Journal of Mechanical Engineering 2013 3

59 (2013) 3

http://www.sv-jme.eu

Strojniški vestnik Journal of Mechanical Engineering

Since 1955

Papers

139

Matej Hudovernik, Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, A. Erman Tekkaya, Karl Kuzman, Janez Marko Slabe 3D Numerical Analysis of 2D Profile Bending with the Torque Superposed Spatial Bending Method

148

Sebastijan Jurendić, Silvia Gaiani Numerical Simulation of Cold Forming of α-Titanium Alloy Sheets

Andreas Schubert, Henning Zeidler, Matthias Hackert-Oschätzchen, Jörg Schneider, Martin Hahn Enhancing Micro-EDM using Ultrasonic Vibration and Approaches for Machining of Nonconducting Ceramics

165

Tomasz Tański Characteristics of Hard Coatings on AZ61 Magnesium Alloys

175

Leszek Adam Dobrzański, Małgorzata Musztyfaga, Aleksandra Drygała

156

Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques

183

Angel Fernández, Manuel Muniesa, Jaime González Characterisation and Processing of Reinforced PA 6 with Halloysite Nanotubes (HNT) for Injection Molding

193

Gašper Gantar, Andrej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik Resource Efficient Injection Moulding with Low Environmental Impacts

Journal of Mechanical Engineering - Strojniški vestnik

Contents

3 year 2013 volume 59 no.

Strojniški vestnik – Journal of Mechanical Engineering (SV-JME) Aim and Scope The international journal publishes original and (mini)review articles covering the concepts of materials science, mechanics, kinematics, thermodynamics, energy and environment, mechatronics and robotics, fluid mechanics, tribology, cybernetics, industrial engineering and structural analysis.

The journal follows new trends and progress proven practice in the mechanical engineering and also in the closely related sciences as are electrical, civil and process engineering, medicine, microbiology, ecology, agriculture, transport systems, aviation, and others, thus creating a unique forum for interdisciplinary or multidisciplinary dialogue.

The international conferences selected papers are welcome for publishing as a special issue of SV-JME with invited co-editor(s).

Editor in Chief Vincenc Butala University of Ljubljana Faculty of Mechanical Engineering, Slovenia

Technical Editor Pika Škraba University of Ljubljana Faculty of Mechanical Engineering, Slovenia Editorial Office University of Ljubljana (UL) Faculty of Mechanical Engineering SV-JME Aškerčeva 6, SI-1000 Ljubljana, Slovenia Phone: 386-(0)1-4771 137 Fax: 386-(0)1-2518 567 E-mail: info@sv-jme.eu, http://www.sv-jme.eu

Print DZS, printed in 450 copies

Founders and Publishers University of Ljubljana (UL) Faculty of Mechanical Engineering, Slovenia University of Maribor (UM) Faculty of Mechanical Engineering, Slovenia Association of Mechanical Engineers of Slovenia

http://www.sv-jme.eu

59 (2013) 3

Chamber of Commerce and Industry of Slovenia Metal Processing Industry Association

Since 1955

Strojniški vestnik Journal of Mechanical Engineering

Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, Karl Kuzman, Janez Marko Slabe ysis of 2D Profile Bending with the Torque Superposed thod

, Silvia Gaiani on of Cold Forming of α-Titanium Alloy Sheets

Hard Coatings on AZ61 Magnesium Alloys

ański, Małgorzata Musztyfaga, Aleksandra Drygała

ng Process of Front Side Metallisation on s Using Conventional and Unconventional Techniques

anuel Muniesa, Jaime González nd Processing of Reinforced PA 6 with Halloysite Nanotubes Molding

drej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik t Injection Moulding with Low Environmental Impacts

Journal of Mechanical Engineering - Strojniški vestnik

, Henning Zeidler, Matthias Hackert-Oschätzchen, artin Hahn EDM using Ultrasonic Vibration and Approaches Nonconducting Ceramics

year

no. 3 2013 59

volume

Cover: The cover picture represents a machine (above) for 3D bending of tubes and profiles with symmetrical or asymmetrical section properties using dynamic Torque Superposed Spatial - TSS bending method (below). This machine allows simultaneous numerical control of several axes - resulting in the feed of the profile, bending in a direction perpendicular to the profile feed, and a change of bending plane. A curvature with continuous or variable radius, in the domain of plane or spatial geometry of final part is, therefore, achieved.

Image Courtesy: Institute for Forming and Lightweight Construction - IUL, Technical University Dortmund

International Editorial Board Koshi Adachi, Graduate School of Engineering,Tohoku University, Japan Bikramjit Basu, Indian Institute of Technology, Kanpur, India Anton Bergant, Litostroj Power, Slovenia Franci Čuš, UM, Faculty of Mech. Engineering, Slovenia Narendra B. Dahotre, University of Tennessee, Knoxville, USA Matija Fajdiga, UL, Faculty of Mech. Engineering, Slovenia Imre Felde, Obuda University, Faculty of Informatics, Hungary Jože Flašker, UM, Faculty of Mech. Engineering, Slovenia Bernard Franković, Faculty of Engineering Rijeka, Croatia Janez Grum, UL, Faculty of Mech. Engineering, Slovenia Imre Horvath, Delft University of Technology, Netherlands Julius Kaplunov, Brunel University, West London, UK Milan Kljajin, J.J. Strossmayer University of Osijek, Croatia Janez Kopač, UL, Faculty of Mech. Engineering, Slovenia Franc Kosel, UL, Faculty of Mech. Engineering, Slovenia Thomas Lübben, University of Bremen, Germany Janez Možina, UL, Faculty of Mech. Engineering, Slovenia Miroslav Plančak, University of Novi Sad, Serbia Brian Prasad, California Institute of Technology, Pasadena, USA Bernd Sauer, University of Kaiserlautern, Germany Brane Širok, UL, Faculty of Mech. Engineering, Slovenia Leopold Škerget, UM, Faculty of Mech. Engineering, Slovenia George E. Totten, Portland State University, USA Nikos C. Tsourveloudis, Technical University of Crete, Greece Toma Udiljak, University of Zagreb, Croatia Arkady Voloshin, Lehigh University, Bethlehem, USA President of Publishing Council Jože Duhovnik UL, Faculty of Mechanical Engineering, Slovenia General information Strojniški vestnik – Journal of Mechanical Engineering is published in 11 issues per year (July and August is a double issue).

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3 Contents

Contents Strojniški vestnik - Journal of Mechanical Engineering volume 59, (2013), number 3 Ljubljana, March 2013 ISSN 0039-2480 Published monthly

Guest Editorial 137 Papers Matej Hudovernik, Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, A. Erman Tekkaya, Karl Kuzman, Janez Marko Slabe: 3D Numerical Analysis of 2D Profile Bending with the Torque Superposed Spatial Bending Method Sebastijan Jurendić, Silvia Gaiani: Numerical Simulation of Cold Forming of α-Titanium Alloy Sheets Andreas Schubert, Henning Zeidler, Matthias Hackert-Oschätzchen, Jörg Schneider, Martin Hahn: Enhancing Micro-EDM using Ultrasonic Vibration and Approaches for Machining of Nonconducting Ceramics Tomasz Tański: Characteristics of Hard Coatings on AZ61 Magnesium Alloys Leszek Adam Dobrzański, Małgorzata Musztyfaga, Aleksandra Drygała: Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques Angel Fernández, Manuel Muniesa, Jaime González: Characterisation and Processing of Reinforced PA 6 with Halloysite Nanotubes (HNT) for Injection Molding Gašper Gantar, Andrej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik: Resource Efficient Injection Moulding with Low Environmental Impacts

139 148 156 165 175 183 193

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3 Guest Editorial

Guest Editorial Special Issue: Industrial Tools and Material Processing Technologies

The present special issue of the Strojniški vestnik - Journal of Mechanical Engineering is strongly connected with the successful 8th International Conference on Industrial Tools and Material Processing Technologies (ICIT&MPT 2011) held October 2-5, 2011, in Ljubljana, Slovenia. ICIT & MPT, which has been organized

by TECOS - Slovenian Tool and Die Development Centre every other year since 1997, has developed into a renowned international scientific conference attracting scientists, researchers, and engineers from all over the world to discuss theoretical and practical issues and to share information on current world trends in the field of industrial tools and material processing technologies.

The 8th ICIT&MPT was focused on current world trends in the field of metal forming processes, materials, laser machining, process simulations & optimisation, surface coatings, non-conventional machining processes, intelligent systems, and technology management. All conference papers on these topics have been collected and published in the Conference Proceedings comprising around 400 pages and presenting a useful source of information for scientists, researchers and engineers in their professional work.

The chairmen of the International Scientific Committee of the 8th ICIT&MPT carefully reviewed all conference papers, selected seven of them, and invited their authors to prepare revised, expanded, and rewritten papers for this special issue. All papers received were then reviewed by distinguished reviewers according to journal procedures and standards.

The selected papers published in this special issue deal with the following topics: (i) numerical simulation of emerging technology of flexible 3D profile bending using the TSS (Torque Superposed Spatial) bending method, (ii) numerical simulation of cold forming of α – titanium alloy sheets, (iii) micro – EDM process with ultrasonic vibration assistance directly applied to the workpiece and indirectly applied high – intensity ultrasonic to the dielectric in metallic materials, as well as in the machining of electrically nonconductive ceramic materials, (iv) characterization of hard coatings on AZ61 magnesium alloys, (v) comparison of a conventional and an unconventional method, i.e., screen printing and selective laser sintering, to improve the quality of forming electrodes of silicon solar cells, (vi) characterization and processing of reinforced PA 6 with HNT (Halloysite Nano Tubes) for injection moulding and (vii) use of LCA (Life Cycle Assessment) and LCC (Life Cycle Cost) methodology to compare and optimize injection mould designs and injection moulding process parameters from economic and environmental point of view as an upgrade to the usual mostly technical aspect.

We sincerely thank all the authors for their valuable contributions in presenting their research achievements and successful industrial applications, as well as their patience with all reviewers’ comments and suggestions.

Our thanks also go to all the reviewers for their effort in reviewing papers, to prof. dr. Vincenc Butala, the Editor in Chief, for giving us the opportunity to prepare this special issue, and to Mrs. Pika Škraba, Technical Editor of the journal, who took care of the coordination between the reviewers and the authors and prepared the papers for publication.

Guest Editors: Asst. prof. dr. Janez Marko Slabe Prof. dr. Janez Grum Prof. dr. Karl Kuzman

137

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 139-147 © 2013 Journal of Mechanical Engineering. All rights reserved.

DOI:10.5545/sv-jme.2012.483

Received for review: 2012-03-16 Received revised form: 2012-05-11 Accepted for publication: 2012-05-16

3D Numerical Analysis of 2D Profile Bending with the Torque Superposed Spatial Bending Method

Hudovernik, M. – Staupendahl, D. – Gharbi, M. – Hermes, M. – Tekkaya, A.E. – Kuzman, K. – Slabe, J.M. Matej Hudovernik1,* – Daniel Staupendahl2 – Mohammad Gharbi2 – Matthias Hermes2 – A. Erman Tekkaya2 – Karl Kuzman3 – Janez Marko Slabe1 1Tecos, Slovenian Tool and Die Development Centre, Slovenia 2Institute of Forming Technology and Lightweight Construction (IUL), Germany Technische Universität Dortmund, Germany 3University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Engineering research, in the field of light weight design, is strongly oriented towards the development of new high strength materials and innovative forming methods, capable of withstanding limitations with regard to the wide variety of technological and economical aspects.

Cost effective lightweight construction, in addition to the reduction of energy/material consumption and overall reduction of weight, also strongly depends on stability, continuity and robustness of production processes. Kinematic solutions for the production of spatial designed structures, in terms of 3D bending of profiles and tubes, show great potential in an increase of efficiency in the field of light weight design. The Torque Superposed Spatial bending method - TSS, developed at the Institute of Forming Technology and Lightweight Construction, Technische Universität Dortmund, represents an innovative, robust and cost effective technical solution for 2D and 3D bending of tubes and profiles and offers a wide spectrum of capabilities, such as process continuity, parameter adaptation and flexibility for spatial bending of profiles with arbitrary cross-sections. In this paper, an introduction to 3D numerical analysis of the 2D profile bending method using TSS method is introduced and presented. The first objective of the work is to establish validity of the numerical model for the bending parameters, such as the bending force and bending momentum. Secondly, further investigations of the state of stresses and strains during load and unload conditions were performed. These are important for any further analysis and understanding of spring-back, residual stresses and cross section deformation of the profiles. The numerical simulations are performed with the use Abaqus software code, with elastic plastic material characteristics, and are, for the purpose of validation, compared to experimental data.

Keywords: TSS, Torque Superposed Spatial, 3D FEM analysis, 2D profile bending, numerical simulations

0 INTRODUCTION Demands for efficient, lightweight, and spatial designed components and parts are nowadays growing rapidly. Decrease of material consumption and an overall reduction of weight are the main reasons for the development of new high strength materials and robust, flexible, and cost effective engineering solutions for the cold forming of tubes and profiles [1] and [2]. Especially three-dimensionally bent profiles show a great potential in an increase of part efficiency by providing optimal material distribution, a reduction of number of parts, and thereby a decrease of assembly procedures.

The need for profile parts with complex 3D contours represents a demanding challenge for any conventional manufacturing method (i.e. [2] to [4]).

Existing innovative solutions (i.e. [5] to [9]) for the production of such spatial profile and tubular components show the importance of and the necessity for robust and adaptive manufacturing techniques.

Studies regarding TSS bending method have shown its vast potential in forming of profiles made of high strength steels (i.e. [10] to [13]). The TSS bending method offers a great number of advantages such *Corr. Author’s Address: TECOS, Kidričeva 25, 3000 Celje, Slovenia, matej.hudovernik@tecos.si

as the increase of process efficiency by decreasing the number of manufacturing steps, and a decrease of production time [10] and [12]. The process parameters, such as bending force and bending momentum, and forming phenomena occurring during the process of bending with a variable bending radius, have not been studied in detail by numerical methods yet. Therefore, computer aided numerical simulations play an essential role in the understanding of the TSS bending method. The existence of complex states of stresses during processing can be analyzed in detail after validation of the bending force and bending momentum parameters. The 3D numerical analysis of TSS profile bending process is given using implicit software code of Abaqus/standard. Results of the simulations were analyzed and validated with experimental data.

1 TSS PROFILE BENDING METHOD The basic mechanical principles of bending with the TSS bending method, are in many ways similar to those of conventional bending technologies. Unwanted phenomena as tearing, necking, lateral buckling and wrinkling are likely to develop during the bending 139

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 139-147

process, if input parameters are not consistent with material characteristics and cross-section properties [14]. Geometrical accuracy, material properties, spring-back, and residual stresses strongly depend on process stability and must be controlled specifically in order to get sufficient results and a high overall product quality [1].

Conventional technologies show many disadvantages regarding the prediction of forming behavior and the overall product quality (i.e. [15] to [20]. The kinematic TSS bending method offers an innovative approach to solving tasks of bending high strength steel profiles by operating in three parallel phases: continuous profile transportation, plasticization during bending, and the application of torque that is needed to perform the change of plane position and realization of 3D geometry.

The basic graphic scheme of the TSS bending process is shown in Fig. 1. The operational functionality of the TSS bending method is built upon the feed of the profile along the c axis with a constant or variable velocity by clamping the profile and pushing it forward by rotation of 6 roll based system. The 2D bending of the profile is realized by the pre-defined numerically controlled movement of the bending head along the x axis, positioned perpendicular to the transportation axis c. The 3D bending contour is accomplished by superposing a torque onto the plane bending process. This is achieved by the numerically controlled rotation of the torsion bearing (α1), mounted around the feeding mechanism. This event causes the change of the bending plane position relative to the profile and creates a 3D shaped contour, as shown in Fig. 1. The two compensation axes (α2- and τ-axis) of the bending head self-align themselves during the process according to the movement of the profile to achieve a tangential run of the bending head relative to the profile [11] and [21].

2 PROCESS ANALYSIS The analysis of 2D and 3D bending of profiles using TSS bending method can be built upon analytical formulations for the bending of tubes and beams with different sections, as presented and described in (i.e. [14] and [21] to 24]). The parameters of the TSS bending process can later be predicted for each individual input material by the use of finite element methods. Spring-back, cross-section deformations, residual stresses, and material properties can therefore be evaluated by using appropriate software code and by comparison with experimental data.

Fig. 2. TSS bending mechanism

El Megharbel et al. [14], describe analytical techniques to determine the bending moment, for 2D bending of square hollow profile cross-section (Fig. 3), and work hardening elastic-plastic material properties. Similarly, this can be applied to the TSS bending method.

Fig. 3. Hollow cross section of the profile [14]

Fig. 1. Graphic scheme for TSS bending method [10]

140

It is assumed, that the yield point ye is located with the definition of ye ≤ B – t [14]. Stress-strain relationship in the field of elastic plastic properties is defined with [14]:

Hudovernik, M. – Staupendahl, D. – Gharbi, M. – Hermes, M. – Tekkaya, A.E. – Kuzman, K. – Slabe, J.M.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 139-147

σel = E·ε for ε ≤ εy , (1)

σpl = C·εn for ε ≥ εy , (2)

ε = y / R . (3)

producing an effective radius trend of 1985 to 416 mm.

These analytical principals are used as the basis for the development of the numerical analysis.

3 MATERIALS AND METHODS All numerical and experimental investigations are based on a 40×40×2.5 mm profile made from air hardening MW1000L Z3 high strength steel [25].

According to the manufacturer, the yield strength of MW1000L is measured at 900 MPa, and the tensile strength at 1100 MPa [24]. The behaviour of this material can be compared to 15CrMoV6 steel, but offers a 14 to 20% higher breaking elongation [25].

The flow curve for MW1000L Z3 high strength steel was described using standard tensile tests and the (Eq. (4)). The flow curve for this specific high strength steel is shown in Fig. 4.

Fig. 5. Investigated profile contour

To generate the axes movements that are necessary to produce the target curvature, digital mock-ups were used (Fig. 6). The mock-up was setup using the kinematics module of CATIA V5 (Dessault Systems). As the profile is fed through the mock-up at a constant velocity, the x-axis and τ-axis movements are generated accordingly.

Fig. 6. Setup of TSS bending mock-up to generate axes movements [12]

4 NUMERICAL SIMULATION SETTINGS Fig. 4. Flow curve for MW1000L Z3 high strength steel Table 1. Material properties for MW1000L steel [25] ν 0.3

E [GPa] 200

G [GPa] 80

Rp0.2 [MPa] 900

Rm [MPa] 1100

The target profile contour for the analysis is shown in Fig. 5. An experimental analysis for the TSS bending process limits using this contour is shown by Staupendahl et al. [12]. The 2D profile bending target contour incorporates variable bending radii ranging from rmin = 400 mm, to rmax = 2000 mm and thus offers the possibility to study a wide range of radii using a single profile shape. To achieve a continuous curvature in the numerical and experimental analyses, the single radii were replaced by a continuous spline

The numerical model for 2D profile bending was tested with the use of implicit Abaqus software code.

The Von Mises yield criteria and isotropic material work hardening rule were assigned to the model.

All tools – the feeding rolls, and bending head – were defined as analytically rigid elements. Influences of elasticity of tools were not considered within the described test model, but these important factors are to be considered in the future research.

4.1 Mesh Properties The numerical model was tested with the use of Abaqus/standard software code. The meshing of the FEM model was performed with the use of quadratic S8R – 8-node doubly curved thick shell elements – with reduced integration. The section was determined

3D Numerical Analysis of 2D Profile Bending with the Torque Superposed Spatial Bending Method

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 139-147

with the Gauss integration rule and 5 integration points over the 2.5 mm thickness of the shell. For an optimal calculation time, approximately 5000 quad shell elements over the surface were evenly distributed.

4.2 Constraints and Loads The bending head of an actual TSS machine consists of 4 rolls. This enables a sufficient transportation support for each outer side of the square profile crosssection. The bending head - in an actual TSS process constantly adapts its position with regard to the local curvature by free rotation around the vertical axis.

The centre of this adaptive vertical rotation is always positioned on x axis according to the pre-defined NC program, and the bending head is always positioned tangential to the local profile curvature as shown in Fig. 7a. For the numerical simulation of 2D profile bending, the bending head is replaced with one single roll as shown in Fig. 7b. The movement is adjusted accordingly.

The loads and boundary conditions are applied with specific amplitude settings by using the displacement/rotation and velocity/angular-velocity functions. The feed rolls are assigned with a constant angular velocity. The bending is applied by the movement of bending roll, as function of time and amplitude as shown in Fig. 9.

Fig. 9. Amplitude functions for bending roll

4.3 Contact Relations and Step Definition

Fig. 7. Replacement of 4 roll bending head with a single bending roll

The FEM model for 2D profile bending using Abaqus/standard code is shown in Fig. 6.

Contact relations between the deformable body, meaning the profile, and all of the analytically rigid bodies such as the bending roll and the set of feeding rolls were set to prevent objects from penetrating each other. In an Abaqus implicit code, contact relationship between shell elements and analytic rigid bodies, are defined with surface to surface discretization method, penalty friction formulation in tangential domain and hard contact in normal behaviour domain. Friction coefficient μ was set to 0.3 concerning contact between feed rolls and shell elements, while frictionless characteristics were assigned to the relation between bending roll and profile.

The dynamic implicit steps, a full Newton solution technique allowing maximum of 106 number of increments and initial increment size of 10-5 , were defined.

5 EXPERIMENTAL ANALYSES

Fig. 8. FEM model for 2D bending of profile using single bending roll

142

The experimental 2D bending of 40×40×2.5 mm profile, made of MW1000L steel, was performed by using the TSS bending machine setup shown in Fig. 10.

During the bending process, the x-component of the bending force was measured using a load cell integrated into the x-axis of the machine. With the

Hudovernik, M. – Staupendahl, D. – Gharbi, M. – Hermes, M. – Tekkaya, A.E. – Kuzman, K. – Slabe, J.M.

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knowledge of the Ď„-axis rotation generated with the digital mock-up, the y-component of the bending force, as well as the actual bending force were calculated

1400 mm of the profile feed, the numerical value of Fbx reaches around 14 kN, while the value of the actual force in x direction is slightly lower, around 13.2 kN. The force slightly decreases at the point of 1600 mm profile feed, where the bending moment is at its maximum value, due to the plasticization of the material of the profile.

Fig. 10. Experimental results of 2D profile bending using TSS bending method

The contour deviations of the profiles were measured using the GOM Atos 3D scanning system.

The best-fit measurement method showed a maximum offset of 0.60 mm. 90% of the profile actually lie in a spectrum of Âą0.3 mm.

Fig. 11. Comparison of bending force component in x axis Fbx

6 RESULTS The FEM model for the analysis of 2D profile bending, defined with 5000 evenly distributed quadratic S8R shell elements, and a reduced integration technique was created in Abaqus standard implicit code. The overall calculation time for a numerical simulation was 23.6 hours.

The experimental analysis was oriented towards the test of repeatibility regarding geometrical accuracy of final product and validation of input parameters, such as bending forces Fbx and Fby, which were meassured during the process. By using these results, the resultant force Fb and the bending momentum Mb, were determined. The FEM model was validated by comparing the values of bending force and bending moment, which were calculated as a reaction to the movement of analyticaly rigid bending roll. Similar validation of parameters is presented in [26] and [27]. The comparison of experimental and numerical results, for the bending force components in relationship to feed of the profile and the original NC input data, are shown in Figs. 11 and 12. The comparison of the numerical and experimental values of the bending moment Mb, with regard to the feed of the profile and the original NC input data is presented in Fig. 13.

The values of numerically calculated force components, compared to experimental data, are slightly higher on the x axis. At approximately

Fig. 12. Comparison of bending force component in c axis Fbc

Fig. 13. Comparison of the bending moment Mb

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The numerical results for Fby, which are shown in Fig. 12, are slightly lower than the actual data, but sufficiently approximate. The accuracy and consistency of bending moment results are shown and compared in Fig. 13.

The FEM model was validated, by bending force and bending momentum, with sufficient approximation to experimental data. Additional analysis procedures were, therefore, undertaken for 2D bending of square profile in order to understand the state of strains and stresses during bending in the profile material.

Results considering stress-strain relationship and related phenomena such as cross section deformations are presented in the following.

The material of the profile tends to plastify in the area shown in Fig. 14, during the application of bending load. In order to gain detailed understanding of how this affects the geometrical change of the profile cross-section, the work was oriented mainly

to this specific area, at the time of maximum bending moment value. In this particular area of interest, the relationship between tensile and compressive stresses were analysed as shown in Fig. 15. The approximate position of neutral stress zone could, therefore, be evaluated.

The forming mechanism for the kinematic TSS 2D bending of square profile, with the specific input NC data, is charachterized by a larger compressive zone over the cross-section surface.

Detailed state of stresses in contrast to the plastic strain equivalent - PEEQ, over the cross-section width, are shown and described in Fig. 16a and b.

Fig. 16a shows the position of elements - el. 1 to 4 across the profile width, for which the stresses and PEEQ’s were calculated and determined, with an average value of stresses and plastic strain equivalent, for the integration points. Numerical results presented in Fig. 16b, show the state of stresses and strains with regard to the conditions mentioned above. The

Fig. 14. Area of plasticity for 2D profile bending

Fig. 15. Tensile/compressive state of stresses in the area of plasticity

a)

b)

Fig. 16. State of stresses and plastic strain equivalent at element node integration points for the detailed area described in Fig. 15

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horizontal doted line marks the position of neutralstress zone, between tensile and compressive fields.

In the field of tensile stresses, these range up to 1200 MPa, while the PEEQ ranges from 0.021; at the points on the most outer contour, and down to around 0.003; at the points closer to the neutral zone.

In the compressive zone, the PEEQ ranges up to 0.045, while stresses range down to –1068 MPa. The forming mechanism is, herefore, from this specific aspect, determined with a smaller tensile field, but higher stresses, compared to the compressive state, where stresses reach lower values, but plastic strains dominate the domain.

The bending load can also directly influence the vertical geometrical properties of the profile crosssection, marked in Fig. 17, which shows the cut of square profile cross-section in the area of plasticity (i.e. Fig. 15).

Fig. 17. The profile slice used for analysis of cross-section deformation

Fig. 17 presents the deformed state of the crosssection, and the average values of the Von Mises stress over the shell surface of the cross-section.

In the tensile field – the outer vertical side of the cross-section – the average value of Von Mises stress amounts to approximately 1100 MPa. In the compressed field – the inner vertical side of the crosssection – this parameter averages to the slightly higher value of around 1200 Mpa. The state of stresses after unloading, due to the unpredictible spring-back effect, are also presented in Fig. 18. This phenomena shows the substantial change in stress orientation.

During unloading, the previously compressive stresses in the inner vertical side of the cross-section, transform into tensile stresses, while those on the outer vertical side, transform into compressive. The stresses after unloading are, however, higher on the inner

vertical side, with average values of approximately 500 MPa, while stresses on the outer side, range at negative values of around 200 MPa.

Fig. 18. The state of stresses over cross-section area for loading and unloading conditions

The state of the unloaded stresses in the upper and lower horizontal side of the cross-section also perform changes, which are not identical to inner and outer sides of the cross-section. The substantial change occurs only in the upper/lower right corner of the cross-section. The significant value of compressive stress there – 1068.3 MPa – transforms into tensile after unloading with a value of 484 MPa. The condition in the other horizontal elements, remain the same with the difference of decrease of stress values.

The average PEEQ values, per each element in the cross-section, are presented in Fig. 19. As mentioned earlier, the values for plastic strain equivalents are higher in the compressive zone, ranging up to 0.045 on the horizontal side and up to 0.051 in the vertical inner side of the profile. The plastic strain values on the tensile side range up to 0.021 – on the horizontal, and up to 0.039 on the outer vertical side.

Fig. 19. PEEQ over cross-section area

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Due to geometrical symmetry of the cross-section, only the upper half of the cross-section is presented.

This specific subject of complex phenomena will also be an important part in the future of numerical research of the TSS bending method.

7 CONCLUSIONS A 3D numerical analysis for 2D bending of square profiles with the TSS – torque superposed spatial – bending method, for the continued spline, defined with a variable bending radius, has been shown and presented in this paper. A numerical model was developed with the use of Abaqus/standard implicit software code. High strength material characteristics, as well as quadratic - S8R, doubly curved thick shell elements, and the Gauss integration rule with 5 integration points over the 2.5 mm thickness of the shell were assigned to the model. The overall time for the simulation was approximately 24 h.

The model was validated by comparison of numerical and experimental results, as shown in Figs.

11 to 13. Due to sufficient approximation of numerical results, regarding the focus on validating the bending force and the bending momentum parameters, in relation to the actual data, additional analysis regarding the state of stresses and strains in the area of plasticity and profile cross-section deformations, were performed. The shown results, represent the foreground for further work, regarding numerical analysis of 2D and 3D bending of symmetrical and asymmetrical profiles and the analysis of spring-back behaviour, the state of residual stresses, and the effect of elasticity of tools on the final profile contour.

8 ACKNOWLEDGEMENTS The research leading to these results has received funding from the European Union’s European Social Fund and from the European Union’s Research Fund for Coal and Steel (RFCS) under grant agreement No. [RFSR-CT-2009-00017]. All experimental work was performed in close collaboration with the Institute of Forming Technology and Lightweight Construction IUL, Technische Universität Dortmund.

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[6] Neugebauer, R., Drossel, W.G., Lorenz, U., Luetz, N. (2002). Hexabend - a new concept for 3D-free-form bending of tubes and profiles to preform hydroforming parts and space-frame-components. Proceedings of the 7th ICTP Yokohama, vol. 2, Advanced Technology of Plasticity, p. 1465-1470.

[7] Murata, M., Kuboti, T., Takahashi, K. (2007).

