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59 (2013) 11

Since 1955

Papers

637

Fantina Rosa Esteves, Tiago Alexandre Carvalho, António Sérgio Pouzada, Carla Isabel Martins: The Influence of Processing on the Aesthetic, Morphological and Mechanical Properties of Structural Foam Mouldings of High-Impact Polystyrene

646

Jan Deckers, Jean-Pierre Kruth, Ludwig Cardon, Khuram Shahzad, Jef Vleugels: Densification and Geometrical Assessments of Alumina Parts Produced Through Indirect Selective Laser Sintering of Alumina-Polystyrene Composite Powder

662

Markus Gottfried Battisti, Walter Friesenbichler: Injection-Moulding Compounding of PP Polymer Nanocomposites

669

Kim Ragaert, Filip De Somer, Stieven Van de Velde, Joris Degrieck, Ludwig Cardon: Methods for Improved Flexural Mechanical Properties of 3D-Plotted PCL-Based Scaffolds for Heart Valve Tissue Engineering

677

Vito Speranza, Umberto Vietri, Roberto Pantani: Monitoring of Injection Moulding of Thermoplastics: Adopting Pressure Transducers to Estimate the Solidification History and the Shrinkage of Moulded Parts

683

Giovanni Lucchetta, Marco Fiorotto: Influence of Rapid Mould Temperature Variation on Appearance of Injection-Moulded Parts

689

Frederik Vogeler, Wesley Verheecke, André Voet, Hans Valkenaers: An Initial Study of Aerosol Jet® Printed Interconnections on Extrusion-Based 3D-Printed Substrates

697

Alejandra Costantino, Valeria Pettarin, Julio Viana, Antonio Pontes, Antonio Pouzada, Patricia Frontini: Polypropylene/Clay Nanocomposites Produced by Shear Controlled Orientation in Injection Moulding: Deformation and Fracture Properties

Journal of Mechanical Engineering - Strojniški vestnik

Contents

11 year 2013 volume 59 no.

Strojniški vestnik Journal of Mechanical Engineering


Strojniški vestnik – Journal of Mechanical Engineering (SV-JME) Aim and Scope The international journal publishes original and (mini)review articles covering the concepts of materials science, mechanics, kinematics, thermodynamics, energy and environment, mechatronics and robotics, fluid mechanics, tribology, cybernetics, industrial engineering and structural analysis. The journal follows new trends and progress proven practice in the mechanical engineering and also in the closely related sciences as are electrical, civil and process engineering, medicine, microbiology, ecology, agriculture, transport systems, aviation, and others, thus creating a unique forum for interdisciplinary or multidisciplinary dialogue. The international conferences selected papers are welcome for publishing as a special issue of SV-JME with invited co-editor(s). Editor in Chief Vincenc Butala University of Ljubljana Faculty of Mechanical Engineering, Slovenia Technical Editor Pika Škraba University of Ljubljana Faculty of Mechanical Engineering, Slovenia 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 440 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 Chamber of Commerce and Industry of Slovenia, Metal Processing Industry Association President of Publishing Council Branko Širok, UL, Faculty of Mech. Engineering, Slovenia Vice-President of Publishing Council Jože Balič, UM, Faculty of Mech. Engineering, Slovenia Cover: Microstructure of injection molded high impact polystyrene structural foams. This is a multilayer material consisting of two solid skins and a foamed core. The cell distribution at the core is dependent of the type of mold (steel or hybrid) and the processing conditions. The more efficient temperature control in the steel mold leads to better uniformity of the cell size which is much smaller than in hybrid moldings. Image Courtesy: Carla Martins, IPC/I3N – Institute for Polymer and Composites, University of Minho.

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 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 Research 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? 5. We kindly ask you to suggest at least two reviewers for your paper and give us their names and contact information (email). 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. - An Experimental section, which should provide details of the experimental set-up and the methods used for obtaining the results. - A Results section, which should clearly and concisely present the data using figures and tables where appropriate. - A Discussion section, which should describe the relationships and generalizations shown by the results and discuss the significance of the results making comparisons with previously published work. (It may be appropriate to combine the Results and Discussion sections into a single section to improve the clarity). - 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. Books: Surname 1, Initials, Surname 2, Initials (year). Title. Publisher, place of publication. [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.


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11 Contents

Contents Strojniški vestnik - Journal of Mechanical Engineering volume 59, (2013), number 11 Ljubljana, November 2013 ISSN 0039-2480 Published monthly

Guest Editorial 635

Papers Fantina Rosa Esteves, Tiago Alexandre Carvalho, António Sérgio Pouzada, Carla Isabel Martins: The Influence of Processing on the Aesthetic, Morphological and Mechanical Properties of Structural Foam Mouldings of High-Impact Polystyrene 637 Jan Deckers, Jean-Pierre Kruth, Ludwig Cardon, Khuram Shahzad, Jef Vleugels: Densification and Geometrical Assessments of Alumina Parts Produced Through Indirect Selective Laser Sintering of Alumina-Polystyrene Composite Powder 646 Markus Gottfried Battisti, Walter Friesenbichler: Injection-Moulding Compounding of PP Polymer Nanocomposites 662 Kim Ragaert, Filip De Somer, Stieven Van de Velde, Joris Degrieck, Ludwig Cardon: Methods for Improved Flexural Mechanical Properties of 3D-Plotted PCL-Based Scaffolds for Heart Valve Tissue Engineering 669 Vito Speranza, Umberto Vietri, Roberto Pantani: Monitoring of Injection Moulding of Thermoplastics: Adopting Pressure Transducers to Estimate the Solidification History and the Shrinkage of Moulded Parts 677 Giovanni Lucchetta, Marco Fiorotto: Influence of Rapid Mould Temperature Variation on Appearance of Injection-Moulded Parts 683 Frederik Vogeler, Wesley Verheecke, André Voet, Hans Valkenaers: An Initial Study of Aerosol Jet® Printed Interconnections on Extrusion-Based 3D-Printed Substrates 689 Alejandra Costantino, Valeria Pettarin, Julio Viana, Antonio Pontes, Antonio Pouzada, Patricia Frontini: Polypropylene/Clay Nanocomposites Produced by Shear Controlled Orientation in Injection Moulding: Deformation and Fracture Properties 697


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11 Guest Editorial

Guest Editorial Special Issue: Polymers and Moulds Innovations This is a special issue associated with PMI 2012, the biannual Polymers and Moulds Innovations conference, which was jointly organized by the Institute for Polymers and Composites of the University of Minho, Portugal, and the Centre for Polymer and Material Technologies of the Ghent University, Belgium. It includes 8 doubled peer reviewed extended and enhanced versions from the 60 papers accepted and presented at the PMI 2012 conference in areas of research ranging from emerging developments in plastics products to practical implementation of new processing concepts. The topics are: • Advances in Polymer Processing • Hybrid Moulds • Mould Simulation and Thermal Control • Polymers & Materials Innovations • Polymer Recycling • Additive Manufacturing • Rapid Prototyping & Tooling and • Trends in Product Development. • • • • • • • • •

Special attention was given to the following areas: the influence of processing conditions on the aesthetic, morphological and mechanical properties of SF mouldings of high-impact polystyrene (HIPS-SF), a powder metallurgy (PM) process to fabricate alumina parts through indirect selective laser sintering (SLS®) of alumina-polystyrene composite powder, a comparison of different compounding techniques and their influence on Young’s modulus for conventional, processing of polymer nanocomposites and for processing in the innovative Injection Moulding Compounder (PNC-IMC), two separate approaches for improving the flexibility of 3D plotted PCL scaffolds for creating heart valve leaflets, a study on a series of injection moulding tests using a general purpose PolyStyrene (PS) as well as an investigation into the effects of changing the holding pressure injection temperature and cavity dimensions on PS, developing innovative technology for the rapid heating and cooling of injection moulds and an analysis of the effect of rapid variations in mould temperature on the improvement of mouldings’ appearance in terms of gloss, research into the creation of AJP-manufactured interconnects on extrusion-based 3D printed substrates, investigation into the effect of distinct morphologies induced by shear controlled orientation in injection moulding (SCORIM) on the mechanical and fracture performance of polypropylene (PP) and PP/nanoclay.

We would like to take this opportunity to thank the editorial staff of Strojniški vestnik - Journal of Mechanical Engineering as well as the conference proceedings editors and scientific committee for their strong support and encouragement of this special issue. We would also like to express our appreciation to all the contributors and anonymous reviewers of the papers for their time and effort. Guest Editors Prof. António Sérgio Pouzada Prof. Ludwig Cardon Prof. Jože Balič 635


636


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 637-645 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.997 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-08-05 Accepted for publication: 2013-08-20

The Influence of Processing on the Aesthetic, Morphological and Mechanical Properties of Structural Foam Mouldings of HighImpact Polystyrene Esteves F.R. – Carvalho T.A. – Pouzada A.S. – Martins C.I. Fantina Rosa Esteves – Tiago Alexandre Carvalho – António Sérgio Pouzada – Carla Isabel Martins* Institute for Polymers and Composites/I3N, University of Minho, Portugal The production of large plastic parts, in small series and at low cost, requires the development of alternative tools and materials. Structural foams (SF) are an appropriate solution, when the production of thicker parts with superior properties is specified. They allow the production of lightweight parts with high stiffness and excellent dimensional stability, due to their sandwich-like structure consisting of a cellular core and two solid skins. The porous core is the result of the addition of a blowing agent in the polymeric matrix. These materials are applied in the urban furniture, automotive, nautical and aerospace industries. The most commonly used process to produce SF is low pressure foam moulding, which is a short-shot process, with impression pressures below 4 MPa. Therefore, the use of hybrid moulds with moulding blocks obtained by rapid prototyping routes is seen as a viable alternative for SF injection moulding. This work reports a study of the influence of processing conditions on the aesthetic, morphological, and mechanical properties of SF mouldings of high-impact polystyrene (HIPS-SF). The results compare the effect of the use of hybrid moulds to conventional moulds on the surface aspect of the circular centre-gated mouldings. Furthermore, the influence of the moulding temperature is ascertained in terms of the resulting cellular morphology, the flexural stiffness of the plate mouldings, and the impact resistance in the instrumented drop weight test. Keywords: structural foam, hybrid mould, skin thickness, mechanical strength, morphology

0 INTRODUCTION Structural foams (SF) are multi-layer materials consisting of two integral skins and a cellular core [1] and [2] This structure combines the lightness of the cellular core with the strength of dense unfoamed skins. An important characteristic of this sandwich structure is its high specific flexural stiffness/strength ratio. For this reason, the main use of this structure is in engineering and load-bearing applications [3]. One well-known use of the SF of glass-filled polypropylene (PP) was in the tanks of domestic washing machines [4]. Thermoplastic SF may be produced by different injection moulding techniques. Commonly, SF are manufactured using a low pressure short-shot process (typically 65 to 90% of the impression volume), in which an amount of melt containing dissolved gas (usually chemical blowing agents (CBA)) is injected to partially fill the impression [5]. The expansion of the blowing agent guaranties the fulfilment of the cavity. This process has advantages including the possibility of production of complex thick and large parts without sink marks; the development of low levels of internal stresses; a lower tendency to warpage and distortion; and the requirement of reduced clamping forces due to the low impression pressures, typically below 4 MPa [2] and [6]. Conversely, the cycle time is longer, the surface finish is poorer, there can be gas emission

during processing, and reprocessing is more complex [3]. The complex inter-relationship between structure properties and moulding conditions have been discussed in depth by Ahmadi et al. [5]. They concluded that the uniformity and fineness of the cellular core are greater when using high injection rates combined with low melt and mould temperatures. The degree and variability of irregularities across the surface of SF mouldings are highly dependent on the extent of flow into the moulding cavity and the processing conditions. Surface roughness is reduced mainly by increasing the shot weight that influences the degree of expansion and, therefore, the overall density of the moulding. Excessive foaming in the cavity results in considerable rupture at the melt flow front and, consequently, a lower surface finish. Increasing the mould temperature allows more time for relaxation of surface defects before cooling is completed; a slightly lower roughness is thus achieved [7]. It is difficult to control the cell size and cell size distribution [8] and [9]. The poor uniformity of cell distribution in mouldings injected with the low pressure process is largely attributed to the deficient mixing of the CBA and the polymer. The size and density of the cells, the skin thickness and the density profile across the thickness have a considerable effect on the mechanical properties of the resulting foams. CBA concentration and injection

*Corr. Author’s Address: Institute for Polymers and Composites/I3N, University of Minho, Campus de Azurem, Guimarães, Portugal, cmartins@dep.uminho.pt

637


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 637-645

pressure are the main factors influencing foam quality. The mould temperature has a significant effect on the skin thickness; back pressure also influences skin thickness, cell size and foam density. Finally, the mould and melt temperatures have negligible effects on the cell density [3]. The flexural stiffness and the impact strength increase with the moulding thickness. The flexural properties depend on whether the SF-layered structure is symmetric or asymmetric [10] and [11] On impact, the energy required for crack initiation and the total energy for failure are substantially reduced by the occurrence of microcells in the outer skins that act as stress raisers [5] and [7] Injection moulding of SF is a low pressure process, ideal for the production of large area parts. Therefore, it is a viable candidate for light moulding tools, as is the case of hybrid moulds [12]. In these tools, the moulding blocks are produced by rapid prototyping and tooling (RPT) with routes such as the vacuum casting of epoxy composites. Typically, these blocks are produced in a composite of epoxy with aluminium powder, which has the advantages of easy manufacturing and short delivery time [13]. The main disadvantages, when compared with conventional steel moulds, are worse thermal and mechanical behaviours, shorter useful lifetimes of the moulding blocks and longer cycle times, which means that this processing option is recommended for a limited number of parts [12]. Acrylonitrile butadiene styrene (ABS) is known for displaying a well-balanced set of properties that ensure correct dimensional stability, excellent surface finish, superior impact strength and metallization characteristics [14]. In the form of SF, applications including furniture, loudspeaker boxes, telephone junction boxes, sprinkler housings, television housings, other housings, or automotive back seats [15]. High-impact polystyrene (HIPS) features many of the ABS properties at a lower price and has a similar processing character. This study focuses on the morphological and mechanical properties of HIPS-SF injection mouldings. The flexural and impact behaviours of HIPS-SF are related to its morphological and physical parameters, such as skin ratio and density. Finally, the experimental results of the flexural stiffness are compared to models proposed by Barzeraghi et al. [16]. 638

1 EXPERIMENTAL 1.1 Raw Materials High impact polystyrene (HIPS) from BASF, Korea, with a specific gravity of 1.05 Mg·m–3 and MFI of 10.95 g / 10 min (200 ºC / 5 kg) was used together with 2 wt% of an endothermic CBA, Tracel IMC 4200SP, from Tramaco, Germany, with a decomposition temperature in the range of 160 to 220 ºC, for the production of HIPS-SF. 1.2 Injection Moulding Centred gated discs of 155 mm in diameter and 5 mm in thickness were processed with an Engel Victory Spex 50 injection moulding machine (Engel, Austria). The mouldings were produced using two mould material combinations for the moulding blocks (core and cavity): a conventional steel mould configuration and a hybrid mould as depicted in Fig. 1. The hybrid mould has an insulating plate (resin filled with glass fibres) in the injection side and a moulding block in the ejection side, made in a composite of epoxy Biresin L74 filled with 60 wt% aluminium powder. This moulding block was produced by vacuum casting and machined to the final geometry of the part, following the standard routine described in [17]. The mouldings were injected with 90% of the mould filling, and the complete filling was promoted by the CBA expansion. Various processing conditions were used, as shown in Table 1.

b)

a)

c)

d)

Fig. 1. a) mould structure, b) moulding, c) steel moulding block, d) hybrid moulding block

Esteves F.R. – Carvalho T.A. – Pouzada A.S. – Martins C.I.


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 637-645

Table 1. Injection moulding processing conditions Parameter

Steel

Tinj

Hybrid 200 220 240

Core

T1

45

75

Cavity

T2 tcool Q

50 180

75 80

Injection temperature [ºC] Mould temperature [ºC] Cooling time [s] Flow rate [cm3/s]

the ratio between the sum of the thickness of the two skin layers and the overall thickness of the samples, expressed as a percentage.

220

60

The moulding code is an alphanumeric value, e.g., H200, where the letter means that the material was injected in the hybrid mould (H) or in the steel mould (S), and the number indicates the injection temperature, in this example 200 ºC. 1.3 Mould Monitoring The two halves of the mould were instrumented with Resitec type J, TC.002 thermo-couples (Resitec, Portugal): one was placed in the injection side at 2 mm from the moulding surface and at 30 mm from the gate (T1), and the other in the ejection side, at the same position (T2) (Fig. 2). The signals were acquired and recorded with a Priamus Multi DAQ 8101A data acquisition system.

Fig. 2. Hybrid mould configuration and setup of thermocouples: 1-steel injection plate, 2-insulating plate, 3-steel ejection plate, 4-moulding block (resin or steel)

These thermocouples were placed in these positions to allow the monitoring of the mould temperature, and to ensure the symmetric development of the moulding structure [10]. 1.4 Morphological Analysis The effect of the processing conditions on the microstructural development of the HIPS-SF parts was assessed using an Olympus BH-2 optical microscope (Olympus, Japan) coupled with a Leica DFC 280 digital camera (Leica, Germany). Thin slices of about 10 µm were cut from samples using a microtome (Leitz, Germany) at 30 mm from the gate as shown in Fig. 3. The slices were placed between a glass slide and cover glass after immersion in Canada balsam. Measurements of the skin ratio were made throughout the section. The skin ratio is defined as

Fig. 3. Sample location for the morphological analysis; insert refers to the slices made along the flow path (S - skin, C - core)

To verify the existence of crushed cells and/or cells of small dimension in the skin, an observation was performed in a NOVA 200 Nano SEM ultrahigh resolution field emission gun scanning electron microscope (FEG-SEM), (FEI, USA). The samples were cryo-fractured, sputter-coated with gold, and fixed in a support with carbon tape adhesive. 1.5 Gloss The gloss measurements were carried with a BYKGardner micro-TRI-glossmeter (BYK, USA) calibrated with a black glass standard according to the ASTM D523-85 standard. A minimum of five specimens for each condition and at three different locations in the sample were used to determine the average gloss and respective standard deviation. For each type of surface, the measuring angle was 85º according to the ASTM standard. The gloss ratings measured by this test method are obtained by comparing the specular reflectance from the specimen to the black glass standard. The aim of gloss measurements was to estimate how the type of the mould (steel moulding block, hybrid moulding block) and processing conditions influence the gloss of injected parts. 1.6 Roughness The topography of the surfaces was assessed with a prototype laser microtopographer [18]. This system does not make contact with the surface during measurements, thereby avoiding any damage. It is based on optical active triangulation with oblique incidence and normal (and/or specular) observation, and mechanical scanning of the sample. The arithmetical average value of roughness, Ra, was

The Influence of Processing on the Aesthetic, Morphological and Mechanical Properties of Structural Foam Mouldings of High-Impact Polystyrene

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 637-645

determined. Five measurements for each condition were carried, and each sample was measured in five different locations. 1.7 Density The density was measured using the impulsion method (Archimedes principle) according to the ASTM standard D 792-00. A Scaltec SBC 31 analytical balance (Denver Instrument, Germany) with capacity of 220 g and accuracy of 0.0001 g was used. A density measurement kit was used together with the balance, and the liquid of reference was isopropanol. Five measurements were performed for each condition, and an average density value was calculated. 1.8 Flexural Testing The flexural tests were performed at room temperature with the three-point support Pouzada and Stevens test method [19]. The testing apparatus was mounted in an Instron 4505 universal testing machine (Instron, USA), in compression mode. The samples were placed on the support base, and the load was applied at the centre of sample using a crosshead speed of 5 mm/ min. A maximum displacement of 5 mm was imposed, guaranteeing that the sample behaved as a plate in the elastic range (Fig. 4). The flexural stiffness data hereafter are the average of five tests.

contraction ratio (ν), the flexural stiffness is defined as [19]: C=

E . (1) (1 −ν 2 )

For small deflections of a circular disc, the flexural stiffness, C, can be analytically expressed in terms of the slope of the F/δ trace and the geometrical parameters of the flexural test. In this case, it is applicable the Bassali equation:

C=

3 R 2 B (v ) S0 , 4π h3

(2)

where h is the sample thickness, R is the radius of the 3-point support circumference (93.5 mm), B(ν) is a function of Poisson’s ratio, which in the range of 0.3 to 0.45 has an average value of 5, for HIPS. The slope, S0, is the corrected slope of the load versus displacement curve when the support points are not on the periphery of the disc. In the cases of samples overhanging the supports, it is necessary to consider the difference between the radius of the sample and the radius of the support circumference. Given the overhang length ΔR, the support diameter D, and the measured slope, S, the corrected slope, S0, is calculated as [19]:

S0 =

S 4.1∆R −   0.59  1 − e D  + 1  

. (3)

1.9 Impact Test The impact tests were performed at room temperature with the CEAST 9350 Fractovis Plus (CEAST, Italy) using the following setup: impact weight of 15.765 kg and drop height of 700 mm, leading to an impact speed of 3.705 m·s-1. The tests were performed according to the European Standard EN ISO 6603-1. 2 RESULTS AND DISCUSSION 2.1 Mould Monitoring

Fig. 4. Schematic of the three-point support flexural test [19]

For isotropic materials that can be mechanically characterized in terms of a modulus (E) and a lateral 640

The mould monitoring data of the hybrid and steel moulds are depicted as examples in Figs. 5 and 6, respectively. The temperatures of the two sides of the mould are different when the hybrid mould is used, whereas in the steel mould they are similar. In both cases, the injection side of the moulding is always hotter due to the proximity of the injection nozzle. An increase

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of temperature after the injection of the material is observed. The temperature starts decaying slowly until the cooling of the sample is completed.

Fig. 5. Evolution of temperature in the core (T1) and cavity (T2) sides of the hybrid mould during the injection cycle of HIPS-SF, for Tinj = 220 ºC The temperatures in the hybrid mould (T1 and T2) have a significant difference when the cycle starts and during cooling. This results from the different thermal conductivities of the mould materials used in the core and cavity, which make the cooling rate more difficult to control in the two sides of the mould. In the steel mould, the temperature in the core and cavity side is similar because the mould materials have the same thermal properties.

Fig. 6. Evolution of temperature in the core (T1) and cavity (T2) sides of the steel mould during the injection cycle of HIPS-SF, for Tinj = 220 °C

2.2 Morphological Characterization The typical structure obtained in the HIPS-SF injection moulding is depicted in Fig. 7. The structure is characterized by two outer solid layers (skin) and a cellular core. Depending on the type of mould

used, different microstructures are developed. With the hybrid mould, there is the development of an asymmetric sandwich structure (Figs. 7a to c), and in the case of the steel mould the structure is symmetric (Fig 7d). This may be related to the variation of the mould temperature during the injection cycle, as shown in Figs. 5 and 6.

a)

b)

c) d) Fig. 7. Polarized light microscopy: influence of the injection temperature on the microstructure of HIPS-SF at a) H200, b) H220, c) H240, d) S220

The mouldings produced in the hybrid mould also have larger cells with a variety of shapes, as compared to the steel mould (Fig. 8). At the core, the cells of the hybrid mouldings are approximately 200 μm in diameter, whereas the cells of the steel mouldings are only 80  μm. The cells dimensions are decreasing from the centre to the skin due to the difference in the melt temperature and are distorted as a result of the fountain flow. With the increase of the injection temperature, this phenomenon becomes more pronounced. Furthermore, the growth of the cell size is evident with the increase of the temperature, due to the lower viscosity and less resistance to the cell growth [10]. It is confirmed in the SEM images the presence of crushed cells and nanopores in the skin (Fig. 9), which affects the mechanical properties, mainly in impact, as is shown next. The crushed cells results from the cell growth at the core that are in development and exert pressure over those close to the wall.

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2.3 Aestethics Analysis The surface roughness and, thus, the gloss of plastics parts are key characteristics that affect the moulding aesthetics. The gloss and roughness (Ra) of the moulding blocks and plastics mouldings are summarised in Table 2 and Table 3, respectively. Table 2. Roughness and gloss of the moulding blocks Mould Ra [μm] Hybrid moulding block 6.96 (1.32) Steel moulding block 1.81 (0.40) Standard deviation values are in parentheses.

Gloss [%] 9.07 (2.24) 59.52 (4.74)

Table 3. Effect of the moulding parameters on the gloss and roughness of HIPS-SF

a)

HIPS-SF HIPS Ra [μm] Gloss [%] Ra [μm] Gloss [%] H200 1.02 (0.24) 1.69 (0.11) 1.07 (0.19) 1.79 (0.15) H220 1.10 (0.06) 1.56 (0.16) 2.22 (0.22) 0.98 (0.07) H240 1.99 (0.38) 0.93 (0.24) 3.70 (0.78) 0.56 (0.17) S220 0.89 (0.11) 14.74 (2.74) 0.60 (0.17) 47.92 (2.56) Standard deviation values are in parentheses.

b) Fig. 8. SEM view of the core of the moulding injected in: a) hybrid mould, b) steel mould at 220 ºC

The surface finish of the moulding blocks and the processing conditions affect the gloss of the corresponding surface of the moulded parts. The hybrid mould blocks have Ra higher than the steel blocks. Consequently, the parts produced in the hybrid mould have higher Ra and display lower gloss due to the replication of the surface roughness of the mould in the plastic part [20]. The roughness increases with the increasing injection temperature. This may result from the lower viscosity of the material enabling better replication of the rough surfaces [20] and [21]. HIPS-SF presents lower Ra and higher gloss because the pressure required to fill the impression with the expansion of the blowing agent is lower than in conventional mouldings of HIPS [6]. Thus, the replication with HIPS-SF is less accurate than non-foamed HIPS. 2.4 Flexural Behaviour

Fig. 9. SEM view of the skin of HIPS-SF (H220) with crushed

cells and nanopores

642

The mechanical properties are dependent on the morphological characteristics, such as density profile and skin ratio [3]. Table 4 shows the flexural stiffness, skin ratio and density data of HIPS-SF for various processing conditions. The data presents an average variation of 5, 9 and 1%, respectively, for the aforementioned properties.

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Table 4. Flexural stiffness, skin ratio and density of HIPS-SF, for various processing conditions H200 H220 H240 S220

Flexural stiffness [MPa] Skin ratio [%] SF 100% SF 2277.36 2614.40 44.86 (119.47) (164.92) (3.7) 2236.76 2568.45 49.36 (88.49) (14.25) (5.6) 2132.03 2542.27 43.73 (135.77) (51.00) (7.5) 2081.14 2326.56 29.55 (190.00) (93.82) (2.6)

A typical fracture of an HIPS-SF moulding is shown in Fig 11.

Density [Mg·m-3] SF 100% 0.94 1.028 (0.014) 0.89 1.029 (0.005) 0.88 1.027 (0.002) 0.86 1.028 (0.004)

Standard deviation values are in parentheses.

In general, increasing the injection temperature decreases the skin ratio and the density. The density, being related with the amount and distribution of material, corresponds to the distribution of the applied loads [22]; therefore, as it decreases, the flexural stiffness also decreases. In non-foamed mouldings, the material available to support the flexural loads is greater than in foamed mouldings, which have a porous core; therefore, its stiffness is higher [11]. The HIPS-SF mouldings in the hybrid mould are slightly stiffer than in the steel mould, due to the larger skin ratio and higher density.

Fig. 11. Fractured HIPS-SF mouldings

The observation of the failed mouldings suggests a ductile fracture, as the crack did not propagate along the radial direction. The effect of the injection temperature on the peak energy is shown in Fig. 12.

2.5 Impact Behaviour A typical drop weight impact test graph for HIPS-SF is depicted in Fig. 10. The peak force, Fp, observed corresponds to the moment the material yields and enters the plastic regime. From this point onward, it is not possible to recover the strain and thus the moulding is no longer useful for service. The area below the first peak corresponds to the energy absorbed up to this moment (peak energy, Up).

Fig. 12. Falling weight impact peak energy of HIPS-SF

As with the flexural properties of HIPS-SF, the resistance to the impact of falling weights is also influenced by the density and the skin thickness. In addition, defects on or near the surface are thought to have a significant role in determining the peak energy. Raising the injection temperature decreases the peak energy, in accordance to the density reduction. The occurrence of microcells near the outer skins may affect the integrity of these outer layers, acting as stress raisers, thus decreasing the impact resistance. The resistance to crack propagation is worsened by the formation of large non-uniform cells at the core of the mouldings, becoming increasingly evident at higher melt temperatures [7]. 2.6 Model Prediction

Fig. 10. Typical output obtained in falling weight tests of HIPS-SF mouldings

There are analytical models for predicting the mechanical properties once the physical and

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morphological characteristics are known. Barzegari et al. proposed a model to predict the flexural behaviour starting from approximations to the density profile [16]. To simplify notations, they used normalized parameters, as relative density R, and relative thickness, r:

ρc = R1 , ρs

ρf = R2 , ρs

Ef

(4)

δc δs δ = r, = 1 − c = 1 − r , (5) δf δf δf

In Model d), the cross-section of Fig. 13d is decomposed into a three-layered structure. It assumes that the skin density decreases linearly to the core density. An intermediate layer, t, is considered, and the core density is uniform. Thus, the normalized flexural modulus is:

where ρf, ρs, ρc are densities of foam, skin and core, respectively, and δf, δs, δc are thicknesses of foam, skin and core, respectively. Some of those density variation approaches are shown in Fig. 6. These cross-sections represent an approximation of the density profile across the thickness. In this study, the models of Fig. 13 c and d are analysed.

