bet een ork operations due to interruptions in the production process, the storage c os t s E [ C as a function of p, q and h, the safety inventory s, the customer buying behaviour E [ X and t he pe na l t i e s Q that e ill have to pay if e fail to meet demand due to a production failure. Fig. 2 sho s the probability density function of inventory fluctuations in the arehouse or bet een t o ork operations in relation to the costs p, q and h. The mathematical algorithm is given by E s. (8) to (10). E C  z    z0    y0  z0 s  0 p zf Z  z  dz   q zf Z  z  dz s  3 CASE STUDY 0 zh f Z  z   Q  , (8) 1 az  2 E C  z    z0   hz0  e 0   a  p  q  aQ  e as  q  h    as  1 az0  4  p  q   3aQ  e as  q  h    as  2  , 2 dE C  dz0 ( )  0, z0 optim  he az0  2   min, To determine the minimum value according to the proposed mathematical method, a simulation model has been developed and is sho n in Fig. 3. The simulation is carried out using the standard soft are tool MS Excel and is therefore easy to use, so that the line management can use it henever there is a change in the input parameters, hich depend mainly on market conditions. The follo ing variables appear in the simulation model, z 0 is initial inventory level P , s safety inventory level P , β risk , E [ C average storage cost as a function of z 0 € , and Q penalties that have to be paid due to unfulfilled re uests to the customer €. 2  3  p  q   2ah  e as  q  h   as  2    as  1    a  p  h  aQ  e as  q  h   as  1  (10) In the selected company, e have carried out a material value flo analysis in the production process of a Basic hinge (Fig. 4). e have focused on the cold forming and surface protection operations. Using the proposed model, e ant to determine ho much inventory should be held to keep the process lean and to identify the risk β that may arise from optimising the inventory level z 0. The capacity of the Basic hinge assembly system is 68,500 P per day, hich corresponds to the customer demand ordering an average of y0 E [ X 50,000 P of hinges. The input data of the model sho n are p, q, h, s, Q , z 0 and x . The parameter a depends on customer buying behaviour, so cannot be influenced, and x represents actual monthly orders. hen e enter the Fig. 3. Mathematical algorithm for simulation in Excel Inventory Risk Decision-Making Techniques Using Customer Behaviour Analysis 321 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 input data into the Excel soft are using our algorithm, e get the output data β, E [ C and z 0. hen e solve the e uation dE [ C (d z 0) 0 using numerical calculation methods, e obtain the optimum z 0opti m, here the cost E [ C is a minimum. In the simulation e use the hat if experiments . In our example e are interested in the function E [ C] ( a, h, p, s, q, Q , z 0, β). β depends indirectly on a and z 0. Using this function, in addition to z 0, e can simulate the behaviour of the cost for different values of a, h, s, q, p, Q and β. If e ant to obtain the optimal value z 0opti m immediately, it can also be obtained by the exact mathematical E . (10) and the given parameters a, h, s, q, p and Q ithout simulation. In our case, s c a n also be referred to as the strategic stock necessary for the smooth running of the production process. In this paper, e have limited ourselves to presenting three scenarios using a practical example. There are many possible combinations, and the choice of the best balance bet een inventory, costs and risks is in the hands of management. e have found some similar studies in the literature that indicate that solving the problem from the point of vie of the balance bet een inventory, costs and risks is important and at the same time interesting for the academic community. e believe that the proposed model is therefore interesting and useful for practice. It can be seen from Fig. 4 that the uantity of products transported daily bet een the t o operations, or the daily uantity demanded by the customer (Q td. ) i s Q td 50,000 P d, here d represents ork day. The uantity of products bet een the t o operations or the uantity of products in transition bet een the t o operations (Q bo ) i s Q bo 466,666 P . ther used variables in Fig.4 are: W represents number of orkers at the orkplace, W P orkplace number, T p process time, T c cycle time, T s setup time, A availability, T ps process time for series, Q s uantity in one series processed and S scrap. The status bet een the t o operations is an interesting data point for us, because it is clearly an overproduction hich represents a cost or aste for the company. e ill verify the claim that aste is in uestion here by applying the proposed mathematical model (E s. (8) to (10)). Let us assume that the daily storage cost of the cup semi-finished product is h 0.5 € P , the cost of storing the safety inventory of the cup semi-finished product q 1 € P , the cost of the cup semi-finished product hen it is not in stock p 2 € P , the uantity of the safety inventory of the cup s 1,000 P , and the penalties to be paid if there is an interruption in the production process and e are not able to fulfil the customer s order Q 10,000 €. The customer s average daily order is E [ X 50,000 P of Basic hinges. The follo ing findings are also important for the simulation, including the data that represent limitations in the optimisation process. The inventory level of the cup semi-finished product z 0 bet een the cold forming and surface protection operations in the existing production process is excessive and is in fact overproduction and aste. This is in contrast to a lean production process. Since the inventory level is excessive, the risk β of production failure is practically non-existent or negligible. The storage costs of excessive uantities of the cup semi-finished product are high and therefore increase the average cost of the cup semi-finished product, raising the price of the finished Basic hinge product and thus eakening the product s position on the market. Excessive inventories generate a range of different astes (transport, additional storage space, redundant processing, increased changeover times, uestionable uality and increased defect volumes). Fig. 4. Material value flow for the Basic hinge 322 Marc, I. – Berlec, T. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 4 ANALYSIS AND DISCUSSION e have selected three scenarios to analyse the lean production process by varying the inventory level bet een the cold forming and surface protection operations. The simulation as carried out in MS Excel environment. Scenario 1: Since the daily demand is 50,000 P , the line manager in the first scenario has decided on of the an initial inventory z 0 30,000 P , i.e., 60 daily safety inventory, hich can be considered as an optimistic decision because in this case there should be no production failure of more than half a day. Scenario 2: In the second scenario, the manager has increased the initial inventory to 60,000 P , hich represents 110 of the daily safety inventory, hich can be considered a plausible scenario. Even if there is a slight change in demand, there is still one day s safety inventory if a production failure occurs. Scenario 3: In the third scenario, the initial inventory as further increased to 100,000 P , hich represents 200 of the daily safety inventory, hich can be considered a pessimistic decision because in this case there is a safety inventory for t o days of production failure. Using the mathematical model and the simulation soft are developed in MS Excel, the line manager in charge of production found out very easily hat the storage costs E [ C (z 0) and the risk β are for each scenario chosen. Fig. 5 sho s a possible choice bet een the optimal inventory z 0, the storage costs E [ C (z 0) and the risk costs β. The coloured boxes sho the values for the selected values for the 3 scenarios previously proposed that could allo a lean production process in the case of reducing the inventory level of the cup semi-finished product bet een the cold forming and surface protection ork operations. Fig. 5. Possible choice between the inventory z 0, the storage costs E [ C (z 0) and risk costs β Fig. 6 sho s the results from Fig. 5 in graphical form. Fig. 6 clearly sho s here the minimum value E [ C (z 0) is, and hence the optimal inventory level z 0 i s Fig. 6. Storage cost fluctuations E [ C (z 0) as a function of the selected inventory level z Inventory Risk Decision-Making Techniques Using Customer Behaviour Analysis 0 323 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 determined. A more detailed analysis of the simulation results for all 3 scenarios is as follo s: Scenario 1: The storage costs are the lo est, E [ C (z 0) 28,247.71 €. As a conse uence, there is a slightly higher risk of production failure β 17 . In Fig. 5, Scenario 1 is indicated by the red box and in Fig. 6 it is indicated by the arro s. Scenario 2: It represents higher storage costs E [ C (z 0) 34, 10. 3 € but a lo er risk of production failure, β 6 . In Fig. 5, scenario 2 is indicated by a black box. Scenario 3: It represents even higher storage costs E [ C (z 0) 51,23 .37 € but the lo est risk of process disruption β 1.7 . iven the target function, the optimal decision is a t m i n z 0. This target is reached in the first scenario, here the selected uantity of the initial inventory is 30,000 P and the cost E [ C (z 0) 28,247.71 €. According to the production management, a risk of 17 is acceptable. Hypothetically, bearing the lean production principles in mind, e should have opted for a production ithout inventories, but in this case the costs are 43,037. 