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[9] Gantner, P., Bauer, H., Harrison, D.K., De Silva, A.K.M. (2005). Free-Bending - A new bending technique in the hydro forming process chain. Journal of Materials Processing Technology, vol. 167, p. 302308, DOI:10.1016/j.jmatprotec.2005.05.052.

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[14] El Megharbel, A., El Nasser, G.A., El Domiaty, A. (2007). Bending of tube and section made of strain-

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[20] Hermes, M., Kleiner, M. (2008). Vorrichtung zum Profilbiegen (device for profile bending). German Patent Application, DE102007013902A1, Registr. Date 20.03.2007, München.

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3D numerical study on wrinkling characteristics in NC bending of aluminium alloy thin-walled tubes with large diameters under multi-die constraints.

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[24] Aimin, Y., Rongqiang, Y., Ying, H. (2009). Theory and application of naturally curved and twisted beams with closed thin walled cross sections. Strojniški vestnik Journal of Mechanical Engineering, vol. 55, no. 12, p.

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[25] DIN EN 10305-1–MW1000L, Precision Tubes acc. For highly stressed components, Salzgitter Mannesmann Precision GmbH technical data sheet 049 R. (2009) (Werkstoffblatt 049 R).

[26] Shariati, M., Sedighi, M., Saemi, J., Allahbakhsh, H.R. (2010). A numerical and experimental study on buckling of cylindrical panels subjected to compressive axial load. Strojniški vestnik - Journal of Mechanical Engineering, vol. 56, no. 10, p. 609-618.

[27] Niţu, E., Iordache, M., Marincei, L., Charpentier, I., Le Coz, G., Ferron, G., Ungureanu, I. (2011). FEModeling of cold rolling by in-feed method of circular grooves. Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 9, p. 667-673, DOI:10.5545/ sv-jme.2010.244.

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DOI:10.5545/sv-jme.2012.415

Received for review: 2012-03-06 Received revised form: 2012-10-02 Accepted for publication: 2013-01-03

Numerical Simulation of Cold Forming of α-Titanium Alloy Sheets

Jurendić, S. – Gaiani, S. Sebastijan Jurendić1, * – Silvia Gaiani1,2 1 Akrapovič d.d., Slovenia 2 University of Modena and Reggio Emilia, Department of Materials Engineering, Italy Despite the generally good cold workability of some α-titanium alloys, their relevant mechanical properties are quite different to those of traditional cold forming materials. The hexagonal close packed (HCP) crystal structure of α-titanium alloys results in a highly textured, highly anisotropic material that exhibits some specifics in its plastic response. A numerical simulation method using the Barlat 1989 material model has been developed to aid in forming tool development and process parameter determination. In order to account for the anisotropic hardening of the material, plastic strain ratios are input into the model as functions of plastic strain and an inversely determined, experimental strain hardening curve is used. The procedure for determining the input data from the tensile test is outlined and demonstrated on the α-titanium alloy 1.2ASN from Kobe Steel. The flow potential exponent m is evaluated via a parametric analysis of the Erichsen test and an appropriate value is determined. The forming limit diagram is adopted as a means for failure prediction and determined using the Nakajima method. Finally, the method is evaluated on an example of a deep drawn part with good correlation to the physical process.

Keywords: α-titanium, HCP metals, numerical simulation, cold forming, anisotropy, deep drawing

0 INTRODUCTION As titanium alloys are increasingly used in ever widening fields of high-performance applications, classical manufacturing techniques are being applied to these high-tech materials. Commercially pure (CP) and near-CP alloys are a special case in this respect, because of their hexagonal close packed (HCP), α-phase crystalline structure. Although these materials generally exhibit good cold workability [1], their relevant mechanical properties are quite different to those of the traditional engineering materials they are replacing, which can pose a problem under mass production conditions.

A comprehensive theoretical investigation of the deformation behaviour of near-CP titanium alloys can be found in [2]. The primary difficulties for cold forming arise from the high level of anisotropy present in these materials, both in yielding and in work hardening.

Given the high price of the raw material and usually relatively low production volumes, numerical simulation can be a very useful tool in this field in helping to establish a reliable production process, both to determine the feasibility of a given process and to optimise the process parameters and tooling geometry beforehand, minimising the need for costly trial and error testing. There have been some attempts in recent years to develop bespoke constitutive models for HCP metals, as found in [3] to [5], however, a key feature of a numerical method applied in an industrial environment is the ability to identify the necessary material parameters promptly and with readily 148

available tests. Biaxial testing needed to determine the input parameters of these constitutive models does not fall into this category, thus the Barlat 1989 material model [6] is used, which allows for input data to be derived from uniaxial tensile testing, while still retaining moderate flexibility.

The relation of the data derived from the standard tensile test to the necessary input parameters is examined and a method for accommodating some specifics of α-titanium plasticity in the Barlat 1989 material model is presented. The Barlat flow potential exponent m is evaluated for α-titanium via a parametric analysis of the standard Erichsen cupping test. From this, a full material characterisation method is defined and used on the near-CP alloy 1.2ASN from Kobe Steel. Finally the method is validated on an example of a deep drawn part.

1 PROPERTIES OF α-TITANIUM ALLOYS AND THE BARLAT 1989 MATERIAL MODEL 1.1 Plasticity of α-Titanium Alloys As with all HCP materials, the crystalline structure of α-titanium dictates some specifics in its plastic response. The shape of the unit cell makes the material prone to texturing during the rolling process and causes the deformation modes to be dependent on the loading direction. The deformation mechanisms of HCP metals are dictated by the c/a ratio of the unit cell [7], α-titanium alloys exhibit a c/a ratio of 1.587, which is lower than the geometrically ideal ratio of 1.663. The slip systems active in this instance are the

*Corr. Author’s Address: Akrapovič, d.d., Malo Hudo 8a, 1295 Ivančna Gorica, Slovenia, sebastijan.jurendic@novelis.com

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prismatic {1010} planes, the basal (0001) planes and the pyramidal {1011} planes, all in the basal direction <1210>. Together, they provide 4 independent slip systems that all occur in the basal direction, as a consequence a deformation system with a non-basal Burgers vector, such as <c+a> pyramidal glide or twinning, must be activated to accommodate an arbitrary plastic deformation (Fig. 1).

Fig. 1. HCP unit cell with slip planes and directions

The fundamental macroscopic plastic properties that follow from the HCP structure and make the material difficult to form are: • high level of anisotropic yielding, • high level of anisotropic hardening, • yielding asymmetry.

Yielding and hardening anisotropy are well illustrated in the engineering σ – ε curves in the longitudinal, diagonal and transverse directions, as shown in Fig. 2.

Yield stress increases significantly from the longitudinal to the transverse direction, while the hardening exponent n falls off, as does the elongation to ultimate tensile strength, although the total elongation stays roughly the same. Also noteworthy is the early and rather gradual onset of localized necking. The total post-necking deformation to break is extensive in the longitudinal direction, while in the transverse direction, the majority of the plastic deformation is achieved prior to necking necking.

The yield asymmetry, also known as the strength differential (SD), is attributed to twinning being activated under compressive loading and is evident in the much lower yield point under compressive loading [8].

All this implies that the yield locus variation with plastic deformation is not only in size, but also in shape, and is asymmetrical with regard to the direction of loading. The former is evident also in the width to thickness plastic strain ratios:

Ri =

ε wpl ε tpl

, (1)

which have been found to exhibit a significant dependence on plastic strain, especially in the diagonal and transverse directions [9].

1.2 Barlat 1989 Material Model The Barlat 1989 material model was chosen at this stage because its input parameters can be derived easily from the standard tensile test and have a welldefined physical relevance. The plane stress yield criterion Φ is defined as: Φ = a K1 + K 2

m

+ a K1 − K 2

m

+ c 2K2

m

= 2σ Y m , (2)

where σY is the yield stress and K1 and K2 are defined as: K1 =

Fig. 2. σ – ε curves in the longitudinal, diagonal and transverse direction typical for α-titanium

σ x + hσ y 2

 σ x − hσ y K 2 =  2  a = 2−2

Numerical Simulation of Cold Forming of α-Titanium Alloy Sheets

, (3)

2

 2 2  + p τ xy , (4) 

R00 R ⋅ 90 , (5) 1 + R00 1 + R90 149

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σx and σy are the stresses in the local x and y directions, and a, c, h, p are material parameters:

h=

c = 2 − a , (6)

Dir.

Elastic modulus [GPa]

R00 1 + R90 ⋅ , (7) 1 + R00 R90

0° 45° 90°

107 109 109

and the p parameter is found by an iterative search from the expression for the plastic strain ratio in an arbitrary direction φ (usually the diagonal direction is used):

Rϕ =

Table 1. General mechanical properties of 1.2ASN alloy

2mσ Y m  ∂Φ ∂Φ +   ∂σ x ∂σ y 

 σϕ  

− 1. (8)

The flow potential exponent m determines the base shape of the yield surface.

The material parameters a, h, c, p are directly related to the R values in the longitudinal, diagonal and transverse directions, thus if these values are input into the model as functions of plastic strain, the shape evolution of the yield locus can be taken into account.

The main drawback of this material model is that it cannot account for yield asymmetry in any way, however, it should still give adequate results for predominantly tensile load paths.

From the mathematical formulation described above, the necessary input data for the constitutive model are: • plastic strain ratios as functions of true plastic strain in three directions: R00, R45, R90 , • yield stress as a function of equivalent plastic strain σ Y ε p in the longitudinal direction, • flow potential exponent m.

The first two are derived from the tensile test, the exponent m, however, does require a biaxial test to evaluate, but it should be similar for all materials with a common crystal structure.

Yield strength [MPa]

Tensile strength [MPa]

323 355 394

457 417 437

Elongation Elongation at tensile at break strength [%] [%] 18.4 32.1 15.7 35.9 9.0 34.5

2.2 Plastic Strain Ratios The plastic strain ratios as functions of true plastic strain should be determined first, as they are needed in the subsequent steps of the characterization procedure.

The calculation procedure defined in the ISO 10113 standard [10] should be followed because of the large measurement error associated with plastic strain ratio measurement, especially at low strains [11]. The true plastic width strain should be plotted as a function of true plastic length strain and a linear regression fit through the data for the range of interest. The plastic strain ratio is calculated as follows from the slope of the linear regression mr:

R=−

mr . (9) 1 + mr

( )

2 ASN MATERIAL CHARACTERIZATION PROCEDURE 2.1 General Mechanical Properties of 1.2ASN

Fig. 3. Plastic strain ratios as functions of true plastic strain in the longitudinal, diagonal and transverse direction

The general tensile properties were measured using the standard tensile test in accordance with the EN ISO 6892:2009 standard with the extensometer gauges at 80 mm. The sheet thickness was 0.9 mm.

Five samples were tested in each direction, the values presented in Table 1 are average values of all the tests in their respective directions.

The plastic strain ratios are determined at intervals of 1% true plastic strain from initial yield to the onset of localized necking. From that point on the curves, high strains are manually extrapolated using linear extrapolation. Fig. 3 shows the final R values as functions of true plastic strain in the longitudinal, diagonal, and transverse direction.

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The variation in the longitudinal direction is nearly negligible while in the diagonal and transverse directions there is a significant initial variation before the values stabilize.

2.3 Yield Curve Determination In contrast to steel, titanium exhibits substantial additional elongation past the point of localized necking with a fairly gradual onset of localization, which is not captured well by any of the traditional hardening laws, thus an experimental true stress-true strain curve was identified using an inverse procedure proposed in [12]. The yield curve is identified iteratively by running numerical simulations of the tensile test and by modifying the yield curve until an acceptable fit between the simulation and the tensile test is achieved. A fit within the scatter between samples of the same batch can be achieved without difficulties.

For this determination a ¼ symmetry model of the parallel section of the tensile specimen was modelled using shell elements. An element size of approximately 0.8 mm was adopted, as it is representative of the element sizes typically used in later forming simulations. Mass scaling was used to maintain a time step of 5·10–7 and the deformation rate was scaled by an order of 103, as this was found to yield satisfactory results.

2.4 Calculation of Barlat 1989 Material Model Parameters The a, c, h and p parameters of the Barlat 1989 material model are generally calculated internally in the numerical simulation software from the plastic strain ratios. Since the plastic strain ratios are functions of plastic strain, the Barlat parameters follow suit. Fig. 5 shows the a, h and p parameters as functions of true plastic strain.

The c parameter is not included in the graph, as it is derived from the a parameter using simple subtraction (Eq. (6)).

Fig. 5. Parameters of the Barlat 1989 material model

2.5 Flow potential exponent m

Fig. 4. Inversely identified yield curve, the power law approximation and the tensile test results

The final identified yield curve for 1.2ASN is shown in Fig. 4, along with the measured true stresstrue strain curve and the functional approximation using a power law. Compared to the power law approximation it is much steeper at high strains, in order to support the extensive post-Rm deformation.

To determine an appropriate value for the m exponent, a parametric analysis was done using the Erichsen cupping test, as defined in the ISO 20482 standard (Fig. 6). In the absence of a viscous pressure bulge test, this test is a suitable alternative for this evaluation as it strains the material with biaxial tension, thus avoiding any SD effects, which pose a problem for the Barlat 1989 material model.

A numerical model of the Erichsen test was built and the input data determined above were used to run the simulations while varying the m parameter. The model is constructed from shell elements, with an average element size of approximately 0.7 mm in the deformation zone. The tools are considered to be rigid.

A penalty contact formulation was used for the blank to tooling interfaces with a friction coefficient of 0.2.

Mass scaling to a time step of 1·10–6 was applied and the punch velocity was scaled by approximately 200 times to shorten calculation times. The resulting

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force-displacement curves from the simulation were compared to the test results (Fig. 7).

pin and a flat pin with a contact area of 1 mm2. The linear pin speed was 15 cm/min, which is in the same order of magnitude as experienced under deep drawing, and the normal load applied to the specimen was 5 N. The duration of the tests was 30 min and the friction coefficient is the average value of the entire test, excluding the initial discontinuities typical of this method.

Combinations of available lubrication conditions were examined and the results are presented in Table 2.

Table 2. Friction coefficients under different lubrication conditions Lubricant / / Oil Oil PVC foil unlubricated PVC foil + oil PVC foil + grease

Fig. 6. Schematic of the Erichsen test tooling

Type of pin Spherical Flat Spherical Flat Flat Flat Flat

μ 0.45 0.50 0.44 0.50 0.38 0.16 0.17

3 LIMITS OF FORMABILITY As the goal of the simulation is to ultimately determine the feasibility of a given deformation process, the forming limit diagram (FLD) was determined for the material. The Nakajima method according to ISO 12004-2 [13] was used with optical strain measurement [14].

The test is based on deforming sheet metal samples of different geometries (Fig. 8) using a hemispherical punch up to the point of fracture.

Fig. 7. Erichsen cupping parametric analysis results

The evaluation shows very good agreement for m = 2. The anomalies in the curves above 7 mm of displacement are attributed to contact instabilities, however, they do not affect the clarity of the results.

2.6 Friction Coefficient Determination The frictional conditions between the blank and the tooling have a major effect on deep drawing processes. In order to evaluate the appropriate values of the friction coefficients, a series of wear tests were performed using a CSM pin-on-disc tribometer, according to the ASTM G99-05 (2010) standard.

In this test, a disc of the test material rotates under a X100 Cr 6 steel pin with a hardness of 58 HRC. Two different pin geometries were used, a spherical 152

Fig. 8. Nakajima test specimen geometries

Different strain paths are achieved by varying the width of the samples and each strain state at fracture corresponds to a point on the major strain – minor strain plot.

The sheet metal used in the determination was 0.9 mm thick. Three samples were tested for seven different geometries in the longitudinal and transverse

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direction, for a total of 42 tests. The resulting forming limit diagram is shown in Fig. 9.

Fig. 10. The part after failure

Fig. 9. Forming limit diagram for 1.2ASN

4 SIMULATION OF A DEEP DRAWN PART 4.1 Physical Part and Numerical Model A motorcycle exhaust end-cap was selected for the simulation. It is a problematic shape to form with its tapering sides and a small top radius. This particular part cannot be drawn from the 1.2ASN material, the material fractures at the leading edge of the punch (Fig. 10). This allows for better evaluation of the method as the limits of formability are surpassed.

A series of drawing tests were carried out with longitudinal and transverse material orientations with regard to the longer axis of the end-cap. The maximum safe drawing depth was established to be 47 mm (the full depth is 90 mm). The thickness distribution was then measured on these samples along the line shown in Fig. 11.

Fig. 12 shows the numerical model of the endcap deep drawing process. The model is comprised of three- and four-node shell elements, only fournode elements are used for the deformable blank and the average initial element size is approximately 4.5 mm. Adaptive remeshing of the blank is adopted, to a minimum element size of 0.9 mm.

Fig. 11. The end-cap drawn to 47 mm with the thickness distribution measurement line

Fig. 12. End-cap drawing process numerical model

The tools are considered rigid and penalty frictional interfaces are prescribed between the tools and the blank. Friction coefficients of 0.2 are used for the lubricated contacts (die-blank, holder-blank), and 0.3 for the unlubricated contact (punch-blank). The die is fixed in all degrees of freedom while the punch and blankholder are free to move only in the z direction.

The punch is prescribed a constant linear velocity and a constant load is applied to the blankholder.

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Mass scaling to a maximum time step of 1·10–6 is used along with time scaling of the punch kinematics by a factor of about 500.

4.2 Results The in-plane strains at 47 mm of drawing depth were plotted on the FLDs for the longitudinal and transverse direction (Figs. 13 and 14). They clearly show the localized deformation in the uniaxial strain region of the FLD that exceeds the forming limit curve. Those points coincide with the leading edge of the punch where fracture occurs.

The simulations predict material fracture on the FLD at around 40 mm of depth for the longitudinal orientations and around 42 mm for the transverse orientation, which is somewhat conservative, however, both the location and shape of the predicted fracture compare well to the actual failed part (Fig.15).

Fig. 15. Failure region predicted on the FLD

Fig. 13. Minor – major strain plot for the longitudinal blank orientation at 47 mm of punch displacement

Fig. 16. Thickness distribution with the longitudinal blank orientation

Fig. 17. Thickness distribution with the transverse blank orientation Fig. 14. Minor – major strain plot for the transverse blank orientation at 47 mm of punch displacement

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The simulated thickness distributions of both blank orientations at the maximum safe drawing

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depth were compared to the experimental thickness distributions measured on the actual parts. The zero point on the horizontal axis corresponds to the die shoulder, with the distance measured along the surface of the part. Figs. 16 and 17 show the results of the comparison.

The simulated thickness distributions compare fairly well to the measurements; the severe discrepancy in the longitudinal direction is due to the simulation predicting a break in the rosette before this depth is achieved. In both cases the simulated thickness lies below the measured distribution up to the point of local failure; it is suspected that the initial sheet might have been somewhat over-gauge.

Overall the results compare well with the actual process even if fracture is predicted prematurely, this might also have to do with the actual thickness of the sheet being slightly more than the nominal thickness used in the simulations.

5 CONCLUSIONS Despite the inherent limitations of the Barlat 1989 material model with regard to the specifics of HPC materials, the simulation results show that this method is entirely adequate for running simulations at this level. The results correlate reasonably well with the experimental data, although the simulation is somewhat conservative. The cause of this could be the lack of strain-rate effects modelling, as the hardening curve is derived directly from the tensile test at low strain rates.

The procedure for acquiring the input data has proved to be robust and reliable in an industrial environment. Its primary drawback is that the inverse curve fitting procedure is time consuming and requires manual alterations to the yield curve between iterations, unless specialized optimization software is used.

While the parametric analysis of the Erichsen cupping test does provide a seemingly conclusive result, a better controlled biaxial stress test is still required to confirm the results. Such a test would provide a definite point on the yield locus that would allow the accurate determination of the m exponent.

6 REFERENCES [1] Lutjering, G., Williams, J.C. (2007). Titanium, 2nd ed.

Springer, Berlin. Heidelberg, New York.

[2] Nixon, M.E., Cazacu, O., Lebensohn, R.A. (2010).

Anisotropic response of high-purity α-titanium: Experimental characterization and constitutive modelling. International Journal of Plasticity, vol. 26, p. 516-532., DOI:10.1016/j.ijplas.2009.08.007.

[3] Plunkett, B., Lebensohn, R.A., Cazacu, O., Barlat, F. (2006). Anisotropic yield function of hexagonal materials taking into account texture development and anisotropic hardening. Acta Materiala, vol. 54, p. 41594169, DOI:10.1016/j.actamat.2006.05.009.

[4] Cazacu, O., Plunkett, B., Barlat, F. (2006). Orthotropic yield criterion for hexagonal close packed metals.

International Journal of Plasticity, vol. 22, p. 11711194, DOI:10.1016/j.ijplas.2005.06.001.

[5] Lee, M.G., Wagoner, R.H., Lee, J.K., Chung, K., Kim, H.Y. (2008). Constitutive modelling for anisotropic/ asymmetric hardening behaviour of magnesium alloys sheets. International Journal of Plasticity, vol. 24, p.

545-582, DOI:10.1016/j.ijplas.2007.05.004.

[6] Barlat, F., Lian, J. (1989). Plastic behaviour and stretchability of sheet metals, Part 1: A yield function for orthotropic sheets under plane strain conditions.

International Journal of Plasticity, vol. 5, p. 51-66, DOI:10.1016/0749-6419(89)90019-3.

[7] Wang, Y.N., Huang, J.C. (2003). Texture analysis in hexagonal materials. Materials Chemistry and Physics, vol. 81, p. 11-26, DOI:10.1016/S0254-0584(03)001688.

[8] Lissenden, C.J., Doraiswamy, D., Arnild, S.M. (2007). Experimental investigation of cyclic and timedependent deformation of titanium alloy at elevated temperature. International Journal of Plasticity, vol.

23, p. 1-24, DOI:10.1016/j.ijplas.2006.01.006.

[9] Chamanfar, A., Mahmudi, R. (2005). Compensation of elastic strains in the determination of plastic strain ratio (R) in sheet metals. Materials Science and Engineering, vol. 397, p. 153-156, DOI:10.1016/j.msea.2005.02.039.

[10] International Standard ISO 10113 (2006). Metallic materials – Sheet and strip – Determination of plastic strain ratio. International Organization for Standardization, Geneva.

[11] ASTM Standard E 517-00 (2000). Standard Test Method for Plastic Strain Ratio r for Sheet Metal.

ASTM International, West Conshohocken.

[12] Koc, P., Štok, B. (2008). Usage of the yield curve in numerical simulations. Strojniški vestnik - Journal of Mechanical Engineering, vol. 54, p. 821-829.

[13] International Standard ISO 12004-2 (2008). Metallic materials – Sheet and strip – Determination of forming limit curves Part2: Determination of forming limit curves in the laboratory. International Organization for Standardization, Geneva.

[14] GOM GmbH Technical Papers (2009). Material Properties: Determination of Process Limitations in Sheet Metal Forming – Forming Limit Diagram (Revision A). GOM GmbH, Braunschweig.

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DOI:10.5545/sv-jme.2012.442

Received for review: 2012-03-23 Received revised form: 2012-12-14 Accepted for publication: 2013-02-04

Enhancing Micro-EDM using Ultrasonic Vibration and Approaches for Machining of Nonconducting Ceramics

Schubert, A. – Zeidler, H. – Hackert-Oschätzchen, M. – Schneider, J. – Hahn, M. Andreas Schubert1,2 – Henning Zeidler1,* – Matthias Hackert-Oschätzchen1 – Jörg Schneider2 – Martin Hahn1 1Chemnitz

University of Technology, Chair Micromanufacturing Technology, Germany Institute for Machine Tools and Forming Technology IWU, Germany

2Fraunhofer

Micro EDM (Electro Discharge Machining) is a known nonconventional process for the machining of hard to cut materials. Due to its ablating nature based on melting and evaporation through heat induced by electrical discharges, it can function independently of the hardness, toughness or brittleness of the workpiece. Because of these benefits, EDM is widely used in tool- and mould making; micro-EDM, with its much lower discharge energies, has been successfully applied to micromachining of high-accuracy parts. The precision manufacturing of high aspect ratio micro geometries such as deep micro bores relies on stable process conditions in the discharge gap. Its minimisation – a precondition for minimal feature size and higher accuracy – limits the effectiveness of conventional flushing techniques, leading to a higher fraction of unwanted discharge states (open and short circuit), lower process speed, and geometrical errors. New hybrid technology approaches, such as ultrasonic or low frequency superposition, significantly raise the process stability and speed.

Another restriction on the use of EDM, the exclusive machinability of electrically conductive materials, is overcome by the application of the assisting electrode method that enables a micro-ED-machining of nonconductive zirconium oxide ceramics.

This paper presents the current status of investigation into the micro-EDM process with ultrasonic vibration assistance – directly applied to the workpiece and indirectly applied high-intensity ultrasonic to the dielectric – in metallic materials as well as in the machining of electrically nonconductive ceramic materials. Using ultrasonically aided micro-EDM, the process speed can be raised by up to 40%, enabling bores of less than 90 µm in diameter with aspect ratios >40 for metallic materials. The modified setup using the assisting electrode principle allows for machining of an aspect ratio >5 for nonconductive ceramic materials, leading to new possibilities for the design and manufacture of complex, high-accuracy micro parts in high-performance engineering materials.

Keywords: electro discharge machining, micro machining, ultrasonic, ceramics

0 INTRODUCTION 0.1 Micro-EDM Electro discharge machining (EDM) is a known and widely used nonconventional machining process for hard to cut materials because of its ability to function independently of the hardness, brittleness or toughness of the workpiece. Using low energies, it has been successfully applied in micro- and precision machining, for example, in the mould making industry.

Higher accuracies in smaller and more complex structures demand an extremely stable process and therefore enhanced flushing techniques.

Additionally, more and more high-tech materials such as ceramics are being used for micro parts; this presents an opportunity to apply µEDM in that emerging material area.

0.2 Process Principle The electro discharge machining process is based on ablation of material through melting and evaporation.

Fig. 1 shows the process principle.

The electrical discharges take place between the tool electrode and the workpiece in a dielectric medium that separates the two. A voltage is applied 156

to both electrodes and, when the breakdown voltage of the medium is reached, a plasma channel allowing for a current flow is established and a discharge takes place. At the base of the plasma channel the temperature can reach T ≥ 10,000 °K, melting and evaporating the electrode material.

When the energy input is stopped the discharge ends, leading to a collapse of the plasma channel and the surrounding gas bubble. The reflow of the dielectric medium flushes liquid material away and cools the electrode surface. Repeating the process, a voltage is attached to the electrodes again and the setup is prepared for the next discharge.

Naturally, the discharge will take place where the breakdown barrier is lowest, that is when the distance between the electrodes is the smallest – in an ideal dielectric – or, in a real dielectric liquid, when the conductivity of the gap between the electrodes is the highest, e.g. when particles or gas bubbles reduce the breakdown voltage of the medium.

By constant repetition of the process, the tool electrode surface is reassembled in the workpiece and, by feeding the tool, a transfer of the geometry takes place. Because of the process nature, the surface is an assembly of single discharges and shows a crater-like topology. The geometrical accuracy and the surface roughness depend on the size and shape of these craters and therefore on the volume that is ablated

*Corr. Author’s Address: Chemnitz University of Technology, Chair Micromanufacturing Technology, 09107 Chemnitz, Germany, henning.zeidler@mb.tu-chemnitz.de

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Fig. 1. EDM process principle

with each discharge. A minimisation of discharge energy is the key to precision and optimal surface characteristics. Consequently, the discharge gap must be minimised, too.

In µEDM, the discharges have a typical duration of te ≈ 100 ns and transfer energy of We ≈ 10 µJ. The resulting crater width depends on the workpiece material properties, but a diameter of dC ≈ 5 µm and depth of ≤1 µm can usually be obtained. The resulting surface roughness can be as low as Rz ≤ 1µm.

A major benefit of the electro discharge machining process, due to its electro-thermal nature of ablation, is independency of material hardness and brittleness. The noncontact nature of the process results in a nearly force-free machining, allowing the usage of soft, easy to machine electrode materials even when shaping very hard workpieces.

This also enables the machining of fragile or thin workpieces. Additionally, there is no limitation to the

angle between the tool and workpiece, so round or irregularly shaped surfaces can be used.

For all those reasons, EDM has been widely used in the generation of micro parts and geometries such as spinnerets (Fig. 2) [1].

Fig. 2. SEM image of non-circular micro bore Ø80 µm in hardened steel, manufactured using vibration-assisted µEDM

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0.3 Limiting Parameters An ideal EDM process delivers identical shape and energy for each discharge. In reality, this is not possible. The estimated efficiency ratio for µEDM is between 30 and 50% [2], meaning that the rest of the discharges are not optimal. Short circuits, open circuits or long arc discharges do not only impact effective machining, but also have negative effects on electrode wear, form distortion, and process speed.

An optimal discharge – in reality – only occurs if the environmental parameters are optimal at the time the discharge takes place. Therefore it is of great importance to deliver, if not optimal, at least constant environmental parameters to be able to find a suitable generator setup.

One way to achieve this is the regulation of the discharge gap width. Only at an optimum value, influenced by the properties of the medium filling the gap, can an efficient discharge take place.

The discharge gap width is commonly regulated based on the mean gap voltage (ue) that delivers different levels for the normal, short or open circuit state. Depending on this and a set target value, the electrode is fed or withdrawn (feed s) (Fig. 3).

Fig. 3. Electrode feed as a reaction to gap voltage

Novel approaches at Fraunhofer IWU and TU Chemnitz aim to develop a current-based predictive

Fig. 4. Flushing options in EDM

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process control to reduce the reaction delay and increase the efficiency ratio of µEDM. However, since the discharge duration is in the range of 100 ns, new algorithms and fast, microprocessor-based systems have to be implemented.

Of course, the more stable the gap conditions are, the less short-circuiting and time-consuming intervention of the gap width regulation occurs. An optimal gap flushing is therefore crucial.