Es

(

)

(

)

= 1 − r 3 1 − R12 − r 2t 2 − R1 − R12 − 3 R 2  t3 −rt 2  − R1 − 1  − (4 − 3R1 − R12 ). (9) 2  10 2

Upon using these models, the predicted flexural stiffness can be calculated. The predictions using Model d) are not too far from the experimental results, with a maximum error of 9%, whereas with Model c), the maximum error is about 14%, as shown in the Table 5. Table 5. Prediction of flexural stiffness Experimental [MPa] H200 H220 H240 S220

2277 2237 2132 2081

CModel c) [MPa] 2582 2529 2490 2235

Predictions ΔCc) CModel d) [MPa] [%] 2489 11.78 2422 11.54 2350 14.38 1987 6.88

ΔCd) [%] 8.51 7.64 9.27 4.51

3 CONCLUSIONS

Fig. 13. Different approaches of density profiles for SF [16]

The core density was calculated using the following equations:

 ρc   ρ f δ st   δ st −  1 −   =   ρs   ρs δ f   δ f

−1

  , (6) 

δ st = δ s1 + δ s 2 , (7)

where δst is the total skin thickness. Model c) assumes that the core density reaches a minimum at the centre of the beam with a linear variation in the core part. Thus, the normalised flexural modulus is obtained by: 644

Ef Es

= 1−

r3 (4 − 3R1 − R1 ) 2 . (8) 10

The processing conditions influence the aesthetic, morphological and mechanical properties of HIPS-SF mouldings injected in hybrid moulds and conventional steel moulds. The mouldings produced with hybrid moulds have reduced gloss due to their higher roughness. The injection temperature has influence on the skin ratio and shape of the cells. With an increase of the injection temperature, there is an increase of the cell size due to the lower viscosity and a decrease of the skin ratio and density, which causes the flexural stiffness also to decrease. The mouldings subjected to impact show a ductile behaviour, and the peak energy in the falling weight test decreases with the increasing injection temperature. The mechanical behaviour of HIPS-SF mouldings can be predicted using analytical models based on morphological properties. The predictions are close to the experimental data, with errors less than 10% with respect to the best model.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 637-645

4 ACKNOWLEDGEMENT This work was developed with the support of the Portuguese Foundation for Science and Technology under the strategic project PEst-C/CTM/ LA0025/2013. The authors acknowledge the support of the Portuguese program QREN, which funded contract 2010/013307 for project ‘Hybridmould 21’. 5 REFERENCES [1] Malloy, R.A. (1994). Plastic Part Design for Injection Molding: An Introduction, Hanser Publishers, New York. [2] Rosato, D.V., Rosato, D.V., Rosato, M.G. (2000). Injection Molding Handbook, 3rd ed. Kluver Academic Publishers, Boston/Dordrecht/London, DOI:10.1007/978-1-4615-4597-2. [3] Barzegari, M.R., Rodrigue, D. (2009). The effect of injection molding conditions on the morphology of polymer structural foams. Polymer Engineering and Science, vol.49, no. 5, p. 949-959, DOI:10.1002/ pen.21283. [4] Oghoubian, R., Smart, J. (1990). Out-of-plane bending of faceted cylinder end plates. The Journal of Strain Analysis for Engineering Design,. vol. 25, no. 2, p. 95101, DOI:10.1243/03093247V252095. [5] Ahmadi, A.A., Hornsby, P.R. (1984) Moulding and Characterization studies with polypropylene structural foam. I. Structure-property interrelationships. Plastics and Rubber Processing and Applications, vol. 5, no. 1, p. 35-49. [6] Nogueira, A.A., Martinho, P.G., Brito, A.M., Pouzada, A.S, (2011). A study on the mouldability of technical parts using hybrid moulds and structural foams, Bártolo, P.J. (ed.): Innovative Developments in Virtual and Physical Prototyping, CRC Press/Balkema, London, p. 399-404. [7] Ahmadi, A.A., Hornsby, P.R. (1985). Moulding and characterization studies with polypropylene structural foam. II. The influence of processing conditions on structure and properties. Plastics and Rubber Processing and Applications, vol. 5, no. 1, p. 51-59. [8] Mark, J.E., (ed.) (2007). Physical properties of polymers handbook. 2nd ed., Springer Science + Business Media, New York. [9] Kamal, M.R., Isayev, A.I., Liu, S.-J. (2009). Injection Molding - Technology and Fundamentals, White J.L. (ed.) Carl Hanser Verlag, Munich, DOI:10.3139/9783446433731. [10] Tovar-Cisneros, C., González-Núñez, R., Rodrigue, D. (2008). Effect of mold temperature on morphology

and mechanical properties of injection molded HDPE structural foams. Journal of Cellular Plastics, vol. 44, no. 3, p. 223-237. [11] Barzegari, M.R., Rodrigue, D. (2009). Flexural behavior of asymmetric structural foams. Journal of Applied Polymer Science, vol. 113, no. 5 p. 3103-3112, DOI:10.1002/app.30335. [12] Pouzada, A.S. (2009). Hybrid moulds: a case of integration of alternative materials and rapid prototyping for tooling. Virtual and Physical Prototyping, vol. 4, no. 4, p. 195-202, DOI:10.1080/17452750903438676. [13] Bareta, D.R., Pouzada, A.S., Costa, C.A. (2007). The effect of rapid tooling materials on mechanical properties of tubular mouldings. International Conference on Polymers & Moulds Innovations. Ghent. [14] Campo, E.A. (2006). The Complete Part Design Handbook for Injection Molding of Thermoplastics. Hanser, Munich. [15] Hough, M., Dolbey, R. (1995). The Plastics Compendium, Rapra Technology, Shrewsbury. [16] Barzegari, M.R., Rodrigue, D. (2007). The effect of density profile on the flexural properties of structural foams. Polymer Engineering & Science, vol. 47, no.9, p. 1459-1468. [17] Vasconcelos, P.V., Jorge Lino, F., Neto, R.J. (2004). Importance of the vacuum in rapid tooling of polymeric-based moulds. International conference on Rapid Product Development, Marinha Grande. [18] Ferreira, E.C., Costa, M. F., Laranjeira, C.R., Oliveira, M.J, Pouzada, A.S. (2004). Comparative study, by optical techniques of the interface polymer/steel in replication conditions. Advanced Materials Forum II, vol. 455-456, p. 467-471, DOI:10.4028/www. scientific.net/MSF.455-456.467. [19] Pouzada, A.S., Stevens, M.J. (1984). Methods of generating flexural design data for injection moulded plates. Plastics and Rubber Processing and Applications, vol. 4, no. 2, p. 181-187. [20] Oliveira, M.J., Brito, A.M., Costa, M.C., Costa, M.F. (2006). Gloss and surface topography of ABS: A study on the influence of the injection molding parameters. Polymer Engineering & Science, vol. 46, no. 10, p. 1394-1401, DOI:10.1002/pen.20607. [21] Vasconcelos, P.V., Lino, F.J., Neto, R.J., Paiva, R. (2006). Design epoxy resins based composites for rapid tooling applications. 5th International Conference on Mechanics and Materials in Design, Porto. [22] Lanz, R.W., Melkote, S.N., Kotnis, M.A. (2002). Machinability of rapid tooling composite board. Journal of Materials Processing Technology, vol. 127, no. 2, p. 242-245, DOI:10.1016/S0924-0136(02)001504.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 646-661 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.998 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-04-19 Accepted for publication: 2013-05-06

Densification and Geometrical Assessments of Alumina Parts Produced Through Indirect Selective Laser Sintering of Alumina-Polystyrene Composite Powder Deckers, J. – Kruth, J.-P. – Cardon, L. – Shahzad, K. – Vleugels, J. Jan Deckers1,* – Jean-Pierre Kruth1 – Ludwig Cardon2 – Khuram Shahzad3 – Jef Vleugels3 1 Catholic

University of Leuven, Department of Mechanical Engineering, Belgium College Ghent, Associated Faculty of Applied Engineering Sciences, Belgium 3 Catholic University of Leuven, Department of Metallurgy and Materials Engineering, Belgium 2 University

A powder metallurgy (PM) process to fabricate alumina parts through indirect selective laser sintering (SLS®) of alumina-polystyrene composite powder particles is presented. The PM process includes powder production through dispersion polymerisation, SLS®, debinding and solidstate sintering. Dimensional changes, which occur during the debinding and solid-state sintering, are assessed. Warm isostatic pressing (WIP) and both pressureless and pressure infiltration are introduced as extra steps of the PM process chain. The influence of WIP and the infiltration steps on the changes in density, geometry and microstructure during the PM process are investigated. Keywords: additive manufacturing, ceramic, indirect selective laser sintering, alumina, polystyrene

0 INTRODUCTION 0.1 AM of Polystyrene Polystyrene is one of the most popular polymers used in additive manufacturing (AM) technology. It has been used to demonstrate newly developed AM technologies, such as layered electro-photographic printing by Cormier et al. [1], selective inhibition sintering by Khoshnevis et al. [2] or various kinds of laminated object manufacturing (LOM) technologies (Brooks and Aitchison [3], de Smit and Broek [4], or Mahale et al. [5]). Moreover, polystyrene has been used to demonstrate that AM technologies can improve the investment-casting (IC) process by reducing tooling costs and production lead-times (Cheah et al. [6]). The combined use of AM technology and investment-casting technology is called ‘rapid investment casting’. As illustrated by Cheah et al. [6], two main application areas exist: 1) AM technology can be applied to produce inserts to injection mould polystyrene parts, as illustrated by Kinsella et al. [7]; 2) AM technologies can be applied to produce polystyrene IC patterns. These patterns can be master patterns (e.g. for silicone rubber moulding), but are mostly sacrificial patterns. Amorphous polystyrene is more suitable than other (semi-crystalline) polymer materials for the production of sacrificial rapid investment casting patterns due to its geometrical stability during the burning out step of the IC process (Kruth et al. [8]). This stability results from the polystyrene patterns’ porosity and low thermal expansion that prevents breaking of the (ceramic) IC mould during burn out. 646

The polystyrene sacrificial patterns are sometimes produced through three-dimensional printing (3DP) (Levy et al. [9]), but mainly through SLS®. In order to increase the polystyrene pattern’s strength, AM of high quality polystyrenes is investigated (e.g. SLS® of high impact polystyrene by Yang et al. [10]) and/ or wax infiltration of the patterns is applied (Ku et al. [11]). Wax infiltration of the polystyrene parts can also seal surface porosities (Cheah et al. [6]). A distinction must be made between the production of metal and ceramic parts through rapid investment casting with (polystyrene) sacrificial patterns. When producing metal parts, the sacrificial polystyrene patterns have the shape of the part to be produced. From the sacrificial pattern, a plaster mould (see Liu et al. [12] or Niino and Yamada [13]), but generally a ceramic moulding shell is sometimes fabricated. Finally, the moulds are used to fabricate the metal parts through a casting process, e.g. vacuum pressure casting of aluminium parts, as applied by Hongjun et al. [14]. Applications of this technology can be found in the production of titanium, aluminium, steel alloys or super alloys for competitive motorsports (Cevolinni et al. [15]). When producing ceramic parts, the sacrificial polystyrene patterns have the negative geometry of the parts to be produced. Through high pressure slip casting, followed by debinding of the polystyrene and a furnace sintering treatment, Si3N4 parts can be obtained (Pfeifer et al. [16]). 0.2 AM of Ceramics through Indirect SLS® of Dry Composite Powders Selective laser sintering (SLS®) is one of the additive manufacturing processes capable of producing

*Corr. Author’s Address: Department of Mechanical Engineering, Celestijnenlaan 300B, 3001 Heverlee, Belgium, Jan.Deckers@mech.kuleuven.be


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macroscopic ceramic parts. SLS® of ceramic components can be done directly or indirectly. Indirect selective laser sintering, the AM technology that is applied in this paper, involves the melting of a sacrificial organic binder phase to produce ‘green parts’, i.e. parts consisting of a binder phase that holds the ceramic particles together. If the binder is organic, it can be removed after the laser sintering step. In this way a ‘brown part’ is obtained. In the last step, the density and strength of the brown part is improved by solid-state sintering (SSS) in a furnace or through another densification processes. Direct selective laser sintering does not involve a sacrificial binder phase and the ceramic material is directly sintered or melted to produce parts (see Kruth et al. [17], Dewidar et al. [18]). Indirect SLS® makes use of either dry powders, deposited by, for example, a counter-rolling roller, or powder-containing slurries that are dried after deposition before laser scanning (Tang et al. [19]). In this paper, indirect SLS® of dry powders is investigated. The counter-rolling roller of a DTM Sinterstation 2000 is used to deposit the produced composite powder. Different types of binders have already been examined to fabricate many ceramic parts via indirect SLS® (Fig. 1). If the binder is organic, it cannot be burned. During the thermal treatment, the inorganic binder (e.g. HBO2) chemically reacts and becomes part of the structural ceramic (e.g. B2O3) (Lee [20] and [21]). Different types of organic binders have already been examined to fabricate ceramic parts via SLS®: • waxes: e.g. stearic acid (Liu et al. [22] and Leu et al. [23]),

thermosets such as phenolic resin (Liu et al. [24]), epoxy resin (e.g. Evans [25] and Stevinson et al. [26]) and others (Agarwala et al. [27]), • thermoplastics: e.g. PMMA (Subramanian et al. [28]) or an acrylic binder (Goodridge et al. [29]). A combination of binders is sometimes used, e.g. a thermoset in combination with the semi-crystalline PA-6 (or nylon 6) to produce graphite (Chakravarthy and Bourell [30]), or a wax in combination with the amorphous thermoplast PMMA to produce the composite ceramic Al2O3-ZrO2-TiC (Bai et al. [31]). In this study, an amorphous thermoplastic polystyrene is chosen to be the binder. Polystyrene is the most commonly used non-semi-crystalline material for SLS®. As already explained in the previous section, it is more suitable than other (semicrystalline) polymer materials for the production of sacrificial rapid IC patterns due to its geometrical stability during the burning-out step of the IC process (Kruth et al. [8]). It is also believed that the parts will benefit from the geometrical stability of polystyrene during the burning-out step (i.e. the debinding step) of indirect laser-sintered alumina-polystyrene parts. Although amorphous thermoplastic polystyrene was already used by Zheng et al. [32] to produce composite alumina-polystyrene parts through SLS®, it has not yet been used to produce pure alumina parts by indirect SLS®. 0.3 Densification Strategies A drawback of producing ceramic parts through indirect SLS® of dry composite powders is the low density of the parts after the SSS step. However, Binder material

Ceramic material

Inorganic

Organic Wax

Thermoset

Thermoplast

       

PMMA [28]     PMMA [31]

Al2O3 Al2O3 Al2O3-B2O3 Al2O3-glass-B2O3 Al2O3-ZrO2-TiC

  HBO2 [21] HBO2 [20]  

stearic acid [22]     unspecified [31]

Apatite-mullite

unspecified acrylic binder [29]

Graphite

phenolic resin [30]

nylon 11 [30]

K2O-Al2O3-SiO2

epoxy resin [24]

SiO2

unspecified [27]

SiC

phenolic resin [25], [26]

ZrO2

 

stearic acid [23]

unspecified [27]

 

ZrB2

Fig. 1. Sacrificial binders used to produce different ceramic parts by the use of a conventional SLS® system Densification and Geometrical Assessments of Alumina Parts Produced Through Indirect Selective Laser Sintering of Alumina-Polystyrene Composite Powder

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densification strategies can be used to further increase the density of the produced parts after SSS. Isostatic pressing techniques (e.g. Deckers et al. [33]) and infiltration methods (e.g. Stevinson et al. [26]) have already been successfully used to increase the final density of parts, produced through indirect SLS®. In this study, warm isostatic pressing of the green laser sintered parts and infiltration (i.e. impregnation of the parts with an ethanol-alumina suspension) have been applied as extra steps of the process to increase the final densities. 0.3.1 Warm Isostatic Pressing (WIP) Different isostatic pressing (IP) techniques exist (all of which differ in the sense that other pressure transferring media are used): hot isostatic pressing (HIP; a heated gas), quasi-isostatic pressing (QIP; powder), cold isostatic pressing (CIP; a liquid at room temperature) and warm isostatic pressing (WIP; a heated liquid). WIP, the IP technique used in this paper, is normally used to produce homogeneous and high density (up to 60%) green powder compacts with increased shape complexity by applying pressure from multiple directions at elevated temperature. During the WIP process, the powder is vacuum packed and immersed in a heated liquid that transmits the pressure uniformly to the powder (Fig. 2). Although WIP has already been used to produce metal (Yang et al. [34]) and ceramic (Galusek et al. [35]) parts, the combination of WIP and indirect SLS® is new (Deckers et al. [36]). During the WIP of SLSed samples, care should be taken when densifying complex geometries with internal cavities, since these cavities might collapse during the pressing process.

Fig. 2. Schematic of Warm Isostatic Pressing (WIP)

0.3.2 Infiltration Besides WIP, the density of the parts can be improved by impregnating them with an ethanol-alumina suspension. Infiltrating SLSed parts is not new. Subramanian et al. [28] reported that green part infiltration with small quantities of alumina colloids 648

largely improves the green part strength during debinding and solid-state sintering. 1 EXPERIMENTAL Fig. 3 schematically presents the main steps of the powder metallurgy process, assessed in this work, to produce alumina parts through AM. In a first step, the composite starting powder was produced. Afterwards, the SLS® parameters were optimized to produce green parts. The final alumina parts were produced by subsequently de-binding (deb.) and SSS of the green parts. Geometrical assessments were used to assess the dimensional changes of the SLSed parts. These changes occurred during the debinding and SSS step. In order to improve the final density of the alumina parts, two possible densification treatments were used: warm isostatic pressing and infiltration (inf.).

Fig. 3. Main steps of the powder metallurgy (PM) process

This paper investigates the quality of the components during the different processing steps through density measurements, geometrical assessment and microscopic imaging. The density was measured with the Archimedes method (Analytical Balance, Sartorius, Germany). The geometrical assessments were realized with a coordinatemeasuring machine (CMM, FN905, Mitutoyo, Japan) or a vernier caliper (Mitutoyo, Japan). The roughness was measured with a Talysurf-120L roughness measuring device (Taylor-Hobson, UK). The microscopic images were assessed with a digital camera, 3D microscopy (Discovery.V20, Carl Zeiss Inc., Germany) or scanning electron microscopy (SEM, XL30 FEG, FEI, The Netherlands). The outer shapes of the parts were captured with a digital camera. Internal, cross-sectional images were taken with the 3D microscope or SEM. In order to take the cross-sectional images, the parts were cut with a diamond blade, embedded in an epoxy resin, and ground. Secondary (SE-SEM), and backscattered electron (BSE-SEM) SEM images were taken.

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1.1 Production of Alumina Parts 1.1.1 Powder Production Zheng et al. [32] used an emulsion polymerization process to produce alumina-polystyrene powder for the SLS® of composite parts. In this paper, a similar dispersion polymerization process was used to produce alumina parts through indirect SLS®. Different batches of powder were prepared in a twolitre three-neck flask equipped with a thermometer and a reflux condenser. The flask was covered with aluminium foil and immersed in a water bath on a heating plate with a magnetic stirrer.

with a magnetic stirrer. The mixture was heated to 65 °C, and finally the polymerization reaction was initiated by adding 2.26 g 2.2-azobisisobutyronitrile (AlBN, Acros Organics, USA). Fig. 4b schematically presents the dispersion polymerization reaction. The polymerization was carried out at 65 °C for 6.5 hours. After reaction, the final product was cooled to room temperature. The next day the mixture was filtered and washed with water three times. The solid product was dried in an oven at 50 °C for 2 hours to remove all solvents. After being dried, the cake material prepared in the 2000 ml flask was ground in a ball mill (Fritsch, Germany) to obtain a fine powder that was sieved (Retsch, Germany) with a mesh of 160 µm. For a more detailed description of the powder production method, see Cardon et al. [37]. 1.1.2 Selective Laser Sintering

a)

b) Fig. 4. Powder production: a) SEM micrograph of alumina starting powder, b) scheme of the dispersion polymerisation process

A mixture of 1134 g ethanol (99.9%, Merck Millipore, USA) and 66 g water was heated above 50 °C, and 222.32 g of the monomer styrene (99.5%, Acros Organics, USA), 2.32 g divinylbenzene to make the styrene reactive (98% DVB, Merck Millipore, USA) and 120.44 g α-alumina powder (Fig. 4a: grade SM8, Baikowski, France) with a mean particle size of 0.3 µm were poured into the solution and stirred

Green samples were fabricated on a Sinterstation 2000 machine (DTM Corporation / 3DSystems, USA) equipped with a 100 W CO2 laser (f100, Synrad, USA) with a wavelength of 10.6 µm, and a laser beam diameter Ø1/e² of 400 µm. Powder layers could be deposited well by a counter current roller. The powder layers were irradiated with the laser beam in N2 (L’Air Liquide, Belgium, [O2] <5 ppm). In order to improve the laser sinterability of the powder, the parts were produced at a powder bed of ~90 °C. The energy required to melt the amorphous polystyrene phase was partly supplied by preheating of the powder bed (distributed cylinder heating and surface IR heating) and by extra laser irradiation, which raised the temperature locally. Besides SLS® tests to investigate the powder production route and to investigate the powder preheating and cooling conditions, a parametrical study was performed to investigate other crucial SLS® parameters. In this parametrical study, 18 cubic parts of 10×10×10 mm³ were produced with a laser power P, scan speed v, scan spacing s varying between respectively 13 to 17 W, 600 to 1200 mm/s, 0.1 to 0.2 mm. The layer thickness l was fixed at 250 µm. The laser energy density e combining these parameters, varied from 0.22 to 0.76 J/mm³.

e=

P s ⋅v ⋅l

. (1)

After SLS®, the relative green density of the parts was measured. The relative density is the ratio of the absolute density and the theoretical density (TD).

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1.1.3 Debinding and Solid-state Sintering In order to obtain the optimal SLS® parameters for producing alumina parts, all green samples obtained during the parametrical study went through at least two furnace treatments: debinding (deb.) and solidstate sintering (see Fig. 3). In the debinding step, the polystyrene was removed from the ‘green’ parts, and a ‘brown’ part was formed. This was done at a heating rate of 0.1 °C/min. with a two-hour dwell time at 600 °C, followed by furnace cooling. Afterwards, the submicrometre alumina particles of the brown part were fully solid-state sintered (full SSS) to form the final part. This means that the brown part underwent the initial, intermediate and final stages of the SSS process. In this second step, a heating rate of 5 °C/min. was applied with a dwell time of two hours at 1600 °C, followed by furnace cooling. Sometimes, initial solid-state sintering (initial SSS) was used to strengthen the brown parts by forming necks between the submicrometre particles. By applying a heating rate of 5 °C/min and a dwell time of two hours at 1050 °C, these brown parts only went through the initial stage of the SSS process. After a debinding and a full SSS furnace treatment, the final densities of the samples of the parametrical study were measured. 1.2 Geometrical Assessments Benchmark parts were used to investigate the percentage linear shrinkage during debinding and SSS. The research did not focus on the dimensional changes that occurred during the SLS® process. This kind of study, which can be used to determine compensation strategies to geometrically match SLSed parts with the corresponding CAD files, was beyond the scope of the research. The percentage linear shrinkage (% linear shrinkage) that occurred during debinding and solidstate sintering has been defined as:

dimension after SSS - green part dimension . (2) green part dimension

In order to investigate the directional dependence of the shrinkage, the scan, cross-scan and build direction were defined as x, y and z direction as illustrated in Fig. 5. Finally, the benchmark part shown in Fig. 9d has been used to investigate the roughness change in the x and y directions during debinding and SSS. Different roughness values (Ra, Rt and Rz) were obtained in the 650

x and y directions. A Gaussion filter with lower (Ls) and higher (Lc) cut-off values of 0.008 and 2.5 mm, respectively, was used to process the measured data.

Fig. 5. Directional dependency of shrinkage during debinding and furnace sintering: definition of x ‘scan’, y ‘cross-scan’ and z ‘build’ direction.

1.3 Densification Strategies 1.3.1 Warm Isostatic Pressing (WIP) Two WIP tests were performed. In the first test, the vacuum packed SLSed part (‘part 2’ in Table 3) was heated in silicone oil to 100 °C, which is above the glass transition temperature of polystyrene. The second vacuum packed sample, which has been SLSed in another run (‘part 3’ in Table 3), was heated in the same silicone oil to 110 °C. On both samples, a uniform pressure of 16.1 MPa was applied for five minutes. 1.3.2 Infiltration Green parts, initial solid-state sintered parts (IS parts) and/or fully solid-state sintered parts (FS parts) were infiltrated with suspensions containing alumina particles (grade SM8, Baikowski, France) with a mean particle size of 0.3 µm to improve the final density of the produced parts (Fig. 3). Pressureless infiltration tests, i.e. dipping where no external pressure is applied to press the suspension into the pores of the part, and pressure infiltration tests were performed. An ethanolbased suspension containing 20 or 30 vol% alumina was used during the pressureless infiltration tests. An ethanol-based suspension containing 40 vol% alumina was used during the pressure infiltration tests, i.e. applying an external pressure to press the suspension into the pores of the part. All suspensions were stabilized with 0.3 wt% citric acid and mixed in a Turbula mixer for 24 hours. Continuous Green Pressureless Infiltration The weight gain during continuous green pressureless infiltration for 30 hours was assessed with a 20 vol% (part 4) and 30 vol% (part 5) alumina suspension. The dried mass (without ethanol) was calculated from the

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wet mass, which had been measured after 1, 2, 3 and 30 hours of infiltration. Stepwise Green Pressureless Infiltration The weight gain during stepwise green pressureless infiltration for some hours was assessed with 20 vol% (part 6) and 30 vol% (part 7) alumina suspensions. After each infiltration step and before measuring the weight, the samples were placed in a drying furnace at 70 °C for two hours to evaporate the ethanol. In total, four infiltration steps were performed.

2 RESULTS 2.1 Production of Alumina Parts 2.1.1 Powder Production The presented powder production route led to a composite powder with 39 wt% polystyrene. As depicted in Fig. 6a, the composite particles were not spherical in shape. Some pristine α-alumina particles, which had a d50 ~ 0.3 µm (Fig. 4a), could still be observed in the powder (Fig. 6b).

Pressureless Infiltration at Different Stages of the PM Process

Pressure infiltration was a final strategy investigated to increase the density of the parts. The pressure infiltration experiments were performed by using an ethanol suspension containing 40 vol% alumina. Eight different pressure infiltration experiments were performed. Four parts (parts 12 to 15) were both green and FS infiltrated. Four other parts (parts 16 to 19) were green, IS and FS infiltrated. During most infiltration experiments, the ethanol suspension was squeezed for five minutes into the open porosity of the parts at a pressure of 1.61, 16.1 or 48.3 MPa. Two parts, parts 15 and 19, were always pressure infiltrated for 30 minutes at 48.3 MPa.

2.1.2 Selective Laser Sintering After SLS®, the relative green density of the parts was measured (bold percentages in Table 1). The relative density is the ratio of the absolute density and the theoretical density (TD). Assuming a TD of 1.05 g/cm³ for polystyrene and 3.92 g/cm³ for Al2O3, the green TD of the SLSed 61wt% alumina (39 wt% polystyrene powder) was 1.90 g/cm³. Relative green densities varied from 52 to 67%, depending on the laser energy density. When scanning with low laser energy densities, the amount of melted polystyrene was too low to consolidate the powder particles. When scanning with too high laser energy densities, the polystyrene could degrade. Table 1. Green Al2O3-PS composite densities after SLS® (top, bold) and final sintered Al2O3 densities after solid-state sintering (bottom); the densities are expressed in % of the theoretical density (TD) 0.1 0.15 0.2 P [W]

900 mm/s 66 67 66 66 64 61 61 62 59 58 66 63 17 15

65 66 60 64 55 60 13

0.1 s [min]

Pressure Infiltration

a) b) Fig. 6. Powder production; a) SE-SEM and b) BSE-SEM micrograph of the produced alumina/polystyrene composite powder

s [min]

As schematically presented in Fig. 3, infiltration can be performed at different stages of the PM process: before the debinding step (green infiltration), after the debinding and an initial solid-state sintering step (IS infiltration) or after fully solid-state sintering (FS infiltration). As described in the section on debinding and solid-state sintering (see section 1.1), the debound part has to be pre-sintered at 1050 °C to give it some strength before IS infiltration is to be applied. After infiltration, the part is further (fully) SSSed at 1600 °C. In order to investigate the combined influence of green and/or IS infiltration and FS infiltration, 4 parts (‘part 8’, ‘part 9’, ‘part 10’ and ‘part 11’) underwent a pressureless infiltration treatment at different stages of the PM process. Each infiltration treatment, in which the 30 vol% suspension was used, lasted for four hours.

0.15 0.2

P [W]

1200 mm/s 64 64 66 63 58 55 62 65 55 52 62 63 17 15

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a) b) c) Fig. 7. Green part after SLS®; a) camera image, b) 3D microscope cross-sectional image and c) BSE-SEM cross-sectional image

As depicted in Fig. 7a, dross was formed at the bottom of the parts during the SLS® process. Due to the relative high laser energies, the polystyrene viscosity decreased and flowed into the underlying powder. Fig. 7b is a cross-sectional image that shows the different layers which were formed during the SLS® process. No cracks are visible. A SEM image of the cross-section (Fig. 7c) illustrates the low green density of the SLSed parts. A large inhomogeneous network of pores is still visible as dark grey (epoxy resin) and black (air) zones, which surround the consolidated powder particles (light zones).

a)

b)

2.1.3 Debinding and Solid-state Sintering As depicted in Table 1, relative densities between 60 and 66% were obtained. The part SLSed with the highest laser energy density (i.e. 0.76 J/mm³) also had the highest final density of 66%. Therefore, the following optimized SLS® parameters of this part were used for all further part production reported: a laser power of 17 W, a scan speed of 900 mm/s, a scan spacing of 0.1 mm and a layer thickness of 250 µm.

c) d) Fig. 9. Small demo (a, b) and large geometrical benchmark parts (c, d) before (top) and after (bottom) debinding and sintering a) b) Fig. 8. Final part with optimized SLS® parameters; a) 3D and b) BSE-SEM image

Unfortunately, all resulting alumina parts of the parametrical study, contained large amounts of cracks after the two furnace treatments (Fig. 8a). Only small porosities could be observed between the cracks (Fig. 8b). 652

Figs. 9a and b illustrate that the smaller parts (i.e. a cross-section <1 cm²) produced through the presented powder metallurgical process did not have surface cracks, despite their internal cracks (see e.g. Fig. 8a). In fact, the larger parts contained large surface cracks (Figs. 9c and d) after debinding and solid-state sintering. The larger parts also curled due to inhomogeneous shrinking.

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Fig. 10. Percentage linear shrinkage of outer and inner dimensions

2.2 Geometrical Assessments Fig. 10 describes the percentage of linear shrinkage of different features of the benchmark parts (i.e. the parts depicted in Figs. 9c and d) as a function of the corresponding green dimension.

to vary significantly for the smaller green dimensions. When comparing outer and inner dimensions, it can be observed that the outer dimensions shrunk more. Fig. 11a describes the percentage of linear shrinkage of the benchmark parts in x ‘scan’ and y ‘cross-scan’ direction during debinding and SSS. Although the measured variation of % shrinkage was slightly larger for the y direction compared to the x direction, no large difference could be observed. In the rest of the paper, no distinction will be made between shrinkages in the x and y directions. The cubic parts produced to study the SLS® parameters, were also used to investigate the percentage of shrinkage in the z direction. In Fig. 11b the percentage of linear shrinkages of the cubic parts are plotted as a function of the laser energy density. It can be observed that the shrinkage in the z direction was systematically larger than the shrinkage in the x/y direction. As depicted in Table 2, the roughness in the x ‘scan’ direction is slightly lower than the roughness in the y ‘cross-scan’ direction. Furthermore, the final part is smoother than the green part. Table 2. Mean roughness values of the benchmark part, depicted in Fig. 9d, after SLS® (green) and after SSS (final) Green part: Ra [µm] Green part: Rt [µm] Green part: Rz [µm] Final part: Ra [µm] Final part: Rt [µm] Final part: Rz [µm]

a)

x direction 18 142 107 19 128 103

y direction 22 202 152 22 167 138

2.3 Densification Strategies 2.3.1 Warm Isostatic Pressing (WIP)

b) Fig. 11. Directional dependency of shrinkage during debinding and furnace sintering; a) percentage of linear shrinkage in x and y directions, b) percentage linear shrinkage in x/y and z direction as a function of laser energy density

By examining the larger dimensions (>20 mm), it can be seen that the percentage of linear shrinkage was about 30%. The percentage of linear shrinkage seemed

The WIP of the green parts resulted in an increase of the green density and of the geometrical shrinkage (not reported in Table 3). For part 2, the WIP increased the green density from 66 to 95% and resulted in a geometrical shrinkage of –10% (x/y direction) and –7% (z direction). Although part 3 was WIPed at a 10 °C higher temperature than part 2, the green density increased from 67 to 80% and the resulting geometrical shrinkage was only –3% (in all directions). After debinding and solid-state sintering, the final densities of part 2 and 3 were 53 to 49%, respectively; the measured shrinkages were about 23% (in all directions): please see Table 3. This is lower than the density and shrinkage of the reference ‘part 1’ (see

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Table 3. Sintered densities and linear shrinkages* of the cubic (10×10×10 mm³) alumina parts after additional densification steps Part

Additional densification steps (besides SLS®, deb. and SSS)

ρ [%]

x/y [%]

z [%]

Part 1

No, i.e. shrinkage resulting from deb. and SSS only

66

-31

-44

Part 2

WIP100 °C

53

-22

-24

Part 3

WIP110 °C

49

-24

-23

Part 4

green continuous pressureless inf.: 20 vol.%, 30 h

54

-19

-17

Part 5

green continuous pressureless inf.: 30 vol.%, 30 h

51

-20

-18

Part 6

4 x stepwise green pressureless inf.: 20 vol.%, 4 h

48

-20

-18

Part 7

4 x stepwise green pressureless inf.: 30 vol.%, 4 h

47

-17

-15

Part 8

FS pressureless inf.: 30 vol%

66

-34

-32

Part 9

IS and FS pressureless inf.: 30 vol%

69

-33

-37

Part 10

green and FS pressureless inf.: 30 vol%

61

-24

-21

Part 11

green, IS and FS pressureless inf.: 30 vol%

61

-23

-20

Part 12

green and FS pressure inf.: 1.61 MPa; 5 min., 40 vol%

64

-17

-16

Part 13

green and FS pressure inf.: 16.1 MPa; 5 min., 40 vol%

63

-21

-17

Part 14

green and FS pressure inf.: 48.3 MPa; 5 min., 40 vol%

62

-21

-16

Part 15

green and FS pressure inf.: 48.3 MPa; 30 min., 40 vol%

63

-20

-18

Part 16

green, IS and FS pressure inf.: 1.61 MPa; 5 min., 40 vol%

71

-17

-17

Part 17

green, IS and FS pressure inf.: 16.1 MPa; 5 min., 40 vol%

85

-18

-17

Part 18

green, IS and FS pressure inf.: 48.3 MPa; 5 min., 40 vol%

84

-20

-16

Part 19

green, IS and FS pressure inf.: 48.3 MPa; 30 min., 40 vol%

76

-19

-17

* The dimensional shrinkages of this table indicate the geometrical changes that appear after the is the reference geometry).

Table 3), which was produced with the optimized SLS® parameter set and not WIPed. 2.3.2 Infiltration Continuous Green Pressureless Infiltration As depicted in Table 4, the largest weight gain occurred during the first hour of infiltration. After 30 hours of infiltration, the 30 vol% alumina suspension led to a higher weight gain than the 20 vol% alumina suspension. Furthermore, comparing ‘part 4’ and ‘part 5’ in Table 3 with ‘part 1’ reveals that 30 hours of green pressureless infiltration decreased the final density and the shrinkage during debinding and solid-state sintering. Moreover, it reduced the difference between the shrinkage in the x/y direction and the shrinkage in the z direction. As depicted in Table 5, the largest increase of weight could be observed after the first infiltration step. The highest weight gain was again observed when infiltrating with the 30% alumina suspension. By comparing ‘part 6’ and ‘part 7’ in Table 3 with ‘part 1’, it can be observed that green pressureless infiltration decreased the final density and the shrinkage during debinding and solid-state sintering. 654

SLS®

process (i.e. the geometry after SLS®

It can also be observed that green pressureless infiltration reduced the difference between shrinkage in the x/y and z directions. Table 4. Measured weight gain during continuous infiltration Part

0 h [g]

1 h [g]

2 h [g]

3 h [g]

30 h [g]

Part 4

1.26

1.48

1.49

1.48

1.53

Part 5

1.28

1.57

1.57

1.58

1.63

Stepwise Green Pressureless Infiltration Table 5. Measured weight gain during stepwise infiltration Part

green [g] step 1 [g] step 2 [g] step 3 [g] step 4 [g]

Part 6

1.28

1.47

1.54

1.57

1.59

Part 7

1.26

1.54

1.61

1.64

1.66

Pressureless Infiltration at Different Stages of the PM Process Tables 6 and 7 describe for each part, at which stages of the PM process the infiltration was performed. Tables 6 and 7 also describe the changes in relative density and geometrical shrinkage during the different steps of the PM process, respectively. Higher densities were obtained when no green infiltration

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Table 6. Relative densities [%] obtained during the different pressureless infiltration strategies* green inf. o o x x

Part Part 8 Part 9 Part 10 Part 11

deb. x x x x

(initial) SSS 1050 °C o 38 o 31

IS inf. o x o x

(full) SSS 1600 °C 65 72 55 58

FS inf. x x x x

(full) SSS 1600 °C 66 69 61 61

Table 7. Percentage shrinkage [%] obtained at different stages of the PM process* Part

green inf.

deb.