5 € and the risk is 40.6 , hich is too high a risk. e have noted that in the mathematical model, and therefore in the simulation, the dominant variable is Z or z 0 (optimal value), since this variable affects both the cost E [ C and the risk β. The larger the inventory Z , the higher the cost and the lo er the risk, and vice versa. Thus, the management s decision depends mainly on the developments on the market. The greater the competition, the leaner the production process must be (because of lo er costs) and the uantity Z must be close to z 0, even if the risk is higher. If the risk is higher, it means that the market situation must be closely monitored, and the simulation must be run again ith the proposed model to choose a ne inventory strategy adapted to the conditions. Since e did not find a similar approach to inventory minimization in the literature, it is not possible to compare the results ith other similar studies. Therefore, the management of the company in uestion has analysed the results in detail and found that they are useful for decision making regarding the level of inventory. 5 CONCLUSIONS Inventory management is an important aspect of managing the production process and the entire logistics chain. This paper presented a mathematical model that allo s fast and easy simulations of inventory management based on consumer buying 324 behaviour. For the mathematical model, the probability density function of demand f X (x ) as chosen hich, in our opinion [1], best approximates the actual reallife customer buying behaviour. Based on the chosen function f X (x ), the changes in the inventory fluctuation z 0, the storage costs E [ C (z 0) and the risk costs β c a n be calculated in the production process as a function of the changed customer demands. This paper is a follo -up to a previous study, hich presented the negative effects of too high inventory levels of semi-finished products per time unit of observation on the average cost of the A P product and thus on the final price of the product. This time, e have extended and additionally focused on the risk that may arise hen deliberately reducing inventory levels bet een the ork operations of a production process ith the aim of approaching the desired lean production process. Logically, reducing the inventory level bet een ork operations increases the risk of interrupting the uantities available to supply the customer and, of course, vice versa. In the practical case under consideration, the proposed model as used to simulate three possible scenarios of a lean production process on an example of inventories, storage costs and risk costs. nly t o ork operations have been chosen, cold forming and surface protection, as an example, but the model could be extended very easily to the hole logistics chain. In the present case, there as clearly overproduction and thus aste. Bet een the t o ork operations observed, an inventory of as many as 466,666 P as determined. By simulating three different scenarios, the optimum inventory level as found to be 30,000 P , at hich the inventory costs are minimal ith a moderate risk of 17 . The presented simulation method, based on a mathematical optimisation model, is a simple tool that provides the line management ith an efficient decision-making support regarding the uantities to be produced in a given time interval, depending on the customer behaviour. The simulation can also be useful for the sales management to analyse, during the negotiations ith customers, hat the desired uantities mean for the production costs and to negotiate the selling price accordingly. The scientific contribution of the presented model and simulation for decision-making support is that by analysing consumer buying behaviour, the changes in the production process can be monitored in terms of inventory costs and the risk that there may be insufficient products available to meet customer needs. There are fe useful solutions in the literature for solving these problems, hich are every-day Marc, I. – Berlec, T. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 317-325 problems in manufacturing. The limitation of the study is that the proposed method is only suitable for batch production. ur further research ill focus on the development of a cost, analysis stream map ( ASM) model that ill offer a detailed overvie of the entire production process in terms of demand, average cost of product A P and risk. By incorporating ASM, the proposed method ill be suitable for the optimisation of lean processes depending on customer behaviour also in industry 4.0. and can be integrated in the E P or MES companies. 8 REFERENCES [1] Tanasić, Z., Janjić, G., Soković, M., Kušar, J. (2022). Implementation of the Lean Concept and Simulations in SMEs - a Case Study. International Journal of Simulation Modelling, vol. 21, no. 1, p. 77-88, DOI:10.2507/IJSIMM21-1-589. [2] Marc, I., Kušar, J., Berlec, T. (2022). Decision-making techniques of the consumer behaviour optimisation of the product own price. Applied Sciences, vol. 12, no. 4, art. 2176, art. ID 2176, DOI:10.3390/app12042176. 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Stochastic Chebyshev goal programming mixed integer linear model for sustainable global production planning. Mathematics, vol. 9, no. 5, art. ID 483, DOI:10.3390/math9050483. [16] Zhang, X., Prajapati, M., Peden, E. (2011). A stochastic production planning model under uncertain seasonal demand and market growth. International Journal of Production Research, vol. 49, no. 7, p. 1957-1975, DOI:10.1080/00207541003690074. [17] Zanjani, M.K., Kadi, D.A., Nourelfath, M. (2013). A stochastic programming approach for sawmill production planning. International Journal of Mathematics in Operational Research, vol. 5, no. 1, p. 1-18, DOI:10.1504/ijmor.2013.050604. [18] Afonso, P., Vyas, V., Antunes, A., Silva, S., Bret, B.P.J. (2021). A stochastic approach for product costing in manufacturing processes. Mathematics, vol. 9, no. 18, art. 2238, DOI:10.3390/math9182238. [19] Goswami, K. (2020). To buy or not to buy: An analysis of the problematic of consumer behaviour. 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Licensee: SV-JME DOI:10.5545/sv-jme.2023.541 Original Scientific Paper Received for review: 2023-01-31 Received revised form: 2023-03-28 Accepted for publication: 2023-05-15 Finite Element Analys is of Notch D epth and Angle in Notch Shear Cutting of Stainless-Steel Sheet Engin, K.E. Kaan Emre Engin* Adiyaman University, Faculty of Engineering, Turkey Piercing is a crucial process in the sheet metal forming industry, and the surface quality of the pierced part is an important parameter that defines the overall quality of the product. However, obtaining a good surface quality is a challenging task that depends on the effect of several process parameters and necessitates the use of non-conventional procedures. Notch shear cutting is a relatively new progressive approach in which the workpiece is indented with a notch form with a predefined notch angle and depth, and then the indented workpiece is subjected to conventional piercing. In this study, conventional piercing and notch shear cutting processes were experimentally performed on 1.4301 stainless steel sheet of 2 mm thickness. Then, finite element (FE) analyses were conducted utilizing Deform-2D software. After ensuring that the experimental and simulation works were consistent with each other, the FE analysis of notch shear cutting was carried out for three distinct notch depths (15 %, 30 %, and 60 % of the workpiece thickness) and six different notch angles (10º, 20º, 30º, 40º, 50º, and 60º). Investigations were performed on shear zone length distributions, which are direct indications of the surface quality on sheared workpieces, crack propagation angles, and required cutting loads. The best surface quality was obtained when the notch angle was set to 50º and the notch depth was set to 15 % of the workpiece thickness. It was also observed that notch angle and notch depth had a certain level of influence on required cutting load. Keywords: metal cutting, notch cutting, piercing, surface quality Highlights The primary objective of the study was to determine the influence of notch angle and notch depth on surface quality and required cutting in notch shear cutting of 1.4301 sheets of 2 mm thickness. Both experimental and finite element (FE) analyses were performed to validate the consistency of two methods. Deform-2D software was utilized for numerical analysis whereas a die set was manufactured for experimental work. The overall best quality which surpassed the surface quality of piercing was achieved with a notch angle of 50º and a notch depth of 15 % of the workpiece thickness. It was determined that when the notch angle increased, the required cutting load decreased. This was also true for increasing notch depth, but the load incurred during the generation of notch depth must be included in the total energy consumption. 