High-aspect ratio micro bores with tool diameters of less than 100 µm, however, present the most difficult flushing conditions. In µEDM, many conventional strategies of gap flushing (Fig. 4), such as high-pressure flushing or flushing through bores in the tool electrode, are, because of the small size and fragility of the tool, not available. The rotational flushing loses efficiency due to the small tool diameter and the resulting low circumferential speed.

Furthermore, as mentioned before, to achieve maximum precision, the discharge gap has to be minimised. Currently, in µEDM, gap widths of wGP ≤ 10 µm are common.

The combination of these factors leads to large difficulties in efficient flushing and new approaches have to be taken to ensure stable process conditions.

One option is the direct vibration of the workpiece, which can be considered a high frequency version of the flushing by lifting principle.

1.2 Low Frequency Vibration Compared to ultrasonic vibration, low-frequency vibration can be easily realised, since there is no need for resonant systems. Tests were conducted to characterise the effect of a LF-vibration on the machining of bore holes.

While the acceleration of process speed is in the region of 10% compared to conventional machining, a larger contribution can be seen in stabilising difficult geometries and reducing geometrical errors and process instabilities, which can be observed by a reduction of the machining time variation on the order of >50% (Fig. 5, Section (2)).

However, the specific machine tool setup has to be taken into consideration when choosing the vibration frequency in order to avoid a negative interaction with the gap width regulation. Section (1) in Fig. 5 shows this phenomenon.

1 VIBRATION SUPERPOSITION 1.1 Effects on the EDM Process The effect of direct vibration superposition on the µEDM process was found to be very beneficial.

The periodic relative movement between tool and workpiece causes a flow of the dielectric and an agitation of the debris particles in the dielectric medium. With this, the settling of debris on the bore ground and the agglomeration of particles are reduced and the state of the gap is equalised.

An improvement in terms of flushing and homogenisation of the dielectric liquid leads to a faster and more stable process with better form accuracy, allowing for higher aspect ratios and more complex structures. Since the vibration frequency directly influences the flow speed in the dielectric, ultrasonic and high-frequency vibrations are especially beneficial. However, since resonant systems are required for these frequencies, they present some difficulties in the implementation if they are to be applied directly to the workpiece.

Fig. 5. Low-frequency vibration aided µEDM; effects on machining time tero and its deviation

1.3 Ultrasonic - Direct Direct ultrasonic vibration of the tool or workpiece (Fig. 6) is regarded as the optimal strategy for improved flushing and stabilised µED-machining of high aspect ratio structures. The high frequency in the range from 20 to 60 kHz with amplitudes of 2 to 10 µm peak-peak significantly influences the state of the frontal discharge gap and therefore the process itself.

Two major effects can be distinguished: 1) the influence of the vibration on the dielectric and 2) the immediate influence to the working gap width through periodic feed-retraction-movements of the vibrated part.

The first effect can be described by a very high velocity movement of the dielectric. Its speed can reach more than 0.5 m/s with accelerations of over 60 km/s², efficiently moving and stirring the dielectric.

An agglomeration of particles is reduced, leading to a

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more uniform gap condition and therefore increasing the amount of efficient discharges.

top radius 8 µm, and its height 800 µm. The distance between the two cones (centre to centre) is 250 µm.

In deep boring experiments, micro bores of diameter 100 µm and depths of 5000 µm could be successfully machined in steel.

Fig. 6. Direct vibration of tool or workpiece in µEDM

The second effect relates to the periodic change in the gap width through vibration. There, a retracting movement ends longer arc discharges that cause geometric deviation and process instabilities [3].

As a result, the process speed is significantly enhanced and, additionally, more complex structures can be reproducibly machined. Fig. 7 shows the tool feed sz for the machining of very deep micro bores with an optimised conventional process and two direct ultrasonically assisted processes. The used tool electrode diameter is 60 µm (a solid tungsten carbide rod, machined by centerless grinding).

Fig. 8. Array of 751 micro bores, diameter 85 µm, depth 1000 µm, grid width 110 µm, steel

Fig. 9. Array of 121 micro cones (segment view); diameter 150/16 µm, height 825 µm, centre-centre-distance 250 µm, steel Fig. 7. Tool feed sz during deep bore µEDM (tool Ø60 µm)

It can be seen that with increasing depth the conventional process (f = 0 Hz) not only slows down, but also becomes unstable, which leads to retracting movements of the tool electrode caused by long short circuit periods. For the direct ultrasonically assisted processes (f = 21.8 kHz), no large retracting movements are observed.

This allows for the machining of high aspect ratio structures such as micro bore arrays (bore diameter 85 µm, depth 1000 µm, grid width 110 µm, steel, Fig. 8) or complex external structures such as the cone array shown in Fig. 9. The cone bottom radius is 75 µm, the 160

1.4 Ultrasonic - Indirect The utilisation of the great advantages of direct ultrasonic vibration induced in the tool or workpiece presents the user with the difficulty of designing a suitable actuation- and clamping setup that can withstand the high acceleration forces. Since these are proportional to the mass of the system, exciting a large part is challenging.

One approach to implementing a similar setup without exciting the workpiece or tool is to integrate an additional vibrating part to overlay the discharge area with a high intensity ultrasonic field.

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Possible setups consist of an ultrasonic actuator/ transducer and a tuned sonotrode that amplifies the vibration amplitude to the targeted value. The sonotrode is immersed in the dielectric and arranged in a way that the high intensity near field of the ultrasonic is aligned to the machining zone (Fig. 10).

Fig. 10. Indirect ultrasonic superposition in µEDM

The usage of external equipment allows for an even higher velocity at the sonotrode tip. The experiments conducted are based on sonotrode tip speeds of up to 2.2 m/s at frequencies of 24 kHz and amplitudes of 30 µm peak-peak which lead to intense movements within the dielectric.

Within the near field, one important effect is clearly visible. High intensity, fast movement causes cavitation, meaning the creation and collapse of small gas bubbles through pressure variation (to below the vapour pressure of the liquid) in the medium.

In addition, cavitation occurs more easily when the homogeneity of the liquid is disturbed by particles.

Within the discharge gap, particles are created during the EDM process and, as a result, cavitation is additionally facilitated. Stuck to the particles, the gas bubbles of the cavitation phenomenon can become stable gas bubbles that rise and create a stream transporting particles out of the discharge gap [4], which contributes to improved flushing and homogenisation.

In practice, a challenge is the alignment and delivery of the high intensity pressure waves to the discharge gap. The angle between the feed direction and the ultrasonic wave front is an important parameter that, for deep structures, should be minimised.

Using the indirect ultrasonic superposition with angles around 60°, a stabilisation of the bore geometry to a more cylindrical shape and a process speed enhancement of 5% can be achieved.

A further improvement is expected when a coaxially induced ultrasonic vibration is applied, as is currently being investigated. The setup used is shown in Fig. 12.

Fig. 12. Setup for coaxial indirect ultrasonic vibration Fig. 11. Equipment for indirect ultrasonic vibration superposition

When applying indirect ultrasonic superposition it has to be considered that, because of the orientation of the sonotrode noncoaxial to the feed direction, the tool electrode is also subject to excitation through the pressure waves within the dielectric. This is especially so in the case of a vibration in the eigenfrequencies of the tool, where electrode resonance movement can adversely affect the process. Consequently, this frequency range should be avoided.

Although in contrast to direct vibration there is no variation in the discharge gap width, the influence on the dielectric can be seen as similar.

The conducted experiments, however, highlight another crucial factor in the micro-EDM of very small bores: the length of the electrode tool. The required mechanical stability can, in principle, be delivered by tungsten carbide tools, but this leads to changed boundary conditions for the micro-EDM: primarily, a decrease in electrical conductivity. In our tests – using Ø 90 and Ø 150 µm tool electrodes and, due to geometrical restrictions, a free tool length of >14 mm – a significant decrease in process speed and stability occurred. Further investigations showed a correlation to the free tool length FTL (distance from clamping/ electrical contact to electrode tip). The process

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changes its characteristics when the FTL exceeds a certain length to diameter ratio. Signal analysis revealed that the maximum peak current decreases substantially with increasing FTL. A second obstacle is the sensitivity of the tool to orthogonal forces such as adhesion. Experiments show that even when using tungsten carbide tools, from a FTL aspect ratio of more than 50 onwards, bore diameter and machining time increase exponentially.

Therefore, adapted machine tool setups have to be designed, allowing for a minimum FTL while still enabling coaxial vibration coupling. One approach is the use of membrane-shaped piezoceramic actuators.

The design for such setup is currently taking place.

2 EDM OF NONCONDUCTING CERAMICS 2.1 Demand and Approach Ceramic materials are, due to their extraordinary properties such as high hardness and biocompatibility, increasingly used in micro parts. Its machining, however, is difficult, and mostly slow and expensive grinding processes are used [5]. Electro discharge machining with its nearly forceless behaviour and independence of the material’s hardness and brittleness seems to be a suitable process, but is mainly limited to conductive materials. In research, several approaches have been taken to machining nonconductive ceramics using EDM, with the ‘assisting electrode’ method developed by Mohri and Fukuzawa [6] and [7] leading to the successful machining of zirconium dioxide materials.

To achieve machinability with µEDM, the process has to be initially started by first delivering an electric circuit through the conductive starting layer as shown in Fig. 13.

While machining the starting layer, the dielectric (hydrocarbon oil) is cracked, providing conductive carbon that settles down onto the ceramic surface.

Thus, a new conductive layer is generated that enables the next discharges to take place.

By tightly controlling the process environment, this sequence of removing the layer, including underlying base material, and creating a new thin layer can be stably repeated.

In previous studies, the assisting electrode approach has been successfully adapted to a vibrationassisted µEDM setup of a biocompatible zirconium dioxide compound in order to examine the effects of different frequencies and amplitudes on process speed and achievable geometries [8] and [9].

2.2 Modifications and Ceramic µEDM Process A standard µEDM machine tool was used for the experiments (Sarix T1-T4) and was modified to comply with the special requirements of µEDM of nonconductive ceramics.

The modifications developed consist of a basin, filled with dielectric oil and equipped with a peristaltic pump for flushing and filtering of the medium independent from the main dielectric circulation, as well as an active workpiece clamping unit that can be excited with a low-frequency vibration of up to f = 1000 Hz at an amplitude of 20 µm peak-peak.

Fig. 13. Assisting electrode scheme for machining nonconductive ceramic materials

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A common silver varnish with 45.9% silver content is used to create the starting layer. It is applied with a paintbrush. The uniformity of the starting layer is very important, therefore, after drying, it is tested by measuring the resistance using a multimeter.

The thickness of the layer is 20 µm.

In previous tests, it was found that machining of nonconducting ceramic ZrO2 material only takes place with a cathodic polarity. The generator should be set accordingly.

The electrode wear can be noticeably reduced by introducing tool rotation; however, with values of ≈ 75% relative wear there is still a high reduction potential.

100 ns long are observed, to machining the base material. There, the discharge shape changes to a short peak followed by a long hold period at low currents.

This type of discharge can last up to 8 µs and its second part is expected to be the period where the carbon layer is deposited.

2.3 Machinability of Nonconductive Ceramics Using the proposed setup, a successful machining of ZrO2 ceramics is performed, even after the starting layer is completely removed (Fig. 14).

Bores with an aspect ratio 5 could be manufactured using Y2O3- (Fig. 15) as well as MgOstabilised ZrO2 material 3 CONCLUSIONS 3.1 Vibration Superposition

Fig. 14. Tool feed (s) during µEDM of nonconducting ZrO2 ceramics (tool Ø120 µm)

Fig. 15. Cross section of µEDM bores in ZrO2 ceramic, diameter 200 µm, tool Ø 120 µm

Pulse analysis shows a change in the discharge shape from machining the starting layer, where short pulses similar to discharges in metal approximately

The µEDM process benefits extraordinarily from vibration superposition.

Where conventional µEDM processes come to their limits and lose their reproducibility, the process is now stabilised in terms of shape accuracy and machining time.

This is already achievable using indirect vibration, induced via the dielectric. By minimising the angle between the tool feed and wave front direction, further improvements are possible. However, new setups have to be designed to reduce the free tool length to a minimum in order to achieve stability and immunity to external forces, especially when using small tool diameters of less than 100 µm. Membrane-shaped piezoceramic actuators using bending mode appear to deliver beneficial properties.

With direct application of vibrations, higher aspect ratio structures can be machined and the process speed is improved. Here, the vibration frequency is the crucial point.

Low-frequency vibration mostly reduces the deviation of the process, both in terms of machining time and shape accuracy. A rise in machining speed in the range of approximately 10% compared to an optimised conventional process is seen.

Using direct ultrasonically aided micro-EDM, the process speed can be raised by up to 40%. The discharge gap state benefits additionally from the cutting off of long arc discharges and geometric deviations are further reduced.

The direct ultrasonic vibration of the tool or workpiece enables the machining of complex structures with very high aspect ratios, as well as

Enhancing Micro-EDM using Ultrasonic Vibration and Approaches for Machining of Nonconducting Ceramics

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enabling bores of less than 90 µm in diameter with aspect ratios >40 for metallic materials.

4 REFERENCES

3.2 Ceramics Using the assisting electrode method and the proposed experimental setup, micro bores can be successfully machined in MgO- and Y2O3-stabilised ZrO2 ceramic materials. The use of tool electrode rotation leads to an increase in machining speed and reduces tool wear. While at the beginning of machining the starting layer discharges similar to those in metalµEDM can be observed, a transition into so-called ceramic discharges takes place that are characterised by a smaller peak current followed by a long constant low current flow. These make up the majority of discharges when machining the ceramic base material.

The process that is used for Y2O3-ZrO2 can also be used for MgO-ZrO2.

3.3 New Boundaries of µEDM Micromachining is facing new challenges with the constantly growing demand for smaller, more accurate structures in new, emerging materials. Electro discharge machining has the potential to fulfil those needs in hard to machine materials.

The application of hybrid processes to enhance stability by influencing the flushing and discharge gap state has shown great potential in the machining of conducting materials. Ultrasonic superposition excels at stabilising and thus accelerating the µEDM of very deep, precise structures.

With the integration of the assisting electrode method, materials considered to be very hard to machine, such as engineered and biocompatible ceramics, have become available for new applications that require geometries and aspect ratios that the existing processes cannot deliver.

Combining the two approaches, new possibilities for the design and manufacture of complex, high-

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accuracy micro parts in high-performance engineering materials can be utilised.

[1] Kunieda, M., Lauwers, B., Rajurkar, K. Schumacher, B. (2005). Advancing EDM through Fundamental Insight into the Process. CIRP Annals - Manufacturing Technology, vol. 54, p. 64-87, DOI:10.1016/S00078506(07)60020-1.

[2] Liao, Y.S., Chang, T.Y., Chuang, T.J. (2008). An on-line monitoring system for a micro electrical discharge machining (micro-EDM) process. Journal of Micromechanics and Microengineering, vol. 18, no. 3, 035009, DOI:10.1088/0960-1317/18/3/035009.

[3] Garn, R., Schubert, A. Zeidler, H. (2011). Analysis of the effect of vibrations on the micro-EDM process at the workpiece surface. Precision Engineering, vol. 35, no. 2, p. 364-368, DOI:10.1016/j. precisioneng.2010.09.015.

[4] Yeo, S.H., Tan, L.K. (1999). Effects of ultrasonic vibrations in micro electro-discharge machining of microholes. Journal of Micromechanics and Microengineering, vol. 9, no. 4, p. 345-352, DOI: 10.1088/0960-1317/9/4/310.

[5] Freiman, S. (ed.) (2007). Global Roadmap for Ceramics and Glass Technology, John Wiley & Sons, Inc., Hoboken.

[6] Mohri, N., Fukuzawa, Y., Tani, T., Saito, N. Furutani, K. (1996). Assisting electrode method for machining insulating ceramics. CIRP Annals - Manufacturing Technology, vol. 45, p. 201-204.

[7] Mohri, N., Chen, S., Fukuzawa, Y., Tani, T., Sata, T. (2002). Some considerations to machining characteristics of insulating ceramics -towards practical use in industry. CIRP Annals - Manufacturing Technology, vol. 51, p. 161-164.

[8] Schubert, A. Zeidler, H. (2009). Manufacturing of Nonconductive ZrO2 Ceramics with MicroEDM. Proceedings of the Euspen 9th International Conference, vol. 2, p. 6-9.

[9] Schubert, A. Zeidler, H. (2010). Implementation of micro-EDM process for machining of nonconductive ceramics. Proceedings of the 16th International Symposium on Electromachining ISEM-XVI, p. 565570.

Schubert, A. – Zeidler, H. – Hackert-Oschätzchen, M. – Schneider, J. – Hahn, M.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 165-174 © 2013 Journal of Mechanical Engineering. All rights reserved.

DOI:10.5545/sv-jme.2012.522

Received for review: 2012-04-11 Received revised form: 2012-07-31 Accepted for publication: 2012-10-25

Characteristics of Hard Coatings on AZ61 Magnesium Alloys Tański, T. Tomasz Tański*

Silesian University of Technology, Institute of Engineering Materials and Biomaterials, Poland Gradient/monolithic coatings (Ti/TiCN/CrN, Ti/TiCN/TiAlN, Ti/DLC/DLC) were deposited onto magnesium alloy (Mg-Al-Zn) substrate by Cathodic Arc Evaporation method and Plasma Assisted Chemical Vapor Deposition method. A thin metallic layer (Ti) was deposited prior to deposition of gradient coatings to improve adhesion. The microstructure wear resistance and adhesion of the investigated coatings were studied. SEM micrographs showed that the deposited coatings are characterized by compact structure without delamination or defects and they closely adhere to each other. The critical load LC lies within the range of 8-17 N, depending on the coating type. The DLC coatings demonstrate the highest wear resistance. The good properties of the PVD gradient coatings make them suitable in various industrial applications.

Keywords: Magnesium alloys, hard coatings, physical vapour deposition (PVD), plasma assisted chemical vapor deposition (PACVD), structure, properties

0 INTRODUCTION Magnesium alloys have excellent physical and mechanical properties for a number of applications. In particular its high strength, weight ratio makes it an ideal metal for automotive and aerospace applications, where weight reduction is of significant concern.

Magnesium is the 6th most abundant element on Earth making up approximately 2.09% by mass of the Earth’s crust. Magnesium alloys, in addition to a low density of (1.7 g/cm3), also have some other advantages like good ductility, noise and vibration dampening characteristics and excellent castability, high stability of the size and shape, low shrinkage, as well recyclability, which makes it possible to achieve recycled alloys with quality and properties very close to primary cast alloys, which enables the application of these materials instead of new manufactured Mg alloy for constructions of less importance [1] to [6].

Due to limited fossil fuel stores and environmental problems associated with fuel emission products, there is a push in the automotive industry to make cars lighter in order to decrease fuel consumption. The use of magnesium alloys can significantly decrease the weight of automobiles without sacrificing structural strength [7] to [10].

For the reason of surface protection of the applied engineering materials, including the discussed magnesium alloys for improving their surface properties, often diverse kinds and types of surface engineering techniques are used, including technologies for coating of surface layers during PVD processes.

The progress in the area of manufacturing and extending the life expectancy of the constructional elements and tools, used in various domains of life, is taking place mostly owing to the more and more common employment of the thin coatings deposition,

made from hard, wear resistant ceramic materials. A big selection of coatings available actually and their deposition technologies are the effects of the recent growing demand for the contemporary modification methods and materials’ surface protection. Currently, the CVD (Chemical Vapour Deposition), and PVD (Physical Vapour Deposition) methods play an important role in the industrial practice among many techniques improving the life of materials [11] to [17]. The PVD techniques are currently commonly used for improving the mechanical and functional properties of a wide range of engineering materials [18] to [20]. Thanks to the relatively low PVD process temperature, other than the CVD processes there is no risk of losing the properties acquired during the heat treatment of the coated materials.

On the leading annual World Congresses on coatings, including ThinFilms2012 in Singapore, among the presented articles several of them concern surface layers with a gradient-, composite-, or nanocomposites structure, obtained using the PVD and CVD technique, replacing older generation coatings, such as simple - monolayer, alloying/multicomponent or multiphase layers, as well as coated on magnesium and aluminium alloys. Progress in the field of coating production obtained by the physical vapour deposition process allows it to achieve hybrid functional coatings - gradient and multi-component, which reveals high mechanical and functional properties. Coatings of this structure have a low friction coefficient (selflubricating coatings) in many working environments, while maintaining high hardness and increased resistance to external factors. The key issue seems to be also to ensure a simultaneous development of both manufacturing technology and treatment of constructional light materials, in particular of magnesium alloys, as well as technology for modelling and protection of the surface, which will finally allow

*Corr. Author’s Address: Institute of Engineering Materials and Biomaterials, Silesian University of Technology, Konarskiego St. 18a, Gliwice, Poland, tomasz.tański@polsl.pl

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to preserve the balance between the modern substrate materials and the new generation coating.

The aim of this innovative work was to obtain best possible hybrid coatings, consisting of – a gradient transition layer, with a continuous change of one or more components reaching from the substrate to the surface top – as well as an outer coating using the Cathodic Arc Evaporation process and Plasma Assisted Chemical Vapor Deposition process on the surface of the cast AZ61 magnesium alloys to increase properties and the low stiffness of the substrate material. This article focuses on the properties and structure analysis in order to evaluate the quality of the obtained coating.

University of Silesia. TEM specimens were prepared by cutting thin plates from the material. The specimens were ground down to foils with a maximum thickness of 80 μm before 3 mm diameter discs were punched from the specimens. The disks were further thinned by ion milling method with the Precision Ion Polishing System (PIPS™), used the ion milling device model 691 supplied by Gatan until one or more holes appeared. The ion milling was done with argon ions, accelerated by a voltage of 15 kV. Table 1. Chemical composition of the investigated magnesium alloys Type of material

1 EXPERIMENTAL PROCEDURE

AZ61

The material used for investigation was the magnesium alloy AZ61. The chemical composition of the investigated alloy is presented in Table 1. The CAE method was employed in own research for depositing the hard, wear resistant PVD coatings (TiCN/CrN, TiCN/TiAlN). The coatings deposition process was carried out in the arc-vacuum chamber based on the arc evaporation method, the so called Cathodic Arc Evaporation in an Ar, N2 and C2H2 atmosphere.

Cathodes containing pure metals (Cr, Ti) and the TiAl (50:50 at. %) alloy were used for deposition of the coatings. The diameter of the used cathodes was 65 mm. After pumping the chamber the base pressure was 5×10-3 Pa (Table 2). The temperature was controlled by thermocouples. Then, the substrates were cleaned by argon ion at the pressure 2 Pa for 20 min. To improve the adhesion of coatings, a transition Cr or Ti interlayer was deposited. The working pressure during the deposition process was 2 to 4 Pa depending of the coatings type. The distance between each of the cathodes and the deposited substrates was 120 mm. Moreover, the DLC coating were deposited using acetylene (C2H2) as precursor and was produced

by PACVD process. The substrates were cleaned by argon ion at the pressure 2 Pa for 20 min in bias voltage 800/200 V. To improve the adhesion of coatings, a transition Ti interlayer was deposited. Conditions of coating deposition are presented in Table 2.

The examinations of thin foils microstructure and phase identification were made on the JEOL 3010CX transmission electron microscope (TEM), at the accelerating voltage of 300 kV using selected area diffraction method (SAD) for phase investigations.

The diffraction patterns from the TEM were solved using a special computer program “Eldyf” software supplied by the Institute of Material Science o the 166

Al 5.92

Mass concentration of the elements [%] Zn Mn Si Fe Mg 0.49 0.15 0.04 0.01 93.33

Rest 0.06

Table 2. Deposition parameters of the investigated coatings

Process parameters

Coating type PVD process Ti/Ti(C,N)Ti/Ti(C,N)gradient/(Ti,Al)N gradient/CrN

PACVD process Ti/DLC/DLC

Base pressure [Pa] Working pressure [Pa] Argon flow rate [sccm (standard cubic centimeter per minute)]

5×10-3

5×10-3

9.0×10-1 / 1.1 to 1.9/2.8 80* 10**

9.0×10-1 / 1.1 to 1.9/2.2 80* 10**

80* -

10***

10***

-

Nitrogen flow rate [sccm]

0→225** 350***

225→0** 250***

-

Acetylene flow rate [sccm]

140→0**

0→170**

230

70* 70** 70*** 60

70* 70** 60*** 60

Substrate bias voltage [V]

Target current [A] Process <150 temperature [ºC] *during metallic layers deposition; **during gradient layers deposition; *** during ceramic layers deposition.

<150

1x10-3 2

500 <180

Microstructure investigation was performed using scanning electron microscope (SEM) ZEISS Supra 25 with a magnification between 10000 and 35000 times.

For microstructure evaluation the Secondary Electrons (SE) detection was used, with the accelerating voltage of 5 to 25 KV. For a complex metallographic analysis of the fractures of the investigated samples, the material with the coated layer was initially cut, and before braking cooled down in liquid nitrogen. Qualitative and quantitative chemical composition analysis in

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micro-areas of the investigated coatings was performed using the X-Ray microanalysis (EDS) by mind of the spectrometer EDS LINK ISIS supplied by Oxrord. This device is attached to the electron scanning microscope Zeiss Supra 35. The investigations were performed by an accelerating voltage of 20 kV. Wear resistance investigations were performed using the ball-on-disk method in dry friction conditions in the horizontal settlement of the rotation axis of the disk. As the counterpart there was a tungsten carbide ball with a diameter of 3 mm used. The tests were performed at room temperature by a defined time using the following test conditions: • load, Fn-5 N, • rotation of the disk 200 turns/min, • wear radius of 2.5 mm, • shift rate of –0.05 m/s.

Tests of the coatings’ adhesion to the substrate material were made using the scratch test on the CSEM REVETEST device, by moving the diamond indenter along the examined specimen surface with gradually increasing load. The device registered the friction force, friction coefficient, indenter penetration depth and acoustic emission along the scratch track.

The tests were made using the following parameters: load range: 0 to 100 N; load increase rate (dL/dt): 100 N/min; indenter’s sliding speed (dx/dt): 10 mm/min; acoustic emission detector’s sensitivity AE: 1.

has allowed it to identify the (Ti, Al)N phase (Figs.

3 and 4) as a Cubic phase of the Fm3m (225) space group with the d-spacing of a = b = c = 0.424173 nm.

For the Ti/Ti(C,N)/CrN coated Mg substrate the CrN phase as a cubic phase of the 225-Fm3m space group with the d-spacing of a = b = c = 0.414 nm (Figs. 5 and 6) was determined.

Fig. 1. Microstructure of the (Ti,Al)N surface layer, bright field, TEM

2 RESULTS AND DISCUSSION Results of diffraction method investigations allow to identify the (Ti,Al)N, CrN and graphit phases occurred in the surface layer (Figs. 4, 6 and 10).

TEM investigation results are presented on Figs.

1 to 10. For analysed cases in the investigated areas a nanocrystalline microstructure of the surface layer was detected. In Fig. 2 the microstructures of the layer (Ti,Al)N are shown and graphite phases, using the dark field technique the size of the subgrains or crystallites can be determined. Also, a globular compact structure was confirmed as well a high uniformity of the crystallites with a uniform size within the range of 10 to 15 nm. According to the literature and based on own investigations it is possible to conclude that the character (unevenly distributed large clusters of reflections from the coating, forming large systematic groups in the polycrystalline diffraction circles) of the electron diffraction pattern of the coating Ti(C,N)/ CrN shows the occurrence of fine-grained crystallite structure of the size of ≤10 nm (Fig. 5). For phase determination of the structure of the surface layer diffraction pattern analysis of the investigated areas

Fig. 2. Microstructure of the (Ti,Al)N surface layer, dark field, TEM

The graphite phase was determined as a hexagonal phase of the 186-P63mc space group with the d-spacing of a = b = 0.2, c = 0.679 nm (Figs. 9

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and 10), identification of the radius of the first circle visible on the obtained diffraction pattern (Fig. 10) allowed unambiguous confirmation of the graphite phase.

and Ti/Ti(C,N)/(Ti,Al)N layers can be determined in the range up to 3.3 µm, while the carbon layer thickness oscillated between 2.5 mm. Fracture investigations confirm also that the Ti/Ti(C,N)/CrN and Ti/Ti(C,N)/(Ti,Al)N layers has a layered structure with a clearly visible transition zone between the gradient layer and the wear resistant coating achieved using separate metals evaporation sources (Figs. 11 and 13).

Fig. 3. Diffraction pattern of the polycrystalline surface layer presented in Figs. 1 and 2

Fig. 5. Microstructure of the CrN surface layer, bright field and diffraction pattern of the polycrystalline surface layer, TEM

Fig. 4. Solution of the diffraction pattern presented in Fig. 3

As a result of the microstructure investigations on scanning electron microscope it has been found that there are no pores or cracks in the produced coating and any defects and failures occurring spontaneously in this single layer are not of significant importance for the properties of the whole layer (Figs. 11, 13 and 15).

Thickness of the obtained coatings was measured using metallographic observations performed on the scanning electron microscope on sample fractures.

Thickness measurements in different locations of the observed fractures have confirmed the uniformity of the coated layers. The thickens of the Ti/Ti(C,N)/CrN 168

Fig. 6. Solution of the diffraction pattern presented in Fig. 7

Due to the same phase composition of the gradient coatings and the wear resistant coating the characteristic separation between these both coatings in the case of carbon coating could not be identified, however characteristic bright, continuous layer Ti in area of the sub-shell, which improves the adhesion (Fig. 15) could be identified. The morphology of the

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fracture of the layers is also characterised by a lack of columnar structure. (Figs. 11 to 15).

irregular shape, slightly flat). There were probably also some hollows formed when the solidified droplets break off after the PVD process has been completed (Figs. 12 and14). The occurrence of such morphology defects is connected to the nature of the cathodic process of the electric arc evaporation. In case of the Ti/DLC/DLC coating there were identified small amounts of droplets (Fig. 16).