Part 8 Part 9 Part 10 Part 11

o o x x

x x x x

(initial) SSS 1050 °C x/y z o o -19 -24 o o -8 -8

IS inf. o x o x

(full) SSS 1600 °C x/y z -32 -31 -32 -37 -24 -23 -23 -22

FS inf. x x x x

(full) SSS 1600 °C x/y z -34 -32 -33 -37 -24 -21 -23 -20

* The performed post-treatments are signed with an ‘x’. The post-treatments that are not performed are marked with an ‘o’.

had been performed. In contrast, IS and FS infiltration increased the final densities of the parts. Furthermore, green infiltration decreased the part shrinkage. By comparing parts 8 to 11 with part 1 in Table 3, it can be observed that infiltration decreased the difference between shrinkages in the x/y direction and z direction. Pressure Infiltration Comparing parts 12 to 15 from Table 8 with part 10 of Table 6 reveals that when only applying green infiltration, the application of pressure did not significantly increase the final densities. Moreover, for both the pressure infiltrated and the pressureless infiltrated parts, the densities after the first SSS step were 52 to 55%. After FS infiltration and again full SSS, the densities were 61 to 64%. The shrinkages in the x/y and z directions after SSS were somewhat less when pressure infiltration was applied. For example, after the first full SSS step, x/y shrinkages of 18 to 21% (Table 9) instead of 24% (Table 7) were observed. By comparing parts 16 to 19 from Table 8 with part 11 of Table 6, it can be seen that applying an extra IS pressure infiltration step had a more pronounced influence on the final densities: densities up to 82% were reached after the first full SSS step (compared to 58% for part 11 in Table 6). After FS infiltration and again SSS, densities up to 85% were assessed (compared to 61% for part 11 in Table 6). Applying the pressure for a longer time decreased the final densities: consider part 19 in Table 8. The shrinkages in the x/y and z directions after SSS were somewhat less when applying pressure infiltration instead of

pressureless infiltration: for example, after the first full SSS step, x/y shrinkages of -19 to -21% (Table 9) instead of -23% (Table 7) were obtained. The part shrinkages also decreased by applying an extra FS pressure infiltration step (Table 9). Finally, by comparing parts 12 to 19 with part 1 in Table 3, it can be once more observed that infiltration decreased the difference between the x/y shrinkage and the z shrinkage. 3 DISCUSSION 3.1 Production of Alumina Parts A disadvantage of the presented PM process is the occurrence of multiple cracks in the final parts when no post-densification step was applied (Fig. 8). Since the green SLSed parts did not contain cracks, they either originated during the debinding treatment or during solid-state sintering. In order to examine when exactly the cracks originated, two green (pressureless) infiltrated parts were further investigated. One part was debound and initially solid-state sintered at 1050 °C to give the brown part some strength without causing too much shrinkage. The other part was debound and solid-state sintered at 1600 °C. Both parts were cut with a diamond blade, and the crosssections were visualized by the 3D microscope. It could be clearly observed that the IS part (Fig. 12a) had some big cracks. This means that the cracks originated during debinding. As shown in Fig. 12b, the cracks were still visible in the FS part after solidstate sintering at 1600 °C.

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Table 8. Relative densities [%] obtained during the different pressure infiltration strategies* Part

Pressure [MPa]

Time [min]

green inf.

deb.

(initial) SSS 1050 °C

IS inf.

(full) SSS 1600 °C

FS inf.

(full) SSS 1600 °C 64

Part 12

1.61

5

x

x

o

O

53

x

Part 13

16.1

5

x

x

o

O

52

x

63

Part 14

48.3

5

x

x

o

O

54

x

62

Part 15

48.3

30

x

x

o

O

55

x

63

Part 16

1.61

5

x

x

31

X

66

x

71

Part 17

16.1

5

x

x

31

X

80

x

85

Part 18

48.3

5

x

x

31

X

82

x

84

Part 19

48.3

30

x

x

32

X

72

x

76

Table 9. Percentage shrinkage [%] obtained at different stages of the PM process* Part

Pressure [MPa]

Time [min]

green inf.

deb.

Part 12 Part 13 Part 14 Part 15 Part 16 Part 17 Part 18 Part 19

1.61 16.1 48.3 48.3 1.61 16.1 48.3 48.3

5 5 5 30 5 5 5 30

x x x x x x x x

x x x x x x x x

(initial) SSS 1050 °C x/y z o o o o o o o o -2 -2 -3 -2 -5 -4 -3 -1

IS inf. o o o o x x x x

(full) SSS 1600 °C x/y z -19 -16 -21 -17 -20 -16 -18 -16 -19 -19 -20 -18 -22 -18 -21 -18

FS inf. x x x x x x x x

(full) SSS 1600 °C x/y z -17 -16 -21 -17 -21 -16 -20 -18 -17 -17 -18 -17 -20 -16 -19 -17

* The performed post-treatments are marked with an ‘x’. The post-treatments that are not performed are marked with an ‘o’.

a) b) Fig. 12. 3D microscope image of a green infiltrated part; a) after initial SSS at 1050°C; and b) after fully SSS at 1600°C

Two strategies were investigated to eliminate the cracks of the final alumina parts. The first strategy was exploring the possibilities of infiltration: see sections 2.3 and 3.3. The second strategy was to investigate the kinetics of the produced powder during SLS® and the debinding cycle. In order to do so, multiple differential scanning calorimetry and thermogravimetry (DSCTGA) analyses (STA 449, Netzsch, UK) and Fourier transform infrared spectroscopy (FTIR) analyses (Bruker, Germany) were performed. It was found that the glass transition temperature Tg of the composite powder after production through the dispersion polymerization process was only 54 °C instead of the 656

expected 110 °C for standard pure polystyrene. This means that the produced polystyrene had a rather low molecular weight or chain length. When using the composite powder for the first time in the Sinterstation 2000 SLS® machine, the powder was preheated to 90 °C. This caused the polymerization process to continue, which resulted in a higher glass transition temperature Tg of 90 °C. During the debinding cycle, a complex degradation process occurred. The degradation process could be summarized as follows: 1. evaporation of unreacted styrene, 2. thermal cracking of polystyrene, 3. reorganization of the main chain to a stable aromatic structure, 4. combustion of the stable aromatic structure. In order to eliminate the cracking during the debinding cycle, an optimized debinding scheme was proposed, which maintained the heating rate of 0.1 °C/min, but introduced a dwell time of 15 minutes at 250 °C (for the reorganization of the main chain) and at the final temperature of 600 °C. Since the final parts still contained cracks after the optimized debinding cycle, it can be concluded that the cracks were probably caused by inhomogeneous distribution of alumina and polystyrene concentrations in the

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composite starting powder. The inhomogeneous concentrations of alumina and polystyrene led to inhomogeneous shrinkage and the resulting cracks in the debinding step. Future research may focus on exploring further the influence of atmosphere conditions on the kinetic behaviour of the produced polystyrene during the debinding. Other powder production routes can also be explored to produce more homogeneous and spherical composite alumina-polystyrene starting powder, e.g. a dissolution-reprecipitation route (Shahzad et al. [38]). 3.2 Geometrical Assessments The benchmark parts depicted in Figs. 9c and d were used to investigate the percentage of shrinkage during debinding and SSS. The measured shrinkages were caused by: • (mostly) attractive capillary forces during debinding (Rahaman [39]; Megias-Alguacil and Gauckler [40]), • attractive Van der Waals forces after debinding and before solid-state sintering, • shrinkage due to atomic diffusion during SSS, • cracks and curling due to inhomogeneous shrinking in the debinding and/or SSS step. Cracks did not contribute to the shrinking process. This means that if no cracks had occurred during the debinding cycle, a linear shrinkage larger than 30% would have been observed.

a) b) Fig. 13. Attractive forces during shrinkage of a) an outer and b) inner geometry

The large variety of the shrinkage for smaller green dimensions (Fig. 10) was probably due to measurement errors, which were relatively larger for smaller dimensions. A possible explanation for the observation that outer dimensions shrunk more than inner dimensions can be found by examining the attractive forces that occurred during shrinkage. As illustrated in Fig. 13a, the attractive forces were not constrained and tended to reduce the outer contours. In this case, the reduction of the outer contours and the shrinkage of the debound part acted in the same

direction. In contrast, the attractive forces tended to increase internal contours (Fig. 13b). The shrinkage of the debound part counteracted this increase, resulting in a lower total shrinkage of the internal geometry. 3.3 Densification Strategies 3.3.1 Warm Isostatic Pressing (WIP) The different densities of part 2 and part 3 after the WIP process might be related to aging of the composite material of the SLSed sample. Although no cracks could be observed after WIP (Fig. 14a), one large internal crack could be observed in the cross-sections of the parts after debinding and solidstate sintering (Fig. 14c). This was in contrast with the large amount of smaller cracks in part 1 (Fig. 8a), which was not WIPed. The large crack might be related to the lower densities and lower shrinkages of the WIPed samples after SSS (see Table 3). Besides the difference in amount and size of cracks, the green and final microstructure of a WIPed part (Figs. 14b and d) was similar to that of a part which was not WIPed (Figs. 7c and 8b). 3.3.2 Infiltration Unfortunately, all the infiltrated parts contained big internal cracks and voids (see e.g. Fig. 15a) after the solid-state sintering process. This was in contrast with the large amount of small cracks, which occurred when no infiltration was applied (see Fig. 8a). As illustrated by Fig. 15b, only small porosities were obtained between the cracked zones. A dense alumina shell was also observed at the edges of the parts (Fig. 15c). This shell was probably created when the alumina suspension was entering the pores and obstructed further infiltration. The shell formation might explain different shrinkage results as: • The decrease of part shrinkage when green (pressureless) infiltration was applied. In this case, the dense shell enlarged the part and prohibited the part to shrink freely. • The decrease of part shrinkage when pressure infiltration instead of pressureless infiltration was applied. In this case, the pressure caused the formation of a thicker shell. • Possibly, shell formation also decreased the difference between shrinkages in the x/y direction and z directions.

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a)

However, the cracks could be filled with alumina, and densities up to 85% were obtained. Fig. 16a shows a 3D microscope image of part 18, which had a density of 84%. A large crack, pressure infiltrated with alumina, is clearly visible. Upon closer examination of the large infiltrated crack, micro-cracks surrounded by dense alumina become visible (Fig. 16b). Smaller micro-cracks, surrounded by dense alumina are also visible in the bulk material, i.e. in the infiltrated zones next to the large crack (Fig. 16c).

b)

4 SUMMARY AND CONCLUSIONS c)

d) Fig. 14. 3D microscope images (a, c), BSE-SEM image (b) and SE-SEM images (d) of a part after WIP (a, b) and part 3 after WIP, debinding and SSS (c, d)

A drawback of the application of the pressure infiltration method was the breaking of the parts. This might be caused by air that was trapped in the part and squeezed during the infiltration process.

A PM process was presented to produce freeform alumina parts with a relative density of 66% through indirect SLS®. The PM process comprised a dispersion polymerization process to produce composite powder particles, as well as a tuned SLS® process, a debinding, and a solid-state sintering sintering step. It was found that the final parts produced through this PM process contained cracks, which were formed during debinding. The larger parts (i.e. a cross-section >1 cm²), contained cracks that were also visible at the

a) b) c) Fig. 15. a) 3D microscope image of part 11, b) BSE-SEM micrograph taken in a non-cracked area of part 8 and c) SE-SEM micrograph taken at the edge of part 9

b) c) a) Fig. 16. a) 3D microscope image of part 18, showing a crack pressure infiltrated with alumina; b) and c) the infiltrated crack and the zone around the infiltrated crack are detailed in BSE-SEM images

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surface. Furthermore, the larger parts tended to curl during the debinding and SSS step. The cracks and the curling were probably caused by inhomogeneous distribution of alumina and polystyrene concentrations in the composite starting powder. Through geometrical assessments, the percentage shrinkage that occurred during the debinding and SSS process was investigated. The measured shrinkages were caused by capillary forces, Van der Waals forces, atomic diffusion, cracking and curling of the parts. The shrinkage was more or less the same in the scan and cross-scan directions (about –31%) but much larger in the building direction (about –44%). Furthermore, the unconstrained attractive shrinking forces seemed to let outer dimensions shrink more compared to inner dimensions. Densification strategies were presented to improve the density of the fabricated parts and reduce the occurrence of cracks in the final parts: WIP and infiltration. Although WIP could increase green densities from 66 to 95%, the final density and shrinkage during debinding and solid-state sintering was lower when WIP was used. This might be due to the presence of large cracks in the final parts. Pressureless and pressure infiltration tests with alumina-ethanol suspensions were performed. After pressureless infiltration, large cracks could be observed in the final parts. During continuous green pressureless infiltration, the largest weight gain was obtained during the first hour using a 30 vol% alumina suspension. During stepwise green pressureless infiltration, the largest weight gain was observed after the first step using a 30 vol% alumina suspension. Furthermore, green pressureless infiltration decreased the part shrinkage during debinding and solid-state sintering. The combined influence of green and/or IS and FS pressureless infiltration with a 30 vol% alumina suspension was also examined. As a result, higher densities were observed when no green pressureless infiltration was performed. In contrast, IS and FS pressureless infiltration increased the final densities of the parts. In general, pressureless infiltration reduced the difference between the shrinkage in the x/y direction and the shrinkage in the z direction. The influence of pressureless infiltration on the geometrical shrinkage might be related to the formation of a dense shell at the edges of the part. Finally, the combined influence of green and FS pressure infiltration with or without IS pressure infiltration was investigated, using 40 vol% alumina suspensions. As with pressureless infiltration, shell formation might be related to the reduction of geometrical shrinkage during the PM process and the

reduced difference between the shrinkage in the x/y and z directions, respectively. Without IS pressure infiltration, the densities of the final parts were similar to the densities obtained by pressureless infiltration. The application of IS pressure infiltration led to an increase of the part densities up to 84%, since the cracks that occurred during the debinding process could be filled with alumina. Nevertheless, microcracks arose. The micro-cracks probably occurred due to non-homogeneous shrinkage during solid-state sintering. 5 ACKNOWLEDGEMENTS This work was financially supported by the Flemish Institute for the Promotion of Scientific Technological Research in Industry (IWT) under project SBO-DiRaMaP. 6 REFERENCES [1] Cormier, D., Taylor, J., Unnanon, K., Kulkarni, P., West, H. (2000). Experiments in Layered ElectroPhotographic Printing. Proceedings of the SFF Symposium, p. 267-274. [2] Khoshnevis, B., Asiabanpour, B., Mojdeh, M., Koraishy, B., Palmer, K., Deng, Z. (2003). SIS-A new SFF method based on powder sintering. Rapid Prototyping Journal, vol. 9, no. 1, p. 30-36, DOI:10.1108/13552540310455638. [3] Brooks, H., Aitchison, D. (2010). A review of state-of-the-art large-size foam cutting rapid prototyping and manufacturing technologies. Rapid Prototyping Journal, vol. 16, no. 5, p. 318-327, DOI:10.1108/13552541011065713. [4] de Smit, B., Broeck, H.J. (2004). Analysing the cutting process of a heated flexible blade in extruded polystyrene foam. Proceeding of the SFF Symposium, p. 591-601. [5] Mahale, T.R., Taylor, J.B., Cormier, D.R. (2000). Fiveaxis freeform fabrication of the thermoplastic parts via SWIFT. Proceedings of the SFF Symposium, p. 289297. [6] Cheah, C.M., Chua, C.K., Lee, C.W., Feng, C., Totong, K. (2005). Rapid prototyping and tooling techniques: a review of applications for rapid investment casting. The International Journal of Advanced Manufacturing Technology, vol. 25, no. 3-4, p. 308-320, DOI:10.1007/ s00170-003-1840-6. [7] Kinsella, M.E., Lilly, B., Carpenter, B., Cooper, K. (2004). Ejection forces and friction coefficients from injection molding experiments using rapid tooled inserts. Proceedings of the SFF Sympsium, p. 669-680. [8] Kruth, J.P., Levy, G., Schindel, R., Craeghs, T., Yasa, E. (2008). Consolidation of polymer powders by selective laser sintering. Proceedings of the 3rd International

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applications. Journal of Applied Polymer Science, vol. 128, no. 3, p. 2121-2128, DOI:10.1002/app.38388. [38] Shahzad, K., Deckers, J., Boury, S., Neirinck, B., Kruth, J., Vleugels, J. (2012). Preparation and indirect selective laser sintering of alumina/PA microspheres. Ceramics International, vol. 38, no. 2, p. 1241-1247, DOI:10.1016/j.ceramint.2011.08.055. [39] Rahaman, M.N. (2003). Chapter 6: Powder consolidation and forming of ceramics. Ceramic Processing and Sintering – 2nd edition, Marcel Dekker Inc., New York. [40] Megias-Alguacil D., Gauckler L.J. (2010). Capillary and van der Waals forces between uncharged colloidal particles linked by a liquid bridge. Colloid and Polymer Science, vol. 288, p. 133-139, DOI:10.1007/s00396009-2106-0.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 662-668 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2012.903 Special Issue, Original Scientific Paper

Received for review: 2012-12-07 Received revised form: 2013-03-15 Accepted for publication: 2013-03-29

Injection-Moulding Compounding of PP Polymer Nanocomposites Battisti, M.G. – Friesenbichler, W. Markus Gottfried Battisti* – Walter Friesenbichler

University of Leoben, Chair of Injection Moulding of Polymers, Department Polymer Engineering and Science, Austria The objective of this study is a comparison of different compounding techniques and their influence on Young’s modulus. Both conventional processing of polymer nanocomposites and processing in the unique Injection Moulding Compounder (PNC-IMC) were evaluated. Additionally, the effects of nanofillers on the thermal conductivity of polymer nanocomposites were investigated at various pressures. In comparison to conventional compounding process, in which the compound must be pelletized and fed into the injection moulding machine for the second plasticizing process, injection-moulding compounding combines these two processing steps. The material compounding and the subsequent injection moulding are done directly with only one plasticizing process, with the use of a heated melt line and a melt accumulator. In this study, both of these techniques were used for the production of polymer nanocomposites. This paper shows the different effects of processing techniques, screw speed, counter-pressure and different extruder length on Young’s modulus and demonstrates that, for the improvements, a compromise between shear energy input and residence time is essential. The increase of thermal conductivity by using nanofillers in comparison to the virgin polypropylene is shown. The investigated increase in thermal conductivity should be extremely appealing for the industry in terms of cycle time reduction in the injection moulding process. These first results give an excellent overview of both the possibilities and the limitations of the innovative concept of the PNC-IMC. Further studies on the detailed understanding of the exfoliation and intercalation of layered silicates in polymer melt will be done. Keywords: injection moulding compounding, polymer nanocomposites, thermal conductivity, exfoliation, intercalation, layered silicates

0 INTRODUCTION 0.1 Injection Moulding Compounding The Injection Moulding Compounder (IMC) concept was first presented by Krauss Maffei at the K1998 for long glass-fibre-reinforced thermoplastics. Its characteristic feature is a twin-screw extruder, which is directly integrated into an injection-moulding system. The injection-moulding compounding process combines two processing steps: the material compounding, which normally takes place at the raw material manufacturer and the injection-moulding process which usually is done at the injection moulder. The final part is processed directly with only one plasticizing process, and the material used does not have to be granulated and plasticized again after the compounding step in the plasticizing unit of an injection-moulding machine. The molten material exits the extruder and is directly fed into a melt accumulator and further into the injection unit. After the specimen volume is reached, a conventional injection-moulding process starts. The accumulator serves as a buffer element between the continuouslyoperating extruder and the discontinuously-operating injection unit. The outcome is greater flexibility for the processor when improving both the component structure and the quality of the moulded part, as well as demonstrably lower production costs [1] and [2]. 662

To date, all presented injection-moulding compounding systems have been designed for the processing of fibres (long glass fibres, natural fibres). The benefit of processing in a “single heat” is the lower working temperature and the shorter residence time in the whole processing facility, as well as the lower reduction of the fibre length. The material’s properties are adjusted by the addition of additives, for instance, to improve bonding of the polar natural fibres to the nonpolar polyolefins, which is also necessary for layered silicates. Another advantage of the one-step process compared to the two-step process is energy savings of about 40%, since the more energy-efficient, continuously operating co-rotating twin-screw extruder replaces the energy-intensive step of plasticating by means of a reciprocating screw [3]. Wobbe [4] states that enormous areas remain unexplored when considering the possibilities and advantages of “single-heat” injection moulding. Examples are: • Long fibres are shortened less (the length is responsible for the stiffness of the fibre composite) because there is no second plasticizing process in the injection unit. • There is less wear and tear on the machine components, because a high level of abrasive fillers can be added to the molten polymer downstream. • It will be easier to produce finished articles from compounds that have a tendency to separate. *Corr. Author’s Address: University of Leoben, Chair of Injection Molding of Polymers, Department Polymer Engineering and Science, Austria, markus.battisti@unileoben.ac.at


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The MUL (University of Leoben) polymer nanocomposite Injection Moulding Compounder (PNC-IMC) described below (Fig. 1) is the world’s first specially adapted IMC for the manufacturing and injection moulding of polymer nanocomposites. It consists of a ZSE 27 MAXX Leistritz-Compounder and an 1800 kN fully-electric injection moulding machine by Engel. A 3-way valve and its particular arrangement allow for using the IMC as three-inone production line either for injection moulding compounding, for injection moulding or as a separately acting twin-screw compounder line. Furthermore, this PNC-IMC can also be used for the homogenization of fillers in high temperature polymers (HTP) in temperatures ranging up to 400 °C.

of polymer nanocomposites and the successful dispersion of silicate plates in a polymer melt. Based on preliminary work at MUL, it is now possible to realize high shear energy as well as long residence time (which are normally mutually exclusive in the extrusion process) by implementing a melt pump in front of the twin-screw extruder [7]. Further possibilities for diversifying the residence time of the polymer melt are the screw length of the extruder, and the processing in an IMC, both of which are possible with the PNC-IMC described. Our aim is the improvement of the mechanical properties [8] and materials with higher thermal conductivity, in order to reduce the cooling period of injection moulded parts and blow-moulded parts with medium-to-high wall thicknesses, respectively. There are many reasons for the use of nanoclays in polymer systems ranging from the enhancement of structural and mechanical properties to functional materials with a tailored performance profile for thermostability, flame retardance, barrier properties or enhancement of thermal conductivity [9] and [10]. 0.2 Thermal Conductivity of Polymers

Fig. 1. System schematic of the MUL Polymer NanoComposite Injection Moulding Compounder (PNC-IMC)

1.2 Compounding of Polymer Nanocomposites Compounding by melt mixing is the most attractive industrial method for the preparation of polymer nanocomposites due to its technological simplicity. Common polymer processing machines are much easier and more pleasant to use than using equipment and procedures in chemical laboratories, which are required for in situ or solution methods [5]. Furthermore, these “laboratory methods” are only useful for limited amounts of compounds; therefore, they are less desirable for many industry applications [6]. From the economical point of view, the use of a co-rotating twin-screw extruder as a continuous processing technology is preferable to melt mixing in a discontinuous kneader. High shear energy (usage of kneading blocks and high screw speed) and long residence time are necessary for the preparation

Thermal conductivity is the most commonly used property that helps to quantify the transport of heat through a material. In polymers, thermal conductivity is based on two different mechanisms: heat is transferred through Van-der-Waals forces, and phonons are stimulated through covalent bonding. By definition, energy is transported proportionally to the speed of sound; therefore, thermal conductivity follows the relationship

λ = c p ⋅ ρ ⋅υ ⋅ l , (1)

where the thermal conductivity λ is directly proportional to the specific heat capacity cp, the density ρ, the speed of sound υ, and the molecular separation l of the elastic waves [11]. Due to the increase in density below the solidification of semi-crystalline thermoplastics, the thermal conductivity is higher in the solid state than in the melt. Furthermore, higher pressure also increases the thermal conductivity. Other factors, such as temperature, morphology and molecular orientation, are also affecting the thermal conductivity of polymers. Highly drawn semi-crystalline polymers like PP can have a much higher thermal conductivity in the direction of orientation of the polymer chains than perpendicular to it.

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It is also beneficial to know the thermal conductivity of filled compositions in order to model and analyse the heat transfer process during moulding. Thermal conductivity can be influenced significantly by the addition of different fillers and additives since the thermal conductivity of most inorganic fillers is around 10 times higher than that of polymers. The increase of thermal conductivity depends on the type, on the amount as well as on the particle shape, the size and the distribution of the used fillers [12]. 1 EXPERIMENTAL 1.1 Materials An isotactic polypropylene (PP) homopolymer Bormed DM55pharm (MFR 2.8 g/10 min; 230 °C/2.16 kg) used for the preparation of nanocomposites was supplied by Borealis. The used nanofiller was Nanofil 5 (montmorillonite intercalated with dimethyl distearyl ammonium chloride) and the compatibilizer Scona TPPP 2112 GA (PP grafted with 1.2 wt% of maleic acid anhydride, PP-g-MA, MFR 6 g/10 min; 190 °C/2.16 kg) was supplied by Kometra. 1.2 Preparation of Polypropylene Nano-Composites The preparation of the polymer nanocomposites had been carried out with the intermeshing, co-rotating twin-screw extruder of the PNC-IMC. The possibility of varying the length of the extruder has opened a large area for scientific development and research, and practical applications. We used 11 or 13 barrel segments for the compounding process in which each barrel segment had a length of four times the screw diameter D. PP and compatibilizer were fed upstream through the main hopper, and the organoclay was fed downstream at the 4th extruder barrel for both equipment configurations. All the components were fed by separately controlled gravimetric dosage units at an overall throughput rate of 6 kg/h. The formulation of the compounded polymer nanocomposites was constant at 90 wt% polypropylene, 5 wt% organoclay and 5 wt% compatibilizer. The screw geometry was held constant. The screw speed and the counterpressure of the melt pump were varied according to Table 1 in order to observe the influence of different processing conditions. It is necessary to know that the pressure gradient in the melt Witte pump is negative to increase the residence time and shear rate of the polymer compound leading to improved dispersion of the nanofiller. The minimal residence time was measured 664

as the time between the colour masterbatch granulate insertion into the hopper and the colour change of the outgoing molten string at the three-way-valve and the injected part, respectively. The extruder temperature profile was set at 170 to 210 °C from hopper up to the EUP 50 underwater pelletizing system, which is a water process & drying system and self-cleaning centrifuge. Table 1. Indication of the prepared nano-composites; extruder length as multiple of the screw diameter D Extruder length D 44 44 44 44 44 44 44 44 44 52 52 52 52 52 52 52 52 52

Screw speed [rpm] 50 50 50 100 100 100 150 150 150 50 50 50 100 100 100 150 150 150

Counter pressure [bar] 50 75 100 50 75 100 50 75 100 50 75 100 50 75 100 50 75 100

In order to monitor the differences to the processing with the PNC-IMC (processing in a single heat), two compounding grades were manufactured, both conventionally with the underwater pelletizing system and in the IMC. These experiments were made both with 5 wt% organoclay and 5 wt% compatibilizer as well as 10 wt% organoclay and 10 wt% compatibilizer, to verify the influence of nanofillers on the thermal conductivity. Another experiment was the manufacturing of polypropylene nanocomposites with a pre-stage masterbatch process. Here, a masterbatch of 50 wt% organoclay and 50 wt% compatibilizer was produced in the twin-screw extruder and then pelletized. The masterbatch pellets were fed into the twin-screw extruder again and then diluted with virgin PP to receive a compound of 5 wt% nanofiller and 5 wt% compatibilizer.

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Table 2. Various options for the production of the polypropylene nanocomposites compounds One-step process Two-step process 3in1 process Master­batch (MB) process

Production of tensile bars in a single heat IMC-process (one plasticizing process) Conventional compounding and pelletizing and subsequent injection moulding of tensile bars • Direct compounding of the organoclay and the compatibilizer into the polymer • Processing with the one- or two-step process • Production of the MB (50 wt% organoclay and 50 wt% compatibilizer) • Dilution of the MB in the compounder with PP to receive a compound with 5 or 10 wt% organoclay and compatibilizer • Processing with the one- or two-step process

1.3 Mechanical Properties – Tensile Test The tensile bars for the investigation of the mechanical properties were produced both directly with the PNCIMC and in the two-step process with the injection moulding machine of the PNC-IMC. To guarantee the best and primarily the same processing conditions for all test specimens, test series were made before the test specimen’s production. In the near future, the optimisation of the manufacturing process with the use of Design of Experiments (DoE) is planned [13]. Table 3. Injection moulding conditions for the production of the test specimens Cylinder temperature [°C] (Z1-Z2-Z3-Z4-ZN) Mould temperature [°C] Holding pressure [bar] Injection speed [cm³/s] Dosing speed [m/s]

190-200-200-210-210 40 800 75 0.075 or IMC

specimen was filled in the cylinder and compressed tightly while filling. Afterwards, the material was compressed with a piston for 10 minutes and then the measurement started at the defined pressure stage. At least three measurements for each temperature were made, with a 12- to 13-minute delay between each measurement. While the specimens were cooling, the waiting time between each temperature range was 20 to 30 minutes. 2 RESULTS AND DISCUSSION 2.1 Effects of Screw Speed, Counter-Pressure and Different Extruder Length on Young’s Modulus As explained earlier in this paper, the shear energy and residence time are the two main influential variables, which normally act opposite in the extrusion process. The effects of screw speed and counter-pressure on Young’s modulus for an extruder length of 44D are plotted in Fig. 2. It is obvious that Young’s modulus decreases with increasing screw speed. The most significant decline occurs from 100 to 150 rpm. That means with higher screw speed and thereby higher shear forces, the quality of the polymer melt is reduced. Increasing counter-pressure of the melt pump shows a similar effect. At 50 and 100 rpm screw speed, Young’s modulus decreases almost linearly; only at 150 rpm is a plateau at 75 bar reached. Young’s modulus values of pure PP result from tensile tests of injection-moulded parts, which were produced conventionally without an injectionmoulding compounding process.

A Zwick Z010 universal tensile testing machine was used to carry out the tensile tests according to standard ISO 527–1. All tests were done at standardized conditions (23±2 °C / 50±5% r.H.). The data was evaluated using the testXpert II software. 1.4 Thermal Conductivity The thermal conductivity measurements based on the Line-Source method according to ASTM D5930-97 in a temperature range from 80 to 240 °C were carried out using a high pressure capillary rheometer with a single channel system having a diameter of 15 mm. At the beginning of each measurement, the high pressure capillary rheometer was heated to the specified temperature for 60 minutes. A defined weight of the

Fig. 2. Effects of screw speed and counter-pressure on Young’s modulus using an extruder length of 44D

Fig. 3 shows the influence of screw speed and counter-pressure on Young’s modulus at an extruder length of 52D. To explore only the effects of longer residence times, the additional 8D were only an

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elongation of the metering zone with conveying elements. In contrast to the results at the 44D extruder length, the processing conditions with a screw speed of 100 rpm delivered the best Young’s modulus values. Furthermore, there is no clear trend visible on the quality related to the counter-pressure. The residence time of the polymer melt varies between 260 s (100 bar), 245 s (75 bar) and 230 s (50 bar). It seems that a counter-pressure of 75 bar leads to the best results.

Fig. 3. Effects of screw speed and counter-pressure on Young’s modulus using an extruder length of 52D

Clearly, the improvement in Young’s modulus compared to virgin polypropylene was approximately 30% for the PNCs investigated, while the influence of processing conditions was low. A consolidated view indicates that longer residence time, which is certainly necessary for the diffusion process of intercalation and subsequent delamination of silicate platelets in the polymer matrix, seems to have a negative effect on the Young’s modulus. The measurements with 52D extruder length led to worse results than the experiments with 44D. The influence of the screw speed controlled shear energy on the improvement of Young’s modulus is not clear; further research is thus necessary. 2.2 Effects of the Production with the Injection Moulding Compounder (IMC) on Young’s Modulus Fig. 4 shows the influence of processing in a single heat. It can be clearly seen that the two-step process generates higher Young’s modulus values than the processing in a single heat. This can have various reasons. The shear energy in the two-step process is higher than the introduced energy in the one-step process, because the pelletized compound has to be molten again in the injection moulding machine and so the injection screw introduces shear energy again. Another reason is that the residence time of the single 666

heat process currently is 2.5 times higher than in the two-step process due to the PNC-IMC layout with its long-melt conveying system. Residence time in IMCs depends on the cycle time of the injection moulding machine and, therefore, it is variable to a certain extent. The two interacting mechanisms of exfoliation and degradation are difficult to separate; therefore, much further work is necessary to increase the degree of exfoliation without simultaneous material degradation.