0 INTRODUCTION Piercing is a fundamental process in the sheet metal forming industry and is defined as the cutting of the sheet metal from the stock ith a die set consisting of a predefined shaped punch and a lo er die. The primary purpose of the piercing process is to avoid orkpiece re ork and maintain good surface uality. Burrs and protruding surfaces create an impediment to achieving this goal. Burrs can cause early tool ear, reduced corrosion resistance and because of their sharp formations, can become the primary source of accidents. These negative impacts result in poor overall process production uality. To remove these formations from the surface, conventional grinding and other cleaning procedures are necessary, resulting in additional labour hours, higher costs, and specific people being ithdra n from the main operating cycle and allocated solely to these deburring and cleaning activities [1]. Due to the nature of the process, numerous parameters should be adjusted and understanding the 326 effect of these parameters is the key step to achieving good cutting results and lo er energy consumption. Adjustments in clearance [2] and [3], cutting speed [4], orkpiece and tool material [5], tip form [6], t o ol coatings [7] and ear characteristics [8] t o [12] have significant effects on both surface uality and energy aspects in conventional piercing procedures. Ho ever, a relatively recent procedure kno n as notch shear cutting has been developed ith the goal of improving the surface uality of sheared sheet components. A notch ith specified characteristics is pushed onto the orkpiece in this progressive method. The operation then proceeds ith the conventional piercing method. Fe studies have investigated the notch shear cutting of sheets and related parameters. Studies investigating different aspects of the notch shear cutting proved that good surface uality ith reduced burr formations can be achieved by using this method. Sachnik et al. [13] used orkpieces D 04, 1.4301, E A -6014 and uSn6 ith 1 mm thickness and performed both closed and open cuts ith a *Corr. Author’s Address: Adiyaman University, Faculty of Engineering, Altinsehir Mah. 3005. Sok. No:13, Adiyaman, Turkey, kengin@adiyaman.edu.tr Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 special progressive die. They mainly investigated the formation of burrs and ho they change according to the notch placement on the orkpiece. Both simulation and experimental orks had been accomplished. The investigations ere done for different values of notch cutting parameters as notch depth (30 to 60 of the sheet metal thickness), notch s radius (0.05 mm, 0.125 mm, and 0.2 mm) and clearance (6 , 8 , and 10 of the material thickness). They sho ed that the burr formation is directly linked ith the position of the notch and notch height. They also stated that around 40 notch heights gave the most burr-free surfaces. Krinninger et al. [14] investigated the effect of notch parameters on the orkpiece by using both experimental and finite element method (FEM) for sheet materials made of aluminium alloys E A 5754 and E A -6014 ith a thickness of 1.0 mm. The clearance bet een the die and punch as 3 of the orkpiece thickness. They designed a progressive die set to perform the experimental ork. otches ere both applied from the topsides and do nsides of the orkpiece; 25 and 40 of the notch depths and a fixed value of 60 as the notch angle ere applied on the orkpieces. They discovered that hile the topside notch could lo er the re uired cutting forces, in general, there as no reduction in cutting forces hen the percentage drop in sheet thickness as considered. ith do nside notches, they observed that a surface distribution ithout burr could be achieved. Feistle et al. [15] performed notch shear cutting studies on press-hardened components (PH ) made of aluminium-silicon-coated manganese-boron steel (22MnB5) sheets at room temperature. The shearing of PH steels at room temperature causes difficulties due to their higher mechanical properties. Moreover, the earing of tools progresses rapidly, hich contributes to the formation of more burrs at the edges of the orkpieces. The orkpieces had 1.5 mm thickness. The die clearance as 15 of the orkpiece thickness. The notch angle as 60 . otch depths ere 16 and 40 of the orkpiece thickness. Three notch variables ( ithout notch, topside notch, and do nside notch) ere examined by both FEM and experimental methods. They measured the zone distributions on the sheared surfaces and calculated the re uired shearing force. They stated that cutting loads ere reduced for notched specimens and provided improved surface uality free of burrs. The existence of a notch on the orkpiece introduces additional parameters to be investigated. The mentioned studies investigated some values of notch depth, but the notch angle values ere generally limited to 60 . Furthermore, none of the studies investigated the effect of these parameters on stainless steels hich are one of the most used materials in industrial applications. The surface uality and re uired cutting load change according to notch depth and notch angle for stainless steels should be investigated ith varied values of these t o parameters. If correct parameter adjustments can be established for 1.4301 sheets during notch shear cutting and high surface uality can be produced, then a technical advantage can be ac uired over conventional piercing techni ues. To achieve this, experimental and computational validations for conventional piercing and notch shear cutting ere conducted initially. Follo ing the consistency of the findings bet een experimental and computational ork, three different notch depths (15 , 30 and 60 of the orkpiece thickness) and six different notch angles (10 , 20 , 30 , 40 , 50 , and 60 ) ere employed to execute the notch shear cutting of 1.4301 stainless steel sheet; the influence of these variables on the surface uality and re uired cutting load ere investigated using finite element analysis. 1 PROCESS PARAMETERS 1.1 Notch Depth, Notch Angle, and Surface Distribution on the Workpiece As an addition to the parameters of conventional piercing, notch shear cutting introduces ne parameters such as notch depth, hich is the percentage of the height of the notch to the thickness of the orkpiece; notch angle, hich is the angle value of the notch indented on the surface of the orkpiece; and notch radius, hich is the radius given to the tip of the notch. otch shear cutting is a progressive process. The related illustration of the steps of the process as given in Fig. 1. First, a notch indenter ith a predetermined angle, depth, position, and radius is manufactured (Fig. 1a). In the second step, the indenter is forced against the orkpiece, resulting in the development of a notch (Fig. 1b). Then, in the last step, the notched orkpiece is subjected to conventional piercing and the process is completed (Fig. 1c). Shearing of the material is a complicated phenomenon. hen a orkpiece is pierced, it goes through a series of stages before being entirely ruptured. ones typical of piercing are generated on the orkpiece s surface during these stages. ollover (die roll) zone, shear (burnished) zone, fracture zone, and burr formation are the four zones mentioned. By measuring these zones and observing ho much Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting of Stainless-Steel Sheet 327 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 they dominate the overall surface length, the surface uality of the sheared orkpiece can be assessed. has the potential to reduce burr forms hile extending the shear zone length. a) b) Fig. 2. Zone distributions on the sheared surface c) Fig. 1. The illustration of the steps of notch shear cutting and related components; a) manufactured indenter with predetermined notch properties, b) formation of notch on the workpiece and, c) conventional piercing of the notched workpiece ollover and burrs are the least desired zones on the surface, and investigations often involve methods and parameter modifications for eliminating these zones. Stahl et al. [16] implemented t o-stage shear cutting instead of one-stage shear cutting of truck frame parts and revealed the effects of different process parameters on the improvement of surface uality and burr formations along ith fatigue enhancement of the sheared parts. Mucha and Tutak [17] studied the impact of clearance on the burr size formed on a thin steel sheet ith a 55 H hardness during the blanking of angled hooks, as ell as the ear characteristics of the punch. ishad et al. [18] collected and compared the methodologies used to improve the process and elaborated on ho the parameters of the process impact the result. A long shear zone length is the most desired situation. The illustration of these zones as given in Fig. 2. Denting the surface ith a notch before piercing 328 otch angle and depth are essential parameters that can influence crack propagation and orientation as ell as zone distributions. Determining the optimal values of these parameters are important. The upper limit of the notch angle value as set using a fine-blanking approach. enerally, a notch is produced on the orkpiece prior to cutting using a v-ring indenter to keep the orkpiece in place and the side angle of the v-ring indenter is normally kept constant as 60 in related procedures like fine blanking. From this point of vie , the upper limit of the notch angle as set at 60 ; moreover, 10 , 20 , 30 , 40 , and 50 ere added to observe ho they influenced the results. From lo to high, notch depths ere set at 15 , 30 , and 60 of the orkpiece thickness, respectively. 2 EXPERIMENTAL METHODOLOGY The first step in the study as to conduct the experiments. Although this study relied heavily on finite element (FE) analysis, computational results should be cross-checked ith experimental data. If computational and experimental ork for piercing and notch shear cutting could be verified to be consistent ith one another, the simulation ork ould be expanded to study the influence of other notch shear cutting parameters. The stock material as 1.4301 stainless steel sheet. The mechanical characteristics of the material have crucial importance in shear-cutting processes. Due to this situation, the mechanical characteristics of 1.4301 sheet ere determined as the beginning phase of the research. If the mechanical characteristics Engin, K.E. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 of the orkpiece material ere defined ith errors in the soft are, the results ould almost certainly differ from those obtained via experimental ork. 1.4301 stainless steel sheet is used in a variety of industrial applications, including mining, maritime and structural applications as ell as the manufacture of bolts, scre s, and containers [19]. Tensile tests ere used to get the stress-strain graphs that the soft are re uired. A stock 1.4301 stainless steel sheet as supplied from local sheet metal vendors. The dimensions of the stock material ere 2000 mm 1000 mm ith 2 mm thickness to satisfy the tensile testing standards. The chemical composition and the mechanical properties of the material according to the manufacturer s sheets are given in Tables 1 and 2. Fig. 4 depicts the metallographic structure of 1.4301 sheet observed by using a scanning electron microscope ( eiss emini SEM 500). A Shimadzu A - Plus tensile testing machine as used to measure the engineering curve of 1.4301 sheet. To reduce the margin of error, the tensile tests ere repeated six times at room temperature for each rolling direction at a strain rate of 0.05 s-1. Table 1. Chemical composition of 1.4301 sheet Composition C Mn P Cr Weight [%] 0.08 2.0 0.04 19.0 Composition Si S Ni Weight [%] 0.75 0.03 9.0 Fig. 4. Metallographic structure of 1.4301 stainless steel sheet Table 2. Mechanical Properties of 1.4301 sheet Properties Tensile strength Yield strength Value 515 MPa 210 MPa Properties Elongation at break Hardness (Rockwell B) Value 60 % 80 Using a ater jet, the specimens ere cut out of the stock according to ASTM E8 standards, considering the rolling directions (0 , 45 , 0 ). The goal of using a ater jet is to lessen the thermal effects that arise on the material s surface. The sides of the specimens ere also polished to prevent the notch effect from occurring during tensile tests. Fig. 3 sho s the dimensions of the test specimen (Fig. 3a) and the actual test specimen (Fig. 3b). a) b) First, engineering stress as measured, and then true stress is obtained by using true stress-strain e uations expressed in E s. (1) and (2) as: σt = σ(1+ε), (1) εt = l n (1+ε), (2) here σt and σ represent true and engineering stresses, εt and ε are true and engineering strains. The obtained engineering and true stress-strain curve of 1.4301 sheet ere given in Fig. 5. Fig. 5. True and engineering stress-strain diagram of 1.4301 stainless steel sheet Fig. 3. a) Dimensions, and b) the actual image of one of the tensile test specimens In the second phase, both piercing and notch shearcutting operations ere carried out experimentally. A manufactured die set as used to perform piercing and notch shear cutting. It consisted of six parts: the Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting of Stainless-Steel Sheet 329 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 punch, upper die block, guides, a fixed-type blank holder, lo er die, and lo er die block. The blank holder and die blocks ere constructed of S235J , hile the remainder of the components ere made of heat-treated 1.7225 steel. The die set (Fig. 6a) and its exploded vie (Fig. 6b) are sho n. Due to the progressive nature of notch shear cutting, hich necessitates the formation of a notch on the orkpiece s surface before conventional piercing can begin, a special hollo punch as produced to be used as an indenter. The manufactured hollo punch s tip sides had a 60º notch angle and a notch tip radius of 0.5 mm. The depth of the hollo as adjusted to 1.2 ± 0.1 mm to prevent the hollo punch to penetrate further than the notch depth of 60 of the orkpiece thickness. The total length of the punch as 80 mm. Fig.7 sho s the illustration of the hollo punch (Fig. 7a) and the formation of the notch indenter (Fig. 7b). learance bet een the die and the punch as calculated according to E . (3) as: C = 100((D m-D p) 2t) , a) (3) b) Fig. 7. a) Illustration and b) dimensions of hollow punch in [mm] here D m is the lo er die diameter, D p is the punch diameter and t is the orkpiece thickness. The lo er die diameter as 10 mm; based on E . (3), the calculated and fabricated punch diameter to perform a conventional piercing process as .80 mm. The total length of the punch as 80 mm. Experiments ere conducted utilizing a hydraulic press ith a capacity of 300 k and under a ram speed of 1 mm s. To keep the punches in place, a punch holder as manufactured and attached to the press s ram. a) a) b) Fig. 6. a) The die set, and b) it’s exploded view The die clearance as fixed at 5 of the thickness of 1.4301 sheet, hich as 2 mm, for both experimental and computational ork. 330 b) Engin, K.E. Fig. 8. a) Notch indentations on the workpiece, and b) conventionally pierced workpiece Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 The piercing operation as carried out using the manufactured die set hile the notch shear-cutting process as carried out in a progressive manner. The notch is first produced on the do nside of the orkpiece. Follo ing that, the standard piercing process as performed. Fig. 8 sho s an example of notch creation on the orkpiece surface (Fig. 8a) and one of the typically pierced orkpieces (Fig. 8b), respectively. For the purposes of numerical analyses, the normalized ockroft and Latham fracture criterion ere utilized. 3 NUMERICAL SIMULATION AND MODEL VERIFICATION Deform-2D soft are as used to execute the computational ork. Sheet metal shearing is one of the many forming techni ues that the soft are can model. The soft are employs axisymmetric modelling hich implies that only half of the setup is re uired to run it. The orkpiece as modelled as a plastic object, hereas the punch, lo er die and blank holder ere defined as rigid bodies. Shear friction occurs hen the orkpiece deforms and is a function of the yield stress. The friction bet een the orkpiece and the shearing tools as assumed to remain constant and as calculated as follo s in E . (4): f s = mk , (4) here ƒ S is the frictional stress, k the shear yield stress, and m the friction factor. According to the soft are s database, shear friction as given a value of 0.12. Large plastic stresses are generated over a narro zone bet een the punch and the die during shearing. At that small deformation zone, the material as considered to be isotropic and yielding follo ed Holloman s e uation hich as stated in E . (5) as:   K n , (5) here σ is the effective stress, ε the effective strain, K is the material constant, and n the strain hardening exponent [20]. Deform-2D database includes the material properties for 1.4301 stainless steel. Fig. demonstrated the consistency of the flo stressstrain curve of 1.4301 stainless steel in Deform-2D s database and the computed flo stress-strain curve obtained from the tensile tests by using Holloman s e uation. Fracture initiation in a piercing process usually begins at the punch s or lo er die s contacting edge. For computational studies, the fraction criterion is critical because good criterion choice brings the outcome values closer to the experimental results. Fig. 9. Flow stress-strain curve of 1.4301 stainless steel The criterion indicates that hen effective strain e uals the critical value (C), fracture initiation begins at the point here the value is reached. The criterion as given in E . (6) as: *      d     f  C, (6) 0 here σ* is the maximum principal stress, ε–f t he fracture strain, C the critical value (damage factor), and σ and ε are the expressions of effective stress and effective strain. To simplify the criterion, the ratio bet een maximum principal and effective stress as assumed to be constant at the shearing zone. This assumption resulted in ε–f = C meaning that crack initiation starts at the point here effective strain and critical values become e ual [21]. The critical value, C, can be evaluated by a tensile test regardless of the orking operation. The critical value of 1.4301 sheet as defined as 0.54, hich as ac uired from the conducted tensile test. The soft are employs a step-based iteration mechanism; at each step, it verifies the orkpiece s values. If the values described by mesh formations on the orkpiece approach the critical value, element deletion is applied to the associated meshes to vie the fracture initiation and propagation. This function is governed by t o parameters: fracture steps and fracture elements. The value of fracture steps sets the step interval at hich the simulation pauses and element deletion is performed. This value as selected as 1, the minimal value. This signifies that element deformation is considered at each step. Fracture elements are the number of elements that must exceed Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting of Stainless-Steel Sheet 331 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 the critical damage value for the simulation to halt and delete elements. This value as set as the system default to 4. Automated mesh generation as used and defined to be executed for every 0.02 mm progression of the top die to provide optimum re-meshing. Element deletion improves visualization coherence ith experimental outcomes. Ho ever, negative aspects of this approach include volume deterioration and excessive mesh deletion if mesh element numbers are maintained lo . Lo mesh numbers result in an imprecise representation of the shear region here the zones inter eave. This circumstance results in inaccurate calculations of the cutting load and interpretations of the zone lengths derived from simulation results. The density of the meshes at the shear zone must be as dense, small, and numerous as possible to prevent volume loss and accurate propagation of the fracture, hich enables consistent visualization of the variation in zone distribution. Initially, the meshes on the orkpiece ere evaluated ith varying element counts; ho ever, as the element count reached 8,100, isoparametric uadratic elements, it as determined that there as no change in cutting load above this element count. Tough, to obtain optimum surface distribution on the sheared orkpieces, 10,000 isoparametric uadratic elements hich are the maximum number of elements that can be