Fig. 7. Microstructure of the DLC surface layer, bright field, TEM

Fig. 9. Diffraction pattern of the surface layer presented in Figs. 1 and 2 Graphite

Fig. 10. Solution of the diffraction pattern presented in Fig. 3 Fig. 8. Microstructure of the DLC surface layer, dark field, TEM

On the basis of the performed observations on scanning electron microscope the coating of the Ti/ DLC/DLC type show an increasing non-homogeneity compared to the Ti/Ti(C,N)/(Ti,Al)N and Ti/Ti(C,N)/ CrN coatings which is connected with the presence of numerous droplet-shaped microparticles and should that fore significantly influence on mechanical properties of the achieved coatings (Figs. 12, 14 and 16). The droplets observed in SEM are noticeably different in terms of size and shape (regular and

As a result obtained by the quantitative X-ray microanalysis using the energy dispersed X-ray EDS spectrometer it the presence of Mg, Al, Zn, Ti, Cr, C, N, as major alloying elements of the cast of magnesium alloys as well the obtained coatings (Fig. 17, Table 4) has been Confirmed. As the size of the particular structure elements is mainly lower than the diameter at the incident electron beam the achieved image in the quantitative EDS analysis can be calculated as the average value, which can lead to partially enhanced values for some element concentrations.

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Fig. 11. Cross-section SEM images of the Ti/Ti(C,N)/(Ti,Al)N coating deposited onto the AZ61 substrate

Fig. 12. Surface topography of the Ti/Ti(C,N)/(Ti,Al)N coating deposited onto AZ61 substrate

Cursor Hight=3.224 mm

Fig. 13. Cross-section SEM images of the Ti/Ti(C,N)/CrN coating deposited onto the AZ61 substrate

Fig. 14. Surface topography of the Ti/Ti(C,N)/CrN coating deposited onto AZ61 substrate

Fig. 15. Cross-section SEM images of the Ti/DLC/DLC coating deposited onto the AZ61 substrate

Fig. 16. Surface topography of the Ti/DLC/DLC coating deposited onto AZ61 substrate

To determine the tribological properties of the investigated coating deposited on the magnesium

alloys substrate, an abrasion test under dry slide friction conditions was carried out by the ball-on-

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disk method. Table 5 presents the friction coefficient and sliding distance results for each type of the investigated substrate.

Under technically dry friction conditions, after the wearing-in period, the friction coefficient recorded for the associations tested is stabilized in the range 0.09 to 0.32 depending on the used substrate and coatings.

Rapid changes of the friction coefficient value are caused by the occurrence of pollutants in form of sample counterface spalling products (balls are made from WC), which disturb the measurement of the friction coefficient. Comparing the friction coefficient results with the friction path length, it has been found that the best wear resistance is characteristic for materials coated with DLC carbon. According to the applied load of 5 N, the average friction coefficients for the DLC coatings with the sliding rate of 0.05 m/s is in the range of 0.09 to 0.19, which is several times lower compared to the friction coefficient values of other examined coatings. However, the results of the friction path length for the DLC coatings were at a level exceeding even 11 times the results of the friction path length achieved for the Ti/Ti(C,N)/CrN coatings. This is characteristic for DLC coatings as they are composed of poorly ordered graphite, which is probably formed by a friction-assisted phase transformation of the surface layer of the DLC matrix and acts as a lubricant at the surface [18]. Accordingly, the high hardness of DLC together with this transfer layer is responsible for the low friction coefficient of the DLC film in comparison with magnesium alloys coated other investigated coatings. At high sliding speed, the transfer layer is more easily formed due to the accumulation of heat, resulting in a lower friction coefficient.

Fig. 17. Cross-section SEM images of the Ti/Ti(C,N)/CrN coating deposited onto the AZ61 substrate

Table 4. The results of quantitative chemical analysis from third 1, 2, 3 areas of coating Ti/Ti(C,N)/CrN deposited onto substrate from AZ61 alloy marked in Fig. 17 Chemical element N Ti Cr Matrix C N Mg Ti Cr Matrix Zn Mg Al Matrix

The mass and atomic concentration of main elements [%] mass atomic Analysis 1 (point 1) 19.64 38.24 07.85 06.71 72.51 55.05 Correction ZAF Analysis 2 (point 2) 27.87 40.83 09.12 17.86 01.83 02.01 52.82 35.01 08.36 4.29 Correction ZAF Analysis 2 (point 3) 03.35 02.21 85.39 87.13 11.26 10.66 Correction ZAF

Table 5. The characteristics of the investigated coatings

Coatings Ti/Ti(C,N)/CrN Ti/Ti(C,N)/(Ti,Al)N Ti/DLC/DLC

Substrate-AZ61 Critical Critical Friction load LC1 load LC2 coefficient [N] [N] 3 8 0.15 to 0.32 3 10 0.17 to 0.22 6 17 0.09 to 0.19

Sliding disance [m] 57 77.7 630

The critical loads LC1 and LC2 were determined in scratch tests. The first critical load LC1 corresponds to the first small jump on the acoustic emission signal, as well as on the friction force curve. The second critical load LC2 is the point at which complete delamination of the coating starts; the first appearance of cracking, chipping, spallation and delamination outside or inside the track with the exposure of the substrate material-the first adhesion-related failure event (Figs. 18b, 19b, 20b). After this point, all the acoustic emission and friction force signals become noisier [14]. The cumulative test results have been listed and compared in Table 3. The highest critical load value LC1 = 6 N and LC2 = 17 N, and that fore the best adhesion of the coating to the substrate was obtained for the Ti/DLC/DLC coating, which is not the best result of coating adhesion to the substrate material, depending mainly on the hardness of the substrate, but for a soft material like magnesium alloy, 17 N is a plausible result.

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a) b) Fig. 18. Scratch failure pictures of the Ti/Ti(C,N)/(Ti,Al)N coating on AZ61 substrate at; a) LC1, and b) LC2

a) b) Fig. 19. Scratch failure pictures of the Ti/Ti(C,N)/CrN coating on AZ61 substrate at; a) LC1, and b) LC2

a) b) Fig. 20. Scratch failure pictures of the Ti/DLC/DLC coating on AZ61 substrate at; a) LC1, and b) LC2

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3 SUMMARY Due to the character of the investigated material (magnesium alloys) and its relatively low melting point, the whole technological PVD and PACVD processes were performed at temperatures up to 150 °C for Ti/Ti(C,N)/CrN coatings and Ti/Ti(C,N)/ (Ti,Al)N coatings, and up to 180° C for Ti/DLC/DLC coating. Investigations carried out using transmission electron microscopy have conformed the (Ti,Al)N, CrN and graphite phases occurred in the surface layer.

In order to evaluate the crystallite size of the obtained phases there was used the dark field technique. The size of measured crystallites is in the range of 15 nm.

Fracture investigations have confirmed the occurrence of a sharp transition zone between the substrate and the coating. Moreover, it can be observed that the obtained coatings show a tight and compact structure with a lack of visible delamination and defects, are uniformly coated and tightly adhere to the substrate.

Comparing the friction coefficient results with the friction path length it has been found that the best wear resistance is characteristic for materials coated with DLC carbon. The scratch tests on coating adhesion reveal the cohesive and adhesive properties of the coatings deposited on the substrate of the AZ61 magnesium alloys. On the basis of the above examinations, it has been found that the critical load of LC2 is between 8 and 17 N. The highest value of the critical load was obtained for the Ti/DLC/DLC coating.

4 ACKNOWLEDGEMENTS Research was financed partially within the framework of the Polish State Committee for Scientific Research Project No. 4688/T02/2009/37 headed by Dr Tomasz Tański.

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[17] Bujak, J., Walkowicz, J., Kusiński, J. (2004).

Influence of the nitrogen pressure on the structure and properties of (Ti,Al)N coatings deposited by cathodic vacuum arc PVD process. Surface and Coatings Technology, vol. 180-181, p. 150-157, DOI:10.1016/j. surfcoat.2003.10.058.

[18] Lukaszkowicz, K., Dobrzański, L.A., (2008). Structure and mechanical properties of gradient coatings

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deposited by PVD technology onto the X40CrMoV5-1 steel substrate. Journal of Materials Science, vol. 43 (2008), p. 3400-3407, DOI:10.1007/s10853-008-25233˝.

[19] Tański, T., Labisz, K. (2012). Electron microscope investigation of PVD coated aluminium alloy surface layer. Solid State Phenomena, vol. 186, p. 192-197, DOI:10.4028/www.scientific.net/SSP.186.192.

[20] Tański, T. (2012). Investigation of the structure and properties of PVD and PACVD-coated magnesium die cast alloys, Magnesium Alloys. Monteiro, W.A. (ed.), New Features on Magnesium Alloys. InTech, Rijeka, p.

29-52.

Tański, T.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 175-182 © 2013 Journal of Mechanical Engineering. All rights reserved.

DOI:10.5545/sv-jme.2012.625

Received for review: 2012-06-02 Received revised form: 2012-10-28 Accepted for publication: 2012-11-29

Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques Dobrzański, L.A. – Musztyfaga, M. – Drygała, A. Leszek Adam Dobrzański – Małgorzata Musztyfaga* – Aleksandra Drygała

Institute of Engineering Materials and Biomaterials, Silesian University of Technology, Poland The paper presents the results of the investigation of the front electrode manufactured using two silver pastes (PV 145 manufactured by Du Pont and another based on nanopowder experimentally prepared) on monocrystalline silicon solar cells in order to reduce contact resistance.

The aim of the paper was a comparison between a conventional and an unconventional method to improve the quality of forming electrodes of silicon solar cells. The Screen Printing (SP) method is the most widely used contact formation technique for commercial silicon solar cells.

Selective Laser Sintering (SLS) is a modern manufacturing technique, which uses a high power CO2 laser to melt or sinter metal powder particles into a mass that has the desired three-dimensional shape, in precisely defined areas. The whole process is controlled by a program that is used for micro-processing. An innovative aspect of this method is the application of a nano silver paste to create the seed layer of the front electrode using the SLS method.

The topography of both co-fired in the infrared belt furnace and melted/sintered in the Eosint M250 Xtended device equipped with CO2 laser were studied. Contacts were investigated using confocal laser scanning microscope (CLSM 5) and a scanning electron microscope (SEM) with an energy dispersive X-ray (EDS) spectrometer for microchemical analysis. Both surface topography and a cross section of the front electrodes were studied using SEM microscope. Phase composition analyses of the chosen front electrodes were done using the XRD method delineated in this paper. Front electrodes were formed on the surface with a different morphology from the solar cells. The median size of the pyramids was measured using the atomic force microscope (AFM). Resistance of the front electrodes was measured using the Transmission Line Model (TLM).

Keywords: electrical properties, solar cell, selective laser sintering, screen printing

0 INTRODUCTION The main focus of this paper was to study the manufacture of the front side metallization of solar cells. Many different methods have been applied to improve the electrical properties of front metallization. The photovoltaic market for silicon solar cells has been growing strongly over the past few years. Scientists are pushed to develop quicker some methods and technologies by the growth, at the same time, achieving both an improvement in efficiency and a reduction in the production costs of photovoltaic cells. Metallization is one of the key process steps in the manufacture of high efficiency solar cells. Fig. 1 presents some methods of manufacturing metal

contacts at the front and back surfaces of solar cells [1] to [5].

The aim of the paper was a comparison of a conventional and an unconventional method to improve the quality of forming electrodes on silicon solar cells. The Screen Printing (SP) method is the most widely used contact formation technique for commercial silicon solar cells. Screen printing is performed on both the front and back sides of solar cell (Fig. 2). Each of the screen printing processes can be divided into three major steps [4] to [6]: 1. Overprinting a collection of back contacts (Al/Ag) and drying, 2. Overprinting a front electrode (Ag) and drying, 3. Co-firing both front and back metal contacts.

Fig. 1. Methods of producing electrical contacts on solar cells [4] and [5] *Corr. Author’s Address: Institute of Engineering Materials and Biomaterials, Silesian University of Technology, Konarskiego St. 18a, Gliwice, Poland, malgorzata.musztyfaga@polsl.pl

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Selective Laser Sintering (SLS) is a modern manufacturing technique, which uses a high power CO2 laser to melt or sinter metal powder particles into a mass that has a desired three-dimensional shape in precisely defined areas. The whole process is controlled by a program that is used for microprocessing. Process technology for the generation of front contacts by SLS is shown in Fig. 3 [4] and [5].

front contact was carried out by experiment and mixtures were prepared using a mechanical mixer.

The technology used to produce the solar cell was developed in the Institute of Metallurgy and Materials Science in Krakow, Poland. The process sequence for manufacturing solar cells into different surface morphologies of silicon (100) is presented in Fig. 4.

Table 1. Basic properties of silicon Type Doped Thickness Area Resistivity

p boron 200 ± 30 μm 5 x 5 cm 1 ÷ 3 Ω•cm

p boron 330 ± 10 μm 5 × 5 cm ~1 Ω•cm

Fig. 2. The screen printing method [4] to [6]

1 EXPERIMENTAL PROCEDURE The studies were carried out on monocrystalline silicon wafers produced by Deutsche Solar AG Germany. The material properties of the silicon used in this paper are presented in Table 1. The SLS method was used on wafers with a thickness ~330 μm, whereas the SP method and co-firing in the furnace was used on wafers with a thickness of ~230 μm.

The silver powder with the following granulation: 70, 60 and 40 nm was applied during the preliminary investigations in order to select both the optimal granulation and the powder. Finally, the nanopowder was applied in investigations based on metallographic observations of front electrodes manufactured using the SLS method. The selection of elements in the contact layer of the solar cell is presented in Table 2. The selection of the chemical composition of the

Fig. 4. The solar cell manufacturing process

The chemical procedure for cleaning the wafers is presented in Table 3. Emitter diffusion is one of the crucial steps in the manufacturing process of solar cells. The CVD method was used to apply the donor source. The emitter was generated at 840 ºC for 40 minutes in an open quartz tube using POCl3 as the doping source. The parasitic junction was removed by means of a special teflon clamp in which the

Fig. 3. Technology for the formation of front contacts by SLS: a) a metallic powder is coated on the wafer, b) a laser beam heats the powder locally and melts metal lines on the top of the wafer, c) the excess metallic powder is moved into the container (1 – p type base material, 2 – n type emitter, 3 – dielectric layer, 4 – textured layer, 5 – powder layer, 6 – passage of focus laser beam, 7 – SLS seed layers on the silicon surface) [4] and [5]

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silicon wafers were stacked using surface EVA foil separation.

Then the clap was immersed in 27%HF:45%HNO3:27%CH3CO-OH solution in the volume ratio 3:5:3 for 40 seconds, followed by rising in DIH2O. After diffusion, the wafers were covered by phosphorous-silicate glass (xSiO2·yP2O5), which was removed by immersion in a bath of 10% HF for two minutes. Surface passivation (SiO2) took place at 800 °C for 10 minutes in a controlled atmosphere containing both O2 and N2. Titanium dioxide (TiOx) as an antireflection coating was deposited by spraying at 300 °C with tetraethylorthotitanat ((C2H5O)4Ti) using purified air as a carrier gas. The important aspect of this paper is the fabrication of the front metallization (the test electrodes system). This front metallization (test electrodes system) was manufactured in order to describe the usefulness of silver pastes using two technologies and to evaluate the contact resistance of the metal-semiconductor junction. Two special test electrode systems were prepared by screen printing method (Fig. 5.): • I sizes of the front paths were: 2 × 10 mm (wide x length), distances between them were: 20, 10, 5 and 2.5 mm, • II sizes of front paths were: 5 × 10 mm (wide × length), distances between them were: 1, 2, 4 and 8 mm.

Fig. 5. Overview of test front electrode system

Many initial series for testing the solar cells were first prepared by laser micro-treatment in order to select conditions for the tests (the feed rate of passage of the laser beam and laser beam) and achieve the smallest resistance value of the connection zone between the electrode and the silicon substrate of the solar cell and a uniform structure. The most advantageous conditions of laser micro-treatment were selected based on the tests and metallographic observations. Table 4 shows the two chosen initial series of solar cell testing, which were then prepared by laser micro-treatment. Then laser micro-treatment

conditions were selected for the test electrodes system I, II (Table 5) based on the results of the metallographic observations. Table 6 presents the general device data.

In the case of co-firing in the conveyor belt IR furnace (Fig. 6), some solar cells were prepared using the test electrodes system I, II (Table 7). The belt IR furnace was equipped with fitted tungsten filament lamps, heating both the top and bottom of the belt.

Fig. 6. A sketch of belt IR (an example) [7] Table 2. Paste properties Mass concentration of elements Solar cells (prepared pastes) [%] with different Basic Organic Ceramic morphology** nanopowder carrier glaze UNCONVECTIONAL METHOD A 83 15 2 1, 2, 3, 4 PV145* 1, 2, 3, 4 CONVECTIONAL METHOD B 85 15 1, 2 C 88.40 11.60 3, 4 PV145* 1, 2, 3, 4 **Where: 1. Non-textured surface with deposited TiOx coating; 2. Non-textured surface without deposited TiOx coating; 3. Textured surface with deposited TiOx coating; 4. Textured surface without deposited TiOx coating; * commercial paste manufactured by the Du Pont company Paste symbol

Table 3. Chemical preparation of silicon wafers Chemical operations Washing in acetone Rinsing Removal of distorted layer Rinsing Removal of metallic contamination Removal of native oxide Rinsing

Chemical recipe CH3COCH3 DIH2O 30% KOH DIH2O 2%HCl

Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques

10%HF DIH2O

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Table 4. Initial conditions of laser micro-treatment for the test electrode system I, II of silicon wafers Series

Solar cell surface

Paste symbol

1 2

Chemically cleaned

PV 145 A

Feed rate of passage of laser beam (ν), [mm/s] 50 to 200 50 to 200

Laser beam, (P) [W] 8.1 to 37.8 21.5 to 48.6

Thickness of printed front electrode by screen printing method [µm] 15, 40, 60 15, 35, 70

Table 5. Conditions of laser micro-treatment testing electrodes of silicon solar cells with test electrode system I, II Series

Solar cells with different morphology*

Paste symbol

The thickness of printed front electrode by screen printing method, [µm]

The feed rate of passage the laser beam (ν) [mm/s]

Laser beam (P) [W]

1

1, 2, 3, 4

A

15, 35

37.8, 40.5, 43.2

50

2

1, 2, 3, 4

A

15, 35

48.5, 51.3, 54

100

*Where: 1. non textured surface with deposited TiOx coating; 2. non textured surface without deposited TiOx coating; 3. Textured surface with deposited TiOx coating; 4. Textured surface without deposited TiOx coating

The following investigations were performed in this paper: • The contact resistance Rc, specific contact resistance ρc, transfer length (LT) of front contact solar cell, and the use of the Transmission Line Model (TLM) method for the measuring position were determined at the Institute of Engineering Materials and Biomaterials.

• The topography of the silicon wafer with texture and antireflection coating (ARC) (this increases the absorption of sunlight and protects against pollution) were determined using an atomic force microscope (Park Systems XE 100) with an uncontacted trybe. The median size of the pyramids was also measured using this microscope.

• Phase composition analysis of the chosen front contacts using the XRD method.

• Microchemical analysis of the front chosen contacts using a scanning electron microscope equipped with an energy dispersive X-ray (EDS) spectrometer.

• The topography of both the surface and cross section of front contacts using:

• Zeiss Supra 35 scanning electron microscope using secondary electron detection with accelerating voltage in the range 5 to 20 kV. • Zeiss confocal laser scanning microscope in which the source of light was a diode laser of approximately 25 mW emitting radiation with a wavelength of approx. 405 nm. The thickness of the profile contact was determined on the basis of six median measurements.

2 RESULTS AND DISCUSSION Based on the metallographic observations (Fig. 7), it was found that is not possible to apply silver pastes 178

to deposit front electrodes (about regular shape and uniform thickness) using EOSINT M 250 Xtended device with SLS technology. Fig. 8a presents test electrode system I deposited using PV145 paste and selectively laser sintered. The observations revealed the occurrence of grains of diversified size and shape, similar to the electrode structure from the same paste before its laser micro-treatment (Fig. 8b), total overbaking, and evaporation of the electrode with increasing laser beam power. Based on the graphs, which present the resistance – distance relationship for the contacts involving parameters: ρc (the specific contact resistance), LT (a track of current impact), and Rc (contact resistance) (Fig. 9), the specific contact resistance defines not only the real joint zone of contact with the Si substrate, but the regions directly under and below surface of phase separation. The specific contact resistance of the front contact was calculated from formula in the literature, but other parameters like Rc and LT were calculated from the linear regression. Contact resistance was measure using the TLM method, which consists of a direct current (I) measurement and voltage (U) measurement between any two separate contacts. In this paper, results were obtained for the specific contacts resistance of the front electrode for a given value of current (10, 30 and 50 mA) depending on the determined conditions of both the laser micro-treatment samples and the co-firing samples in the furnace. Fig. 9 presents an example of the series of test samples co-fired from the PV 145 paste in the temperature range of 830 to 920 °C onto a silicon solar cell with an ARC layer and without texture for a current value of 10 mA. Based on the electrical properties investigations using the TLM method in the first series, it was found that the smallest specific contact resistance values of the test electrodes system I, II are respectively 0.17 to 0.57 Ω·cm2, 0.53 to 1.63 Ω·cm2 for solar cells with

Dobrzański, L.A. – Musztyfaga, M. – Drygała, A.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 175-182

different morphologies. The minimum value of the specific contact resistance was obtained for the test electrodes system I, II (0.17 Ω·cm2; 0.53 Ω·cm2) with a median thickness of 35 µm onto substrate without texture and without TiOx coating using an applied laser beam of 37.8 W and a 50 mm/s feed rate of passage of the laser beam.

Based on the electrical properties investigations using the TLM method in the second series it was found that the smallest specific contact resistance of the test electrode system I, II are respectively 0.01 to 0.77 Ω·cm2, 0.68 to 2.82 Ω·cm2 for solar cells with different morphologies.

Table 6. Technical specification of Eosint M 250 Xtended device Laser CO2 Feed rate of passage the laser beam Output laser beam Diameter of laser beam Wavelength Shielding gas

200 W max. 3.0 m/s 270 W 300 μm 10640 nm nitrogen

The minimum value of the specific contact resistance was obtained for the test electrodes system I (0.01 Ω·cm2, P = 51.3 W), II (0.68 Ω·cm2, P = 48.6 W) about median thickness 35 µm onto the substrate without texture and without TiOx coating with an applied feed rate of passage of the laser beam of 100 mm/s.

The thickness of the test electrodes was determined by checking the height profile of the threedimensional surface topography measured using the confocal laser scanning microscope (Fig. 10).

Table 7. Conditions of co-firing in the furnace test electrodes systems I,II for silicon solar cells Series

Paste symbol

Zone I

1

B, C

530

2

PV 145

530

Temperature [°C] Zone II Zone III 830 860 890 570 920 945 830 860 570 890 920

Based on these series, it was found that silicon substrate morphology has a huge influence on obtaining a minimal resistance value of the electrodes selectively laser sintered from paste A. The minimal

resistance value is greater for the substrate with texture than for the one without texture, which is probably connected to the occurrence of mpty areas under the contacts, since the median thickness of the pyramids’ textured surface for Si (100) is in the range from 3 to 9 µm. The antireflection coating prevents the reflection of rays of sunshine and a loss of energy, as well as creating a barrier to the connection zone, what has an influence on increasing resistance between an electrode layer and a silicon substrate. The thickness of the deposited layer has an influence on the structure of the obtained electrode layer and the resistance value of the resistance electrode. It was found that SLS testing contacts with a median thickness of about 35 µm present both well adhering and condensing layers to the silicon substrate.

Fig. 7. Surface topography of front electrodes system I layer deposited from paste A on the silicon substrate and laser sintered where ν = 100 mm/s and P = 32.4 W

Based on the electrical properties investigations using the TLM method it was found that the smallest specific contact resistance is obtained in a second series of test electrodes system I, II formed from the PV 145 paste. In the temperature range from 860 to 920 °C the minimum value of the specific contact resistance of the test electrode system I, II is equal to 1.08 to 2.73 Ω·cm2; 0.48 to 1.33 Ω·cm2 for solar cells with different morphologies. The highest range of specific contact resistance was found in the first series for front contacts made from C, PV 145 pastes. In the temperature range from 860 to 945 °C the minimum value of the specific contact resistance of test electrodes system I, II is equal to 1.83 to 86.81 Ω·cm2; 1.12 to 71.63 Ω·cm2 for solar cells with different morphologies. Electrical properties investigations using the TLM method confirm that the resistance paste decreases as the co-firing temperature of

Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques

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a) b) Fig. 8. Surface topography of the front electrode system I layer deposited by screen printing from PV 145 paste onto the silicon substrate: a) not sintered, b) SLS with ν=50 mm/s and P=32.4 W

b) a) Fig. 9. a) An example of plot of resistance versus contact distance for the determination of contact parameters (LT - a track of current impact) and Rc -contact resistance, slope= Rp/z, Rp - sheet resistance, Z - contact length); b) typical graphic method from the literature used to determine factors [9] and [10]

Fig. 10. Image of three- and two- dimensional surface topography (CLSM) of front electrode system I from paste A onto the Si substrate without texture or ARC layer by laser micro-machining

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 175-182

electrode increases, this is caused by the decreasing resistance connection between metallic grains. The contact resistance was found to grow with increasing temperature, which was probably caused by melted silver molecules [7], which formed huge agglomerates or where voids would develop under the contacts in textured surfaces.

The contact layers from the preformed PV 145 paste create many homogenous connections with the silicon substrate, while contact layers from the C, PV 145 pastes create point connections. The stability of electrical measurements using the TLM method guaranteed a median thickness of 35 µm of the test electrodes system in the case of paste A, however in a case of the PV 145 paste this influence was not confirmed.

The thickness of the test electrodes co-fired at the highest temperatures with PV 145 and of the test electrodes sintered with a selective laser from the A paste decreases slightly versus the thickness of the electrodes before sintering. As a result of SEM investigations, the examined test front electrodes sintered using a selective laser reveal a diverse crosssection thickness of between 570 nm to 2.3 µm.

The topographies of the silicon wafers with texture and an ARC layer were observed using an atomic force microscope.. The median thickness of the pyramids was determined by atomic force microscope and found to to be equal to 4 µm (Fig. 11).

The qualitative analysis of phase composition (Fig. 12) carried out using the X-ray diffraction method confirms that on the electrodes from sintered pastes, the layers contain the phase Ag, which was generated in congruence with the assumptions (Fig. 13). The presence of reflexes from the SI phase present in the substrate materials was demonstrated in the X-ray diffractograms obtained with the use of the Bragg-Brentano technique.

a)

b)

c) Fig. 12. Images of front electrode system I layer from paste A onto Si substrate without texture or ARC layer by laser micromachinining; a) topography image (SEM); b) fracture image (SEM); c) EDS spectra from X1 area Fig. 11. The topography of monocrystalline silicon wafer with texture (an example)

Fig. 13. X-ray diffraction pattern of front metallization performed from PV 145 paste onto

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silicon surface and co-fired in temperature 830 to 920 °C (where A means solar cells with different morphology) 3 SUMMARY Based on the investigation results, the following optimal process parameters for manufacturing the front electrode of silicon solar cells were selected: selection for the composition of nano silver paste 83%Ag+2%SiO2+15% organic carrier, selection of the median thickness of the deposited layer, laser microtreatment on two test systems – laser beam 37.8 W and feed rate of passage for the laser beam 50 mm/s.

It was found that the silicon substrate morphology has a huge influence on obtaining a minimal resistance value of electrodes using both methods. It is bigger for the substrate with texture than for the one without texture, which is probably connected with the occurrence empty areas under contacts [8], since the median thickness of the pyramids’ textured surface for Si (100) is in the range from 3 to 9 µm.

The antireflection coating prevents the reflection of rays of sunshine and a loss of energy, as well creates a barrier into the connection zone, which increases the resistance between the electrode layer and the silicon substrate. The thickness of the deposited layer has an influence on the structure of the obtained electrode layer and resistance value of the resistance electrode.

It was found that SLS testing of contacts with a median thickness of 35 µm presents both well adhering and condensing layers to the silicon substrate The contact layers performed from the PV 145 paste create many homogenous connections with the silicon substrate, while contact layers from the 85%Ag+15% organic carrier, 88.4%Ag+11.6% organic carrier pastes create point connections.

4 ACKNOWLEDGMENT This research was financed partially within the framework of Scholarship No 51200863 of the International Visegrad Fund, by dr. Małgorzata Musztyfaga.

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5 REFERENCES [1] Gautero, L., Hofmann, M., Rentsch, J., Lemke, A., Mack, S., Seiffe, J., Nekarda, J., Biro, D., Wolf, A., Bitnar, B., Sallese, J.M., Preu, R. (2009). All-screenprinted 120-µm-thin large-area silicon solar cells applying dielectric rear passivation and laser-fired contacts reaching 18% efficiency. Proceedings of Photovoltaic Specialists Conference, 34th IEEE, p.

1888-1893.

[2] Glunz, S.W., Dicker, J., Esterle, M., Hermle, M., Isenberg, J., Kamerewerd, F., Knobloch, J., Willeke, G. (2002). High-efficiency silicon solar cells for lowillumination applications. Proceedings of 29th IEEE Photovoltaic Specialists Conference, p. 450-453.

[3] Green, M.A. (2002). Third generation photovoltaic: solar cells for 2020 and beyond, Physic E, vol. 14, no.

1-2, p. 65-70, DOI:10.1016/S1386-9477(02)00361-2.

[4] Dobrzański, L.A., Musztyfaga, M. (2011). Effect of the front electrode metallisation process on electrical parameters of a silicon solar cell. Journal of Achievements in Materials and Manufacturing Engineering, vol. 48, no. 2, p. 115-144.