Fig. 4. Effects of one-step and two-step process on Young’s modulus using an extruder length of 52D

Another important result of these measurements is that the amount of compatibilizer and nanofiller has remarkably little effect on Young’s modulus. The increase up to 10 wt% improves the compounds stiffness less than 5%. 2.3 Effects of Additional Masterbatch Process on Young’s Modulus The effect of an additional masterbatch process for the preparation of polypropylene nanocomposites is plotted in Fig. 5. Here, 3in1 is the abbreviation for mixing PP, nanofiller and compatibilizer in one compounding process. This compounding procedure is compared with the masterbatch pre-stage process in which a masterbatch of 50 wt% organoclay and 50 wt% compatibilizer was produced in the twinscrew extruder and then pelletized. The masterbatch pellets were fed into the twin-screw extruder again and then diluted with PP to obtain a compound of 5 wt% nanofiller and 5 wt% compatibilizer. This new compound was either transported directly into the injection moulding machine (one-step process) or pelletized again and then injection-moulded (two-step process). It is clear that the two-step process generates better Young’s modulus values than processing in a single

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heat and that the masterbatch process shows immense potential for the production of nanocomposites. The increase in Young’s modulus compared to virgin polypropylene is approximately 30% for 3in1, twostep process and one-step masterbatch process, whereas it is 40% for the masterbatch two-step process.

Fig. 5. Effects of additional masterbatch process in one-step and two-step processing on Young’s modulus using an extruder length of 44D

In conclusion, all investigations of Young’s modulus have shown that the conventional compounding with the pre-stage masterbatch process and the two-step process of the nanocomposite leads to the best values obtained thus far. Nevertheless, one should bear in mind that the 3in1 one-step process generates only 10 % lower Young’s modulus values than the conventional process but with savings in energy (up to 40%) and time. These results have also an essential aspect for the industry, because one of the biggest challenges remains the intercalation and subsequent delamination of silicate platelets in the polymer matrix. With the upstream masterbatch process stage, exploiting the full potential of nanofillers should be more easily achieved. Furthermore, a masterbatch is much easier to handle than a powder with a bulk weight of 210 g/l. 2.4 Effects of Nanofillers at Various Pressures on Thermal Conductivity The increase in thermal conductivity and the influence of pressure on the thermal conductivity are shown in Fig. 6. To maintain an overview, only three different materials with 0, 5 and 10 wt% at 0 and 750 bar were plotted. It can be clearly seen that the thermal conductivity strongly depends on pressure. The increase from 0 to 750 bar is about 10% in the molten state and about 25% in the solid state.

The increase of thermal conductivity through the use of nanofillers in the molten stage amounts about 7.5% with 10 wt% nanofiller and in the molten state about 10% related to the virgin PP. It is striking to see that the effect of 5% nanofiller in the solid state at 0 bar is much higher than at the 750 bar level.

Fig. 6. Effects of nanofillers and pressure on thermal conductivity using an extruder length of 52D

This is not actually a new phenomenon, but should be considered in the production of injection moulded parts regarding cycle time reduction and cost savings. As can be seen in the cooling time in Eq. (2), the thermal conductivity is linked via the effective thermal diffusivity in inverse proportion to the cooling time; therefore, an increase in thermal conductivity by 10% leads to a reduction of the cooling period by 10%.

tk =

 4   T − T ⋅ ln   ⋅  M W aeff (Tw ) ⋅ π  π   TE − TW s2

2

   , (2)  

where tk is cooling time, s wall thickness, aeff(Tw) effective thermal diffusivity, TM melt temperature, TW mean cavity wall temperature and TE demoulding temperature. The increase of thermal conductivity by 10% still has potential to be improved towards 30%. Therefore, focused research in this area towards better intercalation and subsequent delamination of silicate platelets in the polymer matrix will be performed in the near future. 3 CONCLUSIONS The preparation of PP polymer nanocomposites has been carried out with an intermeshing, co-rotating twin-screw extruder of the PNC-IMC in order to improve Young’s modulus and thermal conductivity; 11 or 13 barrel segments were used for the compounding process, and the PNC-IMC was used

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for processing of injected parts in a single heat, single plasticizing process. As shown in this paper, there are many variables influencing Young’s modulus and the thermal conductivity of polymer nanocomposites. Although it is possible to apply high shear forces and long residence time simultaneously, the results show that an extraordinarily long extruder (52D) and the additional use of a melt pump to generate counterpressure seems to have a negative effect on the improvement of Young’s modulus. Additionally, high screw speed is not essential for the preparation of polymer nanocomposites. More effective regarding to the improvement of Young’s modulus is the use of the pre-stage masterbatch process, where a masterbatch of 50 wt% nanofiller and 50 wt% compatibilizer is produced and further diluted with PP. In conclusion, it is crucial to find a compromise between shear energy input and the residence time. Shear energy is necessary for intercalation and subsequent delamination of silicate platelets in the polymer matrix, while residence time is the most crucial parameter for the diffusion process; however, an overly high residence time may be critical for material degradation. Currently, the chosen setup and process parameter set of the PNC-IMC production line enabled reaching the material property improvements of the two-step masterbatch process. To bring the one-step process beyond the two-step process, further research will be done on: • modification of the screw geometry, • perfect adaptation of the mixing ratio between nanofiller and compatibilizer, • use of compatibilizers based on highly viscous PP-types, and • special treatment of the PNC melt between extruder and melt accumulator (flow induced improvement of intercalation and exfoliation). The results concerning the increase of thermal conductivity are encouraging and support further intensive investigations based on industrial interest. 4 ACKNOWLEDGEMENTS We would like to thank Borealis, Linz for the provided PP, Dr. G. Pinter for the provided testing equipment and Dr. G.R. Berger for his valuable comments on the manuscript. 668

5 REFERENCES [1] Jensen, R. (2001). Synergien intelligent genutzt. IMC-Spritzgießcompounder erhöht Wertschöpfung. Kunststoffe, vol. 91, p. 40-45. [2] Sieverding, M., Bürkle, E., Zimmet, R. (2005). IMC Technology opens up new fields of application. Kunststoffe Plast Europe, vol. 8, p. 34-37. [3] Bürkle, E., Scheel, G., Damedde, L. (2009). Energyefficient processing of natural fibre-reinforced plastics. Kunststoffe Plast Europe, vol. 2, p. 39-44. [4] Wobbe, H. (2003). Trends und Visionen der Spritzgießtechnik. Kunststoffe, vol. 10, p. 60-65. [5] Vaia, R.A., Jandt, K.D., Kramer, E.J., Giannelis, E.P. (1996). Microstructural evolution of melt intercalated polymer-organically modified layered silicates nanocomposites. Chemistry of Materials, vol. 8, no. 11, p. 2628-2635, DOI: 10.1021/cm960102h. [6] Mignoni, M.L., Silva, J.V.M., De Souza, M.O., Mauler, R.d.S., De Souza, R.F., Bernando-Gusmao, K. (2011). Polyethylene-montmorillonite nanocomposites obtained by in situ polymerization of ethylene with nickel-diimine catalysts. Journal of Applied Polymer Science, vol. 122, no.3, p. 2159-2165, DOI:10.1002/ app.34358. [7] Kracalik, M., Laske, S., Gschweitl, M., Friesenbichler, W., Langecker, G.R. (2009). Advanced compounding: extrusion of polypropylene nanocomposites using the melt pump. Journal of Applied Polymer Science, vol. 113, no. 3, p. 1422-1428, DOI:10.1002/app.29888. [8] Battisti, M.G., Friesenbichler, W. (2012). Injection molding compounding of PP polymer nanocomposites. Proceedings of International Conference PMI 2012, Ghent, p. 134-140. [9] Sinha Ray, A., Okamoto, M. (2003). Polymer/ layered silicate nanocomposites: a review from preparation to processing. Progress in Polymer Science, vol. 28, no. 11, p. 1539-1641, DOI:10.1016/j. progpolymsci.2003.08.002. [10] Kiliaris, P., Papaspyrides, C.D. (2010). Polymer/layered silicate (clay) nanocomposites: An overview of flame retardancy. Progress in Polymer Science, vol. 35, no. 7, p. 902-958, DOI:10.1016/j.progpolymsci.2010.03.001. [11] Oswald, T.A., G. Menges (1996). Material Science of Polymers for Engineers. Carl Hanser Verlag Munich, Vienna, New York. [12] Laske, S., Duretek, I., Witschnigg, A., Mattausch, H., Tscharnuter, D., Holzer, C. (2012). Influence of the degree of exfoliation on the thermal conductivity of polypropylene nanocomposites. Polymer Engineering & Science, vol. 52, no.8, p. 1749-1753, DOI:10.1002/ pen.23121. [13] Sibalija, T., Majstorovic, V., Sokovic, M. (2011). Taguchi-Based and Intelligent Optimisation of a MultiResponse Process Using Historical Data. Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 4, p.357-365, DOI:10.5545/sv-jme.2010.061.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 669-676 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.1003 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-04-10 Accepted for publication: 2013-05-06

Methods for Improved Flexural Mechanical Properties of 3D-Plotted PCL-Based Scaffolds for Heart Valve Tissue Engineering

Ragaert, K. – De Somer, F. – Van de Velde, S. – Degrieck, J. – Cardon, L. Kim Ragaert1,2,* – Filip De Somer3 – Stieven Van de Velde1 – Joris Degrieck2 – Ludwig Cardon1,2 1 University

College Ghent, Associated Faculty of Applied Engineering Sciences, Belgium University, Faculty of Engineering and Architecture, Belgium 3 University Hospital Ghent, Heart Centre, Belgium

2 Ghent

While porous poly-ε-caprolactone (PCL) scaffolds can be manufactured through 3D plotting with high regularity and reproducibility, it has been a challenge in previous research to mimic the highly flexible behaviour of the natural valve leaflets. In this study, an investigation is made of two separate approaches for the improved flexibility of 3D plotted PCL scaffolds for heart valve leaflets. Firstly, the scaffold geometry is radically altered towards a very open woven-like structure by adequately adapting the processing parameters during 3D plotting. Secondly, the base material itself is altered by blending a fraction of low-molecular weight poly-ethylene-oxide (PEO) into the PCL polymer. The scaffolds are 3D plotted for both series and their flexibility is evaluated in a uni-axial indentation experiment. The results are compared to those of the natural valve tissue and it is found that both approaches result in the desired reduction of the stiffness of the scaffold. Keywords: scaffolds, mechanical properties, flexure, heart valves, tissue engineering, 3D plotting

0 INTRODUCTION A growing application of biodegradable polymers is the manufacture of scaffolds for tissue engineering, an advanced multidisciplinary research field meant to meet the growing demand for donor organs and tissues [1] to [3]. According to the tissue engineering principle, biodegradable scaffolds are used as support structures for the culture of the patient’s harvested cells in an in vitro environment, so as to (re)create healthy tissues meant to replace diseased ones. While this neo-tissue grows, the scaffold slowly degrades into nontoxic components, eventually leaving only the new, functional and healthy tissue behind. This final construct can be implanted into the patient and will not solicit any rejection, because the cells used are the patient’s own. When looking to cardiovascular applications and more specifically to leaflets for heart valve replacement, the elastic-mechanical properties of the scaffold are just as important as the biodegradability and non-toxicity of the material. The leaflet must be strong enough to withstand the blood flow and at the same time be able to follow the elastic movement of a natural valve. In fact, when researching such valve scaffolds, we did not look for the strongest possible scaffold but for the highest possible compliance with the elastic-mechanical behaviour of natural tissue. Adherence to such compliant behaviour will provide the correct mechanical stimuli for the differentiation of the seeded cells.

Poly-ε-caprolactone (PCL) was selected as a base material for the scaffolds in this study because it is relatively flexible in comparison to other aliphatic polyesters and as such is considered more suitable for use with scaffolds for cardiovascular applications [4]. Moreover, it displays very good thermal stability, with a degradation temperature situated in the range of 280 to 330 °C [5], which makes it very suitable for use with a melt processing technique like the micro-extrusion for 3D plotting proposed in this work. Finally, PCL is FDA-approved (by Food and Drug Administration); it is considered to be compatible with both hard and soft tissues [6] to [8] and will degrade slowly in the human body over a period of 24 to 36 months [9] to [11]. The main disadvantage of PCL is the strongly hydrophobic nature of its surface [12] to [14], which leads to nonspecific protein adsorption. In previous research, different series of porous 4-layer PCL scaffolds were manufactured through 3D-plotting, with individual filament sizes from 127 to 410 µm and spacing between filaments 1.8 times the filament size. It was found that while the scaffolds with the thinnest filaments best approximated the uniaxial mechanical properties of the natural valve cusp tissue, they generally remained too stiff to be viable as valve leaflet scaffolds [15] and [16]. Therefore, the focus of the current research is to further reduce the stiffness of the PCL scaffolds, by adapting either the scaffold geometry to a so-called woven-like structure or adapting the scaffold base material by blending PCL with low-molecular weight poly-ethylene-oxide

*Corr. Author’s Address: University College Ghent, Associated Faculty of Applied Engineering Sciences, Voskenslaan 362, 9000 Ghent, Belgium, kim.ragaert@ugent.be

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(PEO), which is deemed an adequate choice for three reasons [17] to [19]: i) The low molecular weight PEO fraction can reduce the mechanical properties of PCL, which will bring them closer to those of the natural valve cusp; ii) PEO is hydrophilic in nature and can attenuate the hydrophobicity of PCL, a property which is known to be unbeneficial to the attachments of cells to the scaffolds due to non-specific protein absorption [20]; iii) PEO is biocompatible and water-soluble. 1 MATERIALS AND METHODS 1.1 Materials For the scaffolds with woven-like geometry, PCL CAPA 6500 from Perstorp (UK) is used. The manufacturer reports a weight-averaged molecular weight of 84500 Da, a polydispersity of 1.78 and a maximum crystalline fraction of 56%. The PCL-PEO blends were developed by the Polymer Chemistry and Biomaterials (PBM) group at Ghent University: three solutions were prepared each with a different concentration of PEO: one, five and ten weight-percent (wt%) PCL (Sigma AldrichBelgium, Mn = 80000 Da) and PEO (Sigma AldrichBelgium, Mn = 2000 Da) were first dissolved in chloroform, after which a precipitation reaction in cooled diethylether was performed. The different blends are hereafter referred to as PCL-PEO1, PCLPEO5 and PCL-PEO10, for their respective PEO contents. Characterisation of the composition of the synthesized PEO-PCL compounds was carried out using proton-nuclear magnetic resonance (1H-NMR), in which 10 mg of the blended polymer was dissolved in deuterated chloroform (CDCl3) for the measurements. Tetramethylsilane (TMS) was used as an internal standard for calibrating the chemical shift. The experiments were performed on a WH 300 MHz 1H-NMR apparatus (Bruker, Belgium) and were conducted in triplicate for each concentration type.

the BioScaffolder (SysEng, Germany), a 3D plotting device. The workings of this apparatus have been previously detailed [21]. In short, the machine melts the thermoplastic polymer in a mobile dispenser head and extrudes thin filaments that are deposited on a plotting table. The final product is built up layer-bylayer. Concerning the porous geometry, the mounted extrusion needle had a diameter of 127 µm and a strand distance (the centre-to-centre distance between two adjacent filaments) of 900 µm was set. The most important processing parameters include a processing temperature of 125 °C, plotting speed of 85 mm/min, and an extrusion screw speed of 11 rpm. From earlier research [22], it had become apparent that, as a rule-of-thumb, the strand distance should remain below twice the filament diameter for the different layers to provide sufficient support for the subsequent filaments to remain taut. By increasing the strand distance well above this practical limit, the filaments will sag into the underlying pores. Additionally, the plotting speed used is in fact too slow in comparison to the extrusion rate provided by the screw, allowing for sufficient material deposition to plot the sagging filaments without thinning or breaking. Together, these altered processing conditions create the woven-like structure of the scaffold, which is shown in Fig. 1. The scaffolds were plotted to a square geometry of 15×15 mm², with three layers stacked according to a 0 to 90 pattern.

1.2 3D Plotting of Woven-Like Scaffolds Compared to the previous research, the porosity of the scaffolds is increased, the number of layers is reduced to three, and the plotting speed is adjusted so that the polymer filaments are placed similarly to those in a woven structure. These scaffolds are characterised by the term “woven-like geometry” for the purpose of this manuscript. They were manufactured using 670

Fig. 1. The woven-like structure of the first series of modified PCL scaffolds

1.3 3D Plotting of PCL-PEO Scaffolds PCL-PEO scaffolds were produced on the same BioScaffolder apparatus. Only limited amounts of

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material were available, about four to five grams per blend type. As an optimization of the parameter set had to be done using up as little material as possible, the larger needle section of 200 µm was chosen in order to avoid possible material flow problems. The strand distance was set at 360 µm and the extrusion temperature at 115 °C. Processing parameters were further fine-tuned for each blend. In general, the plotting speed and extruder screw ratio were reduced for higher PEO contents, which displayed a more viscous behaviour. The dispenser head was disassembled and cleaned for every material changeover. Scaffolds were plotted to a square geometry of 15×15 mm², with four layers stacked according to a 0 to 90 pattern. Some scaffolds were also manufactured using pure PCL to obtain a reference value for the mechanical properties of the unblended PCL. 1.4 Mechanical Properties The experimental setup and the principle of the flexural indentation test are shown in Fig. 2. The tissue is placed over a round hole, where it is held in place by a half-sunk rubber ring and fixed in place by a clamping plate. The tissue is then indented by a descending ball probe, to which the central hole has been aligned. Three tissue properties are derived from the test: i) Extension at break [mm] (EXT): the depth of indentation by the probe at the moment of rupture (defined as a sudden decrease in load by 50%); ii) Maximum load [N] (ML): the maximum load which can be applied to the cusp prior to rupture; iii) Stiffness parameter [N/mm] (ST): the slope of the linear section of the indentation-load curve. This is a measure for the stiffness of the tissue. ML and EXT describe the ultimate properties of the cusps. ST on the other hand describes the flexural response of the leaflet under an applied load and provides the most valuable information when considering the physiological functionality of the tissue. All tests were performed on an LF Plus Universal material tester (Lloyd Instruments, UK), with a load cell of 1 kN, a ball probe of 4.45 mm in diameter and an indentation speed of 25 mm/min. The ball probe was lowered onto the centre of the hole until perforation, which was defined as a sudden decrease in load by 50%. A lower force limit of 0.1 N was set as the zero-indentation point for the scaffold. Force and displacement of the ball probe were recorded and used for the calculation of ML, EXT and ST. Per series,

three scaffolds were tested and their results averaged. All results are represented as the mean ± the standard deviation.

Fig. 2. Principles of the uni-axial indentation test. Reprinted from [23] with permission (Elsevier)

The results are compared to the respective values for the natural valve cusp tissue [23], which can be found in Table 1. Table 1. Mechanical properties for the PCL-PEO scaffolds, compared to those of the natural cusp % PEO 1 5 10 natural cusp

EXT [mm] 2.72±0.10 2.60±0.08 2.57±0.12 3.26±0.64

ML [N] 28.26±0.50 24.88±1.32 24.34±1.58 13.05±4.48

ST [N/mm] 11.65±0.20 11.12±0.02 10.42±0.49 5.97±1.69

1.5 Characterization of PCL-PEO Scaffolds Characterization experiments on the PCL-PEO scaffolds were carried out by the PBM group. A visualization of the scaffold surfaces was carried out using a scanning electron microscope (SEM, JEOL JSM-5600 microscope) with the apparatus in secondary electron mode (SEI). Prior to analysis all samples were coated with a gold layer of approximately 20 nm. X-ray photon spectroscopy (XPS) is a technique that allows the determination of the elemental composition of the outermost 10 to 15 nm of a surface. XPS measurements were performed with an ESCA S-probe VG monochromatised spectrometer (with an Al Kα source X-ray source of 1486 eV) on the scaffolds of all three blend-types. A survey scan spectrum was collected and from the peak-area ratios, the relative elemental composition of the material’s top layer was determined.

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2 RESULTS AND DISCUSSION 2.1 Properties of the woven-like scaffolds The results for the woven-like scaffolds are shown in Fig. 3, in which slanted S-shaped curves may be discerned for all scaffolds. After an initial low-force response, the scaffolds stiffen into a linear flexural behaviour, which is topped off by the final plateau, representative of the rupture of the scaffold by the ball probe. The deduced mechanical properties for the woven-like scaffolds are: EXT = 2.23±0.10 mm, ML = 5.51±0.07 N and ST = 3.79±0.17 N/mm.

around 3.6 ppm, which can be seen to increase with rising amounts of PEO, is attributed to the methylene protons of the –CH2-CH2O- in the PEO chain structure [24].

Fig. 3. Results for the woven-like scaffolds

As described earlier, the stiffness parameter ST is considered to best reflect the flexural behaviour of the leaflet. The ST value is significantly lower than that of the natural cusp tissue. This is, however, considered a positive evolution and an important one at that. In previous research [15], the lowest value achieved (4-layered PCL scaffolds with the same filament size) was 11 N/mm, well above the range of the natural tissue. It was also found that, with the 3D plotted scaffolds, the stiffness of the construct is easily increased by lowering the strand distance or by using thicker individual filaments. The limitation has always been that the stiffness could not be sufficiently reduced. By creating an open woven-like structure, a lower ST value has been realised for the first time. It is expected that this value can easily be raised again by modifying the geometry towards a more dense structure. 2.2 Composition of the PCL-PEO Blends Examples of the 1H-NMR spectra of the three blends are shown in Fig. 4. The peak at a chemical shift of 672

Fig. 4. Sample 1H-NMR spectra; a) PCL-PEO1, b) PCL-PEO5, c) PCL-PEO10

The resulting calculations of the weight percentage of PEO in the different blends from the 1H-NMR spectra yield the following results: 0.90±0.05 % for PCL-PEO1, 4.53±0.42% for

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PCL-PEO5 and 9.21±0.45% for PCL-PEO10. The preparation of the PEO-PCL blends can therefore be considered successful, with the resulting PEO concentrations within 10% of their set value. 2.3 Properties of the PCL-PEO Blended Scaffolds The PCL-PEO scaffolds display a similar type of flexural behaviour, but are clearly more resilient to higher levels of load. The load-displacement curves of PCL-PEO10 are shown in Fig. 5 as an example.

scaffold had been placed within a medium for cell culturing. Disintegration of the PEO fraction will result in a porous PCL scaffold that must also remain structurally intact and fully functional. As such, PEO concentrations can never be too high within the blend. It would be advisable for further research to include a study of the effect of PEO removal from the scaffold on the structure and the mechanical properties.

Fig. 6. Relation between ST and wt% PEO

2.3 Characterization of the PCL-PEO Scaffolds Fig. 5. Results for the load-displacement for PCL-PEO10 scaffolds

The calculated values of EXT, ML and ST are listed in Table 1 per blend type. All three properties can be seen to decrease with an increasing percentage of PEO mixed into the blend. The decrease in ST, the parameter that is most relevant for the flexural behaviour under operating conditions, appears to be directly related to the amount of PEO in the blend. This is illustrated in Fig. 6, where the equation of the linear fit to the curve is also displayed. It implies that for the unblended PCL, the ST value would be 11.79 N/mm. This value matches the ST value found for the unblended PCL scaffolds, which was 11.64±0.20 N/mm. From these results, it is apparent that the addition of a low-molecular weight PEO fraction leads to a proportional reduction in the compound’s – and therefore the scaffold’s – stiffness. It may be possible to use this plasticizing effect of PEO in order to obtain mechanical properties more like those of the natural tissue. It must be remarked that PEO is soluble in water and will dissolve within 24 hours if submerged [17]. This means that the PEO component should not be counted upon for structural integrity once the

SEM-images of the PCL-PEO struts are shown in Fig. 7 to 9 for 1, 5 and 10 wt% of PEO respectively. For the PCL_PEO1 scaffold, a more or less smooth surface is seen, with some protruding material regions in the middle left of the image. This local surface roughness is caused by precipitates of PEO, the amount of which can be observed to increase sharply as the PEO concentration rises; the surface of the PCL_PEO10 scaffolds is covered almost entirely by these precipitates. Because of the roughness the precipitates induce on the surface, a higher signal in the SEM-image is obtained, which results in white pigments on the struts in the images with magnification ×75. This relatively large impact of the PEO fraction on the surface condition of the filaments would suggest that the PEO is mostly located at the surface of the extruded strut and no longer homogenously distributed throughout it. This assumption is confirmed by the XPS results. The different high-resolution carbon XPS spectra of PCL-PEO1, PCL-PEO5 and PCL-PEO10 are shown in Fig. 10 . The spectra of these blends are compared to those of pure PCL and PEO, which have been reported by Manso et al. [25].

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It was therefore concluded that the PEO fraction must have segregated within the blend and moved towards the surface, a phenomenon which is known as viscous encapsulation (phase separation) [27] and has been previously noted in the literature on PCL-PEO polymers [28]. Fig. 7. SEM-images of the PCL-PEO1 scaffolds; a) top view (×75), b) view of a strut (×500)

Fig. 8. SEM-images of the PCL-PEO5 scaffolds; a) top view (×75), b) view of a strut (×500)

Fig. 10. XPS-spectra of; a) PCL-PEO1, b)PCL-PEO5, and c) PCLPEO10

Fig. 9. SEM-images of the PCL-PEO10; (a) top view (×75), (b) view of a strut (×500)

The peak at 289.1 eV, which can be attributed to carboxylic functional groups (O-C=O), is considered indicative of PCL; it is strongly pronounced in the spectrum of pure PCL [26]. The peak at 286.5 eV is very dominant for the spectrum of pure PEO and is considered to be representative of the presence of PEO, as it dominates the C-O bond [26]. Finally, the peak at 284.8 eV is found in the spectrum of both pure materials, as it indicates bonds which both polymers have in their structure (mainly C-H) [26]. With rising amounts of PEO added to the blend, the relative peak area representing PEO can be seen to increase at the cost of the peak indicating the presence of PCL. Table 2 shows the assignments for the quantitative deconvolution of the carbon spectra of the PCLPEO10. The peak at 286.5 eV, which indicates the presence of PEO, corresponds to over 60% of the material surface. This is disproportionately high with regard to the amount of the blended polymer (10%). 674

Table 2. The position and representative atom concentration of the peaks of the XPS-spectrum of PCL-PEO10 (Fig. 10c) Position (eV) 284.8 286.5 289.1

% at conc. 26.74 62.43 10.82

Bond C-O,C-C/C-H C-C/C-H,C-O O-C=O

Presence of PEO, PCL PEO PCL

2.4 Prospects The two approaches of adapting the geometry and the base material have been successful in improving the flexibility of the PCL-based leaflet scaffolds. By creating an extremely open woven-like structure it was even possible to create mechanical properties lower than those of the natural tissue, a feat that had not been accomplished for 3D-plotted PCL-based scaffolds prior to this research. Tried separately, it was evident that addition of a low-molecular weight PEO fraction into the polymer blend also resulted in a reduction of the high ST value of the scaffolds of more conventional and coarser geometries. As such, both approaches have been shown valid for further investigation and the combination of the

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two approaches should allow for a broad range of flexible-mechanical properties to be realized. However, some relevant expected effects of the PEO blending have not been investigated, such as the leaching out of the PEO fraction in a physiological environment. This would result in an additional in vitro/in vivo weakening of the construct and a more micro-porous scaffold surface, the latter of which may actually be beneficial to cell attachment. Furthermore, these experiments have shown that even though blends of PCL-PEO are manufactured as a homogenous mix, processing them into filaments for 3D plotting by micro-extrusion will cause the PEO fraction to segregate out onto the surface of the struts. As hydrophilicity and surface roughness are observed to increase with the inclusion of more PEO, this encapsulation effect may be beneficial to the cellattractiveness of plotted PCL-based scaffolds. 3 CONCLUSIONS In this exploratory study it has been found that the stiffness of PCL-based scaffolds can be reduced by (i) adapting the scaffold geometry to an open woven-like structure, as well as by (ii) altering the base material by blending in lower-strength polymer fractions. By creating a more open structure and allowing the individual filaments to sag into the underlying pores it has – for the first time within the scope of this research – been possible to manufacture PCL scaffolds with flexural stiffness values that are lower than those of the natural tissue, thus overcoming the ongoing problem of PCL’s higher stiffness than the natural cusp. Additionally, the blending of PEO into PCL has been found to further increase the flexibility of the base material at a rate which is proportional to the amount of PEO added. Moreover, the shorter PEO chains have been found to migrate towards the surface of the extruded filaments, where they can contribute to the hydrophilicity of the scaffold surface. In future, combining these aspects of geometry and material will allow for a more exact mimicking of the natural tissue’s flexible behaviour by leaflet scaffolds. 4 ACKNOWLEDGEMENTS The authors wish to acknowledge Professor Peter Dubruel and his research team of the PBM lab at Ghent University for supplying and characterizing the PCL-PEO blends. This research was funded by University College Ghent.

5 REFERENCES [1] Langer, R., Vacanti, J.P. (1993). Tissue Engineering. Science, vol. 260, no. 5110, p. 920-926. [2] Mikos, A.G., Temenoff, J.S. (2000). Formation of highly porous biodegrdable scaffolds for tissue engineering. Electronic Journal of Biotechnology, vol. 3, no. 2, p. 1-6, DOI:10.2225/vol3-issue2-fulltext-5. [3] Isenberg, B.C., Wong, J.Y. (2006). Building structure into engineered tissues. Materials Today, vol. 9, no. 12, p. 54-60, DOI:10.1016/S1369-7021(06)71743-6. [4] Brody, S., Pandit, A. (2007). Approaches to heart valve tissue engineering scaffold design. Journal of Biomedical Materials Research Part B-Applied Biomaterials, vol. 83B, no. 1, p. 16-43, DOI:10.1002/ jbm.b.30763. [5] Sivalingam, G., Madras, G. (2003). Thermal degradation of poly (epsilon-caprolactone). Polymer Degradation and Stability, vol. 80, no. 1, p. 11-16, DOI:10.1016/S0141-3910(02)00376-2. [6] Eshraghi, S., Das, S. (2010). Mechanical and microstructural properties of polycaprolactone scaffolds with one-dimensional, two-dimensional, and three-dimensional orthogonally oriented porous architectures produced by selective laser sintering. Acta Biomaterialia, vol. 6, no. 7, p. 2467-2476. DOI:10.1016/j.actbio.2010.02.002. [7] Guarino, V., Causa, F., Taddei, P., di Foggia, M., Ciapetti, G., Martini, D., Fagnano, C., Baldini, N., Ambrosio, L. (2008). Polylactic acid fibre-reinforced polycaprolactone scaffolds for bone tissue engineering. Biomaterials, vol. 29, no. 27, p. 3662-3670. DOI:10.1016/j.biomaterials.2008.05.024. [8] Wang, Z.W., Su, J.C., Ma, Y.H., Zhang, X., Cao, L.H., Li, M. (2010). Preparation and Properties of Nano Calcium Deficient Apatite/Poly (epsilon-caprolactone) Composite Scaffold. Journal of Inorganic Materials, vol. 25, no. 5, p. 500-506, DOI:10.3724/ sp.j.1077.2010.00500. [9] Stankus, J., Guan, J., Wagner, W. (2004). Biodegradable Elastomers. Wnek, G., Bowlin, G. (eds.), Encyclopedia of Biomaterials and Biomedical Engineering, Marcel Dekker: Richmond. p. 484-494. [10] Lam, C.X.F., Hutmacher, D.W., Schantz, J.T., Woodruff, M.A., Teoh, S.H. (2009). Evaluation of polycaprolactone scaffold degradation for 6 months in vitro and in vivo. Journal of Biomedical Materials Research Part A, vol. 90A, no. 3, p. 906-919, DOI:10.1002/jbm.a.32052. [11] Nair, L.S., Laurencin, C.T. (2007). Biodegradable polymers as biomaterials. Progress in Polymer Science, vol. 32, no. 8-9, p. 762-798. [12] Lee, H., Kim, G. (2010). Biocomposites Electrospun with Poly(epsilon-caprolactone) and Silk Fibroin Powder for Biomedical Applications. Journal of Biomaterials Science-Polymer Edition, vol. 21, no. 13, p. 1687-1699, DOI:10.1163/09205060 9x12548956645680.