defined in Deform-2D and ,786 number of nodes ith 0.03 mm element size ere used. The simulation parameters ere given in Table 3. accurate results. Fig. 10 sho s the steps applied during the computational ork of notch shear cutting. First, it as necessary to push the predefined notch indenter into the orkpiece (Fig. 10a) and produce the indentation (Fig. 10b). The notched orkpiece as then placed in a typical piercing setup (Fig. 10c) and the follo ing steps ere carried out (Fig. 10d). a) b) Table 3. Simulation parameters Parameter No. of elements No. of Nodes Element size Fracture steps No. of fracture elements Remesh criteria c) Value 10,000 9,786 0.03 mm 1 4 Every 0.02 mm of the punch penetration In addition, mesh elements ere stacked as closely as possible ith the use of mesh indo s to raise the total number of elements in the shearing zone. For both piercing and notch shear cutting processes, the distribution of zones and the cutting load values ere estimated after conducting both computational and experimental ork. The numerical analysis phase of the notch shear-cutting process re uired t o se uential steps to duplicate experimental conditions for the most 332 d) Fig. 10. The steps of simulation process; a) formation of notch at the downside of the workpiece, b) notched workpiece, c) conventional piercing, and d) completion of the process 3.1 Ideal Crack Condition During a shearing operation, the fracture of the material begins at the corners of the punch and die that are in contact. The creation of a fracture propagates rapidly, resulting in the separation of Engin, K.E. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 material at the intersection ith the opposite corner. The ideal crack (Φ) occurs hen the re uired angle of crack propagation (θ) is directed to ard the punch and die cutting edges. In many instances, ho ever, the real propagation (β) of a fracture does not follo this anticipated pattern, resulting in secondary cracks that cause surface fla s on the orkpiece. hen the angle values of β and θ are closest, the optimal cutting condition may be attained. This situation can be expressed in E . (7) as: Φ β–θ≅0. Fig. 12 depicts a comparison of the change in cutting load bet een simulation and experimental tests. (7) The ideal crack propagation angle can be expressed in E . (8) as:  c    Arctan  , (8)  t  u  p   here C is the clearance, t the sheet metal ( orkpiece) thickness, and up the punch penetration corresponding to the first crack initiation ithin the sheet. The directions of angle β and angle θ ere illustrated in Fig. 11. Fig. 11. Illustration of Ideal crack propagation angle (θ) and real crack propagation angle (β) Fig. 12. Comparison of required cutting load between computational and experimental work The experimental and numerical findings ere ithin 5 range of each other on average. This as the outcome of the experimental environment s external influences such as dust, additional friction bet een tools and load cell accuracy. The rest of the research as carried out ith computational ork once the measurements ere verified to be identical. Figs. 13 and 14 compare the zone distribution for conventional piercing and notch shear cutting, respectively. Fig. 13. Comparison of zone distributions between computational and experimental work (Piercing) 4 RESULTS AND DISCUSSION Both experimental and numerical ork ere performed under the clearance value of 5 of the orkpiece thickness hich as 2 mm, ith a 10 mm lo er die diameter and .80 mm punch diameter for conventional piercing ith a punch speed of 1 mm s. Throughout the notch formation se uence, the hollo punch ith a 60 notch angle, a notch tip radius of 0.5 mm, and a notch depth of 60 of the orkpiece thickness (1.2 0.1 mm) as utilized. During the second phase (piercing after notch generation) of notch shear cutting, the same settings as ith conventional piercing ere employed. Fig. 14. Comparison of zone distributions between computational and experimental work (Notch shear cutting) The computational ork as completed by follo ing the outlined se uential stages to replicate the experimental ork. First, three values of notch depth Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting of Stainless-Steel Sheet 333 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 (15 , 30 and 60 of the material thickness) and 6 different notch angles (10 , 20 , 30 , 40 , 50 , and 60 ) ith 0.5 mm notch tips ere virtually penetrated to the do nside of 1.4301 sheet ith 2 mm thickness. After the material as notched, the piercing stage of the process as simulated using the configuration in hich the lo er die diameter as 10 mm, the punch speed as 1 mm s, and the clearance as 5 of the material thickness. As in the experimental study, the punch diameter as .80 mm, no radiuses ere specified at the punch s edges, and the tip of the notch as aligned ith the punch s cutting edge at shearing in every simulation. e uired cutting loads to shear the orkpiece, fracture propagation angles and zone distributions ere investigated. In cutting operations, cutting load is an important parameter that specifies the re uired loading capacity of the press and plays a crucial role in overall energy usage. It as discovered that the cutting loads necessary to shear notched materials ere less than the load value for the piercing operation (28.5 k ). The deeper the notch penetrated into the material; the less cutting force as re uired at the se uential piercing step. This outcome as anticipated due to the reduction in material thickness follo ing the creation of a notch in the orkpiece. A 15 notch depth resulted in the highest hereas 60 notch depth resulted in the lo est re uired cutting load for each notch angle. For example, hen the notch angle as fixed to 60 , the re uired cutting loads ere obtained 22.2 k for 15 notch depth, 16.6 k for 30 notch depth, 11.8 k for 60 notch depth, respectively. The effect of the notch depth on the re uired cutting load is uni ue to notch shear cutting, regardless of the orkpiece shape, and is consistent ith the literature [22]. Another observation as made on the effect of notch angles. Apart from the notch depth, the re uired cutting load decreases ith increasing notch angle due to the decrease in the cross-sectional area of the orkpiece bet een the punch and die. The changes in re uired cutting load according to notch depths and angles are given in Fig. 15. Although it might be thought that a deeper notch has the advantage of a lo er cutting load, it is important to remember that t o phases are necessary to complete the entire notch shear-cutting process. If load comparisons ere conducted exclusively for the shearing portion of the operation, it might be argued that a notch could reduce the cutting load. To account for each phase of the operation, ho ever, 15 k should be added to the total load values hich as the average of the loads re uired 334 to form the notch into the surface of 1.4301 sheet for each of the depths listed. Fig. 15. Cutting load change according to notch angle and notch depth Fig. 16 illustrated the disparities bet een the ideal angle (θ) of crack propagation and the real angle (β) of crack propagation. Fig. 16. e a ere e et ee real ra ro a at o a le eal ra ro a at o a le The best surface uality can be achieved hen the difference bet een these t o angles is closest to zero. The distribution of zones and the propagation of cracks are interrelated; therefore, an improper crack propagation difference ill lead to secondary cracks and protruding surfaces. Deform-2D can measure distances on the sheared orkpiece in relation to zone distributions, and the character of the cut (smooth or protruded) can be clearly visualized and estimated by using the soft are s post-processor section once the simulation is complete. Fig. 17 sho s the shear zone percentage change according to notch depth and notch angle. A comparison of Figs. 15 and 16 demonstrated that the difference bet een crack propagation angles and shear zone percentages ere in correlation. For piercing processes, the optimal zone distribution Engin, K.E. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 (best surface uality) occurs hen the rollover and crack zone lengths are minimal and free of burr formation hile the shear zone length is maximal; 15 of the notch depth ith 10º , 20º , and 30º not c h angles resulted in burr formation, hich as the least desirable outcome. The development of protrusions reduced as the notch angle increased, and at 50º t he longest shear zone length and the shortest crack zone ith almost no protrusions ere reached. From Fig. 16, it as observed that the closest values to zero amongst all notch depths and all notch angles ere achieved hen notch angle value as selected as 50 and the notch depth value as selected as 15 of the orkpiece thickness. Fig. 17 confirmed this result, the shear zone percentage, 24.56 , as also the highest at these values. Fig. 17. Shear zone percentage change according to notch depth and notch angle The shear zone length of the specimen measured after piercing (22.6 ) could only be surpassed by using these values. This situation sho ed that a technical advantage can be obtained over conventional piercing ith proper adjustments. The relationship bet een crack propagation angles and shear zone percentages as in accordance ith previous studies. Using the angles of crack propagation, Hambli et al. [23] devised an algorithm for predicting optimal clearance. They established a tolerance convergence percentage of 1 and stated that an optimal clearance could be reached hen the difference bet een crack angles satisfied this value, indicating that no secondary cracks occurred, and good surface uality as ac uired. Engin and Eyercioglu [24] and [25] studied the influence of different process