[5] Vitanov, P., Goranova, E., Stavbrov, V., Ivanov, P., Singh, P.K. (2009). Fabrication of buried contact silicon solar cells using porous silicon. Solar Energy Materials and Solar Cells, vol. 93, no. 3, p. 297-300, DOI:10.1016/j.solmat.2008.10.015.

[6] Dobrzański, L.A., Musztyfaga, M., Drygała, A., Kwaśny, W., Panek, P. (2011). Structure and electrical properties of screen printed contacts on silicon solar cells. Journal of Achievements in Materials and Manufacturing Engineering, vol. 45, no. 2, p. 141-147.

[7] Hamer, D.W., Biggers, J.V. (1976). Technology of SelfContained System, WNT, Warsaw. (in Polish) [8] Panek, P. (2006) Modeling the structure of porous silicon in aspect electrical and optical properties solar cell, PhD Thesis. Institute Library of Metallurgy and Material Science of Polish Academy of Sciences, Krakow. (in Polish) [9] Goetzberg, A., Knobloch, J., Voß, B. (1994). Crystalline silicon solar cells, John Wiley and Sons.

[10] Dobrzański, L.A., Musztyfaga M., Drygała, A., Panek, P. (2010). Investigation of the screen printed contacts of silicon solar cells from Transmissions Line Model. Journal of Achievements in Materials and Manufacturing Engineering, vol. 41, no. 1-2, p. 57-65.

Dobrzański, L.A. – Musztyfaga, M. – Drygała, A.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 183-192 © 2013 Journal of Mechanical Engineering. All rights reserved.

DOI:10.5545/sv-jme.2012.417

Received for review: 2012-03-07 Received revised form: 2012-08-14 Accepted for publication: 2012-09-11

Characterisation and Processing of Reinforced PA 6 with Halloysite Nanotubes (HNT) for Injection Molding Fernández, A. – Muniesa, M. – González, J. Angel Fernández1,2,* –Manuel Muniesa1 – Jaime González2 1 University de Zaragoza, Mechanical Engineering Department, Spain 2 Fundación AITIIP, Parque Tecnológico Cogullada, Spain

The use of nano-scaled reinforcements enchances the mechanical performance of the resulting material and the addition of natural origin nanotubes improve the sustainability and environmental impact of the product. The present work describes the preparation, characterization and processing of nanocomposites based on thermoplastic Polyamide 6 and Halloysite nanotubes (HNT).

The nanocomposites were prepared in several stages and the results ensure the final application at “industrial scale”. The first formulation was a high concentration of Halloysite nanotubes (up to 30% weight content) made with a twin screw extrusion compounder which was used as a masterbatch. Application materials were obtained reducing the masterbatch in raw PA 6 down to 3 or 6% of HNTs weight content. In this stage the influence of use or not an extruder-compounder was analyzed. Finally, a low cost preparation and sustainable development technique as injection moulding was applied to obtain test parts of all developed materials.

Results were analyzed at a microscopic scale (TGA, FTIR, XRD, SEM, TEM) and macroscopic scale (rheology, mechanical properties and flame resistance). TGA and FTIR results show that the content of HNTs was as accurate as expected. XRD results showed a greater interaction between the HNTs and the matrix if the extrusion-compounding process was used before injection moulding. SEM and TEM results showed a better dispersion of HNTs if extrusion was performed at high screw rotation speed. All analysis concluded the nanocomposites can be processed on standard equipment due to the resultant low viscosity. Also, low content of HNTs proved that mechanical properties are highly improved as flame resistance remains equivalent.

Keywords: nanocomposites, Halloysite nanotubes, injection moulding, compounding

0 INTRODUCTION The newest requirements for lightweight automotive materials embrace improved security performance (high impact energy absortion), lower weight (equivalent mechanical properties with lower density) and improved functional specifications (higher flame retardancy and restricted volatile emission). The new hybrid and electric vehicles (HEV) require the key factors indicated to reduce the weight of existing parts to accommodate and new components as the battery pack; all with at least the same security and the greatest possible autonomy.

Nanocomposites emerge as an alternative to existing materials to meet these new requirements. For this reason the research described in this paper focuses on the characterization and processing at the industry scale of nanocomposites based in thermoplastic PA 6 polymeric matrix and Halloysite nanotubes (HNT) as nano-scaled reinforcement. The polyamide 12 nanocomposites based on halloysite nanotubes are promising candidates for structural applications as Lecouvet [1] studied with a mini-compounder.

The use of nano reinforcements extremely increases the interaction between the resin and the charge resulting in stronger and lighter materials with special properties. The use of Halloysite nanotubes improves the mechanical properties of polyamide using a lower amount of charge if compared with *Corr. Author’s Address: University de Zaragoza, Mechanical Engineering Department,-Edif.

Torres Quevedo. María de Luna, 3, 50014 Zaragoza, Spain, angel.fernandez@unizar.es

long fiber glass or glass micro spheres so that same properties are affordable with lower densities.

Nanocomposites made of polyamide and Halloysite nanotubes are of interest because of the natural origin of these nanotubes. These nanotubes not only give good mechanical properties but also low price and sustainability. The Naturalnano Company [2] offers commercial grades of nanocomposites with PP, PA 6 or EPDM matrix. Ulrich’s [3] research described the influence of the molecular weight on thermal, mechanical and rheological properties concluding HNTs were usable and an alternative to Carbon nanotubes (CNT) based nanocomposites because their cost competitiveness.

The methodology to prepare PA/HNT nanocomposites requires the functionalization of nanotubes using amines or block copolymers to improve interfacial interactions with the matrix polymer as studied by Deliang [4]. Previously, the purity of HNT must be the maximum. Functionalized nanotubes have to be solved for different raw materials.

Optimization of this process implies the suppression of agglomerates and aggregates of nanotubes as their exfoliation to improve the maximum contact with polymer getting a real nanocomposite. Previous developments are not limited to polyamide matrix nanocomposites (e.g. PA 6). Also PP and EPDM matrix have been studied in detail by Ismail [5] and Paspakhsh [6]. The application of their results has led 183

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 183-192

to new and commercial products as Pleximer-PP and Nubrid respectively in addition to cited Pleximer-PA [2].

As all nanocomposites based on thermoplastic matrices are challenged to be made in the form of pellets by low cost preparation process like extrusioncompounding and for subsequent use by injection molding, this research used these two technologies.

The novel Polyamide based nanocomposites are well described by Hedicke-Hochstotter [7] as a ready-toprocess material. Obtaining nanocomposites readyto-inject through an extrusion-compounding machine in a one-step-process at industrial scale is extremely difficult with standard equipment. Previous research at lab. scale have demonstrated the need to optimize the extrusion process to disperse enough the nanotubes combined with recirculation process to increase the residence time in order to ensure the correct diffusion of the matrix. Only if the nanotubes are dispersed properly in the thermoplastic matrix, they will remit efficiently its properties to the resulting product.

The complexity of this process is such that it was decided to produce a masterbatch with high concentration of nanotubes with the extruder. So the object of the research was to use the masterbatch to prepare different nanocomposites reducing the weight content of nanotubes. Also, it was studied the feasibility of dispersing masterbatch directly in the injection machine. All extrusion and injection moulding equipments used were industrial-scaled size.

Physicochemical analysis has combined nano and micro techniques as X-ray dispersion (XRD), Fourier transform infrared spectrometry (FTIR), Scanning electron microscopy (SEM) and Transmission electron microscopy (TEM). XRD analysis is focused to previous studies where Baochun [8] and [9] discovered the effect of the halloysite in the crystallization of polyamide. Those revealed that with more interaction acceleration of crystallization and the increment of γ phase occurs. This γ phase appeared against the α phases usually present in PA 6. This research was demonstrated with mini-lab extruders.

Improvements in mechanical behaviour regarding the essential work of fracture (EWF) at low contents of halloysite nanotubes has been studied by Prashantha [10] with additional positive conclusions about the increase in storage modulus, young modulus, tensile strength and notched Charpy impact strength without loss of ductility at laboratory scale.

The fire behaviour of PA 6 + HNT composites was studied by Marney [11] with positive conclusions about the ease of composite preparation as an 184

attractive consideration for further development or study of the systems of flame inhibition of PA 6.

The main goal of this research is to study the effects on the distribution and dispersion of HNT in PA 6 prepared by extrusion and injection molding, as well as to measure the viscosity, stiffness and flame retardancy of the obtained nanocomposites Furthermore, extrusion and injection moulding have been applied with industry scaled equipment.

The conclusions of this research have answered the uncertainities of key factors affecting the design for manufacturing of newer parts made in nanocomposites where feasibility and performance are so close that they fit the newest product requirements.

1 EXPERIMENTAL 1.1 Materials Three materials were used: Raw PA 6 Badadur from BADA Hispanaplast, raw milled Halloysite nanotubes from Naturalnano, Inc, and Pleximer-PA, a masterbatch of PA 6 up to 30% weight content of Halloysite nanotubes also from Naturalnano.

1.2 Processing Equipment A Coperion ZSK 26 extrusion-compounding machine equipped with 29 mm diameter twin-screw and two gravimetric feeders was used for masterbatch preparation. This machine was used also to reduce the content of nanotubes in different blends.

A JSW 85 EL II electrical injection moulding machine was used to inject test specimens in a specific mold.

1.3 Testing Equipment A Gottfert 1500 capillary rheometer was used for rheometry and an Instrom 8032 test machine was used for tensile and flexural test.

For physicochemical analysis the XRD (Rygaku/ Max System) was used to study the crystalline structure of materials. The TGA (Mettler Toledo TGA/STDA 851 SF/1100) was used to measure the variation of properties with temperature in a range from 30 to 800 ºC using. Finally, the FTIR (Bruker Vertex 70 with ATR Golden Gate accessories) was used to analyze the chemical composition of final materials as well as the surface atomic composition was determined using XPS (Kratos Tech Axis Ultra DLD) using AlKα radiation at 15 kV and 10 mA.

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The morphologies of nanocomposites were observed on cryofractured surfaces using a S2300 (Hitachi) scanning electron microscope. SEM equipment works within a range from 0.5 to 30 kV. Specimens were covered with a thin gold layer.

A transmission electron microscope 2000-FXII (JEOL Ltd) with 200 kV acceleration voltage was used to study the dispersion of HNT inside the polyamide matrix. The TEM specimens were cut in ultrathin sections of approximately 100 nm with a microtome, collected on copper grids and covered with a carbon thin layer.

• •

•

1.4 Methodology The methodology included the following stages: • A masterbatch of raw PA 6 with a maximum weight content of 30% of Halloysite nanotubes was made by extrusion-compounding at barrel temperature of 250 ºC and screw speed of 450 rpm. After several trials the results were not as good as expected with several interruptions because power stucked the hopper wall blocking the inlet of the extruder. Due to the difficulties of reliability of the trial the alternative chosen option was to buy the masterbatch containing 30 %wt HNT (Pleximer-PA supplied by Naturalnano, INC.). The specifications of HNTs in the masterbatch were in a range of diameter 8 to 10 nm, mean length of nanotubes 1.5 μm, and a purity of 95%.

• Different blends of PA 6/HNTs were made reducing the masterbatch into Raw PA 6 by extrusion-compounding at a barrel temperature of 250 ºC and two different screw speeds of 450 and 900 rpm for a constant output of 20 kg/h. The weight content of HNTs was selected to 3 or 6% programming the gravimetric feeder.

• After pelletizing, the nanocomposite granules, the raw PA 6 and the masterbatch were tested in a capillary rheometer at temperatures of 240 to 260 ºC and a 1mm diameter capillar.

• For mechanical testing, nanocomposite granules, raw PA 6 and masterbatch were injection-molded into standart test specimen for tensile and flame retardancy test (Fig. 1). The processing conditions were; Barrel temperature profile 245, 245 and 255 ºC (optimized from the hopper to the nozzle); nozzle temperature 255 ºC; screw rotating speed 70 rpm (for a screw diameter of 36 mm); back pressure 4.5 MPa; mould temperature 70 ºC (hot water); filling time 1.2 s; injection speed linearly profiled from 20 to 40%; maximum injection

•

pressure (down 35 MPa for all tests); holding pressure 75 MPa; hold pressure time 8.5 s; cooling time 22 s and total cycle time 40 s.

For physicochemical characterization small pieces of granules and injected parts were analyzed by XRD, TGA, FTIR and XPS. For morphological characterization SEM and TEM microscopic images were taken on cryomicrotomic laminates extracted from the surfaces of the nanocomposite.

For mechanical characterization all compounded PA 6 based materials were moisture conditioned at 25 ºC room temperature and 50% ambient humidity. Flexural modulus calculation was made accordingly to ISO-178 for 3 point bending (thickness-span ratio of 1:16 and cross head displacement rate of 2 mm/min). The dimensions of bending specimens were 80×10×4 mm. Tensile modulus and strength calculation was made accordingly to ASTM D638-72 (crosshead rate of 20 mm/min). These parameters could be directed obtained from σ – ε curves.

Flame resistance evolution was made accordingly to Underwritters Laboratories (UL94-V0, V1 and V2) test for vertical combustion with Bunser flame.

Fig. 1. Injected parts (E-PA 6+30 %wt HNT, B-rawPA 6, C-PA 6+3 %wt HNT, D-PA 6+6 %wt HNT)

2 EXTRUSION-COMPOUNDING AND INJECTION MOULDING TESTS The twin-screw extrusion-compounder was used to prepare different blends of PA 6 and reinforced masterbatch. The pellets obtained were analyzed or injected afterwards. Selected cases are summarized in Table 1. Case R is raw PA 6 and Case M is a Masterbatch of PA 6 with 30% of HNT. Cases E1 to E4 are blends of different proportions of R and M pellets to achieve new nanocomposites reducing the content of HNT to 3% (E1 and E2) and 6% (E3 and E4). The variation of extrusion speed was also studied.

Cases E1 and E2 were performed at high extrusion speed (900 rpm) as E3 and E4 were made slower (450

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rpm). All cases have been characterized at micro and nano scale.

The injection moulding machine was used to inject test specimens using extrusion-compounded materials and others. All these test parts were injected under optimum processing conditions regarding the knowledge of the Technological Centre about injection of reinforced PA 6 and their own methodology to set up the process. Selected cases for analysis and discussion are RI, MI, E2I, E4I, DI1 and DI2. Cases RI, MI, E2I and E4I are injected samples of R, M, E2 and E4 materials. The influence of using the extrusion process to disperse the HNT through the polymer was studied applying Direct Injection (DIi) and comparing both results. DI1 and DI2 are injected samples of manual mixtures of R and M pellets. Their weight content of HNT is 3 and 6% respectively.

All these cases were repeated 10 times after enough neglected cycles waiting for a stabilization of process conditions in the case of Injection molding. All injected specimens have been characterized at macro, micro and nano scale.

Table 1. Extrusion-Compounded materials Case Nº Resultant composition R M E1 E2 E3 E4

PA 6 PA 6+30 %wt HNT PA 6+3 %wt HNT PA 6+3 %wt HNT PA 6+6 %wt HNT PA 6+6 %wt HNT

Extr. speed [rpm] 900 450 900 450

Fig. 2. TGA results for cases E2I and E4I R / M blend [%] 100 / 0 0 / 100 90 / 10 90 / 10 80 / 20 80 / 20

3 RESULTS AND DISCUSSION Different results have been obtained from the analysis of the materials and test parts. These results describe the characterization of the nanocomposites from different points of view: the rheological behaviour, the mechanical properties, the homogeneity of the nanocomposite and the interaction between matrix and filler and the stiffness of the product. Therefore, these results are analyzed at a different size scale depending on the nature of the test.

3.1 Physicochemical Characterization The results of Fourier transform infrared spectroscopy (FTIR), Thermogravimetric analysis (TGA) and X-ray diffraction (XRD) were obtained in order to discuss the chemical composition, degradation temperature and crystallographic structure of the nanocomposites after extrusion-compounding and injection molding.

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The measurement of degradation temperature with TGA analysis showed the differences between different nanocomposites were neglectable and the values were 438 ºC for all cases. This result was independent of HNTs weight content and processing conditions. Fig. 2 shows TGA results for E2I and E4I cases with a residual content similar to the halloysite weight content in the nanocomposite (3 to 6% respectively).

The measurement of chemical composition with FTIR showed clear differences between different samples depending on the weight content of nanotubes.

Fig. 3 shows FTIR results for materials M, E1 and E3.

The curves are quite similar except in the influence area of the characteristic links of the groups present in the nanotubes. A detailed graphic of this area is shown in Fig. 4. At the way number associated to the stretchment of the links of Si (Si-O-Si: 1088 cm-1, Si-O-Al: 1032 cm-1) the curves grow proportionally to the content of nanotubes as expected so it can be concluded that the local presence of dispersed nanotubes has been achieved. Other relevant links are Si-O-Si (open chain siloxanes) at 940 cm-1.

The comparison of material curves with similar content of HNT like, E1, E2 and E2I and DI1 have shown that they are almost equal, so in this case, the results are independent of the extrusion speed (450 or 900 rpm) and the method of preparation of the material (extrusion or extrusion plus injection).

The same conclusions can be drawn in comparisons between samples E3, E4, E4I and DI2.

The measurement of crystallographic structure with XRD has clarified the interaction between different phases and substances in the nanocomposites.

In fact, XRD has been the analytical technique that has detected differences between all the materials and their way of preparation.

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of γ phase crystal take up. The raw Halloysite curve exhibits a peak at 24.89º. Only a high content of Halloysite as case M has shown the same peak, so that the rest of samples do not show the Halloysite characteristic peak due to their low content on nanotubes.

Fig. 3. FTIR results for Cases M, E1 and E3

Fig. 5. XRD results for Cases M, E1, E2I, DI1, R and raw Halloysite

Fig. 4. Detailed FTIR for Cases M, E1 and E3

The results of all materials have been unified in Fig. 5. This graphic compares the WRD result for cases M, E1, E3I, DI1 and R combined also with the raw Halloysite nanotubes. In this graphic the dominant presence of α1 and α2 phases of raw PA 6 (R) can be easily dominant and, in the opposite site the dominant γ phase (metastable) of the masterbacth (M) due to interaction a high halloysite-polymer that influences the acceleration of crystallinity growth and increases the presence of this phase.

Fig. 6 presents a detail of all the above mentioned cases to clarify the discussion. Raw PA 6 (R) exhibits α1 and α2 phases at 20.50º and 22.82º to 23.44º. On the other hand, high concentration HNTs masterbatch (M) exhibits γ phase that appears at 21.26º. This result is as precise as expected in relation to previous research studies cited in the introduction. As HNT act as nucleating agent and accelerate the crystallization, the higher the HNT content is, the larger the presence

Fig. 6. XRD details for Cases M, E1, E2I, DI1, R and raw Halloysite

The analysis of curves E1, E2I and DI1 should be very similar as they all PA 6 + 3 %wt HNT with identical proportions of R and M polymers. However, their XRD curves show differences in polyamide phases distribution.

In the case of E1 material, the resultant curve shows the dominant phase is γ with smaller “shoulders” that reveal a lower presence of α phases.

The E1 material is made by extruding R and M materials at the highest twin-screw rotation speed of 900 rpm. It can therefore be concluded that a good dispersion of phase M has been dispersed in the whole

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resultant nanocomposite and the highest interaction between Halloysite nanotubes with the polymer has been achieved as this blend exhibits the maximum value of γ phase.

The E2I material curve shows peaks of γ and α2 phases to be almost equal. In this case, the material has followed two extrusion processes: The first in the compounding machine, and the second in the injection molding machine during the plastication stage. The injection molding process also substantially increased the residence time of the polymer at high temperature and it should have improved the correct dispersion of nanotubes if their separation and exfoliation were optimal. The main process difference between E1 and E2I is that during extrusion, E2I rotated at half speed than E1, so the friction was quite lower to result in a lack of nanotubes dispersion. This result proves the importance of extrusion speed as a mechanism of friction to ensure the good dispersion of additives including nano scaled reinforcements.

The worst result happened in the case the blend was mixed directly inside the injection machine hopper (DI1). In this case, the mixture of R and M was made manually and not mechanically. Then, α phases appeared roughly revealing a lack of interaction between polymer and reinforcement.

The results obtained in the case E2, E4I and DI2 which correspond to 6% weight content of HNT were similar to those presented for 3% wt HNT.

Fig. 7. Graphical comparison of measured viscosity and regression viscosity curves at 240 ºC

3.2 Rheological Characterisation The results of Capillar rheometry were obtained in order to discuss the rheological behaviour of the blends. The viscosity measurement and the viscous models generation are necessary for the discussion about the processability of the newer materials.

The measurement of apparent viscosity was made in a capillary rheometer for R, E2, E4 and M pellets.

Three temperatures were pre- selected for each set measure, each of them embracing a large range of shear rates and viscosities. After the measurement the Rabinowitsch correction was applied and viscosity and shear stress were determined. These trials were repeated five times for each temperature. Figs. 7 to 9 show the viscosity versus shear rate and the tendency curve for each material at 240, 250 and 260 ºC, respectively.

The results show that the viscosity dependency with shear rate of the nanocomposites in molten state is independent on the weight content of HNTs and remains similar to any thermoplastic. The analysis of the charts proves that the raw PA 6 (case R) is quite 188

Fig. 8. Comparison of viscosities at 250 ºC

Fig. 9. Comparison of viscosities at 260 ºC

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more viscous than the masterbatch (case M). For this reason, the blends of them (R/M) show intermediate values of viscosity, increasing as HNTs weight content is reduced. The reason is that rheological behaviour is highly influenced by the matrix which is a blend of different PA 6 grades.

The fitting of corrected viscosity data returns the coefficients for Carreau-WLF viscous model using the software Cadmould 3D-F SimFit v0.57. These are shown in Table 2. This model is necessary to perform injection moulding simulations with typical softwares like Moldflow, Cadmould 3D-F or Moldex and follows the following expression for the viscosity dependency on shear rate and temperature:

P1 ⋅ aT

η=

, (1)

(1 + γ ⋅ P2 ⋅ aT ) 8.86 (T0 − Ts ) 8.86 (T0 − Ts ) log10 ( aT ) = . (2) ´− 101.6 + (T0 − Ts ) 101.6 + (T0 − Ts ) P3

Table 2. Carreau-WLF model coefficients for R, M, E2 and E4 materials CASE R M E2 E4

P1 [Pa·s] 203.0 81.26 163.9 136.8

P2 [s] 1.11e-3 1.89e-4 1.19e-3 6.94e-4

P3 [-] 0.601 0.921 0.567 0.671

T0 [ºC] 250 250 250 250

TS [ºC] 46.63 126.9 105.0 97.21

The results shown in Figs. 7 to 9 have shown that the viscosity of nanocomposites does not increase dramatically with the content of nanotubes but it is within the order of magnitude of polymer matrix.

The final values of viscosity are correlated to the portion of each component in the final blend. These intermediate values of viscosity could be explained by a good mixture of both materials (R and M) as mechanical and heat treatment is applied in the extruder without additional interaction between the nanotubes and the new matrix material in molten state. Furthermore, good dispersion could be achieved with good interaction of both polyamides.

The results shown in Table 2, after rheological characterization by fitting the measured values, show the feasibility of injection molding process for most applications and nearly any kind of parts with these materials as the viscosity is not so high.

The successful obtainment of the coefficients of Carreu-WLF model for all materials ensured that injection molding simulation could be performed for these materials. In the development of newer parts in the future, the use of process simulation is a critical

stage for the final part development under feasibility criteria, mold design and manufacturing, and robust serial production.

3.3 Mechanical Characterization The results of the tensile test are shown in Table 3.

They represent the averaged values for 10 different specimens for each material. Both elastic modulus and maximum stress increase with HNT content as expected. However, yield stress and elongation at break increase with HNT content up to 3 to 6%, decreasing for higher nanotubes content.

Table 3. Results for mechanical properties of selected cases CASE R E2I E4I M

Elastic Modulus Yield Stress [MPa] [MPa] 2123 48.32 2542 49.80 3054 45.54 3070 36.60

Max. Stress [MPa] 58.00 63.24 69.15 61.39

Elong. at break [%] 45.03 43.75 45.06 16.48

The most important result is the loss of ductility for high HNTs concentration (30%). The most appropriate explanation for the tendency of the evolution of mechanical properties lies in the process, so the specimens were fabricated as explained below.

Cases R, E2I and E4I were previously extruded but masterbatch (M) was injected as it was supplied, without an additional extrusion process.

Fig. 10. Nanotubes aggregates

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3.4 Morphology Characterization The results of SEM microscopy clarified the tensile test results. Fig. 10 shows the tendency of Halloysite nanotubes to create aggregates of 1 μm minimum size. Analyzing the SEM images (Fig. 11) of the masterbatch (M) a lot of aggregates of nanotubes impossible to disperse in the matrix at nanometric scale can be seen. The low values for elongation at break for the masterbatch could be a result of these singularities in the material structure.

Aggregates were transferred to the molded part if there was no pre-extrusion process as shown in Fig. 11. In this case the direct mixing of M and R in the hopper did not prevent this problem.

Fig. 13. TEM image of E2I case

Fig. 11. SEM of M with HNT aggregates

Fig. 14. TEM image of E4I case

Fig. 12. SEM of D2I with HNT aggregates

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Fig. 15. Load-deflection results for bending test

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For this reason, previous extrusion process to injection molding improved the dispersion of nanotubes and resulted in better mechanical performance. Figs. 13 to 14 show the distribution of nanotubes in polymer matrix corresponding to cases E2I and E4I. Also, trials developed at higher rotation speeds of the extruder (E1I and E3I) resulted in better nanotubes dispersion.

3.5 Product Stifness In order to evaluate the mechanical performance of the materials R, E2I, E4I and M, the bending stiffness of the test part was calculated as the ratio load-displacement during bending test (Fig. 15). As expected, stiffness increased with the increase of HNTs content. For small deflections the increments of stiffness were of 9.1, 11.8 and 55.1% for 3, 6 and 30 %wt HNT. 3.6 Flame Resistance In order to evaluate the flame resistance R, E2I and E4I were subjected to UL94 test. All of them passed the 10 second-first application without burning the cotton or reaching the fastening clip. After 10 seconds the second application all of them burned the cotton but did not reach the fastening clip. The analysis of residual flame duration showed it decreased with nanotubes increasing content. Averaged times for 10 trials are 5.12, 2,65 and 2.48 s for R, E2I and E4I respectively. Also, the analysis of residual incandescency duration showed the same tendency with durations of 7.74, 7.55 and 5.68 s for R, E2I and E4I test parts. All nanocomposites were UL94 V2 classified presenting improved flame resistance with increasing weight content of HNTs. 4 CONCLUSIONS A full characterization of the rheological and mechanical behaviour of nanocomposites of polymeric matrix PA 6 based on halloysite nanotubes has been made. All results have been obtained with scaled industry equipment and ensure the feasibility of further development projects.

As the equipment used for injection molding was industry scaled to process large amounts of raw material, a lack of quality and homogeneity was detected. Against the generalized idea all masterbatches are ready-to-use in injection molding process, this research stresses the need to achieve

the best dispersion prior the application in injection molding process.

XRD, FTIR, SEM and TEM techniques used for physicochemical analysis have been shown to be the most useful techniques to study the proper dispersion of nanotubes in the matrix.

The physicochemical characterization revealed that the combination of extrusion-compunding process with injection moulding gave benefits in the mechanical properties of test specimens produced.

These benefits were due to the improved dispersion of nanotubes at high extrusion speeds. The direct injection of masterbatch or a low extrusion speed during mixing led to a lack of uniformity and improper dispersion of nanotubes. This conclusion has been endorsed also by the detection of the maximum values of polyamide γ phase for E1-E3 samples.

Procesability of nanocomposites was demonstrated also with a high content of nanotubes because the measured viscosity remained in the same order of magnitude of raw PA 6 material, so no further difficulties than processing raw materials appeared.

Flame resistance was slightly improved by the use of nanotubes remaining UL94 V2 class.

The product stiffness increased with nanotubes content in all cases. The relevance of this finding is that it measures the response of the part as a whole rather than locally, as the rest of the shown trials do.

All these conclusions have been extracted from the micro and macro developed tests.

5 REFERENCES [1] Lecouvet, B., Gutierrez, J.G., Sclavons, M., Bailly, C. (2011). Structure-property relationships in polyamide 12/halloysite nanotube nanocomposites. Polymer degradation and stability A, vol. 96, no. 2, p. 226-235.

[2] Naturalnano, Inc., from http://www. nanturalnano.com accessed on 2012-03-04.

[3] Ulrich, A., Hedicke, K. (2010). Composites of polyamide 6 and silicate nanotubes of the mineral halloysite: Influence of molecular weight on properties.

Polymer, vol. 51, p. 2690-2699, DOI:10.1016/j. polymer.2010.04.041.

[4] Deliang, L., Quing, L. (2009). Correlation between interfacial interactions and mechanical properties of PA-6 doped with surface-capped nano-silica. Applied Surface Science, vol. 255, p- 2871-2877.

[5] Ismail, H., Pasbakhsh, P. (2008). Morphological thermal and tensile properties of halloysite nanotubes filled EPDM nanocomposites. Polymer Testing, vol. 27, p.

841-850, DOI:10.1016/j.polymertesting.2008.06.007.

[6] Pasbakhsh, P. (2010). EPDM/modified halloysite nanocomposites. Applied Clay Science, vol. 48, no. 3, p.

405-413, DOI:10.1016/j.clay.2010.01.015.

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[7] Hedicke-Hochstotter, K., Teck Lim, G. (2008). Novel polyamide nanocomposites based on silicate nanotubes of the mineral halloysite. Composites Science and Technology, vol. 69, no. 3-4, p. 330-334, DOI:10.1016/j. compscitech.2008.10.011.

[8] Baochun, G., Quanliang. Z. (2009). Crystallization behavior of polyamide 6/ halloysite nanotubes nanocomposites. Thermochimica Acta A, vol. 484, no.