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[13] Shan, X.Q., Yuan, Y., Liu, C.S., Tao, X.Y., Sheng, Y., Xu, F. (2009). Influence of PEG chain on the complement activation suppression and longevity in vivo prolongation of the PCL biomedical nanoparticles. Biomedical Microdevices, vol. 11, no. 6, p. 1187-1194, DOI:10.1007/s10544-009-9336-2. [14] Yao, K.J., Wang, J.F., Zhang, W.J., Lee, J.S., Wang, C.P., Chu, F.X., He, X.M., Tang, C.B. (2011). Degradable Rosin-Ester-Caprolactone Graft Copolymers. Biomacromolecules, vol. 12, no. 6, p. 2171-2177, DOI:10.1021/bm200460u. [15] Ragaert, K., De Somer, F., De Baere, I., Cardon, L., Degrieck, J. (2011). Production & Evaluation of PCL Scaffolds for Tissue Engineered Heart Valves. APEM, vol. 6, no. 3, p. 163-172. [16] Ragaert, K., Cardon, L., De Somer, F., Degrieck, J. (2010). PCL leaflets for tissue engineered heart valves: 3D plotting and mechanical properties. PMI, p. 308312. [17] Kim, G., Park, J.H., Park, S. (2007). Surface-treated and multilayered poly(epsilon-caprolactone) nanofiber webs exhibiting enhanced hydrophilicity. Journal of Polymer Science Part B-Polymer Physics, vol, 45, no. 15, p. 2038-2045, DOI:10.1002/polb.21211. [18] Guarino, V., Netti, P.A., Ambrosio, L. (2008). Development of highly oriented porous structures by PCL/PEO co-continuous blends. Acierno, D., Damore, A., Grassia, L. (eds), IVth International Conference on Times of Polymers. p. 201-203. [19] Reignier, J., Huneault, M.A. (2006). Preparation of interconnected poly(epsilon-caprolactone) porous scaffolds by a combination of polymer and salt particulate leaching. Polymer, vol. 47, no. 13, p. 47034717, DOI:10.1016/j.polymer.2006.04.029. [20] Desmet, T., Billiet, T., Berneel, E., Cornelissen, R., Schaubroeck, D., Schacht, E., Dubruel, P. (2010). Post-Plasma Grafting of AEMA as a Versatile Tool to Biofunctionalise Polyesters for Tissue Engineering. Macromolecular Bioscience, vol. 10, no. 12, p. 14841494, DOI:10.1002/mabi.201000147.

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[21] Ragaert, K., Cardon, L., Dekeyser, A., Degrieck, J. (2010). Machine design & processing considerations for the 3D plotting of thermoplastic scaffolds. Biofabrication, vol. 2, no. 1, DOI:10.1088/17585082/2/1/014107. [22] Ragaert, K. (2011). Micro-Extrusion of Thermoplastics for 3D Plotting of Scaffolds for Tissue Engineering. Ghent University, Department of Materials Science and Engineering, Ghent. [23] Ragaert, K., De Somer, F., Somers, P., De Baere, I., Cardon, L., Degrieck, J. (2012). Flexural mechanical properties of porcine aortic heart valve leaflets. Journal of the Mechanical Behavior of Biomedical Materials, vol. 13, no. 0, p. 78-84, DOI:10.1016/j. jmbbm.2012.04.009. [24] Ma, G., Miao, B., Song, C. (2010). Thermosensitive PCL-PEG-PCL hydrogels: Synthesis, characterization, and delivery of proteins. Journal of Applied Polymer Science, vol. 116, no. 4, p. 1985-1993, DOI:10.1002/ app.31654. [25] Manso, M., Valsesia, A., Lejeune, M., Gilliland, D., Ceccone, G., Rossi, F. (2005). Tailoring surface properties of biomedical polymers by implantation of Ar and He ions. Acta Biomaterialia, vol. 1, no. 4, p. 431-440, DOI:10.1016/j.actbio.2005.03.003. [26] Cheng, Z., Teoh, S.-H. (2004). Surface modification of ultra thin poly ([var epsilon]-caprolactone) films using acrylic acid and collagen. Biomaterials, vol. 25, no. 11, p. 1991-2001, DOI:10.1016/j. biomaterials.2003.08.038. [27] Moroni, L. (2006). A Mechanistic Approach to Design Smart Scaffolds for Tissue Engineering. University of Twente: Enschede. [28] Li, G.M., Cai, Q., Bei, J.Z., Wang, S.G. (2002). Relationship between morphology structure and composition of polycaprolactone/poly (ethylene oxide)/ polylactide copolymeric microspheres. Polymers for Advanced Technologies, vol. 13, no. 9, p. 636-643, DOI:10.1002/pat.287.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 677-682 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.1000 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-04-26 Accepted for publication: 2013-05-17

Monitoring of Injection Moulding of Thermoplastics: Adopting Pressure Transducers to Estimate the Solidification History and the Shrinkage of Moulded Parts Speranza, V. – Vietri, U. – Pantani, R. Vito Speranza – Umberto Vietri – Roberto Pantani*

University of Salerno, Department of Industrial Engineering, Italy In this work, a series of injection moulding tests were conducted using a general purpose PolyStyrene, PS, and changing the holding pressure, injection temperature and cavity dimensions. The pressures at the interface with the mould at several positions along the flow-path were measured by means of pressure-temperature transducers. The samples were measured after moulding to determine dimensional accuracy, which was taken to be the target quality parameter. The pressure profiles obtained were then analysed using a recently developed procedure that is able to estimate the local solidification history from pressure measurements. The local average solidification pressure, namely the average along the thickness direction of the pressures at which each layer solidifies, could thus be estimated. This parameter is known to correlate well with shrinkage and thus a master-curve can be created that can be adopted to monitor the quality of the moulded part on line. Keywords: injection moulding, quality control, shrinkage

0 INTRODUCTION The development of advanced techniques for monitoring and controlling the injection moulding process is a strategic issue for industries involved in polymer processing operations [1]. In-mould sensors can be very helpful for on line measurements and hence for monitoring and control purposes [2]. Ultrasonic [3] and capacitive [4] sensors have been applied to measure the part weight, optical fibers [5] have proven to be able to measure thickness shrinkage, and strain gages [6] and [7] have been adopted to follow the shrinkage evolution from the instant of first solidification. However, these methods are normally limited to scientific purposes: industries are traditionally disinclined to introduce moulds instrumented with a suitable number of sensors that could effectively monitor the injection moulding process. In spite of this, traditional hardware-based temperature and pressure transducers have been widely employed in industry; however, there is no clear correlation between the measured evolution of temperature and pressure and the product quality. [8]. Indeed, it can be easily demonstrated [9] that even the complete pressure curve cannot be adopted as a suitable parameter to fully describe shrinkage, and a criterion based on the reproducibility of the pressure profiles can cause the rejection of parts that are consistent with quality parameters. On the other hand, the local average solidification pressure Ps (the average over the thickness of the pressures at which each layer solidifies locally) was demonstrated to be a suitable parameter for quality part description in the injection moulding process [9]. Determining

the local average solidification pressure Ps requires the determination of both the local pressure history and the local solidification history. In spite of the recent attempts made to experimentally determine the temperature profile along the thickness direction of a moulding [10], the local solidification history is not experimentally obtainable, and thus it is necessary to perform a simulation of the whole injection moulding test in order to obtain it. In a previous work [11], a procedure was introduced which allows the adoption of the measured pressure evolution to make an estimation of the solidification profile and thus of the average solidification pressure. In this work the procedure was applied to injection moulding tests carried out with a general purpose PS. 1 EXPERIMENTAL 1.1 Material The material adopted was a general purpose Polystyrene (Styron PS 678E) supplied by Dow Chemicals. A complete characterization of the resin can be found in the literature [12] to [14]. 1.2 Moulding Conditions Four series of experiments were carried out; each one characterized by the variation of a single parameter (namely, injection temperature and cavity thickness) with respect the reference series (series A). For each series, several holding pressures were adopted from

*Corr. Author’s Address: University of Salerno, Department of Industrial Engineering, via Ponte don Melillo, Fisciano (SA), Italy, rpantani@unisa.it

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80 bar up to more than 1000 bar. A summary of the moulding conditions is reported in Table 1. For each of the moulding conditions, measurements for pressure evolution data were taken at three positions inside the cavity at 15, 60 and 105 mm from the gate. The cavity is 120 mm long and these positions (referred to as P2, P3 and P4, respectively) are located 15 mm from the cavity entrance, in the center of the cavity, and at 15 mm from the cavity tip. The other two transducers were located inside the injection chamber (pos. P0) and just upstream from the gate (pos. P1). A schematic view of the cavity adopted for all the moulding tests is shown in Fig. 1.

solidification conditions and is less sensitive to the presence of constraints with respect to length shrinkage and to mould deformation with respect to thickness shrinkage. The shrinkage was defined as the relative difference between the mould and the product width (both evaluated at 25 °C), as defined by the following equation:

si = (di – dsi) / di ,

(1)

where s was the shrinkage, ds the sample local width, and d the local cavity width (Fig. 1). The subscripts indicate the position inside the cavity where width shrinkage was measured, namely at the positions of pressure transducers inside the cavity (si indicates the transducer position Pi). This means that for each moulding condition, three results for shrinkage were obtained, each one related to a particular local history of temperature and pressure. 2 FROM PRESSURE MEASUREMENTS TO SOLIDIFICATION PROFILE For amorphous polymers, it can be demonstrated [9] that, after the local solidification time tsol, the local pressure profile follows an exponential law, namely:

 t P = P∞ + A exp  −  τ

  , (2) 

where τ is a the characteristic cooling time: 2

Fig. 1. Schematic view of the geometry adopted for all the moulding tests; the dimensions considered for shrinkage are indicated Table 1. Summary of moulding conditions: for each series of experiments, the parameter characterising the series is reported in bold Series A B C D

Ph [bar] 70 to 1300 70 to 1100 140 to 1000 90 to 1200

th [s] 12 12 12 12

tinj [s] 0.45 0.45 0.45 0.45

Tinj [°C] 200 220 240 200

Tmould [°C] 25 25 25 25

Thickness [mm] 2 2 2 4

Ph stands for holding pressure, th for holding time, tinj for injection time, Tinj for injection temperature, Tmould for mould temperature. 1.3 Shrinkage Measurements In this work, we adopted width shrinkage as the quality parameter. This choice was made because width shrinkage is strongly dependent on the local 678

2 2 L . (3) τ =  π  α

In Eq (3), L is the local half-thickness and α is the thermal diffusivity of the material. The parameter P∞ in Eq. (2) is a constant, which represents the pressure that would be reached at very long times, when the polymer reaches thermal equilibrium with the mould. If P∞ is positive, it coincides with the residual pressure. However, it can also be negative, obviously losing any physical meaning, if the solid polymer detaches from the cavity walls. If Eq. (2) holds true, a non-linear regression can be carried out on the experimental pressure curve, aimed at determining the value of tsol (i.e. the time after which the pressure curve is well described by an exponential curve as in Eq. (2)), and of τ (i.e. the characteristic time for that exponential curve). The results of the procedure are shown in Figs. 2 and 3 for some of the moulding tests carried out. These figures show that after a few seconds, the experimental pressure evolutions are well described by the exponential curves at all positions and for all

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conditions. A sensitivity analysis showed that on changing the maximum allowed error between the best fitting curve and the experimental curve, the solidification times found by the procedure change by about 1 s. The procedure captures the fact that by increasing the injection temperature, the solidification times generally increase. Furthermore, differences in the solidification times at the different positions were detected and it was found that, in most cases, the solidification took place at position P4 (at the cavity tip) at earlier times than for positions P2 and P3, as expected on the basis of the fact that the temperature is higher closer to the injection point. As expected, it was also found that by increasing the cavity thickness the solidification times significantly increase. The values found for τ are reported in Fig. 4 for all the pressure curves analysed in this work. All the values found for τ collapse around two numbers (indicated as horizontal dotted lines), one for the series obtained with the thicker cavity and one for the series obtained with the thinner cavity, which differ of a factor of about four. Considering the definition of τ (Eq. (3)), this fact is a confirmation of the reliability of the method since by doubling the thickness of the cavity the value of τ should indeed increase by a factor of four. Furthermore, the values found by the regression procedure are close to what can be calculated by substituting the value of thermal diffusivity (α = 10-7 m2/s [15]) in Eq. (3), shown in Fig. 4 as horizontal solid lines. An advantage of the procedure reported above is that it is possible to estimate the local solidification time without any knowledge of material properties, of moulding conditions, and even of local thickness. Thus, the procedure can be applied to a cavity of unknown or variable thickness. Neglecting the variation in physical properties of the polymer with cooling and the effect of convection [9], the local solidification profile can be obtained:

ys*,long (t )  1 −

allows the calculation of the solidification evolution inside the layers that solidify at short times (Fo < 0.1)

4  t ys*,short (t )  er f −1  exp  − sol π  τ 

2

  2  t   4  π  τ . (5)   

a)

b)

 2  t − tsol   a cos exp    , (4) π  τ  

in which y* is the normalized distance from the skin (y* = 0 at the mould surface and y* = 1 at the midplane) and y*s identifies the layer which is solidifying at time t (y*s = 1 for complete solidification and afterwards). The subscript long indicates that this solution is valid for Fourier numbers (Fo = αt / L2) larger than 0.1, namely for longer times. A solution for the heat conduction for shorter times can be obtained by the penetration theory. This

c)

Fig. 2. Illustration of the exponential fitting on some of the pressure curves analysed in this work; 2 mm thick cavity; a) Tinj = 200 °C; b) Tinj = 220 °C; c) Tinj = 240 °C

An equation that allows us to describe the solidification layer profile over the whole time range can be given as a combination of Eqs. (4) and (5) [9]:

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ys* (t ) = ys*, short (t ) + ξ (t )  ys*, short (t ) − ys*, long (t )  , (6)

in which the function ξ(t) should be zero at low Fourier numbers and 1 at high Fourier numbers. Eq. (7) describes a transition from 0 to 1 in the neighbourhood of Fo = 0.1. One possible expression is:

ξ (t ) = 1 −

1 , (7) 1 + exp [10(t − tc ) ]

with 2

π  tc = 0.1  τ . (8) 2

is worth mentioning that the method does not require knowledge of the initial, mould or solidification temperatures, whose determination presents a certain degree of uncertainty, and it does not require any characterization of the physical parameters of the material. 3 AVERAGE SOLIDIFICATION PRESSURE AND SHRINKAGE The definition of the average solidification pressure arises from considering a viscous – elastic model for shrinkage [16], namely from assuming that the polymer melt turns into an elastic solid as soon as it solidifies. Since solidification proceeds from the mould surfaces to the core, solidification pressure is different for each layer, thus each layer has a different stress-free configuration (larger dimensions for layers solidified under high pressure [16]). On ejection, each layer will experience a different stress so as to bring all of them to the same final length. On the basis of these considerations, the average value over the thickness of the pressures at which each layer solidifies, Ps , was introduced to take into account the effect of pressure on shrinkage.

Fig. 3. Illustration of the exponential fitting on some of the pressure curves analysed in this work; 4 mm thick cavity

Fig. 4. Values of the parameter τ found for all the pressure curves analysed in this work; the horizontal lines identify the theoretical and the average values for τ

Eqs. (4) to (8) show that knowledge of tsol, namely the local solidification time, and of τ, allows the estimation of the whole solidification profile. It 680

1

Ps = ∫ * P(t )dys* (t ). (9) y =0

It was demonstrated [9], [11] and [16], that, for a given polymer, the average solidification pressure is directly related to the local shrinkage. However, the definition of the average solidification pressure requires knowledge of the temperature histories inside the polymer, in order to define y*s(t). The main purpose of the present work is to define a suitable method for obtaining this piece of information using experimental local pressure alone. Once the local solidification profile is known, the local average solidification pressure can be easily calculated by Eq. (9). As reported in the literature [9] and [16], local shrinkage should be directly correlated to Ps . The procedure outlined above was then applied to all the pressure curves of each moulding test: first, the values of tsol and τ were obtained at each position and then the solidification profile was calculated. Eventually, the values of Ps were determined. Fig. 5 shows the experimental values for the shrinkage measured at each cavity position and for all tests carried out on PS versus the local value of the average solidification pressure. Most of the shrinkage data are collected on a single plot, which confirms on the one hand the suitability of Ps in correlating to the quality of the moulded part and on the other

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hand the reliability of the procedure reported in this work in obtaining a single parameter, able to correlate with shrinkage, whose value can be determined by the experimental pressure evolution only. For the same average solidification pressure, the differences in shrinkage were less than 0.2% (it is worth recalling that the accuracy of measurement for shrinkage is ±0.03%). The procedure described in this work is suitable for a master-curve approach: a series of moulding tests can be carried out for the chosen material, recording the pressure curves and measuring the shrinkage close to the pressure transducer; for each pressure curve the value of Ps is calculated and a master curve of shrinkage vs. of Ps is built; afterwards, the procedure is able to automatically associate a value for the shrinkage to each test by calculating on line the value of Ps from each experimental pressure curve.

Fig. 5. Measured width shrinkage vs. average solidification pressure for each cavity position and for all tests carried out in this work

4 CONCLUSIONS In this work a procedure was adopted to calculate the average solidification pressure, a parameter that is critical for the description of local shrinkage, by analysing the local pressure evolution measured using a conventional pressure transducer. The procedure was applied to a general purpose PolyStyrene, which was injection moulded under several processing conditions, where the cavity thickness was also changed. It was shown that all the shrinkage data collect on a single plot when reported versus the average solidification pressure calculated by analysing the experimental pressure curves. The procedure is thus suitable for a master-curve approach in which some data for the shrinkage versus the average solidification pressure can be used as a reference in

order to estimate the shrinkage of the part by analysing the local pressure evolution. The described procedure is particularly suitable for on line monitoring of the chosen quality parameter and does not require either knowledge of local thickness or of the moulding conditions. Furthermore, it can be applied without any characterization of the physical parameters of the material. 5 REFERENCES [1] Chen Z.B., Turng L.S. (2005). Current developments in process and quality control for injection molding. Advances in Polymer Technology, vol. 24, no. 3, p.165182, DOI:10.1002/adv.20046. [2] Michaeli W., Schreiber A. (2009). Online control of the injection molding process based on process variables. Advances in Polymer Technology, vol. 28, no. 2, p. 6576, DOI:10.1002/adv.20153. [3] Visvanathan, K., Balasubramaniam, K. (2007). Ultrasonic torsional guided wave sensor for flow front monitoring inside molds. Review of Scientific Instruments, vol. 78, no. 1, p. 015110, DOI:10.1063/1.2432258. [4] Fung, K.T., Gao, F., Chen, X. (2007). Application of a capacitive transducer for online part weight prediction and fault detection in injection molding. Polymer Engineering & Science, vol. 47, no. 4, p. 347-353, DOI:10.1002/pen.20700. [5] Bur, T. (1998). Optical monitoring of polypropylene injection molding. Journal of Reinforced Plastics and Composites, vol. 17, no. 15, p. 1382-1390. [6] De Santis, F., Pantani, R., Speranza V., Titomanlio, G. (2010). Analysis of shrinkage development of a semicrystalline polymer during injection molding. Industrial & Engineering Chemistry Research, vol. 49, no. 5, p. 2469-2476, DOI:10.1021/ie901316p. [7] Panchal, R.R., Kazmer, D.O. (2010). In-situ shrinkage sensor for injection molding. Journal of Manufacturing Science and Engineering, vol. 132, no. 6, p. 064503, DOI:10.1115/1.4002765. [8] Kurt, M., Kamber, O.S., Kaynak, Y., Atakok, G., Girit, O. (2009). Experimental investigation of plastic injection molding: Assessment of the effects of cavity pressure and mold temperature on the quality of the final products. Materials & Design, vol. 30, no. 8, p. 3217-3224, DOI:10.1016/j.matdes.2009.01.004. [9] Speranza, V., Vietri, U., Pantani, R. (2011). Monitoring of injection molding of thermoplastics: Average solidification pressure as a key parameter for quality control. Macromolecular Research, vol. 19, no. 6, p. 542-554, DOI:10.1007/s13233-011-0610-9. [10] Liu, S., Su, P., Lin, K. (2009). In-situ temperature measurements in the depths of injection molded parts. Measurement, vol. 42, no. 5, p. 771-777, DOI:10.1016/j.measurement.2009.01.002.

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[11] Speranza, V., Vietri, U., Pantani, R. (2012). Adopting the experimental pressure evolution to monitor online the shrinkage in injection molding. Industrial & Engineering Chemistry Research, vol. 51, no. 49, p. 16034-16041, DOI:10.1021/ie302432v. [12] Vietri, U., Sorrentino, A., Speranza, V., Pantani, R. (2011). Improving the predictions of injection molding simulation software. Polymer Engineering & Science, vol. 51, no. 12, p. 2542 -2551, DOI:10.1002/pen.22035. [13] Sorrentino, A., Pantani, R. (2009). Pressure-dependent viscosity and free volume of atactic and syndiotactic polystyrene. Rheologica Acta, vol. 48, no. 4, p. 467478, DOI:10.1007/s00397-009-0348-x. [14] Pantani, R., Speranza, V., Sorrentino, A., Titomanlio, G. (2002). Molecular orientation and

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strain in injection moulding of thermoplastics. Macromolecular Symposia, vol. 185, no. 1, p. 293-307, DOI:10.1002/1521-3900(200208)185:1<293::AIDMASY293>3.0.CO;2-8. [15] Pantani, R., Speranza, V., Titomanlio, G. (2001). Relevance of mold-induced thermal boundary conditions and cavity deformation in the simulation of injection molding. Polymer Engineering & Science, vol. 41, no. 11, p. 2022-2035, DOI:10.1002/ pen.10898. [16] Jansen, K.M.B., Titomanlio, G. (1996). Effect of pressure history on shrinkage and residual stresses – Injection molding with constrained shrinkage. Polymer Engineering & Science, vol. 36, no. 15, p. 2029-2040, DOI:10.1002/pen.10598.

Speranza, V. – Vietri, U. – Pantani, R.


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 683-688 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.1001 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-08-21 Accepted for publication: 2013-09-27

Influence of Rapid Mould Temperature Variation on the Appearance of Injection-Moulded Parts Lucchetta, G. – Fiorotto, M. Giovanni Lucchetta* – Marco Fiorotto

University of Padua, Department of Industrial Engineering, Italy In this work, an innovative technology for the rapid heating and cooling of injection moulds has been developed and used to analyse the effect rapid variations of the mould temperature on the improvement of mouldings’ appearance in terms of gloss. The obtained experimental results show that by maintaining an elevated mould surface temperature, the polymer melt is prevented from solidifying prematurely in the filling and packing stage, thereby improving the replication of the mirror-finished cavity surface. Furthermore, the mould cavity heating combined with the rapid cooling of the moulded part significantly contributes to contrasting the development of surface defects, such as weld line marks. Keywords: injection moulding, rapid heat cycle moulding, gloss

0 INTRODUCTION In recent years, the rapid growth of the computer, communication and consumer electronics industries has driven demand for producing plastic parts with high-quality finishes. The surface gloss of thermoplastic parts is an essential characteristic as it affects the optical behaviour and the aesthetics of the parts [1]. Gloss varies with the refractive index of the polymer, the angle of incidence and the topography of the surface [2]. Previous studies on polymer injection moulding have indicated that gloss improves with increasing mould temperature [3] for some rubber-modified thermoplastics and for pigmented polypropylene [4]. Mould temperature was shown to have a significantly higher influence on gloss than melt temperature, the effect of the latter being negligible for acrylonitrilebutadiene-styrene (ABS) [5]. Increasing the values of the holding time even out the gloss of ABS mouldings after an initial improvement [5], while an increase of packing pressure improves the gloss for the same polymer [6]. Cooling time has a negligible effect on gloss, both for stochastic [7] and geometric micro-structures [8]. Oliveira et al. [9] studied the modifications of the surface morphology and microtopography of injection-moulded ABS parts caused by the processing conditions and they related them with gloss. This work showed that the surface finish and appearance of injection-moulded parts are highly dependent on the process parameters and in particular on the mould temperature. Moulding gloss can be substantially improved by increasing the cavity surface temperature. However, this solution involves higher manufacturing costs due to the significantly longer cooling time. Therefore, technologies for rapid mould temperature variation,

known as rapid heat cycle moulding (RHCM), have been developed especially in the last decade [10]. In this process, a mould is first heated to a pre-set high temperature before melt injection, then kept at the high temperature during filling and packing, and finally cooled to solidify the polymer melt for demoulding. For amorphous polymers, the tool is heated up to a temperature that is 10 °C higher than the glass transition (Tg). The use of RHCM has come to prominence recently, mainly as a result of a commercial incentive to produce high-gloss mouldings without visible weld lines. Among all available heat generation technologies, electrical resistive heating is the most widely used mechanism for rapid mould heating. It is usually accomplished by passing direct or alternating current in cartridge heaters or in a thin electrical conductive layer [11]. The first solution is robust but energy inefficient and slow while the latter is rapid but the thin conductive layer wears out extremely quickly. Alternatively, high-frequency electrical current can be generated at the surface of a large mould mass by the skin effect from a high-frequency electromagnetic field. Two useful technical approaches for implementing this skin effect are induction heating [12] and proximity heating [13]. Induction heating has been effectively used to improve the appearance of moulding surfaces [14] and to assist the micro-injection of high aspect ratio microfeatures [15]. This technology was also employed in compression moulding [16]. Infrared radiation has also been applied for heating the cavity surface [17] and [18]. However, induction, proximity and infrared heating require high initial investments in equipment. Another solution to thermally cycle the mould temperature is by alternating two heating and cooling fluids in the mould [19]. These fluids impose a

*Corr. Author’s Address: University of Padua, Department of Industrial Engineering, via Venezia 1, Padua, Italy, giovanni.lucchetta@unipd.it

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convective heat flux at the fluid-solid interface. The hot fluid may be circulated inside the mould or directly introduced to the mould surface from the cavity. Efficient heating can be achieved using either superheated steam [20] or pressurized hot water circulating in conformal cooling channels [21] or ball bearing-filled slots located in proximity of the cavity wall and connected in a series with the cooling channels [22]. Both the conformal channels and ball bearing filling allow rapid and uniform heating of the cavity and provide mechanical support to contrast high injection pressures. However, conformal cooling channels are expensive to produce while ball bearingfilled slots can be realized only for parts with plane geometry. In this work, an innovative RHCM system has been developed to overcome the limitations of the available technologies [23] and [24], and has been used to analyse the effect of rapid variation of the mould temperature on moulding gloss. The proposed solution is based on the ball filling technology but replaces ball bearings with inserts made of opencell aluminium foam, which possesses high strength for structural applications, and elevated convection coefficients for heat transfer [25].

Fig. 2 is a schematic representation of the new RHCM mould. A K-type thermocouple was placed 1 mm from the cavity surface to detect the mould temperature profile during the moulding cycle. The layout of the heating/cooling system in the cavity insert is reported in Fig. 3.

Fig. 2. Model of the RHCM mould

1 MOULD DESIGN The heating and cooling system geometry is of considerable importance because it affects the heating/ cooling efficiency and temperature uniformity. In this work, a new mould, based on the use of open-cell aluminium foam, was designed. A 64×24×2 mm cover plate was chosen as a test case to study the efficiency of the proposed solution (Fig. 1).

Fig. 3. Layout of the heating/cooling systems in the cavity insert

a) b) c) Fig. 4. Layout of the different RHCM systems: a) 90×33×20 mm aluminium foam, b) 90×30×10 mm aluminium foam and c) 90×30×10 ball filling

Fig. 1. Cover plate used as a case study

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Two slots were realized at a distance of 6 mm from the cavity surface and were filled with three different inserts (Fig. 4). First, two inserts of open cell aluminium foam of 90×33×20 mm were used. The metal foam, supplied by ERG Materials and Aerospace Corporation, has 5 pores/cm (Fig. 5). In the second arrangement, a 90×30×10 mm insert of aluminium foam was used. For comparison, the cover plate was produced replacing the aluminium foam Lucchetta, G. – Fiorotto, M.


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with a ball bearing filling. The balls have a diameter of 10 mm (Fig. 6).

Fig. 5. Insert made of open cell aluminium foam

Fig. 6. Mould slot filled with bearing balls

2 EXPERIMENTAL METHOD The polymer used in this study is an ABS (BASF Terluran KR 2922). The material was characterized by means of a differential scanning calorimeter (TA Instruments Q200). The measured glass transition temperature is 95.35 °C. The injection speed and the melt temperature were set to the highest limits of the moulding window in order to decrease the viscosity of the polymer during the injection phase. The experimental tests were conducted on a 1100 kN injection moulding machine (Wittmann-Battenfeld HM110). Before the injection phase, the mould was heated with water circulating at a temperature of 140 °C until the temperature in correspondence of the thermocouple position reached 100 °C. After the packing phase, it was cooled with water at 30 °C. A high-performance

thermal unit equipped with two conditioning systems and a valve exchange unit (Wittmann-Battenfeld Tempro plus C160 Vario) was used to rapidly vary the cavity surface temperature. The valve exchange device alternates the heating water and the cooling water into the channel system by switching the status of the corresponding control valves. When the mould is still open, the hot water flows into the channel by opening the valve. The energy transferred from the hot water to the mould plate heats the cavity surface. The velocity of the cooling water must be high enough to maintain the turbulent flow for cooling the mould. When the cavity surface is heated to a high temperature, the mould is closed in preparation for the filling process. When the heating process is finished, the temperature at the cavity surface reaches the target value for assistance in the filling and packing of the melt. At this time, the valve control is switched to the cold water flowing into the channels, and the cooling process begins.

Fig. 7. Comparison of temperature profiles relevant to the different heating/cooling systems

As shown in Fig. 7, the first technology, based on the single use of metal foam, allows drastic reductions of the heating and cooling times, by about 7 and 16 s, respectively. Table 1 reports the heating and cooling rate values for the three tested RHCM systems. Table 1. Heating and cooling rate values for the three tested RHCM systems 20 mm foam 10 mm foam 10 mm ball bearing filling

Heating rate [°C/s] 5.0 1.9 1.7

Cooling rate [°C/s] 2.4 2.4 2.0

For RHCM, the required heating and cooling time of the mould mostly depends on the mass of the cavity/core to be heated and cooled. A mould with a low thermal mass exhibits a low thermal inertia and can be rapidly heated and cooled. With an increase of

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the thickness of the metal foam inserts, the volume of the material being heated is reduced with a consistent improvement in cycle time. 3 EFFECT OF RHCM ON THE PRODUCT APPEARANCE All heating/cooling systems allow producing highgloss parts without visible weld lines, which should be located near each hole (Fig. 3) of the cover frame shown in Fig. 8 if this part was moulded using a conventional injection moulding process. This is due to the fact that all of the tested systems were able to reach mould temperature values higher than the glass transition temperature of the ABS used in the experiments.

the perpendicular direction. The diameter of the probe beam of the spectrometer was adjusted to 3 mm. The diffuse reflectance was measured by trapping and removing the specular component. The specular reflectance was calculated by subtracting the diffuse reflectance from the total reflectance.

Fig. 9. Measurement location

Fig. 8. High quality cover plate

Only the cover plates produced using the new RHCM technology with metal foam of 90×33×20 mm were chosen for the gloss analysis. Surface gloss is a subjective impression created by the light flux reflected by a part. In industrial practice, surface gloss is often expressed in relation to the reflection from an ideal polished black surface in the specular direction [26] and [27]. Gloss is measured with a glossmeter, and the results are expressed in gloss units (GU), which are calculated as the reflectometer reading for the surface concerned calibrated with respect to that of a standardized black glass plate with a known refractive index. Spectrophotometers operating in the reflection mode and diffractive optical sensors have also been used to study the gloss differences of injection-moulded plastic products. In this study, the effect of the rapid variation of the mould temperature on surface gloss was analysed using a UV/VIS spectrophotometer (Jasco, V-570) operating in reflection mode. The total reflectance at point A was measured (Fig. 9). The angle of incidence of the spectrophotometer is fixed to 8° to 686

The process variables that were analysed in the experimental design were the mould temperature values measured by the thermocouple at the injection (T1) and the ejection (T2) phases. A simple onefactor-at-a-time design was used, according to values reported in Table 2. Each experiment was replicated four times. Table 2. One-factor-at-a-time experimental design Experiment A B C D E

T1 [°C] 60 80 100 100 100

T2 [°C] 60 60 60 50 70

Keeping heating and cooling water temperatures at constant values of 140 and 30 °C, respectively, T1 and T2 were set by varying heating and cooling times according to the temperature profiles shown in Fig. 7. Fig. 10 shows the comparison of average reflectance results (as mean reflectance values per treatment). A one-way ANOVA conducted on the first three experiments revealed that the mould temperature at injection (T1) has a significant influence on the reflectance (Fig. 11). The results show a significant difference in surface gloss between the conventional injection moulding process (T1 of 60 °C, experiment

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A) and RHCM with mould temperature higher than Tg (experiment C). Such high-glossy surface can be explained by the fact that only at temperatures higher than Tg can the mirror-finished surface of the cavity be entirely replicated onto the moulding surface, because the injected polymer melt cannot solidify at first mould contact.