parameters on 1.4301 stainless steel sheets ith different thicknesses and diameters. They observed that the best surface uality on the sheared orkpieces as achieved hen the difference bet een crack propagation angles as minimal. Fig. 18 displays the surfaces of sheared orkpieces in relation to the notch depth, notch angle, and independently for piercing, as established by the Deform-2D soft are. A further observation as made on another impact of notch depth. It as determined that the length of the shear zone dropped dramatically hen the notch depth increased, independent of the notch angle. This is due to the orkpiece s pre-deformation during notch formation. hen the orkpiece as subjected to the conventional piercing phase, hich as the second step of the process, shear and fracture zones ere generated follo ing the top portion of the notch. In the case of deeper notches, there remained a reduced region for shear zones to develop, and this as the main reason for the drop in shear zone length. Figs. 17 and 18 demonstrated that the crack zone dominated the surface and the shear zone decreased belo 7 for notch angles bet een 40º and 60º and for 60 notch depth. This case revealed that even though pre-deformation (notch) on the orkpiece appeared to make the shearing region less affected during piercing, deeper notches ere not a guarantee of high surface uality and notch angles should be taken into consideration. egardless of notch depth, notch angles determined the crack propagation course and, in turn, affected the shear zone length. It as clearly understood that the obtained results depended on the combined effect of notch angle and notch depth values. For every notch depth value, the notch angle difference altered the crack propagation. As indicated before, for 15 notch depth, lo er notch angle values resulted in burr formation. hen notch angle values increased, burr formations started to reduce up to a limit here the best surface uality could be obtained. In the event of higher notch depths, such as 30 , the shear zone lengths decreased incrementally from 15 to 10 bet een 10º and 30º notch angles but reached their peak at a 40º notch angle ith a length value of 22.43 , and then started to decrease again. This situation sho ed that for every value of notch depth, there exists a notch value at hich the shear zone length could reach its peak height. This criterion as met ith a 50º notch angle at 15 notch depth, a 40º notch angle at a 30 notch depth, a 20º notch angle at a 60 notch depth for 1.4301 sheet, respectively. Ho ever, it should be clearly stated that the general form of the sheared surfaces is also an important parameter. As mentioned before, the primary objective as to obtain the best possible surface uality to transcend conventional piercing. For deeper penetrations such as 30 and particularly Finite Element Analysis of Notch Depth and Angle in Notch Shear Cutting of Stainless-Steel Sheet 335 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 Notch depth [°] 15 Notch angle [%] 30 60 Piercing 10 20 30 40 50 60 Fig. 18. Visual representation of sheared parts as a function of notch depth and notch angle generated by the Deform-2D software 60 notch depths, the increase in the notch angle resulted in the reduction of burr formation; ho ever, the form of the orkpiece gradually deteriorated and took a conical shape at the do nside. This circumstance is e ually undesirable. At high values of notch depth, it must be considered that the use of ider notch angles may result in the deformation of the sheared material s shape. 336 Typically, the outputs of piercing shearing processes are restricted to the parameters applied specifically for that orkpiece material. The same situation applies to this research. In terms of zone distributions, a 50º notch angle and 15 notch depth outperformed the shear zone length value of conventional piercing to produce the greatest surface uality and form. Ho ever, it is limited to Engin, K.E. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 326-338 1.4301 sheets if and only if these conditions are satisfied. Any change in the notch forming position, clearance, punch die diameters, shearing speed and, most importantly, the orkpiece material modifies every effect and may result in varied outputs. Due to the interdependent interactions bet een regulating parameters, it is difficult to establish an optimal setting for shearing and cutting operations that accounts for all process variables and orkpiece materials. 5 CONCLUSIONS The main aim of this study as to determine the effect of notch angle and notch depth on the surface distribution and re uired cutting force during notch shear cutting of 1.4301 stainless steel sheet orkpieces to fill a gap in the literature due to an absence of research on notch shear cutting of commonly used stainless steels. ompared to piercing, notched shear cutting is a relatively novel method for cutting metal sheets. onventional piercing and notch shear cutting ere carried out to compare the results to those obtained from simulations. The remainder of the study is then completed virtually. Using Deform2D soft are, a FEM study of notch shear cutting as performed for three different notch depths (15 , 30 , and 60 of the orkpiece thickness) and six different notch angles (10 , 20 , 30 , 40 , 50 , and 60 ). The results obtained in this study are summarized belo . For specimens ith deeper notches and ider notch angles, the re uired cutting load as observed to decrease. In addition, cutting specimens ith notches re uired less load than conventional piercing. This is a direct byproduct of the material s decreased thickness due to the notch formation on the orkpiece. Due to the progressive nature of the operation, the load applied during notch formation must be added to the total value of the re uired cutting load, resulting in an increase in total energy consumption. In addition, it may take longer to complete the hole process than piercing, hich is not a progressive operation like notch shear cutting. There as an ideal value for each notch depth and notch angle at hich the sheared surface as as good as possible in comparison to the other values for these t o parameters. This condition as achieved at a 50 notch angle for a 15 notch depth, at a 40 notch angle for a 30 notch depth, and at a 20 notch depth for 60 notch angles, respectively. In contrast to the ideal values, ho ever, values that depart from them might lead to increased surface fla s. In addition, the specimen s base became conical due to the increasing notch depths and notch angles (such as 60 notch depth and 60 notch angle) hich is undesirable. The shear zone length of the specimen estimated after piercing (22.6 ) could only be exceeded by notch shear cutting ith a 50 notch angle and 15 notch depth (24.4 ). The remaining shear zone values ere inade uate for shearing 1.4301 sheets hen compared to conventional piercing. If surface uality is of the highest significance and no re ork of the produced orkpieces is necessary during the production, notch shear cutting can produce better results than piercing, particularly in preventing burr forms. Ho ever, to attain the maximum shear zone length ith no burrs, preliminary ork should be conducted prior to application, since any uncorrected value may cause the shear zone to be shorter than piercing. Hence, obtaining the optimal settings is a complex and expensive operation to be accomplished through experimental trial and error, FEM simulations are re uired to conduct the ork. onetheless, it is a time-consuming endeavour. otch shear cutting seems feasible hen utilized for mass manufacturing. All findings are explicitly limited to the shearing of 1.4301 stainless steel sheets and under the specified parameters. 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Journal of Mechanics Engineering and Automation, vol. 6, no. 7, p. 356-363, DOI:10.17265/21595275/2016.07.00-6. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 339-351 © 2023 The Authors. CC BY 4.0 Int. Licensee: SV-JME DOI:10.5545/sv-jme.2022.456 Original Scientific Paper Received for review: 2022-11-19 Received revised form: 2023-05-13 Accepted for publication: 2023-05-31 R ecent Advancement via Exp erimental Investigation of the Mechanical Characteristics of Sisal and Juncus Fibre-R einforced Bio-Composites haouki, B. Abdelkader, K. Mouloud, A. Benchaabane haouki1,2, Kirad Abdelkader1,2 Aissani Mouloud3 1 FU 2 Department DAPL Laboratory, Faculty of Science, University of Blida 1, Algeria of Mechanical Engineering, Faculty of Technology, University of Blida 1, Algeria 3 esearch entre in Industrial Technologies TI, Algeria In this work, the mechanical characteristics of unidirectional bio-composite materials reinforced by two types of natural fibres (sisal and juncus) were studied in order to develop new materials. The effect of the fibres’ extraction methods and their new assembly techniques on the mechanical properties of the elaborated composites was investigated. This is based on three methods of extracting natural fibres: the first uses water treatment alone over a long period, while the second uses alkaline chemical treatment with a sodium hydroxide solution. The last method uses the burial of plant leaves in moist soil. The obtained fibres are assembled according to techniques, such as monolinear fibres, twisting fibres into rope and braiding fibres into rope. The composite materials are produced manually using a pressure-contact moulding process. The outcomes demonstrated that the resulting compounds’ mechanical properties are significantly impacted by the chemical treatment. The sisal/polyester composites exhibit better mechanical tensile test behaviour than those made with juncus fibres. Moreover, contrary to the results of some other studies, the recently developed techniques of assembling with a chemical treatment process enabled the reduction of the bio-composite’s thickness as well as the cost of its preparation. Keywords: natural fibre, sisal, juncus, mechanical properties, bio-composite Highlights Three procedures are used to extract natural fibres: the first involves prolonged water treatment alone; the second involves alkaline chemical treatment using sodium hydroxide solution; the final technique uses burying plant leaves in moist soil. The sisal/polyester composites exhibit better mechanical tensile test behaviour than those made with juncus fibres. Young’s modulus of the composite reinforced with sisal fibres is twice as high as that reinforced with juncus fibres. The twisted rope assembly technique exhibits the highest values of Young’s modulus compared to the other forms of assembly. The recently developed technique of assembling fibres into a rope with a chemical treatment process has contributed to the best reinforcement of the bio-composite materials. 