1-2, p. 48-56, DOI:10.1016/j.tca.2008.12.003.

[9] Baochun, G., Quanliang, Z., Yanda, L. (2009). Structure and performance of polyamide 6 / halloysite nanotubes

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nanocm. Polymer Journal A, vol. 41, no. 10, p. 835842, DOI:10.1295/polymj.PJ2009110.

[10] Prashantha, K., Schmitt, H., Lacrampe, M.F. (2011). Mechanical behaviour and essential work of fracture of halloysite nanotubes filled polyamide 6 nanocomposites. Composites Science and Technology A, vol. 71, no. 16, p. 1859-1866.

[11] Marney, D.C., Russell, L.J., Wu, D.Y. (2008). The suitability of halloysite nanotubes as a fire retardant for PA 6. Polymer degradation and stability, vol. 93, no. 10, p. 1971-1978, DOI:10.1016/j. polymdegradstab.2008.06.018.

Fernández, A. – Muniesa, M. – González, J.

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, 193-200 © 2013 Journal of Mechanical Engineering. All rights reserved.

DOI:10.5545/sv-jme.2012.661

Received for review: 2012-06-15 Received revised form: 2012-10-04 Accepted for publication: 2012-10-15

Resource Efficient Injection Moulding with Low Environmental Impacts

Gantar, G. – Glojek, A. – Mori, M. – Nardin, B. – Sekavčnik, M. Gašper Gantar1,* – Andrej Glojek2 – Mitja Mori3 – Blaž Nardin4 – Mihael Sekavčnik3 Slovenia Slovenian Tool and Die Development Centre, Slovenia 3 University of Ljubljana, Faculty of Mechanical Engineering, Slovenia 4 Gorenje toolmaking, Slovenia 2 TECOS,

1 ENVITA d.o.o.,

The selection of most appropriate design and technological solutions to produce certain mould should capture technical performance, economical issues as well as environmental impacts occurred during the mould life cycle. In the paper an approach is presented to support the selection of alternative mould design solutions in the early design stage. It includes the use of Life Cycle Assessment methodology, Life Cycle Cost methodology and is supported by numerical simulations. The approach is applied to a case study where three mould designs for production of the same plastic product were compared. Finally, general conclusions regarding the resource efficient injection moulding processes are presented.

Keywords: injection moulding, mould, optimisation, life cycle assessment, life cycle cost analysis

0 INTRODUCTION Injection moulding ranks as one of the most widely used processes for producing plastics products. As in other forming processes, the characteristics of injection moulding procedures and products are significantly affected by the quality of moulds used i.e. tools that are mounted into injection moulding machine to produce repeatable products [1].

Once produced, the moulds are used in production for many years. Since raw materials are becoming scarcer and more expensive, and the costs of energy is also increasing the strategy of mould design should not aim only at cost reduction but also at reducing resource consumption and emissions throughout its entire life cycle.

This paper presents an approach to compare and optimize mould design and production process parameters from technical, economic and also environmental point of view. The results are important for the selection of mould design.

A comparison of life cycle performance of different mould designs is presented on the case study.

The environmental impact was quantified by the Life Cycle Assessment (LCA) and the economical by the Life Cycle Cost (LCC) analysis, [2] to [4]. The approach was supported by numerical simulations for the prediction of relevant technical information in the early stage of mould design.

The paper proceeds as follows: Section 1 presents an overview of the case study investigated in the research. Section 2 provides the life cycle model of the mould together with technical background on the injection moulding technology. Section 3 describes

the LCA method and compares environmental impacts arising from different mould designs. Section 4 describes LCC analysis for the selection of the best mould design from the economical point of view.

Section 5 offers discussion and conclusions.

1 OVERVIEW OF CASE STUDY The case study is presented with three alternative mould designs put side by side. They enable an injection of the same plastic product, shown in Fig. 1, but differ in technical solutions, which contribute to the productivity and resource efficiency during their use phase (injection moulding of plastic products). It is expected that one million products will be produced throughout the mould’s life cycle.

Fig. 1. Studied plastic products produced by mould in the use phase (Material: PP 40% GF, d = 160 mm; m = 137 g; Quantity = 1 mio parts)

Three mould designs, which were considered as appropriate by the mould designer, are presented and described in Fig. 2.

*Corr. Author’s Address: ENVITA d.o.o., Tržaška cesta 132, 1000 Ljubljana, Slovenia, gasper.gantar@envita.si

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2 LIFE CYCLE OF A MOULD Life cycle of a mould is presented in Fig 3. Two main stages i.e. mould manufacturing and the use of the mould (injection moulding), are described in details in the following sub- chapters.

It is assumed in our case study that mould components will not be reused in similar moulds in the future; although in general some mould components i.e. guides can be reused.

2.1 Mould Production For smaller moulds, such as the one investigated in the case study, the main standard metal elements are sawed to rough shape from rolled slabs and bars which are then milled, turned and grinded to final tolerance geometry.

Mould components like plates, injectors, ejectors, guiding elements and others are commercially available as standard elements.

The patterns which form the final product geometry within standard plates (sometimes also called active mould surfaces) are usually produced

by CNC milling and/or Electric Discharge Machining (EDM). In our case only the CNC milling process was used. The heat treatment was used to improve the quality of active mould parts.

CNC milling of tool steels requires cutting tools that are strong, precise and expensive, involving special steels and ceramics that are usually coated to improve wear characteristics and it also requires other auxiliary materials; such as cutting fluids.

The main design parameters of the mould which influence properties of the production process and the

Fig. 2. Alternative mould designs

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Fig. 3. LCA model of a mould life cycle

product are the injection system (runners), and the cooling system.

2.2 Injection Moulding Process The injection moulding process begins by feeding the plastic material and appropriate additives from the hopper to the heating/injection cylinder of the injection moulding machine.

The process consumes a great amount of electric energy and generates environmental emissions. Various kinds of chemicals are used both in the protection of moulds and during the injection moulding process. These chemicals are: lubricants, cleaning agents, mould release agents, mould protecting agents and degreasers. Most of the chemicals are packaged in high-pressure containers and applied as atomized spray. The ingredients are mainly petroleum or hydrocarbon by-products (i.e. paraffinic oil, propane and isobutane). If the training of workers is appropriate, the emissions from these chemicals are not substantial.

Minor amounts of particulates come from the polymer feed and mixing in the hopper as well as trimming of the products and grinding of runners and other scrap. Minor amounts of volatile emissions are released from vents in the heating/injection cylinders and the moulds. These emissions consist of traces of monomers, additives and decomposition products of polymers. Decomposition products occur when the processing temperature reaches the temperature where the polymer begins to decompose, which rarely

happens in serial production of polypropylene (PP) products.

The main injection moulding parameters that influence the environmental impact of the production process are: cycle time, mould temperature, clamping force, profile of packing pressure and properties of machine.

3 LIFE CYCLE ASSESMENT (LCA) The most suitable method for evaluating the environmental impact of the product in its entire life cycle is the LCA. The LCA means compiling and evaluation of inputs and outputs and potential environmental impacts of a product system during its lifetime.

The LCA method consists of four stages: definition of the goal and scope, construction of the product’s life cycle model with all mass and energy flows; the so called “Life Cycle Inventory Analysis”, evaluation of the environmental relevance of all flows; the so called “Life Cycle Impact Assessment” and finally, the interpretation of the results.

3.1 Goal and Scope Definition The goal was the assessment of the environmental impact of three alternative mould designs, previously defined and described in Fig. 2. The moulds have different design solutions and are made from different materials and they also differ in their characteristics during the use phase. Therefore, different

Resource Efficient Injection Moulding with Low Environmental Impacts

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environmental impacts are expected throughout their life cycles.

3.1.1 Functional Unit

•

A functional unit is the production of one million plastic products, presented in Fig. 1.

•

3.1.2 Study Boundaries and Simplifications

•

This study includes cradle-to-grave environmental impacts of producing and using an injection mould for production of one million plastic parts, presented in Fig. 1. The system boundaries include: the cradle-togate material and environmental impacts of the energy used, the production of moulds, the use of moulds and disposal of moulds at the end of life cycle. The models include impacts associated with the upstream production of all materials and energy used and endof-life treatment for all materials.

Since the aim of the study is a comparison of mould designs, the use of plastic products was considered out of scope. This means that the flow of plastic material that enters the system and leaves it in the form of a product is not included into the model.

The plastic material of rejected parts and runners are taken into consideration since for the production of the studied products presented in Fig. 1 it was not allowed to grind and reuse the rejected parts and runners. The human labour was also not included into the study.

The retail, internal and external transport are excluded from the study (transport distances of moulds are minimal and transport distances for plastic materials are not known). The production of infrastructure (buildings, machinery, etc.) and maintenance of infrastructure (e.g. replacement of hydraulic oil in machines) are also excluded from the study.

In the first analysis it was assumed that the runners and rejected parts were incinerated. The mould materials were recycled at the end of mould life.

Due to the absence of more accurate data, the following simplifications were also used: • For the materials of moulds and cutting tools, used for CNC machining, we took from the LCA database in GaBi [4] the material with most similar chemical structure to that of the actual materials.

• For all standard elements used in mould production (base plates, injectors, ejectors, guiding elements and other available standard elements) only the environmental impact from

•

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the acquisition of materials was considered but the environmental impact caused by production of standard mould components was neglected.

The environmental emissions associated with the production and uses of auxiliary materials mentioned in Chapter 2.2 were neglected as well.

Cutting fluids were modelled as a mixture of lubricant and process water.

BeCu material is modelled as a combination of 50% Copper, 15% Aluminium and 5% Steel.

It is estimated that the environmental impact of the maintenance level of all mould design solutions is the same and low (mainly manual work is included, which is excluded from our study), therefore it was neglected.

Table 1. Major resource and energy consumption over the mould life cycle Mould 1 Mould 2 Materials for mould Steel [kg] 320 340 BeCu [kg] 3 0 Mould production Design [hours] 120 120 Process planning [hour] 75 80 CNC milling Setup [hour] 2 2 Machining [hour] 75 80 CNC drilling Setup [hour] 2 2 Machining [hour] 36 38 Components fitting and polishing 118 120 – mainly manual labour [hour] Assembly – mainly manual labour 28 30 [hour] Final testing at injection moulding 8 8 machine [hour] Compresses air consumption 0.5 0.5 [Nm3] Energy consumption [kWh] 2900 3000 Injection moulding Injection cooling time [s] 39 34 Injection cycle time [s] 58 54 Clamping force [103 kg] 90 90 Plastic material consumption [g] 137.4 143.9 Expected reject rate [%] 0.5 0.4 Energy consumption for injection 0,138 0,134 moulding [kWh/part] Energy consumption for grinding 0.025 0.025 of runners and reject [kWh/kg]* *If runners and rejected parts are grinded and reused.

Mould 3 320 15 120 78 2 78 2 36 115 28 8 0.5 2900 35,5 54,5 90 137.0 2 0,137 0.025

3.2 Inventory Analysis The production system is the same for all the three studied mould alternatives – Engel injection moulding machine with 900 kN clamping force and peripheral

Gantar, G. – Glojek, A. – Mori, M. – Nardin, B. – Sekavčnik, M.

onal

ss [kg]

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devices. The major resources and energy consumption throughout the mould life cycle are presented in the Table 1.

3.2.1 Data Collection The consumption of material and time for mould production was estimated by industrial experts based on CAD mouldsâ&#x20AC;&#x2122; models.

The injection cooling times and the amount of plastic material consumed (for part and runners) were predicted by numerical simulations (NS). NS is the

most reliable method for predicting relevant technical information of the studied solutions in the early design phase. In our study the full 3D analyses (flow, cool and wrap) with MoldFlow software were used for prediction of output production process parameters based on different input mould design solutions [5].

The total cycle times were estimated based on measurements performed in the industrial environment and suggestions from literature [6].

Consumption of electrical energy during injection mould production phase and the use of the mould were estimated by measurements performed in the

Cutting tools 3 kg Compressed air

0 kg

Process water 1.9 kg

SI: Power grid mix

Lubricant

0 kg

Tool steel 99 kg

0.1 kg

EU-27: Landfill of ferro metals ELCD/PE-GaBi <p-agg>

Mould 1 production <u-so> 15 kg

Steel plate

Landfill

235 kg 3 kg Copper 2.4 kg

Waste water treatment

Aluminum

2 kg

0.459 kg

Injection moulding

Stainless steel 1 kg

Glass fibres 435 kg

323 kg

X 323 kg

Plastic material compaunding

Landfill of ferro metals

Waste incineration of plastics

Waste treatment 652 kg

Polypropylene granulate (PP)

Waste incineration of glass/inert material

1090 kg kg 1.09E003

1090 kg kg 1.09E003

652 kg

435 kg

0.05 kg

0 kg SI: Power grid mix

0.05 kg Process water

Waste water treatment

0 kg SI: Power grid mix

Fig. 4. LCA model from GaBi software Resource Efficient Injection Moulding with Low Environmental Impacts

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industrial environment and suggestions from literature [6]. The potential of different moulds designs to reduce energy consumption during use phase was estimated on the basis of cooling time calculation by FEM and data from literature regarding energy consumption during injection moulding [7] stating that about 50% of energy consumption in a typical injection moulding cycle is required for plasticizing and the rest can be reduced by shortening of the cycle time.

Consumption of compressed air was evaluated indirectly according to measurements of consumption of the entire production facility.

of the mould and its end of life only has a minor environmental impact. It is shown that the design solution 1 is the most effective from the environmental point of view.

3.3 Impact Assessment The LCA model was created using the GaBi software system for life cycle engineering, developed by PE International GmbH [8]. The model is presented in Fig. 4 The impact assessment of all inputs and outputs was performed using the Eco-indicator 99 (EI’99), which consist of several environmental impact categories, aggregating all emissions and resources consumption into three areas: Human Health (HH), Ecosystem Quality (EQ) and Resources (R).

Afterwards, the methodology weights the scores into a single value. The weighting coefficients were applied according to the hierarchic/average perspective (H/A).

The EI’99 was also used by other authors studying the environmental impact of moulds, therefore it was selected owing to a simple comparison of results [9].

3.4 Interpretation of Results Table 2. Environmental impact of the alternative mould design solutions (EI’99 points) Mould 1 Mould 2 Mould 3 Environmental impact 4556 6031 4972 (EI’99 points) Mould end of life 4 4 4 Mould material and production 108 110 115 Plastic material production*,** 241 1637 608 Injection moulding process 4203 4280 4245 * If runners and rejected parts are not granulated and reused. **Only the material that is disposed in the form of runners and rejected parts was included into the calculation (material included into products that leave the production facility was not included for greater transparency).

The EI’99 score obtained for the production of 1 mio parts (Fig. 5, Table 2) shows that the major environmental impact is caused during the injection moulding mainly due to energy consumption of the injection moulding machine. The production 198

Fig. 5. Environmental impact of the alternative mould design solutions (EI’99 points)

4 LIFE CYCLE COST The LCC generally refers to all the foreseen costs associated with the product throughout its life from “cradle to grave”. The LCC is an estimate of total costs from inception to disposal for both equipment and projects related to the products life cycle. The objective of LCC analysis is to choose an alternative where the lowest possible long term costs of ownership are achieved, [2] and [4].

As with the calculations of the environmental impacts, technological characteristics of the mould considerably influence the use-phase (maintenance level and characteristics of injection moulding process such as cycle time, amount of material wasted, reject rate, closing pressure, required temperature of plastic material, etc.).

The data about all the relevant production steps for mould production, the material consumed, the labour involved, equipment used, cutting tools used and other consumables enable the estimation of the costs of the mould. The comparison of mould designs should also include the subsequent mould life phases like injection moulding together with mould maintenance and mould end of life scenario (see Fig. 3). The most relevant inputs required to perform the LCC analysis are presented in Tables 1 and 3.

Some simplifications were assumed. The operator’s and engineer’s wages, energy costs, consumables and material costs, annual available production time of equipment are kept constant for all mould design solutions. The calculation was done with the assumption that the runners are not reused.

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Table 3. Cost input values Mould 1 Mould 2 Mould 3 Costs of standard components [€] 5000 4500 5000 Costs of disposable cutting tools [€] 700 600 700 Costs of other consumables [€] 100 100 100 Costs of items for all mould designs on functional unit / Cost general data Materials* Tool steel [€/kg] 7.8 BeCu [€/kg] 11.5 Employees Engineer wage [€/hour] 20 Operator wage [€/hour] 12 Machines – process costs**,*** HSM – CNC milling machine [€/hour] 40 CNC drilling machine [€/hour] 22 Injection moulding machine [€/hour] 17 Electricity [€/kWh] 0.1 Injection moulding general data Plastic material [€/kg] 1.7 Setup time of injection moulding machine [hour] 3 Batch size during injection moulding [part] 16000 Maintenance of moulds**** Basic cleaning [€/year] 150 Adjustment of the splitting line [€/year] 1070 Adjustment of ejectors [€/year] 600 Lubrication of guiding system [€/year] 60 Polishing [€/year] 120 Hot runner repair [€/year] 540 Average cost of materials for non-standard components only. * ** The input values are average values gathered from practice. *** Process costs include machine and labour (operator of the

machine is included). **** The mould will be in use for 3.2 years.

additional processing costs; the design solution 2 is also promising.

5 RESULTS AND CONCLUSIONS This paper presents an approach to compare and optimize mould design and production process parameters from technical, economic and also environmental point of view. The results are important for the choice of mould design and also for further optimisation.

5.1 Comparison and Ranking of Mould Designs The mould designs can be compared if the most important results from the previous sections of the paper are presented in one table (see Table 5).

Table 5. Comparison of mould designs Costs of the mould [€] LCC [€] Environmental impact (EI’99 points)

Mould 1 2nd 18445 3rd 570427 1st 4556

Mould 2 1st 18369 1st 558742 3rd 6031

Mould 3 3rd 18727 2nd 560108 2nd 4972

The costs of mould design 2 are the lowest, and also the total costs during entire life cycle are the lowest for mould design 2. On the contrary the mould design 1, which is not outstanding regarding costs, enables production with the lowest environmental impact.

5.1.1 Optimisation of Mould and Production Process

Table 4. LCC results Total costs [€] Costs of the mould [€] Plastic material production [€] Injection moulding [€] Maintenance costs [€] End of life costs [€] Costs per part [€]

Mould 1 570427 18446 234748 306015 11218 0 0.570

Mould 2 558742 18369 245438 285448 9487 0 0.559

Mould 3 560108 18728 237558 292604 11218 0 0.560

LCC was used to compare economic viability of alternative mould design solutions. The results are presented in Table 4. It is shown that design solutions 2 and 3 are more viable from the economical point of view. Taking into account that the runners and rejected parts can be ground and reused with minimal

The efficient optimisation of the process from the environmental aspect can also be performed based on results from Table 2. As we have seen from the analysis, the two main causes of the environmental burden during the production of plastic products are: • electric energy consumed during injection moulding process; • environmental impact caused by production and disposal of plastic material that finishes in the form of waste runners or rejected products.

Since the available machinery for injection moulding in the company was limited, the replacement of the hydraulic Engel injection moulding machine with an energy efficient one was not possible.

For the production of the studied products presented in Fig. 1 it was not allowed to grind and reuse the rejected parts and runners. But the

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company found the possibility of using the rejected material PP+40%GF for another product with lower requirements (inner plastic parts will low load carrying requirements). The company introduced the grinder into the production process. Additional electric energy is used for grinding the runners and the reject parts, but the benefit is the reduced amount of material needed for production of other products. It was calculated with the modified LCA models that by using this solution in combination with mould design 2 (which is most cost effective), the company could achieve only a minimal reduction of costs (total costs per part are reduced by 1.25%) but a significant reduction of the environmental impact (the total environmental impact can be reduced by approximately 20%; the exact value depends on further reuse of rejected material and allocation procedure used in LCA calculations).

On the contrary changes in mould design that do not affect its performance during injection moulding or improvement of the end of life strategy cannot reduce the total environmental burden significantly.

In general, the following conclusions can be derived from the study: • the potential of the mould design to influence the environmental impact of the injection moulding phase is high, • in mass production, better production process usually consumes fewer resources (energy and material), provides • the presented methodology manufacturing companies with feasible means to assess their environmental performance.

In the paper, it is assumed the injected parts produced comply with the confirmed quality in all three mould solutions. In case where technical performance (reliability of technology, time-toplastic parts, number of production steps, etc.) would be considerably different, it is reasonable to include technical performance into the decision-making system.

Finally, it should be noted that the presented approach can be incorporated into the standard service of the advanced mould manufacturing companies.

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6 ACKNOWLEDGEMENT To company Daplast for supplying the relevant industrial data. The presented research work was cofinanced by European Social Fund within the project No. JR-2-2009 (150/2010) at national agency JAPTI. Their contribution is greatly acknowledged.

7 REFERENCES [1] Tang, S.H., Kong, Y.M., Sapuan, S.M., Samin, R., Sulaiman, S. (2006). Design and thermal analysis of plastic injection mould. Journal of Materials Processing Technology, vol. 171, no. 2, p. 259-267, DOI:10.1016/j.jmatprotec.2005.06.075.

[2] Horne, R., Grant, T., Verghese, K. (2009). Life Cycle Assessment: Principles, Practice and Prospects, CSIRO Publishing, Collingwood.

[3] Folgado, R., Pecas, P., Henriques, E. (2010). Life cycle cost for technology selection: A Case study in the manufacturing of injection moulds. International Journal of Production Economics, vol. 128, p. 368-378, DOI:10.1016/j.ijpe.2010.07.036.

[4] Dhillon, B.S. (2010). Life Cycle Costing for Engineers, 1st ed., CRC Press, Taylor and Francis Group, Boca Raton.

[5] Autodesk Simulation Moldflow Insight (2012).

Autodesk Simulation Moldflow plastic injection molding simulation software, Autodesk, San Rafael.

[6] Lucchetta, G., Bariani, P.F. (2010). Sustainable design of injection moulded parts by material intensity reduction. CIRP Annals - Manufacturing Technology, vol. 59 , p. 33-36, DOI:10.1016/j.cirp.2010.03.092.

[7] Morrow, W.R., Qi, H., Kim, I., Mazumder, J., Skerlos, S.J. (2007). Environmental aspect of laserbased and conventional tool and die manufacturing.

Journal of Cleaner Production, vol. 15, p. 923-943, DOI:10.1016/j.jclepro.2005.11.030.

[8] PE International: Software and Database for Life Cycle Engineering (2010). Leinfelden-Echterdingen.

[9] Pecas, P., Ribeiro, I., Folgado, R., Henriques, E. (2009). A life cycle engineering model for technology selection: a case study on plastic injection moulds for low production volume. Journal of Cleaner Production, vol. 17, p. 846-856, DOI:10.1016/j. jclepro.2009.01.001.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3 Vsebina

Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 59, (2013), številka 3 Ljubljana, marec 2013 ISSN 0039-2480 Izhaja mesečno

Gostujoči uvodnik Razširjeni povzetki člankov Matej Hudovernik, Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, A. Erman Tekkaya, Karl Kuzman, Janez Marko Slabe: 3D-numerična analiza 2D-krivljenja profilov z metodo prostorske superpozicije navora Sebastijan Jurendić, Silvia Gaiani: Numerična simulacija preoblikovanja v hladnem pločevin iz α-titanovih zlitin Andreas Schubert, Henning Zeidler, Matthias Hackert-Oschätzchen, Jörg Schneider, Martin Hahn: Izboljšanje postopka mikroelektroerozijske obdelave z ultrazvočnimi vibracijami in pristopi k obdelavi neprevodne keramike Tomasz Tański: Lastnosti trdih prevlek na magnezijevih zlitinah AZ61 Leszek Adam Dobrzański, Małgorzata Musztyfaga, Aleksandra Drygała: Konvencionalni in nekonvencionalni postopki izdelave elektrod na silicijevih sončnih celicah Angel Fernández, Manuel Muniesa, Jaime González: Karakterizacija in priprava za brizganje PA 6, ojačenega s halojzitnimi nanocevkami (HNT) Gašper Gantar, Andrej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik: Učinkovita raba virov in majhen vpliv na okolje pri brizganju plastike Osebne vesti Doktorske disertacije, magistrska dela 2. stopnje, diplomske naloge

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3 Gostujoči uvodnik

Gostujoči uvodnik Tematska številka: Industrijska orodja in izdelovalne tehnologije Pričujoča tematska številka Strojniškega vestnika - Journal of Mechanical Engineering je nastala v povezavi z uspešno 8. mednarodno konferenco ICIT & MPT 2011 (International Conference on Industrial Tools and Material Processing Technologies), ki se je odvijala med 2. in 5. oktobrom 2011 v Ljubljani. TECOS – Razvojni center orodjarstva Slovenije organizira konferenco ICIT&MPT na vsake dve leti že od leta 1997. V tem času se je konferenca razvila v mednarodno prepoznavno znanstveno konferenco, ki vsakokrat privabi številne znanstvenike, raziskovalce in inženirje iz celega sveta in jim ponudi odlično okolje za strokovne diskusije, izmenjavo teoretičnih in praktičnih znanj ter izmenjavo informacij o najnovejših svetovnih trendih na področju industrijskih orodij in izdelovalnih tehnologij.

Na 8. ICIT & MPT konferenci so bili obravnavani svetovni trendi na področju preoblikovanja, materialov, laserskih obdelav, simulacije in optimizacije procesov, površinskih prevlek, nekonvencionalnih obdelovalnih postopkov, inteligentnih sistemov in managementa tehnologij. Vsi konferenčni članki, ki obravnavajo zgoraj navedene tematike, so zbrani in objavljeni v zborniku konference, ki obsega okoli 400 strani in predstavlja uporaben vir teoretičnih in praktičnih znanj ter informacij, ki jih znanstveniki, raziskovalci in inženirji lahko koristno uporabijo pri svojem strokovnem delu.

Predsedujoči mednarodnega znanstvenega odbora 8. ICIT & MPT konference so skrbno pregledali vse konferenčne članke in izbrali sedem izmed njih ter povabili njihove avtorje, da pripravijo razširjene ter ustrezno spremenjene in predelane članke za to tematsko številko. Vsi članki so bili nato recenzirani v skladu z običajnimi procedurami in standardi, ki veljajo pri reviji.

Izbrani članki objavljeni v tej tematski številki obravnavajo: numerične simulacije prihajajoče fleksibilne tehnologije 3D krivljenja profilov z uporabo TSS (Torque Superposed Spatial) metode, numerične simulacije hladnega preoblikovanja pločevine iz α – titanove zlitine, mikro – EDM postopek s pomočjo ultrazvočne vibracije – ultrazvok z visoko intenziteto neposredno apliciran na obdelovanec in posredno apliciran preko dielektrika – za obdelavo kovinskih materialov kot tudi za obdelavo električno neprevodnih keramičnih materialov, vrednotenje trdih prevlek na AZ61 magnezijevih zlitinah, primerjavo konvencionalne in nekonvencionalne metode, t.j. sitotiska in selektivnega laserskega sintranja, z namenom izboljšanja kakovosti izdelave elektrod silicijevih sončnih celic, vrednotenje in injekcijsko brizganje PA 6 ojačanega z HNT (Halloysite Nano Tubes) in uporabo LCA (Life Cycle Assessment) in LCC (Life Cycle Cost) metodologije za primerjavo in optimizacijo konstrukcij orodij za injekcijsko brizganje in parametrov injekcijskega brizganja z ekonomskega in okoljskega vidika kot nadgradnjo običajnemu predvsem tehničnemu vidiku.

Zahvaljujemo se vsem avtorjem za njihove cenjene prispevke, v katerih so predstavili njihove raziskovalne dosežke in uspešne industrijske aplikacije ter za njihove napore vložene v okviru recenzijskega postopka.

Zahvaljujemo se tudi vsem recenzentom za njihov prispevek, prof. dr. Vincencu Butali, glavnemu uredniku za dano možnost priprave te tematske številke in ga. Piki Škraba, tehnični urednici, ki je skrbela za koordinacijo med recenzenti in avtorji ter pripravila članke za objavo.

Gostujoči uredniki: Doc. dr. Janez Marko Slabe Prof. dr. Janez Grum Prof. dr. Karl Kuzman

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Prejeto v recenzijo: 2012-03-16 Prejeto popravljeno: 2012-05-11 Odobreno za objavo: 2012-05-16

3D-numerična analiza 2D-krivljenja profilov z metodo prostorske superpozicije navora Hudovernik, M. – Staupendahl, D. – Gharbi, M. – Hermes, M. – Tekkaya, A.E. – Kuzman, K. – Slabe, J.M. Matej Hudovernik1,* – Daniel Staupendahl2 – Mohammad Gharbi2 – Matthias Hermes2 – A. Erman Tekkaya2 – Karl Kuzman3 – Janez Marko Slabe1 1Tecos, Razvojni center orodjarstva Slovenije, Slovenija 2Institut za preoblikovalne tehnologije in lahke konstrukcije (IUL), Tehniška Univerza Dortmund, Nemčija 3Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija

Področje lahkih konstrukcij je močno usmerjeno k razvoju novih materialov visokih trdnosti. Le-ti v domeni obstoječih tehnologij prinašajo številne tehnološke, procesne in ekonomske ovire. Tehnološke rešitve za preoblikovanje v hladnem morajo torej ažurno slediti trendom uporabe omenjenih materialov.

Ekonomska učinkovitost lahkih konstrukcij je odvisno od mnogih dejavnikov. Izstopajo standardi in direktive glede zmanjšanja porabe energije, materiala in skupne teže končnih produktov. Prostorsko oblikovane komponente iz cevi in profilov različnih prerezov, ki imajo na omenjenem področju pomembno vlogo, morajo torej izpolnjevati visoke zahteve in zagotavljati kakovost. Kakovostna izdelava tovrstnih produktov pa je močno odvisna od stabilnosti, robustnosti in prilagodljivosti preoblikovalnih procesov. Obstoječe tehnološke rešitve za preoblikovanje cevi in profilov v domeni 3D-krivljenja nakazujejo velik potencial pri povečanju ekonomske učinkovitosti prostorskih konstrukcij, saj omogočajo optimalno razporeditev materiala, zmanjšanje števila sestavnih delov in montažnih postopkov ter zmanjšanje skupne teže.