Fig. 10. Specular reflectance from the test location

Fig. 11. Influence of the mould temperature at injection (T1) on reflectance

A one-way ANOVA conducted on the last three experiments revealed that the mould temperature at ejection (T2) has a negligible influence on the reflectance. This setting is, however, essential for keeping cycle time values low. 5 CONCLUSIONS In this work, an innovative RHCM system has been developed and used to analyse the effect of the rapid variation of the mould temperature on appearance of injection-moulded parts. The proposed solution is based on the ball filling technology (BFMOLDTM) but replaces ball bearings with inserts made of opencell aluminium foam, which allow increasing the heat exchange surface as compared to the traditional

cooling channels. A prototype mould was designed and realized. Two inserts of open-cell aluminium foam were placed just under the cavity surface and were integrated in the conditioning system of the mould. The metallic foam allows the flowing of a fluid at a controlled temperature and generates a cavity structure that favours an evenly spread tempering system immediately below the mould surface. The temperature profile near the cavity surface was detected by a thermocouple and was compared with experimental values measured in the same reference point, using thinner metallic foam inserts and ball bearing filling. The comparison between the ball bearing filling technology and the RHCM system with metal foams shows that the solution proposed in this work allows an improvement in the cycle time of about 16 s. Furthermore, the mould cavity heating combined with the rapid cooling of the moulded part contributed to eliminating the development of visible weld lines and to producing high-gloss parts. The innovative RHCM system was used to analyse the effect of rapid variations of the mould temperature on the specular reflectance of the moulding surface. The rapid heating and cooling to the mould surface leads to a significant improvement in surface gloss when heating the mould up to a temperature higher than Tg. This is mainly because the high cavity surface temperature can prevent the melt from freezing prematurely in the filling and packing stage and improve the replication of the mirrorfinished cavity surface, resulting in superior aesthetics (no visible weld line) and higher gloss. 5 REFERENCES [1] Berger, G.R, Friesenbichler, W., Teichert, C. (2010). Replication of stochastic and geometric micro structures: Effects on functionality and visual appearance. 26th Annual Meeting of the Polymer Processing Society, Banff. [2] Ignell, S., Kleist, U., Rigdahl, M. (2009). Visual perception and measurements of texture and gloss of injection-molded plastics. Polymer Engineering and Science, vol. 49, no. 2, p. 344-353, DOI:10.1002/ pen.21279. [3] Dawkins, E., Engelmann, P., Horton, K. (1998). Color and Gloss - The Connection to Process Conditions. Journal of Injection Molding Technology, vol. 2, no. 1, p. 1-7. [4] Pisciotti, F., Boldizar, A., Rigdahl, M., Arino, I. (2005). Effects of injection-molding conditions on the gloss and color of pigmented polypropylene. Polymer Engineering and Science, vol. 45, no. 12, p. 1557-1567, DOI:10.1002/ pen.20358.

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[5] Koppi, K.A., Ceraso, J.M., Cleven, J.A., Salamon, B.A. (2002). Gloss modeling of injection molded rubbermodified styrenic polymers. SPE ANTEC Technical Papers, vol. 48, p. 184-188. [6] Edwards, S.A., Choudhury, N.R. (2004). Variations in surface gloss on rubber-modified thermoplastics: relation to morphological and rheological behavior. Polymer Engineering and Science, vol. 44, no. 1, p. 96112, DOI:10.1002/pen.20009. [7] Schauf, D. (1988). Application Technology Information (ATI) 584e: Reproducing Textures from the Cavity Surface to the Surface of the Thermoplastic Moulding. Bayer Material Science AG, Leverkusen. [8] Berger, G.R., Gruber, D.P., Friesenbichler, W., Teichert, C., Burgsteiner M. (2011). Replication of stochastic and geometric micro structures: aspects of visual appearance. International Polymer Processing, vol. 26, no. 3, p. 313322, DOI:10.3139/217.2451. [9] Oliveira, M.J., Brito, A.M., Costa, M.C., Costa, M.F. (2006). Gloss and surface topography of ABS: a study on the influence of the injection molding parameters. Polymer Engineering and Science, vol. 46, no. 10, p. 1394-1401, DOI:10.1002/pen.20607. [10] Guilong, W., Guoqun, Z., Huiping, L., Yanjin, G. (2010). Analysis of thermal cycling efficiency and optimal design of heating/cooling systems for rapid heat cycle injection molding process. Material and Design, vol. 31, no.7, p. 3426-3441, DOI:10.1016/j.matdes.2010.01.042. [11] Yao, D., Kim, B.Y. (2002). Development of rapid heating and cooling systems for injection molding applications. Polymer Engineering and Science, vol. 42, no. 12, p. 2471-2481, DOI:10.1002/pen.11133. [12] Chen, S.C., Jong, W.R, Chang, J.A., Chang, Y.J. (2006). Dynamic mold surface temperature control using induction and heater heating combined with coolant cooling. International Polymer Processing, vol. 21, no. 5, p. 457-463. [13] Yao, D., Kimberling, T.E., Kim, B. (2006). Highfrequency proximity heating for injection molding applications. Polymer Engineering and Science, vol. 46, no. 7, p. 938-945, DOI:10.1002/pen.20548. [14] Chen, S.C., Jong, W.R., Chang, J.A., (2006). Dynamic mold Surface Temperature Control Using Induction Heating and its Effects on the Surface Appearance of Weld Line. Journal of Applied Polymer Science, vol. 101, no. 2, p. 1174-1180, DOI:10.1002/app.24070. [15] Chen, S.C., Jong, W.R., Chang, Y.J., Chang, J.A., Cin, J.C. (2006). Rapid mold temperature variation for assisting micro injection of high aspect ratio microfeature parts using induction heating technology. Journal of Micromechanics and Microengineering, vol. 16, no. 9, p. 1783-1791, DOI:10.1088/0960-1317/16/9/005.

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[16] Miller, K., Ramani, K. (1998). Analysis of an inductively heated compression molding process. Advances in Polymer Technology, vol. 17, no. 3, p. 251-257, DOI:10.1002/(SICI)1098-2329(199823)17:3<251::AIDADV5>3.0.CO;2-R. [17] Yu, M.C., Young, W.B., Hsu, P.M. (2007). Micro injection molding with the infrared assisted heating system. Materials Science and Engineering: A, vol. 460, no. 461, p. 288-295, DOI:10.1016/j.msea.2007.02.036. [18] Berger, G.R., Roock, S., Gießauf, J., Gruber, D.P., Friesenbichler, W., Steinbichler, G. (2011). Improving the polymer surface quality by infrared radiation driven dynamic mold temperature control. 27th Annual Conference of Polymer Processing Society, Marrakech. [19] Fu, G., Loh, N.H., Tor, S.B., Tay, B.Y., Murakoshi, Y., Maeda, R. (2005). A Variotherm mold for micro metal injection molding. Microsystem Technologies, vol. 11, no. 12, p. 1267-1271, DOI:10.1007/s00542-005-0605-6. [20] Zhao, G.-Q., Wang, G.-L., Li, H.-P., Guan, Y-J. (2009). Research and application of rapid heating cycle molding technology. Journal of Plasticity Engineering, vol. 16, no. 1, p. 190-195. [21] Jeng, M.-C., Chen, S.-C., Minh, P.S, Chang, J.A., Chung, C.-S. (2010). Rapid mold temperature control in injection molding by using steam heating. International Communications in Heat and Mass Transfer, vol. 37, no. 9, p. 1295-1304, DOI:10.1016/j. icheatmasstransfer.2010.07.012. [22] Bariani, P.F., Lucchetta, G., Fiorotto, M. (2011). System for rapid heating and cooling of moulds for polymer materials (Sistema di riscaldamento e raffreddamento rapido di stampi per materiali polimerici). Italian Patent Application PD2011A000117, Italian Patent and Trademark Office, Padova. [23] Lucchetta, G., Fiorotto, M., Bariani, P.F. (2012). Influence of rapid mold temperature variation on surface topography replication and appearance of injection-molded parts. CIRP Annals - Manufacturing Technology, vol. 61, no. 1, p. 539-542, DOI:10.1016/j. cirp.2012.03.091. [24] Ozmat, B., Leyda, B., Benson, B. (2004). Thermal applications of open-cell metal foams. Materials and Manufacturing Processes, vol. 19, no. 5, p. 839-862, DOI:10.1081/AMP-200030568. [25] ASTM Standard D2457 (2008). Standard Test Method for Specular Gloss of Plastic Films and Solid Plastics. ASTM International, West Conshohocken. [26] ASTM Standard D523 (2008). Standard Test Method for Specular Gloss, ASTM International, West Conshohocken.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 689-696 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2013.999 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-03-07 Accepted for publication: 2013-03-13

An Initial Study of Aerosol Jet® Printed Interconnections on Extrusion-Based 3D-Printed Substrates Vogeler, F. – Verheecke, W. – Voet, A. – Valkenaers, H. Frederik Vogeler* – Wesley Verheecke – André Voet – Hans Valkenaers 1 Thomas

More University College Mechelen, Belgium

The combination of different additive manufacturing techniques to produce freeform products with multifunctional properties is gaining increasing popularity. In the research presented, Aerosol Jet® Printing (AJP) is combined with extrusion-based 3D printing. AJP starts with an ink to create micro-tracks. These tracks commonly have widths ranging from a few micrometers up to several millimeters and track heights ranging from a few tenths of a micrometer up to several micrometers, unlike extrusion-based 3D printers with which the extruded material usually has a resolution of tenths of millimeters. AJP can therefore be a complementary technique for extrusion-based 3D printing; in this manner, fine high resolution features can be added onto relatively rapidly produced extrusion-based 3D printed parts. Furthermore, AJP can be used to produce electrically conductive tracks to create interconnections, inductors, capacitors, strain gauges, etc. In this paper, the creation of AJP-manufactured interconnects on extrusion-based 3D printed substrates is investigated. The relevant AJP process parameters to take into account are the flow rates of the aerosol, the flow rate of the sheath gas, the temperature settings of the ink and substrate, and the platform speed and nozzle-to-substrate distance. To obtain reliable results, the AJP process parameters are optimized for printing single-layered and multilayered silver ink tracks on extrusion-based 3D-printed surfaces. Important quality output parameters include the dimensions and the electrical properties of the printed interconnects. Keywords: aerosol jet® printing, extrusion-based 3D printing, silver ink, interconnects, printed tracks, conductive tracks

0 INTRODUCTION 0.1 Key Benefits of Hybrid Additive Manufacturing Additive Manufacturing (AM) [1] is no longer merely used as a production method for creating prototypes; today, it is more often used to create products with a high added value. Although there is still room for the optimization of individual AM processes, it is most likely that the next trend in AM is the combination of AM techniques, which can have a number of advantages, e.g. improving the accuracy of an AM process or creating multi-material products. In the presented research, extrusion-based 3D printing is combined with Aerosol Jet® Printing (AJP). Extrusion-based 3D printing is typically used to create functional plastic (commonly ABS, PC or PLA) parts with sizes ranging from a cubic centimetre up to a cubic metre. AJP, in contrast, is used to create parts with micro-sized features. It is also frequently used to create printed electronics, including interconnects, resistors, inductors, capacitors, etc. In our research, AJP is used to print conductive tracks on extrusionbased 3D-printed parts. A fundamental reason combining these technologies has so much potential is because of the feasibility of combining both techniques on the same machine. Both techniques are categorized as direct write technologies [2], they only require an xyz manipulation stage and a material deposition head.

0.2 Extrusion-Based 3D Printing Extrusion-based 3D printing [3] is a generic term for all AM processes that utilize a head that deposits material by extruding viscous materials through a nozzle. The most commonly known extrusionbased 3D-printing technique is Fused Deposition Modeling™ (FDM®) [4] and [5]. This technique heats up a filament before pushing it through a fine opening. Extrusion-based 3D printing usually has a resolution ranging from hundred to several hundred micrometres. 0.3 Aerosol Jet® Printing AJP uses an ink as a starter material to build up products [6] to [9]. The ink is transformed into an aerosol by the atomizer. There are several ways of atomizing liquids, depending on the type of atomizer used [10] and [11]. In this research, a pneumatic atomizer is used. It creates a nebula of the ink by sending carrier gas through a venturi. The venturi creates a vacuum that sucks the ink through a narrow tube of a few millimetres in diameter. The ink is then sent through an opening and turned into droplets. The amount of carrier gas that is fed to the atomizer is called the ‘atomization gas flow’. In the final stage, the aerosol is focused onto the substrate by adding a concentric flow of gas, which is called ‘sheath gas’. Adding the sheath gas to the aerosol occurs inside the print nozzle, by sending

*Corr. Author’s Address: Thomas More University, College Mechelen, J. De Nayerlaan 5, 2860 Sint-Katelijne-Waver, Belgium, frederik.vogeler@lessius.kuleuven.be

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the gas along a conical shaft [12] to [16]. In the center of the conical shaft, the aerosol is added. An overview of the AJP process is given in Fig. 1.

Fig. 2. Isometric view of the substrate; all dimensions are displayed in millimetres

Fig. 1. Overview of the AJP process

Because the flow coming directly out of the pneumatic atomizer is too high for the nozzle to handle, an extra element is placed between the atomizer and nozzle. This element is called the ‘Virtual Impactor’ (VI). It reduces the aerosol flow by taking away some of the carrier gas (and small aerosol droplets) [17] to [23]. Inside the VI, the aerosol tube is interrupted, and gas is sucked out of the VI. Due to the mass inertia of the droplets, the larger aerosol droplets continue a straight trajectory and are not sucked out of the VI; smaller droplets and carrier gas are sucked out. The flow that is taken away from the aerosol is called the ‘VI exhaust’. The difference between the atomizer flow and VI exhaust flow is fed to the nozzle; this is referred to as the ‘VI difference’. Depending on the processed material, substrate material and conditions, process settings and nozzlesize single-printed track widths can vary from just a few micrometres up to several millimetres [24] to [26]. Track heights can range from a few tenths of a micrometre up to several micrometres.

The samples were built up layer by layer in the z direction indicated by Fig. 2; they were filled with cross hatching. On the top surface, indicated by hatching in Fig. 2, test circuitry is printed by means of AJP. The top surface was first sanded down for half a minute, with silicon carbide P220 sand paper on a horizontal disk sander. Untreated samples from the used 3D printer have a surface roughness that is too high for the robust printing of silver ink tracks. This problem can be resolved by printing on parts with a finer cross hatching or using an ink-substrate combination with better wetting properties. In all the used samples, the sanding resulted in an average Ra value of 1.42 µm, with a standard deviation of 0.73 µm. This is a relatively large surface roughness compared to other substrates for printed electronics (e.g. glass or polyimide foil). The influence of the level of abrasion on the AJP result has not been investigated and will be included in future research. Before the samples were placed inside the AJP machine, the surface was treated with isopropanol to cleanse the surface and reinsure a repeatable substrate surface. The samples were first sanded down because preliminary tests showed that untreated samples gave insufficient adhesion between the printed ink and substrate. To investigate the conductivity of AJP tracks, circuitry was designed, shown in Fig. 3. Four square probing patches were to perform a 4-point probing test over a track length of 10 mm.

1 SETUP OF THE EXPERIMENT The substrate samples in the following experiments are produced on a Stratasys Dimension® SST1200es machine, from ABSplus™-P430 plastic. The test samples have a rectangular shape with a height of 8 mm. An isometric sketch of the substrate samples is shown in Fig. 2. 690

Fig. 3. Dimensions of the test circuitry; all dimensions are displayed in millimetres

The circuitry was printed with Cabot CSD-32® material, a commercially available ink based on silver

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nanoparticles (45 to 55 wt%), used as conductive filler material in ethylene glycol. The silver particles have a size of less than 60 nm and are wrapped in a polymer coating. To reinsure conductivity, the material supplier suggests sintering the printed material. Greer et al. 2007 [27] showed that sintering temperatures of nanoparticle silver inks should at least reach 150 °C in order to reach adequate conductive behavior. Because the glass transition temperature of ABS is about 100 °C, printed samples were cured in an oven at 80 °C for only 1.5 hours.

average of the detected edges was calculated. In Eq. (1), the average edge smoothness is represented by ESa in µm, n represents the number of detected points on an edge and xi is the shortest distance in µm from a detected point to the line fitted through all detected points. The edge smoothness will become salient when several tracks need to be printed side by side. 1 n ESa = ∑ x . (1) n i =1 i

Fig. 4. Photographs of a finished sample

During the experiments, the hardware setup of the AJP machine was unchanged. A nozzle with an opening of 300 µm was used to produce all the samples. Although smaller nozzle sizes are available and would provide smaller print resolution, such nozzles are probably not appropriate for this initial study. By having a smaller print resolution, tracks will have smaller cross sections, making it more difficult to obtain conductive behavior. 2 ANALYSIS OF THE PRINTED CIRCUITRY To analyze the electrical conductivity of the produced samples (Fig. 4), a Keithley® 580 micro-ohmmeter was used. The maximum resistivity that could be measured is 200 kΩ. Samples that have a higher resistivity are considered to have no conductive properties. The dimensions of the tracks were quantified with both a Digi-Microscope 500X Series vision system and a Taylor-Hobson® tactile 2D profile measuring device. The vision system consists of a USB-microscope placed perpendicular to the substrate. Unfortunately, this method cannot give information about the height (z direction) of a printed track. Important information that it does provide is the track width and that smoothness of the track edges. To determine these parameters, the edges of the track were first determined with the algorithm described by [28], a few examples of the result of this edge detection algorithm are given in Fig. 5. For the edge smoothness, the arithmetic

Fig. 5. Examples of the edge detection algorithm from the vision system for various types of printed AJP tracks

A 2D profile system was used to create an intersection profile of the track. This system gives information about the track width, height and cross section area. The downside of this system is that when printing only one layer of material it is impossible to differentiate the printed track from the substrate roughness profile. Consequently, this system was only used for analyzing the circuitries produced with multiple layers. Fig. 6 shows an example of an intersection measured with the 2D profile system.

Fig. 6. Examples of a measured intersection of a track, produced by printing 5 tracks side by side and 5 layers

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3 INVESTIGATION OF THE PROCESS PARAMETERS 3.1 VI Difference and Temperature Settings In a first set of tests, the VI difference and temperature settings of the AJP process were investigated. For this test, the previously presented circuitry was printed with one track and one layer. Regarding the AJP process, only the VI exhaust gas, printing head temperature and substrate temperature were changed. The atomizer gas flow was set to 133.33 × 10-7 m3/s (800 cm3/min) and the sheath gas flow was set to 10.83 × 10-7 m³/s (65 cm3/min); the printing speed was set to 8 mm/s and the nozzle-to-substrate distance was set to 3 mm.

smoothness. A higher temperature of the nozzle will evaporate more solvent in the aerosol droplets before they reach the substrate, changing the concentration of silver particles in the aerosol droplets. This higher concentration of silver particles causes the ink to act in a more viscous manner; once deposited, less ink will flow on the substrate. A higher temperature of the substrate will increase the effect of the reduced flow of the ink droplets on the substrate. The resistivity of the samples was too high to measure. 3.2 Sheath Gas Setting The second set of tests concerned the sheath gas settings. In this series of tests, the sheath gas was altered. Furthermore, two sets of VI difference where tested: one at 4.17 × 10-7 m3/s (25 cm3/min) and one at 2.50 × 10-7 m3/s (15 cm3/min). The atomizer gas flow and printing speed was set the same as in the tests described in Section 3.1. The temperature of the nozzle was set at 35 °C and the substrate temperature at 60 °C. Only one track was printed in one layer.

Fig. 7. The VI difference and temperature settings

Fig. 7 shows the results of the VI difference and temperature settings tests. As expected, the width of the track rises as the VI difference increases, causing the nozzle to deposit more aerosol per time unit. The high possibility of a (statistical) relationship between track width and VI difference is also confirmed by the correlation coefficients. For printing at low temperatures, an r-value (correlation coefficient) of 0.96 was obtained and at high temperatures an r-value of 0.98. In practice, attempting to control the track width using the VI difference is recommended. Concerning the edge smoothness, correlation with the VI difference is exceptionally low: for the test at low temperatures, an r value of –0.10 was obtained; for the test at high temperatures, an r value of 0.25 was obtained. Processing temperatures do seem to have an influence on both the track width and the edge smoothness. Printing at higher temperatures generates narrower tracks but generates a larger edge 692

Fig. 8. The sheath gas setting

Fig. 8 shows the results of the sheath gas and VI difference tests. As demonstrated by the previous test, the tracks become wider when the VI difference becomes larger; only the first point in the left graph of Fig. 8 shows a smaller track for a larger VI difference. It is most likely that the first value of the VI difference of 4.17 × 10-7 m3/s (25 cm3/min) is an outlier. The correlation coefficient between the width and sheath gas setting is 0.45 for the higher VI difference setting and –0.84 for the lower VI difference setting. These values indicate that using the sheath gas to control the track width is not the best option. According to the results of Section 3.1, it would be

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better to use the VI-difference for controlling the track width. The edge smoothness, however, seems to correlate better with the sheath gas settings. For a VI difference of 4.17 × 10-7 m3/s (25 cm3/min), the correlation is –0.88, for a VI difference of 2.50 × 10-7 m3/s (15 cm3/min), the correlation is 0.97. The fact that correlation is negative for a higher setting of VI difference compared to the lower settings of VI difference indicates that the edge smoothness is influenced by the combination of VI difference and sheath gas. In practice, sheath gas is a good candidate for controlling the edge smoothness. The resistivity of the samples was, as in the tests of Section 3.1, too high to measure. 3.3 Nozzle-to-Substrate Distance In the next test, the nozzle-to-substrate distance was investigated. The aerosol can only be focused within certain boundaries of the nozzle-to-substrate distance. For this test, the atomizer gas flow and printing speed was set the same as in Sections 3.1 and 3.2; sheath gas was set at 9.17 × 10-7 m3/s (55 cm3/min), VI difference at 3.33 × 10-7 m3/s (20 cm³/min), the temperature of the nozzle was set at 50 °C and the temperature of the substrate was set at 75 °C. As in the previous tests, the samples consisted of one printed track in one layer.

4 mm, while the edge smoothness seems to become bigger with a larger nozzle-to-substrate distance. As with the samples in Sections 3.1 and 3.2, none showed signs of conductivity. Consequently, several tracks side by side or several layers of material are needed in order to create functional interconnections. In the next set of tests, several tracks are printed side by side and several layers of material are deposited. 3.4 Multiple Tracks Side by Side During this test, the circuitry is printed with several tracks side by side. To reinsure overlap between the tracks, the parallel central axes of the tracks are printed 50 µm apart. Atomizer gas flow and printing speed were set as before (Sections 3.1 to 3.3), nozzle temperature was set to 50 °C, substrate temperature to 80 °C, sheath gas was set at 10.83 × 10-7 m3/s (65 cm3/min) and VI difference was set at 5.00 × 10-7 m3/s (30 cm3/min).

Fig. 10. Number of tracks side by side

Fig. 9. The nozzle to substrate distance

This test shows the importance of the nozzle-tosubstrate distance. Fig. 9 shows that the ideal focusing position is around 3 mm; not only is the track width that smallest, but also the edge smoothness is smallest for this position. Furthermore, the track width seems to stagnate for a nozzle-to-substrate distance larger than

Fig. 10 shows an obvious linear behavior for the width in function of the number of tracks side by side (r value of 1.00). The edge smoothness does not show a trend (r value of –0.46). The edge smoothness for multiple tracks side by side is significantly lower than the edge smoothness for samples consisting of only a single track. The first sample of only five tracks side by side gave no conductivity. The last sample of 20 tracks side by side gave a resistivity of 67.4 kΩ; this is probably an outlier. The remaining two samples gave a resistance of 99.61 Ω (10 tracks side by side) and 11.24 Ω (15 tracks side by side).

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3.5 Multiple Layers In the final test, the number of layers in the functions of the width, edge smoothness, resistivity, track height and cross section of the track were examined. The same AJP settings were used as in Section 3.4. The circuitry was printed with five tracks side by side, with the parallel central axes of the tracks printed 50 µm apart. Only two samples were created: one with three layers and one with five layers.

In practice, choosing the number of tracks side by side and number of layers will be a consideration based on electrical properties, track width and processing time. A simple alternative for obtaining conductivity is merely using a nozzle with a bigger opening. This will allow the creation of wider tracks and will require less processing time. Multiple layers can be distinguished from the roughness profile of the substrate. Therefore, it was possible to measure the track profile. The track profile was measured at three locations. From each measurement, the track height and cross section area was extracted. The results of these measurements are displayed in Fig. 12. The figure shows the average value and, minimum and maximum measured value (indicated by the error bars). The graphs show no surprises, both the track height as well as the cross section increase with the number of layers.

Fig. 11. The number of layers

Fig. 11 shows that the width of the track increases with the number of printed layers. This suggests that not all material printed onto an existing layer remains upon that layer, but tends to run down the side of the layer. For printing mere interconnections, this is not a problem; this might, however, cause a problem when trying to use AJP as a 3D printing technique. Just as in the early days of direct 3D inkjet printing, getting droplets to remain upon a new layer proved to be one of the most challenging tasks. As with multiple tracks printed side by side, no trend was observed (r-value of 0.38) for the edge smoothness. As mentioned in Section 3.4, one layer of five tracks side by side gave no conductivity. Three and five layers gave conductivity of 24.75 and 20.61 Ω, respectively. However, one layer of 15 tracks printed side by side (Section 3.4) gave a lower resistance than five layers of five tracks printed side by side. Therefore, considering processing time, it would be more beneficial to print several tracks side by side in order to obtain better electrical properties. However, printing several tracks side by side does result in broader tracks. 694

Fig. 12. The number of layers

What might seem strange is that there is no oneto-one relationship between the number of layers and the track height; furthermore, the left graph of Fig. 12 seems to indicate a growing increase in height in relation to the number of layers (3 layers give a height of 4.4 µm; 5 layers give a height of 9.7 µm). There are a number of probable causes for this phenomenon, which need to be investigated in the future. First, the wetting behavior of the first layer is different than that of the following layers as the first layer is deposited onto the ABS substrate, in contrast to the other layers that are deposited onto printed material. Other possible causes could be the slumping or mixing of not fully dried layers, causing an increase of the width of the previous printed track.

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4 CONCLUSION AND DISCUSSION In this paper, circuitry for testing the conductivity was printed with silver ink, by means of AJP, onto a substrate manufactured with extrusion-based 3D printing. Although the substrate was first sanded down in order to generate the proper surface characteristics, this preliminary research mentions some key aspects towards combining extrusion-based 3D printing and AJP. The main focus of the research was understanding the relationship between the AJP process settings and the dimensional and electrical properties of the printed interconnects. To date, only a limited set of experiments have been performed; future research will also have to include the repeatability of the process, as the presented tests are limited. Other aspects to obtain better and repeatable printing and conductive behavior will also have to be examined, e.g. other hardware setups, proper sintering methods, other conductive inks, surface activation, etc. Nevertheless, this work will be a good starting point for future research in this area. Based on correlation studies, when printing single tracks, the VI-difference is probably the best choice for controlling track width; the edge smoothness is best controlled by the sheath gas settings. Printing single tracks, however, gave no electrical conductive properties to the tracks. As a result, multiple tracks or layers of material are required in order to generate sufficient conductive properties. The presented research shows that, considering processing time, printing several tracks side by side is more beneficial than printing multiple layers. Although in practice making a choice in number of tracks side by side and layers will also consider the required resolution of the interconnects. 5 ACKNOWLEDGEMENTS The authors would like to thank the IWT for their support in the PO7810-EUR-ERA-01 project within the framework of the TETRA/ EraSME 3DAMEEA. 6 REFERENCES [1] Kruth, J.-P., Leu, M.C., Nakagawa, T. (1998). Progress in additive manufacturing and rapid prototyping. CIRP Annals - Manufacturing Technology, vol. 47, no. 2, p. 525-540, DOI:10.1016/S0007-8506(07)63240-5. [2] Hon, K.K.B., Li, L., Hutchings, I.M. (2008). Direct writing technology - Advances and developments. CIRP Annals - Manufacturing Technology, vol. 57, no. 2, p. 601-620, DOI:10.1016/j.cirp.2008.09.006.

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of Aerosol Science, vol. 31, no. 12, p. 1421-1431, DOI:10.1016/S0021-8502(00)00048-3. [18] Haglund, J.S., McFarland, A.R. (2004). A circumferential slot virtual impactor. Journal of Aerosol Science and Technology, vol. 38, p. 664-674, DOI:10.1080/02786820490486015. [19] Lee, P., Chen, D.-R., Pui, D.Y.H. (2003). Experimental study of a nanoparticle virtual impactor. Journal of Nanoparticle Research, vol. 5, no. 3-4, p. 269-280, DOI:10.1023/A:1025538930994. [20] Lim, K.S., Lee, K.W. (2006). Collection efficiency and particle loss of virtual impactors with different methods of increasing pressure drop. Journal of Aerosol Science, vol. 37, no. 10, p. 1188-1197, DOI:10.1016/j. jaerosci.2005.11.011. [21] Loo, B.W., Cork, C.P. (1988). Development of High Efficiency Virtual Impactors. Journal of Aerosol Science and Technology, vol. 9, no. 3, p. 167-176, DOI:10.1080/02786828808959205. [22] Noone, K.B., Heintzenberg, J. (1991). On the determination of droplet size distributions with the counterflow virtual impactor. Journal of Atmospheric Research, vol. 26, no. 5, p. 389-405, DOI:10.1016/01698095(91)90060-A. [23] US Patent No. 4,767,524 (1988). Virtual Impactor. United States Patent and Trademark Office, Alexandria.

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, 697-704 © 2013 Journal of Mechanical Engineering. All rights reserved. DOI:10.5545/sv-jme.2012.1002 Special Issue, Original Scientific Paper

Received for review: 2013-01-22 Received revised form: 2013-06-06 Accepted for publication: 2013-06-12

Polypropylene/Clay Nanocomposites Produced by Shear Controlled Orientation in Injection Moulding: Deformation and Fracture Properties

Costantino, A. – Pettarin, V. – Viana, J. – Pontes, A. –Pouzada, A. – Frontini, P. Alejandra Costantino1 – Valeria Pettarin1,* – Julio Viana2 – Antonio Pontes2 – Antonio Pouzada2 – Patricia Frontini1 1 National

University of Mar del Plata, Institute of Science and Technology of Materials, Argentina 2 University of Minho, Institute for Polymers and Composites, Portugal

The effect of distinct morphologies induced by shear controlled orientation in injection moulding (SCORIM) in the mechanical and fracture performance of polypropylene (PP) and PP/nanoclay mouldings is examined in this work. The effect of high shear conditions applied during processing was assessed. Samples exhibited a range of fracture stability ranging from a modest non-linearity to a quasi-stable regime depending on material type and injection conditions. Neat PP showed non-linear brittle behaviour while nanocomposites exhibited quasi-stable behaviour induced by the large deformation capability of the skin layer. Despite the fracture initiating at practically the same loading levels, the propagation energy varied with processing conditions and nanoclay content. The reduction of the core layer achieved by the SCORIM processing along with the differences between the skin and core favoured by the presence of nanoclay are responsible for the toughening of the SCORIM PP/nanoclay thick mouldings. Keywords: nanocomposites, polypropylene, SCORIM, fracture

0 INTRODUCTION Clays containing polymer nanocomposites are an alternative to conventional microcomposites due to improved mechanical, thermal, and processing properties as well as enhanced flammability resistance and barrier properties at very low filler loadings (<5 wt%). However, nanocomposites need sufficient stiffness, strength and toughness to meet particular design requirements. In structural and semi-structural applications of these materials, adequate fracture toughness, in addition to high stiffness and mechanical strength, is often required. The design of structural parts, their connecting and assembly may be based on fracture mechanical approaches. From an industrial perspective, the preparation of thermoplastic polymer nano-composites by melt blending using conventional plastics compounding equipment remains the solution of choice, especially when commodity polymers like polypropylene (PP), which is of great interest for the packaging and automotive industries, are used [1] to [4]. A lot of effort has been focused on the performance of PP nanocomposites prepared by melt compounding [5] to [8]. Manufacturing of PP/ nanoclay composites using conventional injection moulding (CIM) is not an easy task. Mouldings usually displayed poor nanoclay dispersion since the imposed thermomechanical conditions do not generate large enough stress fields to deagglomerate/ exfoliate the clay agglomerates completely [7] and [8]. Besides poor dispersion, the presence of voids and defects in thick specimens is a common feature in

many industrial mouldings. In fact, when PP/nanoclay thick specimens are obtained by CIM, large voids are generated during processing. These samples show an apparently stable fracture with a crack arrest regime induced by the presence of the large voids. These voids show partially relaxed triaxiality so that the samples reach different maximum load values in each case (see examples shown in Fig. 1 for the single edge notched bending (SENB) fracture test). As a consequence, it is very difficult to establish a reliable fracture parameter to be used in design calculations and prediction of part performance. Alternatively to CIM, the shear control orientation in injection moulding (SCORIM) process has been shown to offer potential benefits to the injection moulding process by controlling the microstructure of the moulded materials [9] and [10]. Significant and simultaneous improvements in the stiffness, strength and toughness, the better alignment of fillers, fibres and macromolecules, the moulding to closer tolerances, the enhancement of dimensional stability and reproducibility, the improvement in the aesthetic of the parts, the elimination of microcracks and voids in the mouldings, and the elimination of internal weld lines and sink marks have been reported for PP and other semicrystalline polymers [11]. However, little research effort has been devoted to the moulding of polymer nanocomposites by SCORIM. In fact, the high shear levels applied during SCORIM is expected to lead to better in-process dispersion and orientation of the fillers in the nanocomposite. In addition, there are previous promising findings on the use of SCORIM to manufacture

*Corr. Author’s Address: INTEMA, CONICET – UNMdP, Juan B. Justo 4302 – B7608FDQ Mar del Plata, Argentina, pettarin@fi.mdp.edu.ar

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PP/nanoclay composites [12] to [15], which have encouraged us to conduct the present research. Our results showed that for PP, which is a semicrystalline polymer, nanoclay acts as a morphology director, and the concomitant use of SCORIM and nanoclays results in distinct morphologies [16]. Nanoclays act as α- and γ-phase nucleating agents, increasing the degree of epitaxiality in the skin of the PP/nanoclay mouldings. When compared with CIM, SCORIM results in a higher degree of crystallinity of the skin layer. SCORIM mouldings exhibit a well-defined ‘shear’ -core structure: a centre-core that occupies less space in comparison with typical CIM moulded parts and a larger shear zone. When the nanoclay is added, an interesting multilayered structure occupies more than two times the thickness of the shear zone of the PP/nanoclay pieces, i.e. there is even less core. It is therefore expected that these changes will be reflected in the mechanical and fracture properties of the mouldings.