0 INTRODUCTION The increasing interest in plant fibres is evident due to the proliferation of documents dealing ith the use of these natural or modified fibres in composite materials. These plant fibres are already occupying an important place in the composites industry thanks to their best mechanical and physicochemical properties [1]. They are used in various fields of application, such as transport, construction, medical and leisure [2] and [3]. Among the main advantages of these natural fibres are their availability in several countries, regeneration, bio-degradability, and the possibility of extraction by various methods ithout damaging the fibres. They can also play a remarkable role in the development of ne bio-degradable green materials ith desirable characteristics. This can be used to solve some of the current ecological and environmental problems. atural fibres have become a viable, eco-friendly, and plentiful alternative to expensive, non-rene able synthetic fibres [1]. They present a promising reinforcement for composites for some applications due to their lo cost, lo density, and relatively good mechanical properties [4] and [5]. In addition, they are also characterized by the absence of health risks, easy handling, high flexibility, and sound insulation [6]. atural fibres such as sisal, linen, kenaf, Alfa and jute have been used as reinforcement in bio-composites [7] t o [9]. ur interest in this ork is oriented to ards the search for plant fibres that are generally the most abundant during the year. Furthermore, research is moving to ards the development of bio-composites ith the best possible mechanical properties at a lo er cost. enerally, the mechanical properties of biocomposites are often influenced by several factors, and they heavily depend on the fibre content and, therefore, on the nature and uality of the i m pl e m e nt a t i on [10]. These factors can be grouped into t o types based on their origin: constitutional and structural. Among the elements of the constitutional *Corr. Author’s Address: FUNDAPL Laboratory, Faculty of Science, University of Blida 1, Algeria. chaoukiben27@yahoo.com 339 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 339-351 type is the nature of the plant (kind of fibre), the nature of the resin, the physicochemical characteristics of these components (fibre resin), the geometry of the fibre cells and their porosities, etc. The main elements of the structural type are the fibre-resin structure, the fibres orientation, the size of the fibres, the mass fraction of the fibres, the fibre matrix interface uality, the extraction methods and the manufacturing process of the composite (contact moulding, vacuum pressure moulding, etc.) [10]. Improving the properties of a bio-composite or its performance comes do n to improving the various elements mentioned above. Therefore, among the methods of extracting plant fibres, the method of chemical treatment ith sodium hydroxide ( a H), hich, according to the literature [11] t o [13], facilitates the separation of fibres from the leaf matrix of their plant, by reducing impurities such as pectin, ax, and lignin around the outer surface of the fibres. It also improves the bond bet een the fibres matrix of the composite to give the best mechanical results [14]. Belaadi et al. [15] indicated that the mechanical properties of polymer composites reinforced ith sisal fibres are largely influenced by the mechanical properties of these fibres. Through their studies, they analysed the mechanical behaviour of sisal fibres and compared them ith other bio-composites made of jute and juncus fibres; they observed the existence of the same influence of the fibres on the overall behaviour of the composites. Joseph et al. [14] analysed the effect of ater absorption on the tensile strength of sisal polypropylene (PP) composites at different volume fractions. They found that the maximum tensile stress decreased as a function of the immersion duration. In addition, they found that sisal PP composites ith treated fibres exhibit higher strength than those of composites ithout treated fibres. Also, the mechanical properties of the polymer laminates reinforced ith sisal fibres ere studied, and they sho ed that the fracture stress increased ith the immersion time in ater during the extraction phase of these fibres [16]. In addition, the mechanical properties of the composites (epoxy type) reinforced ith sisal fibres ere examined. The mechanical decortication method as used to extract these fibres, hich ere subjected to chemical treatment ith alkalis and binding agents. These pre-treated fibres sho ed improved mechanical and hydrophilic tendencies compared to untreated fibres. Thus, they resulted in efficient bonding at the fibre polymer matrix interfaces [17]. The researchers also sho ed that the mechanical properties of the developed composites depend on various parameters, 340 such as fibre length, fibre orientation, and fibre volume fraction [17]. Uppal et al. [18] observed that composites containing short shredded sisal fibres have high mechanical properties compared to those formed from sisal fabric. The ageing of these fibres also changes the mechanical properties of the bio-composites compared to those having ne ly chopped sisal fibres. Maurya et al. [19] studied the mechanical properties of the epoxy composite reinforced ith sisal fibres by varying the length of the fibre and keeping a constant eight percentage of the fibres. It as concluded that the tensile strength of the composite as not improved by reinforcing the length of the sisal fibres. In this ork, the influence of three extraction methods ( ith various physical aspects) of sisal and juncus fibres, as ell as the ne techni ues for assembling these fibres, on the mechanical tensile properties of composites ill be studied in order to deduce the best-elaborated bio-composite and propose it to the industry. Although they can be found on other continents as ell, the sisal and juncus fibres used in this study ere taken from orth African plants. 1 EXPERIMENTAL 1.1 Choice of Plants and Natural Fibre Extraction Methods The natural fibres studied ere extracted from sisal and juncus plants, as illustrated by Fig.1. The choice of plants is dictated by the availability in our region of orth Africa and by the rapid rene ability factor during the year. There are several methods for extracting plant fibres [20], mainly according to mechanical, chemical, and biological processes. The first fibre extraction method is mechanical, under ater alone ( ater treatment method T t). It consists of cutting the juncus stalks longitudinally into t o slices length ise and in the same ay for the sisal leaves; then, these sisal leaves and the cut juncus stalks are immersed in a large container of ater. They ill be kept for 4 days to 5 days at room temperature until fully saturated ith ater. The recovered sheets are then ashed and dried, then mechanically tapped and manually brushed to extract the fibres. The obtained fibres from plants are sho n in Fig. 2 (bundles no. 1 and 4). The second method is a chemical method (the alkaline treatment method, TAlk). After immersing the stems or the sliced leaves of the plants in ater for 24 hours, a solution of sodium hydroxide ( a H) at a 7 concentration is added for 48 hours at Chaouki, B. – Abdelkader, K. – Mouloud, A. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 339-351 ambient temperature. ext, they are ell rinsed, and the traces of a H are neutralized in these fibres ith a 2 distilled ater solution of sulphuric acid (H2S 4) immersed for 30 minutes [14]. After ards, they are immersed in distilled ater for one hour to have a neutral potential of hydrogen (PH). The leaves obtained are placed on soft ground, and ith a ooden stick, these leaves and stems are struck and shaken until the fibres are separated. Finally, they are dried in a room at room temperature for 24 hours, Fig. 2, (bundles no. 2 and 3). The third method is biological under et ground (T r). It consists of burying the leaves and stems of sisal or juncus in the et ground for 55 days to 57 days [21]. These sheets must be cut longitudinally beforehand into t o or three parts; this promotes the biodegradation of these leaves, thus making the extraction of the fibres easier. Then, the obtained fibres are ashed ith ater and then dried at room temperature, Fig. 2 (bundle no. 5). or torsion, the second one forms ropes by t isting, and finally, it forms unidirectional straight fibres ithout overlapping, so e named this last assembly monolinear . The strings contain six fibres each. 1.2.1 Assembly in Ropes by Braiding A braid is an assembly of bundles of fibres ith a total of six threads. The different icks of fibres pass alternately bet een them (Fig. 3a), such that the left strand is passed over the neighbouring strand and belo the follo ing one. It proceeds in this ay until the last strand. The crossing of the icks is done at a right angle and offers an obli ue checkerboard pattern (Fig. 3a). 