Metodo krivljenja, ki temelji na osnovi prostorske superpozicije navora, angl. Torque Superposed Spatial – TSS, razvijajo na institutu za preoblikovalne tehnologije in lahke konstrukcije IUL na Tehniški Univerzi v Dortmundu. Je inovativna, robustna in fleksibilna tehnološka rešitev, s katero je možno izdelovati ravninske in prostorske komponente iz dolgih profilov, s simetričnimi ali nesimetričnimi geometrijami prečnih prerezov, in iz materialov visoke trdnosti. Izdelovati jih je mogoče v enem kontinuiranem koraku brez uporabe posebnih dragih orodij. Metoda TSS temelji na kontinuiranih dinamičnih karakteristikah, kjer je podajanje profila zagotovljeno s konstantno ali spremenljivo hitrostjo, proces krivljenja pa se izvaja z nenehnim podajanjem upogibnega elementa v smeri pravokotno na os podajanja. S spreminjanjem naklona ravnine, kjer je nameščen šestvaljni mehanizem za podajanje profila, ob sočasnem krivljenju oblikujemo 3D-konturo profila. Dosedanje študije metode TSS za 2D- in 3D-krivljenje profilov so bile izvedene predvsem v domeni eksperimentalnih in analitičnih metod. Procesni parametri in preoblikovalni mehanizmi med samim procesom krivljenja s spremenljivim polmerom pa do sedaj še niso bili numerično preučeni. Numerične simulacije, ki so predstavljene v tem članku, imajo torej pomembno vlogo pri razumevanju preoblikovalnih mehanizmov metode krivljenja TSS. Glavni namen predstavljenega dela je izdelava dovršenega numeričnega modela za postopek 2D-krivljenja kvadratnega profila z metodo TSS z uporabo implicitnega programskega orodja Abaqus/standard, ter vrednotenje rezultatov izbranih parametrov, kot sta upogibna sila in upogibni moment, z eksperimentalnimi rezultati pridobljenimi na IUL v Dortmundu. Na podlagi potrjenega numeričnega modela so izvedene analize napetostnih in deformacijskih stanj v kritičnih področjih med in po procesu krivljenja. V diskusiji je obravnavana problematika pojava deformacij prečnega prereza profila med krivljenjem, ki sočasno nakazuje potencialno področje nadaljnjih raziskav za vrednotenje elastičnega izravnavanja in pojava zaostalih napetosti. S tem je podan prispevek k izboljšavi razumevanja kompleksnosti preoblikovalnih mehanizmov med krivljenjem profilov z metodo TSS. Na osnovi predstavljenega članka so podane smernice za delo v prihodnje. Načrtuje se izvedba dodatnih analiz postopkov ravninskega krivljenja ob različnih nastavitvah parametrov. Po potrditvi vseh pomembnih procesnih parametrov za 2D-krivljenje pa bodo končno opravljene še numerične analize napetostnih in deformacijskih stanj pri postopku prostorskega oz. 3D-krivljenja profilov.

Ključne besede: prostorska superpozicija navora, TSS, upogibanje, 3D MKE, 2D-krivljenje profilov

*Naslov avtorja za dopisovanje: TECOS, Razvojni center orodjarstva Slovenije, Kidričeva 25, 3000 Celje, Slovenija, matej.hudovernik@tecos.si

SI 29

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 30 © 2013 Strojniški vestnik. Vse pravice pridržane.

Prejeto v recenzijo: 2012-03-06 Prejeto popravljeno: 2012-10-02 Odobreno za objavo: 2013-01-03

Numerična simulacija preoblikovanja v hladnem pločevin iz α-titanovih zlitin Jurendić, S. – Gaiani, S. Sebastijan Jurendić1, * – Silvia Gaiani1,2 1 Akrapovič d.d., Slovenija 2 Univerza v Modeni in Reggio Emilia, Oddelek za inženiring materialov, Italija

S širjenjem uporabe titanovih zlitin se tudi klasični proizvodni procesi vedno pogosteje uporabljajo pri proizvodnji izdelkov iz teh zlitin. To je izziv za serijsko proizvodnjo, saj se osnovne mehanske lastnosti titanovih zlitin lahko bistveno razlikujejo od lastnosti klasičnih konstrukcijskih materialov, ki jih te zlitine zamenjujejo, čeprav je v splošnem preoblikovalnost titanovih zlitin dobra. Numerične analize so postale nepogrešljivo orodje na področju snovanja, vrednotenja in optimiranja proizvodnih procesov, saj omogočajo vpogled v mehaniko preoblikovanja, ki bi ga bilo sicer nemogoče pridobiti. Zato je bila razvita metodologija popisa in modeliranja pločevin iz α-titanovih zlitin za numerične simulacije preoblikovalnih procesov v hladnem, in predstavljena na primeru simulacije globokega vleka. Analizirana je bila mehanika plastifikacije α-titanovih zlitin s heksagonalno kristalno strukturo in njen odraz na makroskopskih mehanskih lastnostih pločevine. Identificirane so bile poglavitne karakteristike elasto-plastičnega odziva materiala, ki so pomembne za fenomenološki popis makroskopskih mehanskih lastnosti teh zlitin. Ob upoštevanju naštetega je bil izbran anizotropen materialni model, ki omogoča dovolj natančen popis materiala in je hkrati še dovolj praktičen za uporabo v industrijskem okolju, torej omogoča določitev potrebnih materialnih parametrov z razpoložljivimi sredstvi in v doglednem času.

Opredeljeni in analizirani so potrebni materialni podatki za izbrani ortotropni Barlatov (1989) materialni model, prav tako so razdelane metode za njihovo določevanje. Večino podatkov se da pridobiti iz standardnega nateznega preizkusa z vzdolžnim in prečnim merjenjem raztezkov. Za določanje parametrov anizotropije materiala so potrebni podatki v vzdolžni, diagonalni in prečni smeri glede na smer valjanja pločevine. Ker pločevine iz α-titanovih zlitin kažejo anizotropijo tako v meji tečenja kot tudi v utrjevanju materiala, so parametri anizotropije določeni kot funkcije plastičnega raztezka. Utrjevalna krivulja je določena z inverznim numeričnim postopkom, ki zajema numerično simuliranje nateznega preizkusa in iterativno spreminjanje utrjevalne krivulje, dokler se simulirani odziv ne ujema z izmerjenim. Tak postopek je nujen za določevanje odziva materiala po začetku lokalnega tanjšanja vzorca, saj pločevine iz α-titanovih zlitin po začetku lokalnega tanjšanja nateznega vzorca dosegajo visoke raztezke. Za določitev Barlatovega eksponenta krivulje plastičnega tečenja je potreben dvoosni preizkus. V ta namen je bil uporabljen standardni Erichsenov tehnološki preizkus, eksponent pa je bil določen s parametrično analizo numerične simulacije Erichsenovega preizkusa. Tako določeni materialni podatki so bili uporabljeni v simulaciji postopka globokega vleka. Uporabljena je bila oblika, ki zagotavlja porušitev materiala, kar omogoča vrednotenje simulacije do mej preoblikovalnosti. Za merilo porušitve je bila uporabljena krivulja mejnih deformacij. Rezultati simulacij kažejo razmeroma dobro ujemanje s preizkusi pri predvidevanju mej preoblikovalnega postopka, vendar so simulacije nekoliko konzervativne.

Glavna omejitev te metode je materialni model, ki ne upošteva asimetrije med mejo tečenja v nategu in tlaku α-titanovih pločevin, hkrati pa je težavno oceniti, v kolikšni meri to vpliva na netočnost rezultatov, saj so podatki o tlačnih lastnostih pločevin izredno težko merljivi. Članek predstavlja praktično uporabo simulacije preoblikovanja α-titanove pločevine z inovativno uporabo Barlatovega (1989) materialnega modela, ki zaobide nekatere omejitve tega materialnega modela in na razmeroma preprost način, z uporabo preprostih in uveljavljenih preizkusov, omogoča uporaben numerični popis teh materialov.

Ključne besede: α-titan, heksagonalna kristalna struktura, numerična simulacija, preoblikovanje v hladnem, anizotropija, globoki vlek

SI 30

*Naslov avtorja za dopisovanje: Akrapovič, d.d., Malo Hudo 8a, 1295 Ivančna Gorica, Slovenija, sebastijan.jurendic@novelis.com

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 31 © 2013 Strojniški vestnik. Vse pravice pridržane.

Prejeto v recenzijo: 2012-03-23 Prejeto popravljeno: 2012-12-14 Odobreno za objavo: 2013-02-04

Izboljšanje postopka mikroelektroerozijske obdelave z ultrazvočnimi vibracijami in pristopi k obdelavi neprevodne keramike

Schubert, A. – Zeidler, H. – Hackert-Oschätzchen, M. – Schneider, J. – Hahn, M. Andreas Schubert1,2 – Henning Zeidler1,* – Matthias Hackert-Oschätzchen1 – Jörg Schneider2 – Hahn, Martin1 1Tehniška

univerza v Chemnitzu, Katedra za mikroproizvodne tehnologije, Nemčija institut za obdelovalne stroje in tehnologije preoblikovanja IWU, Nemčija

2Fraunhoferjev

Mikroelektroerozijska obdelava je dobro znan nekonvencionalni postopek za obdelavo materialov, ki jih je težavno obdelovati z odrezavanjem. Ta postopek odvzemanja materiala s taljenjem in uparjanjem, kjer se toplota dovaja z električnimi razelektritvami, ni odvisen od trdote, žilavosti ali krhkosti obdelovanca. Elektroerozijska obdelava se zato pogosto uporablja v orodjarstvu, mikroelektroerozijska obdelava z bistveno manjšimi energijami razelektritev pa se je uveljavila pri mikroobdelavi delov visoke natančnosti.

Natančna izdelava mikrogeometrij z velikim razmerjem dimenzij, kot so npr. globoke mikroizvrtine, je odvisna od stabilnih pogojev procesa v reži. Minimalna reža je pogoj za izdelavo miniaturnih oblik z visoko natančnostjo, ki hkrati omejuje učinkovitost običajnih tehnik čiščenja, povzroča večji delež neželenih razelektritvenih stanj (prekinjen tokokrog in kratek stik), upočasnjuje proces in uvaja geometrijske napake. Novi pristopi hibridnih tehnologij, npr. s superpozicijo ultrazvoka ali nizkofrekvenčnega zvoka, občutno izboljšajo stabilnost in hitrost procesa.

Naslednja omejitev postopka, t. j. zmožnost obdelave samo električno prevodnih materialov, je odpravljena s pomožno elektrodo za mikroelektroerozijsko obdelavo električno neprevodne keramike iz cirkonijevega oksida.

V članku je predstavljeno stanje razvoja na področju mikroelektroerozijske obdelave kovinskih materialov in električno neprevodnih keramik s podporo neposrednih ultrazvočnih vibracij obdelovanca in posrednih visokointenzivnih ultrazvočnih vibracij dielektrika.

Mikroelektroerozijska obdelava s podporo ultrazvoka je lahko hitrejša tudi do 40 % ter omogoča vrtanje lukenj premera pod 90 mm in z razmerjem dimenzij nad 40 pri kovinskih materialih. Prilagojena zasnova naprave s pomožno elektrodo omogoča obdelavo oblik z razmerjem dimenzij nad 5 pri neprevodnih keramičnih materialih, s čimer se odpirajo nove možnosti za snovanje in proizvodnjo zahtevnih visokonatančnih mikroizdelkov iz viskozmogljivih tehničnih materialov.

Področje mikroobdelave se sooča z vedno novimi izzivi zaradi vse večjega povpraševanja po manjših in natančnejših konstrukcijah iz novih materialov. Elektroerozijska obdelava je lahko pravi odgovor na te zahteve pri materialih, ki so zahtevni za obdelavo. Superpozicija vibracij pa danes še ni popolnoma prilagojena zahtevam industrijske obdelave in jo je treba pripraviti za vsak obdelovanec posebej. Potrebnih je več raziskav, ki bodo privedle do uporabnih rešitev. Elektroerozijska obdelava je danes uporabna le za omejeno število neprevodnih keramičnih materialov. Nujna je podrobna analiza mehanizmov procesa obdelave, ki bo izboljšala njihovo razumevanje in dala tudi rešitve elektrod za industrijsko okolje.

Velik potencial pri obdelavi prevodnih materialov imajo hibridni procesi, ki izboljšujejo čiščenje in razmere v reži.

Ultrazvočna superpozicija se odlično odreže pri stabilizaciji procesa ter lahko pospeši mikroelektroerozijsko obdelavo zelo globokih in natančnih struktur.

Nove aplikacije zahtevajo geometrije in razmerja dimenzij, ki jim obstoječi postopki niso kos, z vključitvijo pomožne elektrode pa bodo za nove aplikacije uporabni tudi materiali, ki jih je zelo težko obdelovati, tehnični materiali in biozdružljive keramike.

Z združitvijo obeh pristopov je mogoče izkoristiti nove priložnosti za snovanje in izdelavo kompleksnih in visokonatančnih mikroizdelkov iz visokozmogljivih tehničnih materialov.

Ključne besede: elektroerozijska obdelava, mikroobdelava, ultrazvok, keramika, natančna obdelava, visokozmogljivi materiali

*Naslov avtorja za dopisovanje: Tehniška univerza v Chemnitzu, Katedra za mikroproizvodne tehnologije, 09107 Chemnitz, Nemčija, henning.zeidler@mb.tu-chemnitz.de

SI 31

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Prejeto v recenzijo: 2012-04-11 Prejeto popravljeno: 2012-07-31 Odobreno za objavo: 2012-10-25

Lastnosti trdih prevlek na magnezijevih zlitinah AZ61 Tański, T. Tomasz Tański*

Šlezijska tehniška univerza, Institut za tehnične materiale in biomateriale, Poljska

Cilj tega inovativnega dela je iskanje najboljših hibridnih prevlek, sestavljenih iz prehodnega sloja z zvezno spremembo ene ali več komponent od substrata do vrha površine, ter iz zunanjega sloja. Prevleke so nanesene po postopku PVD ali CVD na površino ulitka iz magnezijeve zlitine AZ61 ter izboljšujejo lastnosti in majhno togost substrata. Članek obravnava lastnosti in strukturo prevleke kot osnovo za vrednotenje kakovosti prevleke.

Struktura substrata in prevleke je bila analizirana z uporabo vrstičnega elektronskega mikroskopa ZEISS SUPRA 35 in presevnega elektronskega mikroskopa STEM TITAN 80-300, ter s preizkusom protiobrabne obstojnosti in adhezivnosti prevleke. Uporabljen je bil postopek naparevanja plasti s katodnim lokom v atmosferi Ar, N2 in C2H2.

Za nanos prevlek so bile uporabljene katode iz čiste kovine (Cr, Ti) in TiAl (50:50 atom. %) premera 65 mm.

Osnovni tlak po izčrpavanju v komori je bil 5x103 Pa. Substrat je bil nato 20 minut čiščen z ioni argona pri tlaku 2 Pa. Za izboljšanje adhezije prevlek je bil nanesen vmesni prehodni sloj Cr ali Ti. Delovni tlak med postopkom nanašanja je bil 2 do 4 Pa, odvisno od vrste prevleke. Razdalja med katodo in substratom je bila 120 mm.

Kot prekurzor pri nanosu prevleke DLC po postopku PACVD je bil uporabljen acetilen (C2H2). Substrat je bil 20 minut čiščen z ioni argona pri tlaku 2 Pa in nosilni napetosti 800/200 V, za izboljšanje adhezije prevlek pa je bil nanesen vmesni prehodni sloj Ti. Nanesene prevleke imajo eno-, dvo- ali večslojno strukturo, posamezni sloji pa so enakomerni ter se dobro držijo substrata in drug drugega. Rezultati difrakcijske analize omogočajo identifikacijo faz TiAlN, CrN in grafita v površinskem sloju. Analiza PVD- in CVD-prevlek na površini ulitka iz magnezijeve zlitine je jasno pokazala, da je kritična obremenitev Lc v območju 8 do 17 N, odvisno od vrste prevleke. Največjo obrabno obstojnost imajo prevleke DLC. Kljub temu, da članek sodi v obširno in sodobno raziskovalno področje, pa rezultati zadevajo samo izbrano skupino in metodologijo preskušanja magnezijevih zlitin Mg-Al-Zn. Za doseganje novih operativnih ter funkcijskih značilnosti in lastnosti pogosto uporabljenih materialov, kamor sodijo tudi zlitine Mg-Al-Zn, se pogosto izvaja toplotna obdelava, t. j. izločevalno utrjanje in/ali površinska toplotna obdelava, površine pa se nato dodatno obdelajo z različnimi postopki za oplemenitenje površin.

V članku je predstavljena raziskava PVD- in CVD-prevlek na magnezijevih zlitinah kot nekonvencionalnem substratu. Sodobni materiali morajo imeti za dolgo in zanesljivo delo odlične mehanske, fizikalne, kemične in tehnološke lastnosti. Zlitine neželeznih kovin, kamor spadajo tudi magnezijeve zlitine, izpolnjujejo zahteve in pričakovanja glede sodobnih materialov.

Ključne besede: proizvodnja in obdelava, tanke in debele prevleke, magnezijeve zlitine, prevleke PVD in CVD, struktura, lastnosti

SI 32

*Naslov avtorja za dopisovanje: Šlezijska tehniška univerza, Institut za tehnične materiale in biomateriale, Konarskiego St. 18a, Gliwice, Poljska, tomasz.tański@polsl.pl

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 33 © 2013 Strojniški vestnik. Vse pravice pridržane.

Prejeto v recenzijo: 2012-06-02 Prejeto popravljeno: 2012-10-28 Odobreno za objavo: 2012-11-29

Konvencionalni in nekonvencionalni postopki izdelave elektrod na silicijevih sončnih celicah

Dobrzański, L.A. – Musztyfaga, M. – Drygała, A. Leszek Adam Dobrzański – Małgorzata Musztyfaga* – Aleksandra Drygała Šlezijska tehniška univerza, Institut za tehnične materiale in biomateriale, Pojska Članek podaja primerjavo konvencionalnih in nekonvencionalnih postopkov oblikovanja elektrod na silicijevih sončnih celicah z namenom izboljšanja kakovosti elektrod. Kontakti komercialnih silicijevih sončnih celic se najpogosteje oblikujejo po postopku sitotiska. Selektivno lasersko sintranje (SLS) je sodoben proizvodni postopek, kjer se delci kovinskega prahu na točno določenih mestih talijo ali sintrajo v želeno trodimenzionalno obliko s pomočjo CO2-laserja velike moči, celoten postopek pa nadzoruje program za mikroobdelavo. V članku so predstavljeni rezultati preiskave sprednjih elektrod, izdelanih iz dveh vrst srebrove paste (PV 145 proizvajalca Du Pont in eksperimentalnega nanoprahu) na monokristalnih silicijevih sončnih celicah, cilj pa je bil zmanjšanje kontaktne upornosti.

Jedro raziskave torej predstavlja iskanje optimalne zgradbe stika, ki mora zagotavljati trajno povezavo med elektrodo in silicijevim substratom, ne sme imeti por ali mikrorazpok ter daje minimalno upornost sistema. Pri tem so bili preučeni vplivi parametrov obdelave na strukturo in lastnosti sprednje elektrode. Minimalna upornost stika med silicijem in sprednjo elektrodo fotonapetostne celice je ključen element pri snovanju fotonapetostnih celic, ki neposredno vpliva na vzporedno upornost fotonapetostne celice ter določa vrednost faktorja polnjenja, zato je neposredno povezan z izkoristkom fotonapetostne pretvorbe. Dejanski cilj projekta je priprava smernic za izbiro primernega postopka izdelave sprednje elektrode ter orodja za samodejno oblikovanje elektrode. Upornost sprednje elektrode je bila določena po metodi prenosnih vodov (TLM) na preizkuševališču, ki je bilo razvito in zgrajeno na Institute za tehnične materiale in biomateriale. Metoda TLM vključuje neposredno merjenje toka (I) in napetosti (U) med dvema ločenima kontaktoma.

Raziskana je bila topografija vzorcev, izdelanih z žganjem v infrardeči peči oz. taljenih/sintranih v napravi Eosint M250 Xtended s CO2-laserjem. Kontakti so bili pregledani s pomočjo konfokalnega laserskega mikroskopa (CLSM 5) in vrstičnega elektronskega mikroskopa (SEM) z energijsko disperzijskim spektrometrom (EDS) za mikrokemijsko analizo. Tako površinska topografija kot prerez sprednje elektrode sta bila preiskana z mikroskopom SEM. Analiza fazne sestave sprednjih elektrod je bila opravljena po metodi XRD. Sprednje elektrode so bile oblikovane na površini sončnih celic z različno morfologijo, srednja velikost piramid pa je bila izmerjena z mikroskopom na atomsko silo (AFM).

Ugotovljeno je bilo, da ima najboljše električne in strukturne lastnosti med analiziranimi sprednjimi elektrodami lasersko sintrana elektroda, izdelana iz nanopaste, medtem ko daje med konvencionalnimi postopki najboljše električne in strukturne lastnosti lasersko sintranje elektrode iz paste PV 145. Podana so eksperimentalno določena tehnološka priporočila za lasersko mikroobdelavo sprednje elektrode silicijevih sončnih celic, ki vključujejo optimalno sestavo paste, moč in hitrost skeniranja laserskega žarka ter morfologijo silicijevega substrata, zagotavljajo pa enakomerno pretalitev materiala, ki se dobro drži substrata, kakor tudi majhno upornost stika sprednje elektrode in substrata. V tem delu je bila uporabljena tudi inovativna rešitev nanosa srebrove nanoplasti za začetni sloj sprednje elektrode pri metodi SLS. Kontaktna upornost izdelane sprednje elektrode je odvisna od sestave paste, morfologije površine silicijeve rezine ter pogojev žganja in laserske mikroobdelave.

Ključne besede: električne lastnosti, sončna celica, selektivno lasersko sintranje, sitotisk, kontaktna upornost, model prenosnega voda (TLM), monokristalni silicij

*Naslov avtorja za dopisovanje: Šlezijska tehniška univerza, Institut za tehnične materiale in biomateriale, Konarskiego St. 18a, Gliwice, Poljska, malgorzata.musztyfaga@polsl.pl

SI 33

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 34 © 2013 Strojniški vestnik. Vse pravice pridržane.

Prejeto v recenzijo: 2012-03-07 Prejeto popravljeno: 2012-08-14 Odobreno za objavo: 2012-09-11

Karakterizacija in priprava za brizganje PA 6, ojačenega s halojzitnimi nanocevkami (HNT) Fernández, A. – Muniesa, M. – González, J. Angel Fernández1,2,* –Manuel Muniesa1 – Jaime González2 1 Univerza v Zaragozi, Oddelek za strojništvo, Španija 2 Tehnološki center AITIIP, Tehnološki park Cogullada, Španija

Nanoojačitve izboljšajo mehanske lastnosti materiala, uporaba nanocevk naravnega izvora pa tudi izboljšuje trajnostnost in zmanjšuje vpliv izdelka na okolje. V članku je predstavljena priprava, karakterizacija in predelava nanokompozitov na osnovi termoplastičnega poliamida 6 in halojzitnih nanocevk (HNT). Raziskava je bila usmerjena v dokazovanje primernosti uporabe teh materialov za brizganje izdelkov. Ugotavljanje potenciala za uporabo teh materialov v serijski proizvodnji zahteva uporabo industrijske opreme za pripravo nanokompozitov in izdelavo preskusnih izdelkov. Cilj raziskave je pripraviti in analizirati polimerni material PA6 z dispergiranjem dodanih nanocevk s tremi ali šestimi utežnimi deleži za brizganje vzorcev za natezni in upogibni preiskus ter za izdelavo vzorcev za preiskus gorenja.

Nanokompoziti so bili pripravljeni v več korakih, rezultati pa omogočajo končno uporabo v industriji. V prvi formulaciji je bil uporabljen velik delež halojzitnih nanocevk (do 30 % mase), pripravljena pa je bila v dvopolžnem ekstrudorju. Končno je bila uporabljena cenena in trajnostna tehnika brizganja, s katero so bili pripravljeni preizkušanci iz vseh razvitih materialov.

Rezultati so bili analizirani na mikroskopski ravni s termogravimetrično analizo (TGA), Fourierjevo infrardečo spektroskopijo (FTIR), rentgensko difrakcijo (XRD), vrstično elektronsko mikroskopijo (SEM) in presevno elektronsko mikroskopijo (TEM), na makroskopski ravni pa s pomočjo kapilarne reometrije, nateznega in upogibnega preizkusa ter požarnega preizkusa.

Pri uporabi industrijske opreme za brizganje se izkaže pomanjkanje homogenosti končnega materiala, v nasprotju s splošno veljavnim prepričanjem, da naj bi bili vsi masterbatchi pripravljeni za brizganje. Rezultati dokazujejo potrebo po doseganju najboljše disperzije dodatka z ekstruzijskim kompaunderjem v prvi stopnji.

Tehnike XRD, FTIR, SEM in TEM so se izkazale za najuporabnejše pri preučevanju disperzije nanocevk v matriksu. Razkrile so, da prinaša kombinacija procesa ekstruzijskega kompaundiranja in brizganja izboljšane mehanske lastnosti zaradi izboljšane disperzije nanocevk pri visokih hitrostih ekstruzije. Takšen sklep potrjuje tudi odkritje maksimalne vrednosti poliamidne faze g v rezultatih FTIR. Obdelovalnost nanokompozitov je bila dokazana tudi pri visoki vsebnosti nanocevk, saj izmerjena viskoznost ostaja približno enaka kot pri surovem materialu PA 6. Požarna obstojnost se je z uporabo nanocevk nekoliko izboljšala in material ostaja v razredu UL94 V2. Togost izdelkov z vsebnostjo nanocevk se je v vseh primerih povečala.

Prihodnje raziskave bi morale biti usmerjene v uporabo drugih nanoojačitev in kombiniranje z drugimi polimernimi matriksi. Uporaba teh materialov za izdelavo zahtevnejših izdelkov bi tudi omogočila napovedovanje vpliva geometrije na zmogljivost celotnega izdelka.

Vrednost tega članka je v ovrednotenju vpliva metod in pogojev predelave na materialne lastnosti in zmogljivost preizkušancev. Novost pri tej raziskavi je uporaba industrijske opreme za pripravo materiala in proizvodnjo. Znanstvene metode, uporabljena oprema in rezultati so zanimivi za praktično uporabo v industriji predelave plastike, s posebnim poudarkom na avtomobilski industriji in drugih visokozahtevnih področjih.

Ključne besede: nanokompoziti, halojzitne nanocevke, brizganje, kompaundiranje

SI 34

*Naslov avtorja za dopisovanje: Univerza v Zaragozi, Oddelek za strojništvo, Edif. Torres Quevedo. María de Luna, 3, 50014 Zaragoza, Španija, angel.fernandez@unizar.es

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 35 © 2013 Strojniški vestnik. Vse pravice pridržane.

Prejeto v recenzijo: 2012-06-15 Prejeto popravljeno: 2012-10-04 Odobreno za objavo: 2012-10-15

Učinkovita raba virov in majhen vpliv na okolje pri brizganju plastike

Gantar, G. – Glojek, A. – Mori, M. – Nardin, B. – Sekavčnik, M. Gašper Gantar1,* – Andrej Glojek2 – Mitja Mori3 – Blaž Nardin4 – Mihael Sekavčnik3 Slovenija Razvojni center orodjarstva Slovenije, Slovenija 3 Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija 4 Gorenje Orodjarna, Slovenija 2 TECOS,

1 ENVITA d.o.o.,

Brizganje je eden najbolj razširjenih postopkov za proizvodnjo plastičnih izdelkov. Kakor pri ostalih postopkih preoblikovanja, so tudi lastnosti postopkov brizganja in izdelkov močno odvisne od kakovosti orodij, ki jih vgradimo v stroje za brizganje in morajo proizvajati ponovljive izdelke. Proces izbire najprimernejše zasnove in tehnološke rešitve za proizvodnjo nekega izdelka mora vključevati tehnične zmogljivosti, ekonomske vidike in obremenitev okolja med celotnim življenjskim ciklom orodja. V članku je predstavljena primerjava in optimizacija zasnove orodij za brizganje plastike ter njihovih tehničnih parametrov s tehničnega, ekonomskega in okoljskega vidika.

Pripravljena orodja so del proizvodnega procesa mnogo let. Ker surovine postajajo vse redkejše in dražje, stroški energije pa naraščajo, mora biti vsaka strategija za snovanje orodij usmerjena ne le v nižanje stroškov, temveč tudi v zmanjševanje porabe virov in emisij v celotnem življenjskem ciklu orodja. Študija primera vzporedno obravnava tri različne konstrukcije orodja. Orodja so namenjena brizganju istega plastičnega izdelka, razlikujejo pa se po tehničnih rešitvah, ki prispevajo k produktivnosti in učinkovitosti rabe virov v fazi eksploatacije (t. j. brizganja plastičnih izdelkov). Predstavljena je tudi analiza življenjskega cikla.

Vpliv na okolje je kvantificiran z analizo življenjskega cikla (LCA), ekonomski vidik pa je obravnavan z analizo stroškov v življenjskem ciklu (LCC). Pristop je podprt z numeričnimi simulacijami za napovedovanje pomembnih tehničnih parametrov v zgodnjih fazah konstrukcije orodja.