Fig. 1. Load vs. displacement curves for equivalent PP/nanoclay mouldings processed by CIM; differences in the flexural loaddisplacement curves for equivalent fracture samples are clearly seen

The present work investigates the effect of distinct morphologies induced by SCORIM in the mechanical and fracture performance of the PP/ nanoclay mouldings. 1 MATERIALS AND METHODS 1.1 Sample Manufacturing Characteristics

and

Microstructural

Compounds were based on Domolen 1100L homopolymer PP and 6, 10 and 14% of P-802 Nanocor masterbatch of 50% PP / 50% nanoclay 698

(MB). The materials were first blended in a drum rotating at 60 rpm. Double edge gated parallelepiped bars of 130×13×8 mm were injection moulded using a Ferromatik Milacron K-85A injection moulding machine, equipped with a SCORIM mould with a manifold and hydraulic system, which manipulates the melt inside the mould impression. The hydraulic pistons operating in various modes displace the melt during the holding phase of the moulding process, shearing the solid/melt interface and aligning fillers and polymer molecules. Table 1. Variable injection processing set-up for SCORIM Run 1 2 3 4 5 6 7 8 9

Stroke time 1 (1 s) 1 1 2 (2 s) 2 2 3 (3 s) 3 3

Melt temperature 1 (200 °C) 2 (220 °C) 3 (240 °C) 2 3 1 3 1 2

Stroke number 1 (3) 2 (7) 3 (13) 3 1 2 2 3 1

The moulding programme was defined according to a design of experiments (DOE) based on a 3-level factorial array (Table 1). The processing parameters considered were melt temperature (200, 220 and 240 °C), stroke time (i.e., piston movement time of 1, 2 and 3 s) and stroke number (i.e., number of piston movements: 3, 7 and 13). The other injection moulding parameters were kept constant: a mould temperature of 30 °C, an injection pressure of 15 MPa, an injection velocity of 10 mm/s, a holding pressure of 5.3 MPa and a cooling time of 30 s. Details are as reported in [16]. A complete characterisation of the microstructure of these mouldings has already been reported [16]. An intercalated PP/nanoclay structure was obtained with an intergallery space of 13.4 nm for all systems, as calculated from the (001) peak of X-ray diffraction patterns using the Bragg’s law (λ = 2d001 sinθ001). Several characteristics of PP morphology were analyzed and two of them are crucial: the degree of orientation of α-PP crystallites and the amount of γ-PP phase in the skin layer. The degree of orientation of α-PP was quantified by means of A110 = Iα(110) / (Iα(110) + Iα(111) + Iα(131)+(041)), which gives an indication of the orientation roughly parallel to the c-axis of the crystallites, which in this case corresponds to the flow direction in the injected samples. Ii is the i peak height of XRD after background subtraction. A strong increase in the

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Iα(040) / Iα(110) ratio for the SCORIM PP-nanoclay samples was found. This indicates that in SCORIM mouldings the nanoclay induced the orientation of the α crystalline-phase of PP in the (040) direction – i.e. increased the degree of epitaxiality [17] – and imparted a low level of crystalline phase orientation in the flow direction. Values of the A110 index are presented in Table 2. Regarding the effect of the SCORIM processing conditions on the morphology development, it was found that A110 increases with high shearing, and decreases with low shearing and low melt temperatures. The content of the crystalline γ-phase was determined using the following equation G  =  Iγ(117) / (Iγ(1117) + (Iα(130))) and the results are shown in Table 2. The amount of γ increases with MB content in an asymptotic way to a maximum value, which is in agreement with other authors who have claimed that nanoclay [18] to [22] and SCORIM [23] promote the formation of γ-phase PP crystallites. The nanoclay changes the equilibrium state of the polymer (conformation) and it provides favorable sites for possible epitaxial growth of the γ-phase because the lattice mismatch is less than 10% [19]. Moreover, γ crystals are nucleated and grow epitaxially on the lateral (010) faces of the α crystal [24] and appear to be favored by, or linked to, the absence of chain folding [25]. The mobility of the PP-MA matrix is significantly reduced in the presence of maleic anhydride grafting in the main chain (present in MB), which causes a decrease in chain folding. In the presence of clay particles in polymer nanocomposites, the movement of polymer chains inside the clay particles is restricted. The correlation length of the clay particles is roughly the same as that of the radius of gyration of the matrix. Thus, the formation of the γ-phase is enhanced in the presence of clay particles [21] and [22]. Table 2. Orientation indexes and γ-PP content for SCORIM samples Run 1 2 3 4 5 6 7 8 9

A110 (PP/MB-6) 0.158 0.165 0.186 0.187 0.194 0.200 0.189 0.190 0.189

Material PP PP/MB-6 PP/MB-10 PP/MB-14

G 0 0.48 0.52 0.55

For more information about the macro and microstructure of these mouldings readers are referred to [16]. 1.2 Mechanical Performance The yield stress of the material in the mouldings was determined in compression using prismatic specimens of rectangular cross section (h = 10 mm, h/L =  1.5 h, and L being the height and the width of the sample). These samples were cut out from moulded parts. Smooth and parallel faces were obtained by machining. The tests were performed using an Instron 4467 universal testing machine, at 1 mm/min and room temperature. To avoid the eventual influence of voids developed during injection moulding, the start of the plastic domain, i.e. the start of yielding, was considered to be the end of the linear portion of the stress-strain plot [26]. The flexural modulus was measured according to the ASTM D 790-03 standard, at 5 mm/min and room temperature. At least five samples of each processing condition were tested. 1.3 Fracture Characterisation The fracture characterisation was performed on SENB deformed in 3-point bending at room temperature at a constant crosshead velocity of 5 mm/min in the Instron 4467. Specimens with span, S, and width, W, of 52 and 13 mm, respectively, were prepared by machining from the injected bars. Sharp notches were introduced using a Notchvis Ceast machine with a sharp fly cutter (tip radius less than 12 mm) to reach a crack-to-depth ratio (a/W) equal to 0.5. The fracture initiation was evaluated from the critical stress intensity factor in mode I, KIC, (Eq. (4)) using load-line displacement plots and the crack length: P K IC = f (a / W ), (1) BW 1/ 2 where P is the load, B is the sample breadth, W is the sample width, a is the crack length and f  (a/W) is a dimensionless function of a/W according to ASTM D 5045. Due to the non-linearity of the load-displacement plots, P was determined at the intersection of a straight line with a compliance of 5% greater than that of the initial compliance and the load-displacement trace. In every case the geometry validity requirements were met according to [27]:

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 K B, a,(W − a ) ≥ 2.5 ⋅  σy 

2

  . (2) 

The value of the J integral was computed from:

J=

2 ⋅U . (3) B (W − a )

The energy, U, was computed by the integration of the load-displacement traces up to the maximum load point (Jmax) and up to the instability load point (Jc) [28]. This value of J was computed to designate the point of unstable fracture, or loss of structural integrity, in samples displaying quasi-stable crack propagation. 2 RESULTS AND DISCUSSION 2.1 Deformation Performance

thereafter. Besides the amount of nanoclay, the intercalated nanoclay morphology may also restrict the mobility of the polymer chains and contribute to improvement in the modulus. Furthermore, in flexural tests, the mechanical response is governed by the most exterior layer, i.e. the skin of injected samples. In these samples, the PP microstructure of the skin layer changed with the addition of nanoclay, although the microstructural features are independent of the nanoclay content [16]. Due to the segregation effect of the nanoclay in the skin, it seems that the effective content of nanoclay is similarly independent of MB concentration [16], and hence the modulus. This can be corroborated by observing the skin zone using Transmission Electron Microscopy (TEM) (as shown in Fig 3). A marked nanoclay orientation along the flow direction in the skin layer of the injected samples is visible, but complete exfoliation cannot be observed either in the skin or the core zones.

The yield stress of PP nanocomposites was slightly larger than in neat PP, but varied very little with the nanoclay content (Fig. 2). The trends reported in the literature are contradictory, reporting either an increase or a decrease in properties with the addition of nanoclay [29].

a)

Fig. 2. Mechanical properties of PP nanocomposites

A quite significant increase in the flexural modulus of nanocomposites in comparison with neat PP was also found, on the order of 35 to 40% (Fig. 2). From theoretical predictions using the MoriTanaka composite micromechanical model, it is possible to estimate an increase in the elastic modulus of PP with 3 to 6% of nanoclay of the same order [30]. For lower incorporation of nanoclay, in the range between 1 and 6%, the variations in the elastic modulus are negligible, an increase being observed 700

b)

Fig. 3. TEM images of PP/MB-5: a) skin, and b) core

2.2 Fracture Performance The samples exhibited different fracture stabilities ranging from a modest non-linearity to a quasi-stable regime. The fracture regime depends on the type of

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material and the injection moulding conditions (Fig. 4). PP shows a non-linear behaviour and unstable brittle fracture after reaching the maximum load. The load-displacement curves dropped to zero immediately after reaching the maximum load with relatively short displacement (Fig. 4).

macromolecules (fostered by the oriented nanoclays), thus improving the deformation capability. In fact, we must point out that the occurrence of necking is likely a combined effect of: the skin/core ratio, the skin orientation level, and possibly the morphology rearrangement in the skin region under different loading directions, which are also influenced by the heat evolution and the dissipation processes. Since SCORIM induces a thicker multi-layered skin in the PP/nanoclay mouldings [16], the ‘skin effect’ is favoured. This phenomenon leads to differences in the load-displacement curves, promoting a more stable crack propagation followed by a necking phase with a much larger extension [8], [32] and [33]. When voids appear in the fracture surface, they inhibit or instabilize the ‘skin-effect’ due to stress concentration.

Fig. 5. PP/MB-6 sample after the fracture test: note the highly deformed skin without fracture

Fig. 4. Fracture load vs displacement curves for SCORIM injected a) PP and b) PP/MB-6 respectively

The composite mouldings produced by SCORIM were of acceptable quality and showed various loaddisplacement patterns. Most of the samples exhibited a quasi-stable behaviour: the load increased nonlinearly, and then it kept constant up to a certain displacement when a drastic drop occurred (Fig. 4). Moreover, in most of the samples a large deformation of the skin layer was observed (Fig. 5). It should be mentioned here that this ‘skin effect’ is characterised by tail instability load-displacement curves, which have already reported for PP based systems [31] and [32]. This large deformation of the skin layer may be related to the high level of molecular and nanoclay orientation (Fig. 3) that favours the sliding of the

For these PP/nanoclay composites, small differences were seen in the KIC and Jmax (which are related to the fracture initiation) (Table 2), while large differences were seen in the Jc (which is related to fracture propagation) (Table 3), in comparison with neat PP samples. Replicas of each condition show significant scatter originating from a combination of different causes. The differences observed are typical of semicrystalline polymers injection mouldings that are influenced by microstructural parameters, which include crystalline structure, degree of crystallinity, supermolecular structure, and skin-core configuration [34]. In addition, PP systems are considered to be in the ductile-to-brittle transition regime at room temperature, which implies a larger scatter [35] due to dissimilar contributions of the plastics melt flow and the stable crack growth seen in different specimens of the same set.

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Table 2. Initiation fracture parameters of the mouldings Run 1 2 3 4 5 6 7 8 9

KIC [MPa·m1/2] PP PP/MB-6 1.4±0.2 1.6±0.3 1.45±0.3 1.5±0.3 1.25±0.1 1.3±0.2 1.5±0.3 1.6±0.1 1.3±0.2 1.4±0.3 1.3±0.1 1.7±0.2 1.3±0.2 1.6±0.1 1.6±0.4 1.8±0.1 1.2±0.1 1.4±0.3

Jmax [N/mm] PP PP/MB-6 6.7±1 10.25±3.7 6±1 7±0.8 7±1 5±2 11±8.7 7.75±1 7.2±1.6 5±1.2 8.4±1.9 7.5±2 8.8±3.8 7.3±1.5 8.4±2.1 11±2.5 6±1.2 6.5±2

The ANOVA analysis of the results indicates that the increment of the shear level (higher stroke number and longer time) leads to a decrease in Jc. The correlation of these results with the results of the processing induced morphology [16] seems to indicate that the most relevant morphological feature is the level of orientation at the skin in the flow direction: higher orientation levels lead to lower toughness (see Fig. 6), keeping in mind that higher orientation implies higher γ-PP content.

Table 3. Jc data for the mouldings Run 1 2 3 4 5 6 7 8 9

Jc [N/mm] All samples Samples with no visible defects PP PP/MB-6 PP PP/MB-6 6.7±1 58.2±35.9 6.7±1 94.5±7.8 6±1 20±13.7 6±1 * 7±1 54±25.8 7±1 65.5±3 11±8.7 10.8±3.5 7.2±0.9 8.3±1.2 7.2±1.6 14.8±8 6.5±0.6 23.5±2.1 8.4±1.9 10.4±3.3 8.4±1.9 10.4±3.2 8.8±3.8 11.3±4.7 7±0.01 9±1.7 8.4±2.1 40.4±35.4 8.4±2.1 78.5±13 6±1.2 16.4±11.9 6±1.2 *

Fig. 6. Influence of skin orientation of PP crystallites (A110) on toughness of PP/MB-6

* Neglected behaviour

The addition of nanoclay improved the fracture performance of the SCORIM PP/nanoclay samples thanks to the differences induced by the fillers in the microstructure: as we stated previously [16], the fillers refine the crystal structure and orient the α-PP crystals thus promoting epitaxial growth of the γ-phase, which improves toughness [18]. With regard to this last point, several studies have demonstrated the coexistence and concurrent growth of α and γ lamellae, and that there is a continual decrease in lamellar thickness with γ content [36]. It has also been shown that toughness decreases linearly with the reciprocal value of lamellar thickness [37]. Moreover, polypropylenes with high contents of the γ phase show behavior typical of stiff-plastic materials, i.e. a high value of the elastic modulus but with high ductility [38]. Therefore, it is concluded that the generated γ-phase in the skin of the samples induces a large-scale plastic deformation of the skin with consequent tearing of matrix ligaments leading to fibrillation and enhanced toughness. The best global results (skin effect with a high value of J and acceptable reproducibility) were achieved using processing conditions that imply low injection temperature and shear forces. 702

3 CONCLUSIONS The characterisation of PP/nanoclay mouldings produced by SCORIM leads to the following conclusions: The SCORIM mouldings of neat PP showed a nonlinear brittle behaviour, whereas PP nanocomposites exhibited a quasi-stable behaviour induced by a larger deformability of the skin layer. Despite the fracture initiating at practically the same loading levels, the overall crack propagation energy values varied as a function of the processing conditions. The statistical analysis indicates that the reduction in core size achieved in the SCORIM processing, along with the differences between the skin (or shear zone) and core zones, and the favourable effect of the presence of nanoclays, are responsible for the improvement in the toughness of the SCORIM PP/ nanoclay thick samples. 4 REFERENCES [1] Alexandre, M., Dubois, P. (2000). Polymer-layered silicate nanocomposites: preparation, properties and

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Costantino, A. – Pettarin, V. – Viana, J. – Pontes, A. –Pouzada, A. – Frontini, P.


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11 Vsebina

Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 59, (2013), številka 11 Ljubljana, november 2013 ISSN 0039-2480 Izhaja mesečno

Gostujoči uvodnik

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Razširjeni povzetki člankov Fantina Rosa Esteves, Tiago Alexandre Carvalho, António Sérgio Pouzada, Carla Isabel Martins: Vrednotenje zgoščevanja in geometrije aluminijastih delov, izdelanih po postopku posrednega laserskega sintranja kompozitnega prahu aluminija in polistirena Jan Deckers, Jean-Pierre Kruth, Ludwig Cardon, Khuram Shahzad, Jef Vleugels: Vrednotenje zgoščevanja in geometrije aluminijastih delov, izdelanih po postopku posrednega laserskega sintranja kompozitnega prahu aluminija in polistirena Markus Gottfried Battisti, Walter Friesenbichler: Predelava nanokompozitov PP z brizgalnim kompaunderjem Kim Ragaert, Filip De Somer, Stieven Van de Velde, Joris Degrieck, Ludwig Cardon: Metode za izboljšanje upogibnih lastnosti 3D-natisnjenih ogrodij iz PCL za tkivni inženiring srčnih zaklopk Vito Speranza, Umberto Vietri, Roberto Pantani: Nadzor brizganja termoplastov: uporaba tlačnih pretvornikov za ocenjevanje poteka strjevanja in krčenja brizgancev Giovanni Lucchetta, Marco Fiorotto: Vpliv hitrih sprememb temperature v orodju na videz brizganih izdelkov Frederik Vogeler, Wesley Verheecke, André Voet, Hans Valkenaers: Prva študija izdelave povezav po postopku AerosolJet® na substratih, izdelanih z ekstruzijskim 3D-nalaganjem Alejandra Costantino, Valeria Pettarin, Julio Viana, Antonio Pontes, Antonio Pouzada, Patricia Frontini: Nanokompoziti polipropilena/gline, izdelani po postopku brizganja s strižno nadzorovanim usmerjanjem: deformacijske in lomne lastnosti

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Osebne vesti Doktorske disertacije, diplomska dela

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SI 133 SI 134 SI 135 SI 136 SI 137 SI 138 SI 139


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11 Gostujoči uvodnik

Gostujoči uvodnik Tematska številka: Inovacije v polimerih in orodjih Tematska številka je posvečena bienalni konferenci o inovacijah v polimerih in orodjih PMI 2012. Skupaj jo organizirata Institut za polimere in kompozite pri Univerzi v Minhu na Portugalskem ter Center za polimere in materiale pri Univerzi v Ghentu, Belgija. V njej je osem dvojno recenziranih razširjenih in dopolnjenih člankov, izbranih med več kot 60 članki, sprejetimi in predstavljenimi na konferenci PMI 2012. Ti članki pokrivajo področja raziskav, ki segajo od novosti pri razvoju plastičnih izdelkov do praktične implementacije novih konceptov predelave. Obravnavane so bile naslednje teme: • razvoj pri predelavi polimerov, • hibridna orodja, • simulacija in toplotni nadzor orodij, • inovacije pri polimerih in materialih, • recikliranje polimerov, • dodajalne izdelovalne tehnologije, • hitra izdelava prototipov in orodij ter • trendi pri razvoju izdelkov. • • • • • • • •

Posebna pozornost je bila posvečena naslednjim področjem: Preučitev vpliva pogojev predelave na estetske, morfološke in mehanske lastnosti brizgancev iz strukturne pene polistirena, odpornega na udarce (HIPS-SF); Postopek metalurgije prahov za izdelavo aluminijastih delov s posrednim selektivnim laserskim sintranjem (SLS®) kompozitnega prahu aluminija in polistirena; Primerjava različnih tehnik kompaundiranja in njihov vpliv na modul elastičnosti pri konvencionalni predelavi polimernih nanokompozitov in pri predelavi v edinstvenem brizgalnem kompaunderju (PNC-IMC); Preučitev dveh različnih pristopov za izboljšanje fleksibilnosti 3D-natisnjenih ogrodij iz PCL za lističe srčnih zaklopk; Študija vrste postopkov brizganja polistirena (PS) za splošne namene ter spreminjanja zadrževalnega tlaka, temperature brizganja in dimenzij gnezd; Razvoj inovativne tehnologije za hitro ogrevanje in ohlajevanje orodij za brizganje ter uporaba te tehnologije za analizo vpliva hitrih sprememb temperature orodja na izboljšanje videza (sijaja) brizgancev; Raziskava izdelave povezav po postopku AJP na ekstrudiranih 3D-natisnjenih substratih; Raziskava vpliva različnih morfologij, ki nastanejo ob strižno nadzorovanem usmerjanju pri brizganju (SCORIM), na mehanske in lomne lastnosti polipropilena (PP) ter PP/nanogline.

Zahvaljujemo se uredništvu Strojniškega vestnika – Journal of Mechanical Engineering, urednikom zbornika konference in znanstvenemu odboru za njihovo podporo in spodbudo pri pripravi te tematske številke. Zahvaljujemo se tudi vsem avtorjem in anonimnim recenzentom člankov za njihov čas in trud. Gostujoči uredniki Prof. António Sérgio Pouzada Prof. Ludwig Cardon Prof. Jože Balič

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 133 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-08-05 Odobreno za objavo: 2013-08-20

Vpliv predelave na estetske, morfološke in mehanske lastnosti brizgancev iz strukturne pene polistirena z visoko odpornostjo proti udarcem Fantina Rosa Esteves ‒ Tiago Alexandre Carvalho ‒ António Sérgio Pouzada ‒ Carla Isabel Martins* Institut za polimere in kompozite/I3N, Univerza v Minhu, Portugalska

Poceni proizvodnja manjših količin velikih plastičnih izdelkov zahteva uporabo alternativnih materialov in orodij. Strukturne pene (SF) so rešitev pri debelejših izdelkih z dobrimi lastnostmi. Zaradi njihove sendvič strukture, sestavljene iz celičnega jedra in dveh polnih plasti na površju, je iz njih mogoče izdelovati lahke dele z veliko togostjo in dobro dimenzijsko stabilnostjo. Porozno jedro se oblikuje z dodajanjem posebnega sredstva za penjenje v polimerni matriks. Ti materiali se uporabljajo pri urbanem pohištvu, v avtomobilski, navtični, letalski in vesoljski industriji. Za izdelavo SF se najpogosteje uporablja postopek nizkotlačnega brizganja, proces s kratkim brizgom in tlaki pod 4 MPa. Hibridna orodja z gnezdi, izdelanimi po postopkih hitre izdelave, so alternativa brizganju velikih plastičnih delov v cenejših orodjih. V članku je predstavljen vpliv pogojev obdelave in uporabljenega orodja na estetske, morfološke in mehanske lastnosti brizgancev iz strukturne pene polistirena, odpornega na udarce (HIPS). Rešitev s hibridnim orodjem je primerjana s konvencionalnim orodjem. Brizganci iz materiala HIPS (specifična teža 1,05 Mg.m–3 in MFI 10,95 g/10 min (200 ºC/5 kg) proizvajalca BASF) z 2 utež.% endotermnega CBA (Tracel IMC 4200SP proizvajalca Tramaco in Nemčije) so bili izdelani v stroju za brizganje plastike Engel Victory Spex 50. V študiji sta bili uporabljeni dve gnezdi, eno konvencionalno iz jekla in drugo hibridno iz epoksija Biresin L74, polnjenjega s 60 utež. % aluminijastega prahu, izdelano z vakuumskim litjem in obdelano na končno geometrijo izdelka. Gnezda so bila napolnjena s talino do 90%, po ekspanziji CBA pa so se izpolnila do konca. Temperatura brizganja se je spreminjala (200-220-240 ºC). Ostali pogoji predelave so bili prilagojeni glede na uporabljeno orodno gnezdo (jekleno ali hibridno) – temperatura orodja 55 °C za hibridno orodje in 75 °C za jekleno orodje, s čimer je bila upoštevana različna toplotna prevodnost obeh materialov. Pri preučitvi vpliva pogojev predelave je bila upoštevana nastala celična morfologija (ta se preučuje z optično in vrstično elektronsko mikroskopijo), upogibna togost orodnih plošč (ta se preskuša s tritočkovno metodo po Pouzadi in Stevensu) in udarna žilavost (merjeno z udarnim preskusom po standardu EN ISO 6603-1). Ovrednotene so bile tudi druge lastnosti kot so sijaj (ASTM D523-85), hrapavost (s pomočjo prototipne laserske naprave za mikrotopografijo) in gostota (ASTM standard D 792-00). Pogoji predelave vplivajo na estetske, morfološke in mehanske lastnosti brizgancev HIPS-SF, narejenih v hibridnih in konvencionalnih jeklenih orodjih. Tipična morofloška struktura, ki jo pridobimo s postopkom brizganja HIPS-SF, ima dve zunanji trdni plasti (lupino) in celično jedro. Odvisno od vrste uporabljenega orodja se oblikujejo različne mikrostrukture. Pri jeklenem orodju je struktura simetrična s celicami povprečne velikosti 80 μm po razsežnosti debeline, pri hibridnem orodju pa se oblikuje asimetrična sendvič struktura z večjimi celicami (povprečno 200  μm). V obeh primerih se velikost celic zaradi razlik v temperaturi taline zmanjšuje od sredine proti lupini. S povečevanjem temperature brizganja se povečuje tudi velikost celic zaradi manjše viskoznosti in manjšega upora pri rasti celic. Zmanjšuje se tudi delež debeline lupine v debelini brizganca, kar povzroči manjšo gostoto in zmanjšanje upogibne togosti. Brizganci HIPS-SF iz hibridnega orodja so zaradi večjega deleža debeline lupine in večje gostote nekoliko bolj togi od brizgancev iz jeklenega orodja. HIPS-SF ima manjšo hrapavost in večji sijaj kot HIPS, saj je tlak, potreben za izpolnitev kalupa z ekspanzijo sredstva za penjenje, manjši kot pri konvencionalnih brizgancih iz nepenjenega HIPS. Preslikava površine s HIPSSF je zato manj natančna kot pri nepenjenem HIPS. Brizganci iz hibridnih orodij imajo večjo hrapavost in zato manjši sijaj. Hrapavost se povečuje z višanjem temperature brizganja. Brizganci pri udarnem preskusu kažejo duktilne lastnosti in vršna energija pri testu s padajočo utežjo se zmanjšuje s povečevanjem temperature brizganja. Obstojnost proti napredovanju razpok se poslabša zaradi oblikovanja velikih neenakomernih celic v jedru brizganca, kar postaja vse bolj izrazito pri višjih temperaturah taline. Mehansko vedenje brizgancev HIPS-SF je bilo napovedano z modeli po Barzegariju. Model je rezultate eksperimenta napovedal z manj kot 10% napake pri najboljšem modelu. Ključne besede: strukturna pena, hibridno orodje, debelina lupine, mehanska trdnost *Naslov avtorja za dopisovanje: Institut za polimere in kompozite/I3N, Univerza v Minhu, Campus de Azurem, Guimarães, Portugalska, cmartins@dep.uminho.pt

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 134 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-04-19 Odobreno za objavo: 2013-05-06

Vrednotenje zgoščevanja in geometrije aluminijastih delov, izdelanih po postopku posrednega laserskega sintranja kompozitnega prahu aluminija in polistirena Deckers, J. – Kruth, J.-P. – Cardon, L. – Shahzad, K. – Vleugels, J. Jan Deckers1,* – Jean-Pierre Kruth1 – Ludwig Cardon2 – Khuram Shahzad3 – Jef Vleugels3

2 Raziskovalna

1 Oddelek za strojništvo, KU Leuven, Belgija skupina CPMT, Pridružena fakulteta za aplikativne tehniške vede, Univerzitetni kolidž v Ghentu, Belgija 3 Oddelek

za metalurgijo in materiale, KU Leuven, Heverlee, Belgija

Cilj tega članka je ovrednotenje novega postopka na področju metalurgije prahov za izdelavo aluminijastih delov s posrednim selektivnim laserskim sintranjem (SLS). Definicija problema: Glavni namen tega dela je izdelava tehničnih keramičnih delov visoke gostote po postopku selektivnega laserskega sintranja. V ta namen so uvedene strategije zgoščevanja kot dodatni koraki procesa PM. Strategije zgoščevanja vključujejo vroče izostatsko stiskanje (WIP) in infiltracijo. Pri slednji se za zapolnjevanje odprtih por uporablja suspenzija aluminija v etanolu. Metodologija: Deli so izdelani s sintezo aglomeratov kompozita aluminija in polistirena z disperzijsko polimerizacijo, SLS, odstranjevanjem veziva in sintranjem v trdnem stanju (SSS). Vroče izostatsko stiskanje in različne strategije infiltracije so bile uporabljene kot dodatni koraki postopka PM za povečanje končne gostote izdelkov. Opravljeni so bili eksperimenti tako s tlačno infiltracijo, torej z uporabo zunanjega tlaka za potiskanje suspenzije v pore izdelka, kakor tudi z breztlačno infiltracijo. Preučena je bila tudi infiltracija zelenih kosov; kosov, začetno sintranih v trdnem stanju in kosov, popolnoma sintranih v trdnem stanju. Po različnih fazah procesa so bile opravljene meritve gostote, geometrijsko vrednotenje in mikrostrukturna analiza z vrstičnim elektronskim mikroskopom (SEM). Rezultati, ugotovitve: Če niso bili uporabljeni dodatni koraki zgoščevanja, je bilo linearno krčenje delov SLS med odstranjevanjem veziva in SSS približno 30-odstotno. Končni izdelki so imeli veliko malih razpok, gostota pa je znašala 66%. Z vročim izostatskim stiskanjem je bilo mogoče povečati gostoto zelenih delov, ne pa tudi končne gostote. Vsak končni izdelek, obdelan z WIP, je imel eno veliko razpoko. Z infiltracijo se je v splošnem zmanjšalo krčenje lasersko sintranih delov med odstranjevanjem veziva in sintranjem v trdnem stanju. Večina infiltriranih kosov je imela eno veliko razpoko. Uporaba tlačne infiltracije na začetno v trdnem stanju sintranih delih je omogočila povečanje gostote delov do 84%, saj so se razpoke, ki so nastale med postopkom odstranjevanja veziva, napolnile z aluminijem. Kljub temu pa so nastale mikrorazpoke, verjetno zaradi nehomogenega krčenja med sintranjem v trdnem stanju. Omejitve raziskave, implikacije: V sintranih delih so ostale razpoke. Homogenizacija mikrostrukture delov z optimizacijo začetnega kompozitnega prahu, nalaganje med selektivnim laserskim sintranjem, parametri sintranja, parametri vročega izostatskega stiskanja in parametri infiltracije so ključni za odpravo teh pomanjkljivosti. Prispevek, novosti, vrednost: Amorfni termoplastični polistiren prej še ni bil uporabljen za izdelavo delov iz čistega aluminija po postopku posrednega laserskega sintranja. Kombinirana uporaba SLS in WIP je novost na področju posrednega laserskega sintranja keramike. Inovacija je tudi kombinirana uporaba SLS in različnih tehnik infiltracije (tlačna ali breztlačna, infiltracija zelenih kosov; kosov, začetno sintranih v trdnem stanju in kosov, popolnoma sintranih v trdnem stanju). Ključne besede: dodajalne izdelovalne tehnologije, keramika, posredno lasersko sintranje, aluminij, polistiren

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*Naslov avtorja za dopisovanje: Oddelek za strojništvo, KU Leuven, Celestijnenlaan 300B, 3001 Heverlee, Belgija, Jan.Deckers@mech.kuleuven.be


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 135 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2012-12-07 Prejeto popravljeno: 2013-03-15 Odobreno za objavo: 2013-03-29

Predelava nanokompozitov PP z brizgalnim kompaunderjem Battisti, M.G. – Friesenbichler, W. Markus Gottfried Battisti* – Walter Friesenbichler

Univerza v Leobnu, Katedra za injekcijsko brizganje polimerov, Oddelek za znanost in inženiring polimerov, Avstrija

Cilj študije je primerjava različnih tehnologij kompaundiranja ter določitev njihovega vpliva na modul elastičnosti. Ovrednoteni so bili konvencionalni postopki predelave nanokompozitov in edinstveni brizgalni kompaunder (PNC-IMC). Raziskan je bil tudi vpliv nanopolnil na toplotno prevodnost polimernih nanokompozitov pri različnih tlakih. Za razliko od konvencionalnih postopkov kompaundiranja, kjer je treba iz kompaunda pripraviti pelete in jih dovajati v brizgalni stroj za drugi proces plastifikacije, sta v brizgalnem kompaunderju oba koraka predelave združena. Kompaundiranje materiala in nato brizganje se izvajata neposredno v enem samem procesu plastifikacije s pomočjo ogrevanega voda za talino in zbiralnika taline. V tej študiji sta bili obe tehnologiji uporabljeni za proizvodnjo polimernih nanokompozitov. V članku so predstavljeni vplivi tehnologij predelave, hitrosti polža, protitlaka in dolžin ekstruderja na modul elastičnosti. Dokazano je, da je za izboljšavo procesa nujno iskanje kompromisa med vnosom strižne energije in časom zadrževanja. Prikazano je povečanje toplotne prevodnosti pri uporabi nanopolnil v primerjavi s čistim polipropilenom. Povečana toplotna prevodnost je zelo zanimiva za industrijo, saj omogoča skrajšanje delovnega cikla pri injekcijskem brizganju. Prvi rezultati dajejo dober pregled zmožnosti in omejitev pri inovativnem konceptu PNC-IMC. Nadaljnje študije bodo usmerjene v mehanizme eksfoliacije in interkalacije silikatov v plasteh v polimerni talini. Ključne besede: brizgalni kompaunder, polimerni nanokompoziti, toplotna prevodnost, eksfoliacija, interkalacija, silikati v plasteh