1.2.2 Assembly in Ropes by Torsion In this mode of assembly, the six fibre threads are grouped in parallel and t isted in pairs (2 2 2 threads), then t isted together so as to form the rope by torsion (Fig. 3b). In this second case, the strands of t isted fibres are placed in a spiral hose centre is that of the rope. 1.2.3 Straight Unidirectional Assembly of the Fibres (Mono Linear) Fig. 1. Studied plants in their natural habitat; a) sisal and b) juncus In this last and usual mode of assembly, the fibres are grouped in parallel, uasi-straight and unidirectional lines ithout overlapping (Fig. 3c). Fig. 3. Assembly of fibres into a) ropes by braiding, b) ropes by twisting, and c) monolinear Fig. 2. Some natural fibres obtained by different methods of extraction; no. 1, 2, 5 sisal’s fibres, and no. 3 and 4 juncus’ fibres 1.2 Forming the New Assembly of Plant Fibres The assembly of both studied fibres in the structure of composite materials is manually placed into t o ne rope shapes, ith the addition of the usual form of linear assembly. The first one forms ropes by braiding 1.3 Fabrication of Laminated Composite Plates and Single Resin Plates In order to fabricate our plates of bio-composite materials, e first prepared the plant fibres and then assembled them. Secondly, a polyester-type resin based on three components as also prepared. This type of polyester as chosen because it is the most used and the cheapest. The polyester matrix is therefore obtained by mixing a primary resin ith a Recent Advancement via Experimental Investigation of the Mechanical Characteristics of Sisal and Juncus Fibre-Reinforced Bio-Composites 341 Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 339-351 hardener (7 ) and an accelerator (3 ) by volume. Polyester-only (resin) tensile specimens are obtained from manually prepared plates. These specimens are used as a reference for the bio-composite specimens. Unidirectional sisal polyester and juncus polyester laminated plates (bio-composite) are made in mono-layers, such that the mass of fibres in a plate is approximately 30 0.5 g. The ends of the fibres are attached by double-sided adhesive tape to the edges of the mould during their deposits. This is done to ensure that the directions of the fibres are parallel and straight hen casting the resin manually. The composite plates ere fabricated using the contact moulding method. The thicknesses of the plates are checked using metal edges (2 0.1mm). The bio-composites are impregnated at room temperature (25 1 ). The polyester resin obtained is catalysed and hardened in proportions of bet een 2 and 3 by mass of the primary resin. nce the plates are reticulated, they all undergo a 24-hour polymerization cycle at room temperature before demoulding. In order to have a total polymerization of the polyester resin, the laminated plates are left in the open air for 3 days before being cold cut into test specimens. The bio-composite plates produced in dimensions of 180 mm 170 mm are cut into specimens according to the ASTM 303 standards, as cited by Hassan and Abdullah [22]. In order to estimate the fibres mass rate, each plate as eighed ith a precision electronic balance (0.01 g). The tensile specimens (Fig. 4a) are in accordance ith the ASTM 303 standard [23] and have the follo ing dimensions: Total length L 175 mm, thickness h 2 mm and idth b 25 mm. Fig. 4b sho s some tensile specimens of biocomposites ith edges in their ends. b) Fig. 5. a) Tensile testing machine with zoom under the test, and b) zoom after tensile test bservation analyses of sisal and juncus fibres ithout resin and others into resin ere carried out using optical microscopy ( M) and scanning electron microscopy (SEM). 2 RESULTS AND DISCUSSIONS 2.1 Mechanical Properties of Polyester Resin Fig. 4. a) Schematic specimen with its dimensions; and b) some tensile specimens of bio-composite with wedges in their ends 1.4 Mechanical and Structural Characterization of Biocomposites To characterize the laminated plates produced, the mechanical properties and microstructure of all biocomposites elaborate are studied. The determination of the main mechanical properties is carried out via 342 monotonic tensile tests under a universal machine of the ick- mbH type (Fig. 5a). The tests are carried out ith a pre-load of 1 until failure. To ensure good reproducibility of the results, at least three specimens for each category ere tested at the same speed, hich as uasi-static on the order of 1 mm min. The direction of the tensile forces is the same as the orientation of the fibres. The length of the metal standards ( edge) placed at the end of the test specimens is e ual to 25 mm. The latter serves to prevent crushing under the ja s of the universal testing machine (UTM). The longitudinal direction of the specimens is chosen to be the same direction as that of the fibres. Fig. 5b sho s a bio-composite specimen under axial tensile loading ith a zoom lens sho ing crack initiation. Fig. 6 sho s the curves of tensile testing of specimens in polyester-alone resin, and an example of a broken tensile specimen is illustrated. Each time, three specimens are tested. It is noted that the shapes of these curves are very close to each other except for a slight shift in elongation during the rupture phase. The overall mechanical behaviour of the resin specimens sho s a bilinear stress-strain relationship, follo ed by sudden rupture. The elastic limit is approximately 5 0.5 MPa (Fig. 6). oung s modulus, the tensile strength, and the deformation of each specimen are grouped together in Table 1, associated ith mean values. Chaouki, B. – Abdelkader, K. – Mouloud, A. Strojniški vestnik - Journal of Mechanical Engineering 69(2023)7-8, 339-351 according to the extraction method, is classified by E . (1): ERes Fig. 6. Curves of tensile tests of polyester specimens alone and an example of the specimen after the test Table 1. Summary of the mechanical properties of the polyester resin specimens Specimens Tst1 Tst2 Tst3 Average values Young’s modulus E [GPa] 1.701 1.533 1.811 1.681 ±0.148 Plastic zone modulus E 2 [GPa] 0.871 0.805 0.805 0.827 ±0.044 Maximum tensile strength R m [MPa] 18.25 15.12 16.41 17.33 ±2.21 Rate of deformation [%] 1.80 1.52 1.67 1.66 ±0.16 Table 1 sho s the average maximum stress reaching a value of 17.33 MPa before the specimen breaks. This value is comparable to those given in references [24] and [25]. oung s modulus (elastic zone) has an average value of 1.681 Pa, and that of the plastic zone is around 0.827 Pa, ith an error of 0.15 Pa. Ho ever, the rate of deformation is of the order of 1.66 0.16 . These results ill serve as a reference for the composites to be developed. 2.2 Young’s Modulus of Developed Bio-composites The results of oung s modulus of the specimens of the different cases developed ith the sisal fibres (Fig. 7a) and the juncus fibres (Fig. 7b) are presented in histograms. The different cases studied concern the three methods of extraction (chemical treatment TAlk, ater treatment T t, and stripping in the et ground T r) and the three forms of fibre assembly (assembly in cords by t isting 1 , in cords by braiding 2 , and fibres in straight lines noted mono-linear 3 ). verall, it can be seen that oung s modulus E of the composites is higher than that of the pure resin, and its value has been amplified several times (from 2 to 7 times that of the resin, depending on the nature of the fibres; Fig. 7) by the addition of fibres. From this Fig. 7, the effect of the extraction method on oung s modulus of the composites can therefore be classified according to their values as elasticity modulus E Ref < ETwt < ETGr < ETAlk . (1) This result is valid for both cases of plants regardless of the type of assembly form. From Fig. 7a, oung s modulus of composites reinforced by sisal fibres and obtained by chemical treatment (S 1 TAlk) is compared to the modulus of the same composite obtained ith ater treatment (S 1 T t); it sho s an increase of approximately 70.01 . Ho ever, the comparison of oung s modulus of (S 1 T r) ith that of (S 1 T t) gave an increase in value of about 13.60 . The same observation of an increase in modulus is observed for the other cases of S 2 and S 3, but ith lo percentage values. It is concluded that the extraction method by treatment ith a H has increased oung s modulus value of the composite compared to the other extraction methods. This observation is in accordance ith some of the literature [26] and [27]. oung s modulus of composites reinforced ith juncus fibres (Fig. 7b) evolves in the same ay as that of sisal, according to the different extraction methods, but ith a difference in value. This reveals the importance and interest in using chemical treatment of fibres. Ho ever, there is no case for under et ground (T r) extraction for this juncus plant. ote: this last method of extraction by the under et ground techni ue on juncus fibres as not successful during the tests, so its effect as not studied. This is due to the very fast degradation and deterioration of these fibres during the execution of this method because of the lo amount of pectin in this plant. It can be summarized that for the same nature of fibre and for a given extraction method, oung s modulus E of the composite evolves according to the three forms of assembly as indicated in E . (2). It is noticed that the case of t isting assembly ( 1) presents the highest values of the modulus compared to the other forms of assembly ( 2, 3). They can therefore be classified by E .(2): E #3