Analiza LCA je pokazala optimizirane rešitve, ki omogočajo določeno znižanje stroškov (celotni stroški na izdelek so manjši za 1,25%) in pomembno zmanjšanje vpliva na okolje (celoten vpliv na okolje je lahko manjši za približno 20%). Izkazalo se je, da spremembe konstrukcije orodja, ki ne vplivajo na zmogljivost procesa brizganja in na izboljšanje strategije ravnanja ob izteku življenjske dobe, ne morejo pomembneje zmanjšati celotnega vpliva na okolje. V splošnem lahko povzamemo, da je zasnova orodja pomemben dejavnik vpliva na okolje v fazi brizganja. Samoumevno je, da boljši proizvodni proces porabi manj virov v serijski proizvodnji (energije in surovin).

V študiji je bilo privzeto, da zahteve glede kakovosti izpolnjujejo brizganci, izdelani v vseh treh orodjih.

V primeru, da bi bila tehnična zmogljivost določenega orodja bistveno drugačna (zanesljivost tehnologije, časi ciklov, število izdelovalnih operacij itd.), bi bilo v sistem odločanja smiselno vključiti tudi tehnično zmogljivost.

Predstavljeni pristop je enostavno integrirati v standardno ponudbo storitev naprednih orodjarskih podjetij.

Pristop omogoča optimizacijo z ozirom na tehnične lastnosti procesa in zasnovo orodja, stroške in vpliv na okolje.

Ključne besede: brizganje plastike, orodje, optimizacija, analiza življenjskega cikla, analiza stroškov v življenjskem ciklu

*Naslov avtorja za dopisovanje: ENVITA d.o.o., Tržaška cesta 132, 1000 Ljubljana, Slovenija, gasper.gantar@envita.si

SI 35

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 36-38 Osebne objave

Doktorske disertacije, magistrska dela 2. stopnje, diplomske naloge

DOKTORSKE DISERTACIJE Na Fakulteti za strojništvo Univerze v Ljubljani je obranil svojo doktorsko disertacijo: ● dne 13. februarja 2013 Primož KRŽIČ z naslovom: »Izdelava in optimiranje računalniških algoritmov za dosego boljše kakovosti NC programov« (mentor: prof. dr. Janez Kopač); Z razvojem industrijskih robotov se njihova uporaba širi na nova tehnološka področja od katerih je zelo razširjeno tudi frezanje s pomočjo industrijskega robota. Za nadaljnjo popolno uveljavitev frezanja s pomočjo industrijskih robotov je ena od najpomembnejših omejitev pomanjkanje programske opreme, ki bi uporabnikom omogočala hitrejšo in bolj intuitivno evalvacijo kinematičnih omejitev industrijskih robotov. Doktorsko delo predstavi pomen robotskega frezanja in programske opreme zanj v sodobnem industrijskem okolju. Nadalje je predstavljen razvoj kinematičnega modela za industrijski robot KUKA KR150 L110-2 ter njegova razširitev na področje frezanja.

Opravljena je analiza možnih metod za razreševanje kinematičnih omejitev ter njihova aplikacija na kinematični model robota. Razvit je evalvacijski sistem za razreševanje kinematičnih omejitev pri frezanju, kateri s pomočjo treh filtrirnih algoritmov omogoča občutno skrajšanje računskih časov, na osnovi evalvacijskih algoritmov pa olajša izbiro primerne konfiguracije za uspešno izvedbo želene poti orodja. S pomočjo vizualizacijskega sistema so rezultati grafično predstavljeni uporabniku. Analiza delovanja filtrirnih algoritmov in evalvacijskega sistema pokaže, da rešitev s filtriranjem lahko izjemno pripomore k skrajšanju časa izračuna.

Z integracijo evalvacijskega in vizualizacijskega sistema v programski paket Mastercam/Robotmaster pa ponuja uporabniku boljše možnosti izbire pri nastavljanju parametrov obdelave ter hitrejši izbor najugodnejše rešitve kinematičnih omejitev in končno iz tega razloga pripomore k večji kakovosti NC programov. * Na Fakulteti za strojništvo Univerze v Mariboru sta obranila svojo doktorsko disertacijo: ● dne 15. februarja 2013 Matej FIKE z naslovom: »Eksperimentalna in numerična raziskava tokovnih SI 36

pojavov v aksialnem ventilatorju« (mentor: prof. dr.

Aleš Hribernik); Aksialni ventilatorji so zasnovani tako, da delujejo v stabilnem področju dušilne krivulje.

Prostorske oziroma druge omejitve lahko povzročijo, da delovanje ventilatorja preide iz stabilnega v nestabilno področje, kjer se razvijejo kompleksni 3D časovno odvisni tokovni pojavi, ki negativno vplivajo na karakteristiko ventilatorja. Da bi razumeli kompleksne pojave v medlopatičnem kanalu ventilatorja, je bila najprej narejena eksperimentalna in numerična analiza toka okoli osamljenega krila.

V nadaljevanju je predstavljena eksperimentalna in numerična raziskava tokovnih pojavov v aksialnem ventilatorju s poudarkom na nestacionarnem delovanju. V okvirju eksperimentalne raziskave je bila posneta dušilna krivulja, ki je bila razdeljena na stabilni in nestabilni del. S PIV merilno metodo so bila posneta tokovna polja pred rotorjem ventilatorja in v medlopatičnem prostoru v različnih delovnih točkah.

Raziskovanje je usmerjeno v analizo vrtečega zastoja, ki se razvije pri prehodu iz stabilnega v nestabilno delovanje. Razvita je bila metoda rekonstrukcije vrtečega zastoja, s katero je bila v izbranih delovnih točkah opravljena rekonstrukcija vrtečega zastoja pred rotorjem in v medlopatičnem kanalu. Prav tako je bil izračunan delež povratnega toka v medlopatičnem kanalu. Rezultati meritev, integralni in lokalni, so bili primerjani z rezultati numeričnih simulacij. Izvedene so bile stacionarne in časovno odvisne simulacije z uporabljenima k-ε in SST turbulentnima modeloma. Stabilni del dušilne krivulje je bil primerjan z rezultati stacionarnih simulacij, nestabilni del pa s časovno odvisnimi simulacijami.

Ugotovljeno je bilo, da stacionarne simulacije dobro napovejo časovno povprečena tokovna polja in potek stabilnega dela dušilne krivulje. Časovno odvisne simulacije z uporabljenima k-ε in SST turbulentnima modeloma slabo napovejo potek nestabilnega dela dušilne krivulje. Oba omenjena turbulentna modela uspeta napovedati pojav vrtečega zastoja, ne uspeta pa napovedati časovno odvisnega tokovnega polja v primeru osamljenega krila kot tudi v primeru ventilatorja; ● dne 18. februarja 2013 Janez LUPŠE z naslovom: »Numerični model robnih elementov za nestacionarne turbulentne tokove« (mentor: prof. dr.

Leopold Škerget); V doktorski disertaciji obravnavamo razvoj numeričnega algoritma za reševanje turbulentnih tokov, osnovanega izključno na metodi robnih

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3 SI 36-38

elementov (MRE). Rešujemo Navier-Stokesov sistem enačb, zapisan v hitrostno-vrtinčnem zapisu.

Sestavljen je iz enačb kinematike toka, s katerimi poiščemo neznane vrednosti vrtinčnega polja na robu območja reševanja in neznane vrednosti hitrostnega polja v območju, ter enačb kinetike vrtinčnosti, temperature in modelov turbulence.

Enačba, s katero poiščemo vrednosti vrtinčnega polja na robu območja, je diskretizirana s standardno, enoobmočno MRE, vse ostale enačbe pa so zapisane v diskretni obliki s pomočjo MRE s podobmočji, kar omogoča velike prihranke pri računalniškem spominu.

Ker je v večini realnih primerov turbulentnih tokov nepraktično ali celo nemogoče direktno reševati Navier-Stokesov sistem enačb, ga poenostavimo. V delu obravnavamo predvsem poenostavitev vodilnih enačb s pomočjo Reynoldsovega povprečenja (RANS). Z ohranitvijo lokalnega časovnega odvoda prenosnih spremenljivk časovno povprečenje omejimo na določen časovni interval ter tako dobimo neastacionarne-RANS enačbe (URANS).

Razviti algoritem smo preverili na testnih primerih, katerih analitične rešitve poznamo.

Naslednji korak je bil preračun laminarnih tokov, kjer smo izračunali tok v kanalu, tok v gnani kotanji in naravno konvekcijo v zaprti kotanji. Na koncu smo izračunali še vrsto turbulentnih tokov: turbulentni tok v kanalu, turbulentni tok preko stopnice v kanalu, turbulentni tok v kanalu s periodičnimi zožitvami in tok v kanalu s kvadratno oviro. Z naštetimi primeri smo potrdili pravilnost razvitega numeričnega algoritma in implementacije modelov turbulence, tako za časovno povprečene kot tudi nestacionarne izračune turbulentnega toka.

MAGISTRSKA DELA 2. STOPNJE Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv magister inženir strojništva: dne 27. februarja 2013: Špela BRGLEZ z naslovom: »Numerična analiza tokovnih in toplotnih razmer pri zaviranju zavornega diska« (mentor: prof. dr. Leopold Škerget); dne 28. februarja 2013: Karl MLAKAR z naslovom: »Vplivi prostorskega modeliranja izdelkov na računalniško integrirano proizvodnjo« (mentor: prof. dr. Jože Balič); * Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv magister gospodarski inženir: dne 27. februarja 2013:

Miha HRIBAR z naslovom: »Razvoj modela centralnega nadzora in spremljanja razvojnih projektov v Skupini Gorenje d.d.« (mentor: doc. dr.

Iztok Palčič, somentor: prof. dr. Anton Hauc); dne 28. februarja 2013: Karmen KOSTANJŠEK z naslovom: »Vpliv znanja na inovativnost in produktivnost v industrijskem okolju avtokonfekcije« (mentor: doc. dr. Marjan Leber, somentor: prof. dr. Vojko Potočan); DIPLOMSKE NALOGE Na Fakulteti za strojništvo Univerze v Ljubljani sta pridobila naziv univerzitetni diplomirani inženir strojništva: dne 28. februarja 2013: Miha HRIBAR z naslovom: »Priprava deionizirane vode za farmacevtsko industrijo« (mentor: prof. dr. Iztok Golobič); Mitja HVALA z naslovom: »Avtomatizirana kontrola kakovosti grafitnih komutatorjev na osnovi merjenja upornosti« (mentor: prof. dr. Peter Butala). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 28. februarja 2013: Simon ČREŠNAR z naslovom: »Konstruiranje ekso-skeleta« (mentor: doc. dr. Aleš Belšak, somentor: izr. prof. dr. Miran Ulbin); Damjan GOBEC z naslovom: »Inovativna zasnova zložljive kuhinje za majhna stanovanja« (mentor: izr. prof. Vojmir Pogačar); Jernej MRVIČ z naslovom: »Zasnova hidravlične opreme za stiskanje bučnih semen« (mentor: doc. dr.

Darko Lovrec); Tomaž PLANINŠIČ z naslovom: »Osnovne smernice za razvoj sodobnih modulnih trezorskih prostorov« (mentor: izr. prof. dr. Bojan Dolšak, somentor: asist. dr. Jasmin Kaljun). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv univerzitetni diplomirani gospodarski inženir: dne 28. februarja 2013: Primož GAJŠEK z naslovom: »Projekt vzpostavitve proizvodnje za izdelavo centrirnih izvrtin v podjetju Impakta metal d.o.o.« (mentor: doc. dr. Iztok Palčič, somentorica: doc. dr. Karin Širec); SI 37

Strojniški vestnik - Journal of Mechanical Engineering 59(2013)3, SI 36-38

* Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva: dne 6. februarja 2013: Miha KOPRIVEC z naslovom: »Optimizacija in dodelava hladnega obloka za injekcijsko brizganje gume« (mentor: doc. dr. Davorin Kramar, somentor: prof. dr. Janez Kopač); Jernej MARZEL z naslovom: »Karakterizacija obdelave titanove (Ti6Al4) in nikljeve (Inconel718) zlitine z uporabo različnih hladilno-mazalnih sredstev« (mentor: doc. dr. Franci Pušavec, somentor: prof. dr. Janez Kopač); Nejc JEZERŠEK z naslovom: »Merjenje cilindričnosti z linijskim laserjem« (mentor: doc. dr.

Joško Valentinčič, somentor: doc. dr. Henri Orbanić); Patrik MARKIČ z naslovom: »Energijska obremenjenost delovnega mesta in metoda za ocenjevanje tveganj nastanka poklicnih bolezni« (mentor: prof. dr. Vincenc Butala, somentor: doc. dr.

Boris Jerman); Rok TURŠIČ z naslovom: »Vsebnosti vodika v zvarnih spojih« (mentor: prof. dr. Janez Tušek); dne 11. februarja 2013: Alen ČERNELČ z naslovom: »Izraba toplote destilacije s toplotno črpalko« (mentor: prof. dr. Alojz Poredoš); Jan LEGAN z naslovom: »Analiza delovanja toplotne črpalke zrak-voda« (mentor: prof. dr. Alojz Poredoš); Maja LOVKO z naslovom: »Analiza poškodb udara strele v kompozitno oplato letala splošne kategorije« (mentor: izr. prof. dr. Tadej Kosel, somentor: doc. dr. Aleš Berkopec), Anja POPOVIČ z naslovom: »Izkoriščanje odvečne toplote v industriji za ogrevanje stavb« (mentor: prof. dr. Alojz Poredoš). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva: dne 28. februarja 2013:

SI 38

David BRODEJ z naslovom: »Sprememba konstrukcije orodja za brizganje kabelskega kanala« (mentorica: viš. pred. dr. Marina Novak); Matej ČUJEŠ z naslovom: »Pogon izvozne mize preše« (mentor: doc. dr. Janez Kramberger); David KNEZ z naslovom: »Kombinirani toplotni prenosniki dimnih plinov podjetja Eurovartrade d.o.o.« (mentor: doc. dr. Matjaž Ramšak, somentor: izr. prof. dr. Jure Marn); Josip MATANOVIĆ z naslovom: »Optimizacija procesa tlačnega litja s pomočjo jet cool sistema« (mentor: izr. prof. dr. Borut Buchmeister, somentor: doc. dr. Marjan Leber); Milan PECL z naslovom: »Razvoj orodja za stiskanje izdelkov« (mentor: izr. prof. dr. Stanislav Pehan); Matjaž ROZMAN z naslovom: »Konstruiranje hidravlične vpenjalne priprave za nosilec motorja« (mentorica: viš. pred. dr. Marina Novak, somentor: izr. prof. dr. Bojan Dolšak). * Na Fakulteti za strojništvo Univerze v Mariboru je pridobil naziv diplomirani inženir strojništva (UN): dne 28. februarja 2013: Jernej HUDOURNIK z naslovom: »Izdelava ohišja s postopkom brizganja plastičnih mas« (mentor: izr. prof. dr. Igor Drstvenšek, somentor: mag. Tomaž Brajlih); Aljaž MAJCEN RAVNAK z naslovom: »Razvoj in izdelava vodila za namestitev pedikularnih vijakov« (mentor: izr. prof. dr. Igor Drstvenšek, somentor: mag.

Tomaž Brajlih). * Na Fakulteti za strojništvo Univerze v Ljubljani je pridobil naziv diplomirani inženir strojništva (VS): dne 11. februarja 2013: Erik POLJANEC z naslovom: »Analiza sistema fotonapetostnih toplotnih sončnih sprejemnikov v kombinaciji s sezonskim senzibilnim hranilnikom toplote« (mentor: doc. dr. Andrej Kitanovski).

Strojniški vestnik – Journal of Mechanical Engineering (SV-JME) Aim and Scope The international journal publishes original and (mini)review articles covering the concepts of materials science, mechanics, kinematics, thermodynamics, energy and environment, mechatronics and robotics, fluid mechanics, tribology, cybernetics, industrial engineering and structural analysis.

The journal follows new trends and progress proven practice in the mechanical engineering and also in the closely related sciences as are electrical, civil and process engineering, medicine, microbiology, ecology, agriculture, transport systems, aviation, and others, thus creating a unique forum for interdisciplinary or multidisciplinary dialogue.

The international conferences selected papers are welcome for publishing as a special issue of SV-JME with invited co-editor(s).

Editor in Chief Vincenc Butala University of Ljubljana Faculty of Mechanical Engineering, Slovenia Technical Editor Pika Škraba University of Ljubljana Faculty of Mechanical Engineering, Slovenia Editorial Office University of Ljubljana (UL) Faculty of Mechanical Engineering SV-JME Aškerčeva 6, SI-1000 Ljubljana, Slovenia Phone: 386-(0)1-4771 137 Fax: 386-(0)1-2518 567 E-mail: info@sv-jme.eu, http://www.sv-jme.eu Print DZS, printed in 450 copies Founders and Publishers University of Ljubljana (UL) Faculty of Mechanical Engineering, Slovenia University of Maribor (UM) Faculty of Mechanical Engineering, Slovenia Association of Mechanical Engineers of Slovenia

http://www.sv-jme.eu

59 (2013) 3

Chamber of Commerce and Industry of Slovenia Metal Processing Industry Association

Since 1955

Strojniški vestnik Journal of Mechanical Engineering

Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, Karl Kuzman, Janez Marko Slabe ysis of 2D Profile Bending with the Torque Superposed thod

, Silvia Gaiani on of Cold Forming of α-Titanium Alloy Sheets

Hard Coatings on AZ61 Magnesium Alloys

ański, Małgorzata Musztyfaga, Aleksandra Drygała

ng Process of Front Side Metallisation on s Using Conventional and Unconventional Techniques

anuel Muniesa, Jaime González nd Processing of Reinforced PA 6 with Halloysite Nanotubes Molding

drej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik t Injection Moulding with Low Environmental Impacts

Journal of Mechanical Engineering - Strojniški vestnik

, Henning Zeidler, Matthias Hackert-Oschätzchen, artin Hahn EDM using Ultrasonic Vibration and Approaches Nonconducting Ceramics

year

no. 3 2013 59

volume

Cover: The cover picture represents a machine (above) for 3D bending of tubes and profiles with symmetrical or asymmetrical section properties using dynamic Torque Superposed Spatial - TSS bending method (below). This machine allows simultaneous numerical control of several axes - resulting in the feed of the profile, bending in a direction perpendicular to the profile feed, and a change of bending plane. A curvature with continuous or variable radius, in the domain of plane or spatial geometry of final part is, therefore, achieved.

Image Courtesy: Institute for Forming and Lightweight Construction - IUL, Technical University Dortmund

International Editorial Board Koshi Adachi, Graduate School of Engineering,Tohoku University, Japan Bikramjit Basu, Indian Institute of Technology, Kanpur, India Anton Bergant, Litostroj Power, Slovenia Franci Čuš, UM, Faculty of Mech. Engineering, Slovenia Narendra B. Dahotre, University of Tennessee, Knoxville, USA Matija Fajdiga, UL, Faculty of Mech. Engineering, Slovenia Imre Felde, Obuda University, Faculty of Informatics, Hungary Jože Flašker, UM, Faculty of Mech. Engineering, Slovenia Bernard Franković, Faculty of Engineering Rijeka, Croatia Janez Grum, UL, Faculty of Mech. Engineering, Slovenia Imre Horvath, Delft University of Technology, Netherlands Julius Kaplunov, Brunel University, West London, UK Milan Kljajin, J.J. Strossmayer University of Osijek, Croatia Janez Kopač, UL, Faculty of Mech. Engineering, Slovenia Franc Kosel, UL, Faculty of Mech. Engineering, Slovenia Thomas Lübben, University of Bremen, Germany Janez Možina, UL, Faculty of Mech. Engineering, Slovenia Miroslav Plančak, University of Novi Sad, Serbia Brian Prasad, California Institute of Technology, Pasadena, USA Bernd Sauer, University of Kaiserlautern, Germany Brane Širok, UL, Faculty of Mech. Engineering, Slovenia Leopold Škerget, UM, Faculty of Mech. Engineering, Slovenia George E. Totten, Portland State University, USA Nikos C. Tsourveloudis, Technical University of Crete, Greece Toma Udiljak, University of Zagreb, Croatia Arkady Voloshin, Lehigh University, Bethlehem, USA President of Publishing Council Jože Duhovnik UL, Faculty of Mechanical Engineering, Slovenia General information Strojniški vestnik – Journal of Mechanical Engineering is published in 11 issues per year (July and August is a double issue).

Institutional prices include print & online access: institutional subscription price and foreign subscription €100,00 (the price of a single issue is €10,00); general public subscription and student subscription €50,00 (the price of a single issue is €5,00). Prices are exclusive of tax.

Delivery is included in the price. The recipient is responsible for paying any import duties or taxes. Legal title passes to the customer on dispatch by our distributor.

Single issues from current and recent volumes are available at the current single-issue price. To order the journal, please complete the form on our website. For submissions, subscriptions and all other information please visit: http://en.sv-jme.eu/.

You can advertise on the inner and outer side of the back cover of the magazine. The authors of the published papers are invited to send photos or pictures with short explanation for cover content.

We would like to thank the reviewers who have taken part in the peerreview process.

ISSN 0039-2480 © 2013 Strojniški vestnik - Journal of Mechanical Engineering. All rights reserved. SV-JME is indexed / abstracted in: SCI-Expanded, Compendex, Inspec, ProQuest-CSA, SCOPUS, TEMA. The list of the remaining bases, in which SV-JME is indexed, is available on the website.

The journal is subsidized by Slovenian Book Agency.

Strojniški vestnik - Journal of Mechanical Engineering is also available on http://www.sv-jme.eu, where you access also to papers’ supplements, such as simulations, etc.

Instructions for Authors All manuscripts must be in English. Pages should be numbered sequentially. The maximum length of contributions is 10 pages. Longer contributions will only be accepted if authors provide justification in a cover letter. Short manuscripts should be less than 4 pages. For full instructions see the Authors Guideline section on the journal’s website: http://en.sv-jme.eu/.

Please note that file size limit at the journal’s website is 8Mb. Announcement: The authors are kindly invited to submitt the paper through our web site: http://ojs.sv-jme.eu. Please note that file size limit at the journal’s website is 8Mb. The Author is also able to accompany the paper with Supplementary Files in the form of Cover Letter, data sets, research instruments, source texts, etc. The Author is able to track the submission through the editorial process - as well as participate in the copyediting and proofreading of submissions accepted for publication - by logging in, and using the username and password provided.

Please provide a cover letter stating the following information about the submitted paper: 1. Paper title, list of authors and affiliations.

2. The type of your paper: original scientific paper (1.01), review scientific paper (1.02) or short scientific paper (1.03).

3. A declaration that your paper is unpublished work, not considered elsewhere for publication.

4. State the value of the paper or its practical, theoretical and scientific implications. What is new in the paper with respect to the state-of-the-art in the published papers?

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Every manuscript submitted to the SV-JME undergoes the course of the peer-review process.

THE FORMAT OF THE MANUSCRIPT The manuscript should be written in the following format: - A Title, which adequately describes the content of the manuscript.

- An Abstract should not exceed 250 words. The Abstract should state the principal objectives and the scope of the investigation, as well as the methodology employed. It should summarize the results and state the principal conclusions.

- 6 significant key words should follow the abstract to aid indexing.

- An Introduction, which should provide a review of recent literature and sufficient background information to allow the results of the article to be understood and evaluated.

- A Theory or experimental methods used.

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- A Results section, which should clearly and concisely present the data using figures and tables where appropriate.

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- Conclusions, which should present one or more conclusions that have been drawn from the results and subsequent discussion and do not duplicate the Abstract.

- References, which must be cited consecutively in the text using square brackets [1] and collected together in a reference list at the end of the manuscript.

Units - standard SI symbols and abbreviations should be used. Symbols for physical quantities in the text should be written in italics (e.g. v, T, n, etc.).

Symbols for units that consist of letters should be in plain text (e.g. ms-1, K, min, mm, etc.)

Abbreviations should be spelt out in full on first appearance, e.g., variable time geometry (VTG).

Meaning of symbols and units belonging to symbols should be explained in each case or quoted in a special table at the end of the manuscript before References.

Figures must be cited in a consecutive numerical order in the text and referred to in both the text and the caption as Fig. 1, Fig. 2, etc. Figures should be prepared without borders and on white grounding and should be sent separately in their original formats.

Pictures may be saved in resolution good enough for printing in any common format, e.g. BMP, GIF or JPG. However, graphs and line drawings should be prepared as vector images, e.g. CDR, AI. When labeling axes, physical quantities, e.g. t, v, m, etc. should be used whenever possible to minimize the need to label the axes in two languages.

Multi-curve graphs should have individual curves marked with a symbol. The meaning of the symbol should be explained in the figure caption.

Tables should carry separate titles and must be numbered in consecutive numerical order in the text and referred to in both the text and the caption as

Table 1, Table 2, etc. In addition to the physical quantity, e.g. t (in italics), units (normal text), should be added in square brackets. The tables should each have a heading. Tables should not duplicate data found elsewhere in the manuscript.

Acknowledgement of collaboration or preparation assistance may be included before References. Please note the source of funding for the research.

REFERENCES A reference list must be included using the following information as a guide. Only cited text references are included. Each reference is referred to in the text by a number enclosed in a square bracket (i.e., [3] or [2] to [6] for more references). No reference to the author is necessary.

References must be numbered and ordered according to where they are first mentioned in the paper, not alphabetically. All references must be complete and accurate. All non-English or. non-German titles must be translated into English with the added note (in language) at the end of reference. Examples follow.

Journal Papers: Surname 1, Initials, Surname 2, Initials (year). Title. Journal, volume, number, pages, DOI code.

[1] Hackenschmidt, R., Alber-Laukant, B., Rieg, F. (2010). Simulating nonlinear materials under centrifugal forces by using intelligent crosslinked simulations. Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 7-8, p. 531-538, DOI:10.5545/sv-jme.2011.013.

Journal titles should not be abbreviated. Note that journal title is set in italics.

Please add DOI code when available and link it to the web site.

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[2] Groover, M.P. (2007). Fundamentals of Modern Manufacturing. John Wiley & Sons, Hoboken.

Note that the title of the book is italicized.

Chapters in Books: Surname 1, Initials, Surname 2, Initials (year). Chapter title. Editor(s) of book, book title. Publisher, place of publication, pages.

[3] Carbone, G., Ceccarelli, M. (2005). Legged robotic systems. Kordić, V., Lazinica, A., Merdan, M. (Eds.), Cutting Edge Robotics. Pro literatur Verlag, Mammendorf, p. 553-576.

Proceedings Papers: Surname 1, Initials, Surname 2, Initials (year). Paper title. Proceedings title, pages.

[4] Štefanić, N., Martinčević-Mikić, S., Tošanović, N. (2009). Applied Lean System in Process Industry. MOTSP 2009 Conference Proceedings, p.

422-427.

Standards: Standard-Code (year). Title. Organisation. Place.

[5] ISO/DIS 16000-6.2:2002. Indoor Air – Part 6: Determination of Volatile Organic Compounds in Indoor and Chamber Air by Active Sampling on TENAX TA Sorbent, Thermal Desorption and Gas Chromatography using MSD/FID. International Organization for Standardization. Geneva. www pages: Surname, Initials or Company name. Title, from http://address, date of access.

[6] Rockwell Automation. Arena, from http://www.arenasimulation.com, accessed on 2009-09-07.

EXTENDED ABSTRACT By the time the paper is accepted for publishing, the authors are requested to send the extended abstract (approx. one A4 page or 3.500 to 4.000 characters). The instructions for writing the extended abstract are published on the web page http://www.sv-jme.eu/ information-for-authors/.

COPYRIGHT Authors submitting a manuscript do so on the understanding that the work has not been published before, is not being considered for publication elsewhere and has been read and approved by all authors. The submission of the manuscript by the authors means that the authors automatically agree to transfer copyright to SV-JME and when the manuscript is accepted for publication. All accepted manuscripts must be accompanied by a Copyright Transfer Agreement, which should be sent to the editor. The work should be original by the authors and not be published elsewhere in any language without the written consent of the publisher.

The proof will be sent to the author showing the final layout of the article. Proof correction must be minimal and fast. Thus it is essential that manuscripts are accurate when submitted.

Authors can track the status of their accepted articles on http://en.svjme.eu/.

PUBLICATION FEE For all articles authors will be asked to pay a publication fee prior to the article appearing in the journal. However, this fee only needs to be paid after the article has been accepted for publishing. The fee is 300.00 EUR (for articles with maximum of 10 pages), 20.00 EUR for each addition page.

Additional costs for a color page is 90.00 EUR.

59 (2013) 3

http://www.sv-jme.eu

Strojniški vestnik Journal of Mechanical Engineering

Since 1955

Papers

139

Matej Hudovernik, Daniel Staupendahl, Mohammad Gharbi, Matthias Hermes, A. Erman Tekkaya, Karl Kuzman, Janez Marko Slabe 3D Numerical Analysis of 2D Profile Bending with the Torque Superposed Spatial Bending Method

148

Sebastijan Jurendić, Silvia Gaiani Numerical Simulation of Cold Forming of α-Titanium Alloy Sheets

Andreas Schubert, Henning Zeidler, Matthias Hackert-Oschätzchen, Jörg Schneider, Martin Hahn Enhancing Micro-EDM using Ultrasonic Vibration and Approaches for Machining of Nonconducting Ceramics

165

Tomasz Tański Characteristics of Hard Coatings on AZ61 Magnesium Alloys

175

Leszek Adam Dobrzański, Małgorzata Musztyfaga, Aleksandra Drygała

156

Final Manufacturing Process of Front Side Metallisation on Silicon Solar Cells Using Conventional and Unconventional Techniques

183

Angel Fernández, Manuel Muniesa, Jaime González Characterisation and Processing of Reinforced PA 6 with Halloysite Nanotubes (HNT) for Injection Molding

193

Gašper Gantar, Andrej Glojek, Mitja Mori, Blaž Nardin, Mihael Sekavčnik Resource Efficient Injection Moulding with Low Environmental Impacts

Journal of Mechanical Engineering - Strojniški vestnik

Contents

3 year 2013 volume 59 no.