*Naslov avtorja za dopisovanje: Univerza v Leobnu, Katedra za injekcijsko brizganje polimerov, Oddelek za znanost in inženiring polimerov, Avstrija, markus.battisti@unileoben.ac.at

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Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 136 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-04-10 Odobreno za objavo: 2013-05-06

Metode za izboljšanje upogibnih lastnosti 3D-natisnjenih ogrodij iz PCL za tkivni inženiring srčnih zaklopk

Ragaert, K. – De Somer, F. – Van de Velde, S. – Degrieck, J. – Cardon, L. Kim Ragaert1,2,* – Filip De Somer3 – Stieven Van de Velde1 – Joris Degrieck2 – Ludwig Cardon1,2 1 Univerzitetni

kolidž v Ghentu, Pridružena fakulteta za aplikativne tehniške vede, Belgija v Ghentu, Fakulteta za inženiring in arhitekturo, Belgija 3 Univerzitetna bolnišnica v Ghentu, Kardiološki center, Belgija

2 Univerza

Poli-e-kaprolakton (PCL) je znan polimer za 3D-tiskanje ogrodij, ki se uporabljajo pri tkivnem inženiringu. Cilj raziskave je izdelava fleksibilnejših ogrodij na osnovi PCL za aplikacije kot so lističi srčnih zaklopk. Obstoječa ogrodja iz PCL za tkivni inženiring lističev srčnih zaklopk so v primerjavi z naravnimi preveč toga, zato prihaja do elastomehanskega neujemanja vsadka z okolico ter do suboptimalne (mehanske) stimulacije celic. Raziskani sta bili dve metodi za izboljšanje fleksibilnosti 3D-natisnjenih ogrodij iz PCL za lističe srčnih zaklopk. Geometrija ogrodja je bila prvič radikalno spremenjena v zelo odprto, tkanini podobno strukturo, in sicer z ustrezno prilagoditvijo parametrov 3D-tiskanja. Sam osnovni material je bil spremenjen z dodatkom deleža polietilen oksida (PEO) z majhno molekulsko maso v polimer PCL. V tej eksploratorni študiji je bilo ugotovljeno, da je togost ogrodja na osnovi PCL mogoče zmanjšati (i) s prilagoditvijo geometrije ogrodja v odprto, tkanini podobno strukturo, ter (ii) s spremembo osnovnega materiala z dodatkom deleža polimera z manjšo trdnostjo. Z bolj odprto strukturo in povešanjem posameznih filamentov v pore pod njimi je bilo prvič v okviru te raziskave mogoče izdelati ogrodja PCL z upogibno togostjo, ki je manjša kot pri naravnem tkivu. S tem je bil odpravljen dolgotrajni problem večje togosti PCL v primerjavi z naravno loputko. Z dodatkom PEO v PLC se še dodatno poveča fleksibilnost osnovnega materiala s stopnjo, ki je sorazmerna količini dodanega PEO. Krajše verige PEO migrirajo proti površini ekstrudiranih filamentov, kjer prispevajo k hidrofilnosti površine ogrodja. V prihodnosti bo mogoče s kombinacijo geometrije in materiala ogrodja lističev natančneje posnemati upogibno vedenje naravnega tkiva. Članek podaja prispevek k razvoju ogrodij za inženiring tkiva srčnih zaklopk z identifikacijo dveh metod za povečanje fleksibilnosti konstrukcij na osnovi PCL. Raziskava omogoča fino prilagoditev fleksibilnosti s kombiniranjem obeh rešitev, rezultat pa je ogrodje s podobnim odzivom na obremenitve kot pri naravnem tkivu. Ključne besede: ogrodje, mehanske lastnosti, upogibanje, srčne zaklopke, tkivni inženiring, 3D-tiskanje

SI 136

*Naslov avtorja za dopisovanje: Univerzitetni kolidž v Ghentu, Pridružena fakulteta za aplikativne tehniške vede, Voskenslaan 362, 9000 Ghent, Belgija, kim.ragaert@ugent.be


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 137 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-04-26 Odobreno za objavo: 2013-05-17

Nadzor brizganja termoplastov: uporaba tlačnih pretvornikov za ocenjevanje poteka strjevanja in krčenja brizgancev Speranza, V. – Vietri, U. – Pantani, R. Vito Speranza – Umberto Vietri – Roberto Pantani*

Oddelek za industrijski inženiring, Univerza v Salernu, Italija

Pojav krčenja, ki je popisan z odstotno razliko v merah brizganega izdelka in gnezda, v katerem se polimer strdi, je izjemno pomemben za vse tiste, ki se ukvarjajo z brizganjem plastike, in je ključen parameter za vrednotenje kakovosti. Tradicionalni temperaturni in tlačni pretvorniki so v industriji brizganja plastike sicer zelo razširjeni, kljub temu pa korelacija med izmerjenim potekom temperature in tlaka ter kakovostjo izdelka ni neposredna. Niti celotna tlačna krivulja namreč ne zadošča za popoln popis krčenja izdelka. Po drugi strani je iz literature dobro znano, da je primeren parameter za popis krčenja v procesu brizganja lokalni povprečni tlak strjevanja oz. povprečje tlakov pravokotno na steno izdelka, pri katerih se strjujejo posamezne plasti polimerov. Poznavanje lokalnega tlaka strjevanja zahteva določitev lokalne zgodovine tlaka in lokalne zgodovine strjevanja. Slednja se običajno pridobi s simulacijo celotnega postopka brizganja, saj trenutno ni pretvornikov, ki bi lahko merili porazdelitev temperature v polimernem delu med brizganjem. V tem delu je bil uporabljen postopek, ki omogoča ocenitev profila strjevanja in s tem povprečnega tlaka strjevanja z analizo poteka tlaka v eksperimentalnem gnezdu. Postopek zahteva merjenje poteka tlaka v gnezdu s tradicionalnim tlačnim pretvornikom. Tlačni profil se nato uporabi za analizo z enačbami prenosa toplote, rezultat pa je več relevantnih parametrov (lokalni čas strjevanja in karakteristični čas hlajenja). Ti parametri so nato uporabljeni za popis zgodovine strjevanja, ki v povezavi s tlačnim profilom daje lokalni tlak strjevanja. Postopek je bil preverjen v vrsti preizkusov s polistirenom za splošne namene pri različnih procesnih parametrih: tlak zadrževanja, temperatura brizganja in mere gnezda. Vzorci so bili po brizganju premerjeni in izračunani so bili skrčki. Postopek je bil nato uporabljen pri vseh preskusih brizganja in določena je bila glavna krivulja krčenja v odvisnosti od povprečnega tlaka strjevanja. S tem je bila potrjena veljavnost metode, vsaj za amorfne polimere. Opisani postopek je še posebej primeren za medprocesni nadzor izbranega parametra kakovosti ter ne zahteva niti poznavanja lokalne debeline niti pogojev brizganja. Uporaben je tudi brez karakterizacije fizikalnih parametrov materiala. Ključne besede: predelava polimerov, brizganje, kontrola kakovosti, krčenje, strjevanje, povprečni tlak strjevanja, tlak v gnezdu, čas strjevanja

*Naslov avtorja za dopisovanje: Oddelek za industrijski inženiring, Univerza v Salernu, via Giovanni Paolo II, I-84084 Fisciano (SA), Italija, rpantani@unisa.it

SI 137


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 138 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-08-21 Odobreno za objavo: 2013-09-27

Vpliv hitrih sprememb temperature v orodju na videz brizganih izdelkov Giovanni Lucchetta* ‒ Marco Fiorotto Univerza v Padovi, Oddelek za industrijski inženiring, Italija

V članku je predstavljena inovativna tehnologija za ogrevanje in hlajenje orodij za brizganje plastike. Tehnologija je bila uporabljena za analizo vpliva hitrih sprememb temperature v orodju na izboljšanje sijaja brizgancev. Površinski sijaj termoplastičnih brizgancev se povečuje z naraščanjem temperature v orodju. Takšna rešitev pa pomeni tudi večje proizvodne stroške zaradi bistveno daljšega časa hlajenja. V zadnjem desetletju so zato razvijali tehnologije za hitro spreminjanje temperature v orodju oz. brizganje s hitrim ciklom segrevanja in ohlajanja (RHCM). V predstavljenem delu je bil eksperimentalno preizkušen vpliv hitrih sprememb temperature v orodju na videz brizganih izdelkov. Zasnovano in izdelano je bilo prototipno orodje. Dva vložka iz odprtocelične aluminijeve pene sta bila postavljena tik pod površino gnezda in integrirana v sistem za uravnavanje temperature orodja. Kovinska pena omogoča pretok tekočine pri nadzorovani temperaturi in ustvari strukturo gnezda, ki omogoča enakomerno razporejen sistem za temperiranje tik pod površino orodja. Temperaturni profil blizu površine gnezda je bil izmerjen s pomočjo termopara in primerjan z eksperimentalnimi vrednostmi, izmerjenimi v isti referenčni točki s tanjšimi vložki iz kovinske pene in kroglično polnitvijo. Vpliv hitrih sprememb temperature orodja na sijaj površine je bil analiziran s spektrofotometrom UV/VIS, ki je deloval v načinu odboja. Primerjava med tehnologijo kroglične polnitve in sistemi RHCM s kovinsko peno je pokazala, da predlagana rešitev omogoča skrajšanje cikla za približno 16 sekund. Kombinacija ogrevanja orodnega gnezda s hitrim hlajenjem brizganca je prinesla odpravo vidnih zvarnih linij in dele visokega sijaja. Hitro segrevanje in ohlajevanje površine orodja omogoča bistveno izboljšanje sijaja površine, ko se orodje segreje na temperaturo nad Tg. Visoka temperatura površine gnezda namreč preprečuje prezgodnjo strditev taline v fazi polnjenja in zgoščevanja ter izboljša preslikavanje zrcalno gladke površine gnezda za odličen videz (brez vidnih zvarnih linij) in večji sijaj. Prihodnje raziskave bodo usmerjene v izdelavo orodij za dele zahtevnejših oblik in v uporabo izolacijskih sistemov za dodatno izboljšanje učinkovitosti ogrevanja/hlajenja. Preučiti je treba tudi termomehanske utrujenostne lastnosti kovinske pene in s tem primernost tehnologije za masovno proizvodnjo. Gre za prvo poročilo o vplivu inovativne tehnologije RHCM na videz brizgancev (sijaj in zvarne linije), pri čemer so bili uporabljeni orodni vložki z odprtocelično aluminijevo peno. Članek podaja tudi nove informacije o učinkovitosti metod RHCM s primerjavo predlagane tehnologije in konvencionalne rešitve s kroglično polnitvijo. Ključne besede: injekcijsko brizganje, brizganje s hitrim ciklom ogrevanja in ohlajevanja, sijaj, zvarne linije, odprtocelična kovinska pena

SI 138

*Naslov avtorja za dopisovanje: Univerza v Padovi, Oddelek za industrijski inženiring, via Venezia 1, Padua, Italija, giovanni.lucchetta@unipd.it


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 139 © 2013 Strojniški vestnik. Vse pravice pridržane. Tematska številka

Prejeto v recenzijo: 2013-01-22 Prejeto popravljeno: 2013-03-07 Odobreno za objavo: 2013-03-13

Prva študija izdelave povezav po postopku AerosolJet® na substratih, izdelanih z ekstruzijskim 3D-nalaganjem Vogeler, F. – Verheecke, W. – Voet, A. – Valkenaers, H. Frederik Vogeler1,2,* – Wesley Verheecke1 – André Voet1,2 – Hans Valkenaers1,2 2Katoliška

1Thomas More Mechelen – Campus De Nayer, Belgija univerza v Leuvnu, Tehniška fakulteta, Oddelek za strojništvo, Belgija

Eden od novejših izzivov pri uporabi dodajalnih izdelovalnih tehnologij je proizvodnja večfunkcijskih izdelkov z električnimi in termičnimi funkcijami po hibridnih postopkih. Možna aplikacija hibridnih dodajalnih izdelovalnih tehnologij je tudi vdelovanje senzorjev, anten in značk RFID v izdelke prostih oblik. V članku je predstavljen predlog kombinacije ekstruzijskega 3D-nalaganja in tiskanja po postopku Aerosol Jet® (AJP). Preizkusi se trenutno izvajajo na ločenih strojih, v praksi pa bi bilo mogoče obe tehniki integrirati tudi v isti stroj. Postopki ekstruzijskega 3D-nalaganja se trenutno uporabljajo za izdelavo funkcijskih termoplastičnih komponent z ločljivostjo približno 100 mm. Postopek AJP se je uveljavil pri izdelavi elektronskih komponent: povezav, uporov, tuljav, kondenzatorjev itd, uporablja pa črnilo ter je podoben postopku brizgalnega tiskanja. Omogoča tiskanje mikrosledi širine od 10 mm pa do več milimetrov, ter višine od 0,025 do 10 mm; odvisno od strojne opreme, delovnih parametrov ter interakcije med črnilom in substratom. Večina dosedanjih preizkusov postopka AJP je bilo opravljenih na steklenem ali poliamidnem substratu, zelo malo pa je znanega o izdelavi električno prevodnih vodov na komponentah, izdelanih s postopki 3D-tiskanja. Raziskava se je začela s preučevanjem vpliva parametrov AJP na kakovost potiska substratov, izdelanih z ekstruzijskim 3D-nalaganjem. Eksperiment je bil opravljen s tiskanjem vezja za štiritočkovne meritve upora na vzorcih ABS (akrilnitrilbutadien-stiren), izdelanih po postopku ekstruzijskega 3D-nalaganja. Za tiskanje vezja je bilo uporabljeno običajno prevodno srebrno črnilo. 3D-vzorci so bili pred tiskanjem pobrušeni in očiščeni z izopropanolom, s čimer so bili zagotovljeni ponovljivi pogoji površine. Prihodnje raziskave bodo osredotočene na potisk neobdelanih vzorcev. Dimenzijske lastnosti potiskanih vzorcev so bile preverjene z digitalnim mikroskopom. Izmerjena je bila širina vodov in kakovost robov, z napravo za 2D-otipavanje profila pa je bil ugotovljen tudi prerez voda. Prevodne lastnosti potiskanih vzorcev so bile preverjene s štiritočkovnim mikroohmmetrom. Najprej je bil preučen vpliv parametrov postopka AJP na dimenzije posameznih natisnjenih vodov. Med preizkušanimi parametri AJP so bili pretok aerosola skozi šobe, pretok nosilnega plina, temperaturne nastavitve ter oddaljenost šobe od substrata. Izdelani in pregledani so bili tudi vzorci z več vzporednimi vodi in večslojnimi natisnjenimi vodi. Korelacijske analize kažejo, da je pretok aerosola na šobi verjetno najboljša izbira za nadzor širine voda pri tiskanju posameznih vodov. Kakovost robov se najboljše nadzoruje z nastavitvami nosilnega plina. Zgolj enosledni nanos voda nima električno prevodnih lastnosti, zato je za zahtevane električne lastnosti nujen večslojni nanos materiala. Neugodne prevodne lastnosti je mogoče pojasniti tudi z nizko temperaturo strjevanja vzorcev. Substrat je izdelan iz ABS-a, zato mora biti temperatura strjevanja oz. sintranja črnila nižja od 100 °C. Raziskava je pokazala, da je tiskanje vzporednih sledi z ozirom na čas obdelave primernejše od večslojnega tiskanja. V praksi bo izbira števila vzporednih sledi oz. slojev odvisna tudi od zahtevane ločljivosti povezav. Nabor opravljenih eksperimentov je trenutno še majhen in preizkuse bo zato treba ponavljati tudi v prihodnjih raziskavah. Predstavljeno delo pa je vseeno dobra osnova za prihodnje raziskave na tem področju. Ključne besede: tiskanje Aerosol Jet®, ekstruzijsko 3D-nalaganje, srebrno črnilo, povezave, tiskani vodi, električno prevodne sledi, hibridna izdelava

*Naslov avtorja za dopisovanje: Thomas More Mechelen – Campus De Nayer, J. De Nayerlaan 5, 2860 Sint-Katelijne-Waver, Belgija, frederik.vogeler@lessius.kuleuven.be

SI 139


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 140 Prejeto v recenzijo: 2013-01-22 © 2013 Strojniški vestnik. Vse pravice pridržane. Prejeto popravljeno: 2013-06-06 Tematska številka Odobreno za objavo: 2013-06-12

Nanokompoziti polipropilena/gline, izdelani po postopku brizganja s strižno nadzorovanim usmerjanjem: deformacijske in lomne lastnosti

Costantino, A. – Pettarin, V. – Viana, J. – Pontes, A. –Pouzada, A. – Frontini, P. Alejandra Costantino1 – Valeria Pettarin1,* – Julio Viana2 – Antonio Pontes2 – Antonio Pouzada2 – Patricia Frontini1 1 Nacionalna

univerza v Mar del Plati, Institut za znanost in tehnologijo materialov, Argentina v Minhu, Institut za polimere in kompozite, Portugalska

2 Univerza

Nanokompoziti gline in polimerov so alternativa za konvencionalne mikrokompozite, saj omogočajo izboljšane mehanske, toplotne in obdelovalne lastnosti, zmanjšujejo gorljivost ter izboljšujejo zaporne lastnosti že pri zelo majhnih vsebnostih polnila (<5 utež. %). Nanokompoziti pa morajo imeti za izpolnjevanje konkretnih konstrukcijskih zahtev tudi ustrezno togost, trdnost in žilavost. Priprava nanokompozitov s termoplastičnimi polimeri s pomočjo konvencionalne opreme za kompaundiranje ostaja prednostna rešitev za industrijo, zlasti pri uporabi osnovnih polimerov kot je polipropilen (PP), ki se intenzivno uporablja v industriji embalaže ter v avtomobilski industriji. Izdelava kompozitov PP/glina po konvencionalnih postopkih brizganja (CIM) ni enostavna naloga. Nanoglina je običajno slabo dispergirana po brizgancu, pri debelih izdelkih pa je pogosta nepopolna izpolnitev gnezda in druge napake. Postopek SCORIM (strižno nadzorovano usmerjanje pri brizganju) je alternativa za CIM, ki v proces brizganja prinaša določene prednosti z nadzorom mikrostrukture brizganih materialov. Brizganju polimernih nanokompozitov po postopku SCORIM pa je za zdaj posvečenih le malo raziskav. V tem delu so raziskani vplivi različnih morfologij, ki nastanejo ob strižno nadzorovanem usmerjanju pri brizganju (SCORIM), na mehanske in lomne lastnosti polipropilena (PP) ter brizgancev iz PP/nanogline. Za mehansko karakterizacijo so bili opravljeni enoosni tlačni in upogibni preizkusi pri sobni temperaturi in v kvazistatičnih pogojih. Za karakterizacijo lomnih lastnosti je bil opravljen tritočkovni upogibni test SENB, podatki pa so bili analizirani po pristopih lomne mehanike. Iniciacija loma je bila vrednotena po faktorju intenzivnosti kritične napetosti v načinu I, KIC, vrednost integrala J pa je bila izračunana za točko nestabilnega loma oz. izgubo integritete konstrukcije pri vzorcih s kvazistabilnim napredovanjem razpoke. Za določanje srednjih mehanskih in lomnih lastnosti je bila opravljena tudi popolna karakterizacija mikrostrukture brizgancev z optično polarizacijsko mikroskopijo (POM), rentgensko difrakcijo (XRD) in diferencialno skenirno kalorimetrijo (DSC), o čemer so avtorji že poročali. Pri vzorcih je bil ugotovljen cel razpon lomne stabilnosti, od zmerne nelinearnosti do kvazistabilnega režima, odvisno od vrste materiala in pogojev brizganja. Pri čistem PP je bilo ugotovljeno nelinearno krhko vedenje, medtem ko so imeli nanokompoziti kvazistabilno vedenje po zaslugi velike deformacijske zmogljivosti vrhnjega sloja. Kljub temu, da je do iniciacije razpoke prišlo praktično pri enakih obremenitvah, pa je bila energija napredovanja odvisna od pogojev obdelave in od vsebnosti nanogline. Statistična analiza nakazuje, da je za izboljšanje žilavosti debelih vzorcev SCORIM kompozita PP/nanoglina skupaj z mikrostrukturnimi razlikami med vrhnjo plastjo (ali strižno cono) in jedrom odgovorno zmanjšanje velikosti jedra pri obdelavi SCORIM, na katerega ugodno vpliva prisotnost nanogline. Ključne besede: nanokompoziti, glina, polipropilen, SCORIM, lomna mehanika, deformacije

SI 140

*Naslov avtorja za dopisovanje: Nacionalna univerza v Mar del Plati, Institut za znanost in tehnologijo materialov, Juan B. Justo 4302 – B7608FDQ Mar del Plata, Argentina, pettarin@fi.mdp.edu.ar


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 141-145 Osebne objave

Doktorske disertacije, diplomske naloge

DOKTORSKE DISERTACIJE Na Fakulteti za strojništvo Univerze v Ljubljani so obranili svojo doktorsko disertacijo: ●    dne 1. oktobra 2013 Janez KOGOVŠEK z naslovom: »Tribološki učinki nanodelcev v mazivih« (mentor: prof. dr. Mitjan Kalin, somentorica: prof.dr. Maja Remškar (IJS)); V doktorski nalogi smo obravnavali tribološke učinke in mehanizme mazanja MoS2 nanocevk, dodanih olju. Učinke teh nanodelcev smo primerjali z učinki različnih delcev trdnih maziv. Ovrednotili smo vpliv lastnosti delcev na znižanje trenja in obrabe v režimu mejnega mazanja. Preučevali smo mehanizme mazanja z MoS2 nanocevkami. Na podlagi obnašanja MoS2 nanocevk dodanih olju pri mazanju jeklenih in DLC-prevlečenih kontaktnih površin v vseh režimih mazanja smo določili vplive hrapavosti površin, utekanja in kontaktnih materialov na takšno mazanje. Raziskali in primerjali smo oblikovanje tribofilma iz MoS2 nanocevk na jeklenih in DLC-prevlečenih površinah; ●    dne 10. oktobra 2013 Edvard HÖFLER z naslovom: »Računsko hidrodinamično oblikovanje nove diagonalne turbine« (mentor: izr. prof. dr. Anton Bergant, somentor: prof. dr. Branko Širok); Predstavljen je hidrodinamični razvoj turbine, ki je hibrid med Francisovo turbino visoke specifične hitrosti in vertikalno cevno turbino. V tezi je razvit izviren teoretični model določitve ožjega optimalnega področja delovanja nove turbine, ki temelji na integralnem koeficientu tangencialnega vzgona celotnega rotorskega lopatja. Računsko hidrodinamično oblikovanje lopatja gonilnika je izvedeno z avtorjevim programskim orodjem, katerega osnova je v tem delu nadgrajena metoda ukrivljenih tokovnic (SCM) z upoštevanjem sil lopate ter vključenimi dodatki inverznega oblikovanja. Z orodji računalniške dinamike tekočin so raziskane značilke celotnega stroja in overjena energetska učinkovitost lopatja turbine. Predlagana diagonalna turbina je novost v segmentu specifično hitrih turbin z enojno regulacijo in je učinkovita alternativa konvencionalnim Francisovim turbinam; ●    dne 16. oktobra 2013 Seid ŽAPČEVIĆ z naslovom: »Model samoučečega proizvodnega delovnega sistema« (mentor: prof. dr. Peter Butala); Za preživetje v konkurenčni globalni ekonomiji morajo biti proizvodni sistemi sposobni prilagajanja novim okoliščinam. Pomemben pogoj za prilagajanje proizvodnih sistemov na spremembe je njihova

sposobnost učenja in avtonomnost v širšem smislu. Glavno vprašanje je, kako avtonomne proizvodne sisteme usposobiti za učenje na podlagi lastnih izkušenj in za uporabo naučenega znanja za prilagajanje na visoke zahteve, ki se postavljajo pred adaptivne distribuirane proizvodne sisteme. Disertacija temelji na hipotezi, da se lahko razvoj sposobnih samoučečih adaptivnih distribuiranih proizvodnih sistemov doseže s sodobnimi metodami sistemskega in sistematičnega zbiranja podatkov, z inteligentno analizo zbranih podatkov in z metodami odkrivanja znanja ter z uporabo najnovejših sistemov upravljanja znanja s ciljem integracije naučenega znanja za podporo vodstvu pri sprejemanju odločitev v okviru funkcij vodenja, načrtovanja in razporejanja v krmiljenju procesov. Za rešitev hipoteze je bil razvit koncept samoučečega avtonomnega delovnega sistema, v katerem so predstavljeni: struktura sistema, zanki za zbiranje podatkov in učenje ter struktura modelov znanja. Razvit je kibernetski model ter podana definicija samoučečega proizvodnega sistema. Razviti so modeli znanja na treh ravneh: 1) generični referenčni model znanja (meta-meta-model), 2) procesno specifični referenčni model znanja (metamodel) in 3) model znanja, ki bazira na realnih podatkih in določenem algoritmu rudarjenja podatkov. Za uporabo znanja je razvit kibernetski model uporabe znanja na podlagi matematičnih modelov znanja, pridobljenih v zanki za samoučenje, obenem pa je bil razvit tudi inovativen model adaptivnega krmiljenja procesa. Podrobno proučevanje primera na področju odkrivanja in uporabe znanja za podporo odločanju pri vodenju in v adaptivnem krmiljenju procesa je bilo izvedeno na realnih industrijskih podatkih s področja visokotlačnega litja v kalupe; ●    dne 18. oktobra 2013 Vid NOVAK z naslovom: »Bliskovni vlakenski laserji za mikroobdelave v elektroniki« (mentor: prof. dr. Janez Možina, somentor: doc. dr. Rok Petkovšek); Uvedba fleksibilnih laserskih mikroobdelovalnih postopkov v maloserijsko proizvodnjo in na področje hitrega razvoja prototipov zahteva uporabo laserskih virov s širokim parametrskim prostorom ter visoko svetilnostjo. V diodno vzbujanem vlakenskem laserju, tipa oscilator-ojačevalnik, sta oscilator in ojačevalnik laserja strukturno in funkcionalno ločena. Izhod direktno modulirane polprevodniške laserske diode je ojačen v enorodovnem vlakenskem ojačevalniku. Laserski vir te vrste omogoča medsebojno neodvisno izbiro vrednosti komponent parametrskega prostora - časa bliskov, njihove frekvence repeticije in SI 141


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amplitude. Vrednosti komponent so elektronsko nastavljive v realnem času. Energijski izkoristek, stabilnost izhoda in kvaliteta žarka so visoki v celotnem razponu parametrskega prostora. V okviru tega dela sta predstavljeni eksperimentalni postavitvi bliskovnega in kontinuirnega laserja. Parametrski prostor je bil razširjen v subnanosekundno področje. V območju repeticijskih frekvenc od 10 kHz do 2 MHz so bile dosežene dolžine bliskov med 600 ps in 100 ns. Energija laserskega bliska znaša do 680 µJ, vršna moč pa do 54 kW. Izhod kontinuirne izvedbe laserja je primeren za frekvenčno transformacijo iz bližnjega infrardečega spektra v vidni oz. ultravijolični del; ●    dne 21. oktobra 2013 Aljaž POGAČNIK z naslovom: »Vplivi fizikalnih parametrov na tribološke lastnosti polimerov za zobnike« (mentor: prof. dr. Mitjan Kalin); Doktorska naloga obravnava vpliv fizikalnih parametrov (obremenitev, drsna hitrost, temperatura) na trenje, obrabo in obrabne mehanizme polimernih materialov, ki se pogosto uporabljajo za zobnike. S pomočjo obsežnih triboloških analiz osmih različnih kontaktnih kombinacij treh materialov (PA, POM in jeklo) so bile narejene obrabne mape in določeni kritični kontaktni pogoji. Na osnovi analiz teoretičnih kontaktnih temperatur, obrabnih površin in sprememb triboloških mehanizmov so v doktorski nalogi definirani in predlagani kriteriji mehanskih in toplotnih obremenitev, s katerimi lahko vrednotimo vpliv fizikalnih parametrov na kritične spremembe triboloških mehanizmov polimernih materialov. Na osnovi analize mikro-deformacijske krivulje vršičkov in plastične deformacije površin je predstavljena tudi pričakovana realna kontaktna površina v triboloških kontaktih polimernih materialov; ●    dne 22. oktobra 2013 Marina Borisovna GERGESOVA z naslovom: »Karakterizacija časovno odvisnih lastnosti polimerov preko reševanja inverznih problemov« (mentor: prof. dr. Igor Emri); Lastnosti polimerov se lahko s časom drastično spremenijo, kar vpliva na njihovo dolgoročno vedenje in trajnost. V praksi to pomeni, da se funkcionalnost in posledično uporabnost polimernih produktov po določenem času zelo spremenita, pojavijo se poškodbe in izdelek postane trajno neuporaben. Novi standardi, ki so trenutno v pripravi, bodo za napovedovanje trajnosti polimernih konstrukcij zahtevali karakterizacijo časovno odvisnih lastnosti polimerov v daljšem časovnem obdobju preko meritev lezenja in relaksacije. Standardizirana oprema, ki je trenutno na voljo za karakterizacijo, ne omogoča direktnih meritev želenih časovno odvisnih materialnih lastnosti. Informacijo o tem je mogoče pridobiti iz eksperimentalnih podatkov le z uporabo matematičnih sredstev. SI 142

Doktorsko delo predstavlja vzpostavitev numerične metodologije, ki je sestavljena iz treh ključnih korakov, za napoved časovno odvisnih lastnosti polimernih materialov na osnovi eksperimentalnih podatkov, pridobljenih s standardiziranimi eksperimenti. ●    dne 22. oktobra 2013 Ivan Vladimirovich SAPRUNOV z naslovom: »Balistične lastnosti termoplastičnih materialov, izpostavljenih udarni obremenitvi v širšem temperaturnem območju« (mentor: prof. dr. Igor Emri); V tem delu smo raziskovali, kako spreminjanje različnih parametrov vpliva na udarno žilavost polimerov. V sklopu dela je bila analizirana balistična zmogljivost polimernih diskov premera Ø90 mm, izdelanih iz polietilena nizke gostote (LDPE), poliamida-6 (PA6) in termoplastičnega poliuretana (TPU), ki smo jih izpostavili udarnim obremenitvam jeklenih izstrelkov v hitrostnem razponu 10 do 136 m/s. Analizirana je bila balistična meja in udarna žilavost vzorcev treh debelin (2, 4 in 8 mm), za dve različni masi izstrelka (17 in 45 g), dveh različnih oblik konice izstrelka (oster in polkrožen) in pet poljubno izbranih temperatur (-37, 0, 30, 60 in 90 °C). Potrjeno je bilo, da hitro povečanje odpornosti proti udarnim obremenitvam ustreza prehodu od krhke k duktilne vrsti odpovedi materiala. Poleg vpliva temperature materiala je prehod iz krhkega k duktilnemu vedenju lahko povzročen s spreminjanjem debeline vzorca in celo mase izstrelka. Ugotovljeno je bilo, da povečanje mase izstrelka povzroči prehod od krhkega k duktilnemu vedenju, kar je nepričakovan in pomemben rezultat. Narejena je bila napoved temperaturne odvisnosti udarne žilavosti, ki temelji na poznavanju (i) temeljnih viskoelastičnih lastnosti in uporabe (ii) simulacije udarnega preizkusa z uporabo metode končnih elementov. Obe metodi uporabljata kot primarne vhodne podatke osnovne viskoelastične materialne lastnosti v smislu bodisi časovno- ali frekvenčnoodvisnega strižnega modula in voljnosti. Oba pristopa pa ne omogočata napovedovanja temperaturno odvisnega trenda udarne žilavosti. * Na Fakulteti za strojništvo Univerze v Mariboru je obranil svojo doktorsko disertacijo: ●    dne 8. oktobra 2013 Marko KLANČIŠAR z naslovom: »Eksperimentalno numerična analiza tokovno reaktivnih veličin večplamenskega gorilnika« (mentor: prof. dr. Niko Samec); Optimizacija zgorevalnih naprav v smislu povečanja učinkovitosti in zmanjšanja obremenitve okolja s polutanti, je danes glavno vodilo pri


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raziskavi ter snovanju novih projektov povezanih z zgorevanjem. To je mogoče doseči s pomočjo točnih zasnov pomembnih komponent gorilnika ter kurišča, kar dodatno zajema pojave prenosa mase in toplote. V zadnjih nekaj letih je računalniška dinamika tekočin (CFD) postala popularna metoda za uporaben pristop k pridobivanju preliminarnih informacij in napovedovanju obnašanja komercialnih zgorevalnih zasnov. Vendar so pojavi, ki potekajo v gorilnikih zelo kompleksni; recirkulacija dimnih plinov, prenos energije, turbulentna kemijska kinetika in razmerje toka-kemijskih reakcij so v primeru vrtinčnih zgorevalnih naprav še ojačeni. Z omenjenega vidika je CFD analiza vrtinčnih ne-predmešanih reaktivnih tokov ena najpomembnejših in zahtevnih področij v moderni CFD. Za primerjavo so dokumentirani eksperimentalni podatki, ki predstavljajo uporabno informacijo za kontrolo in testiranje CFD. Za popis dinamike tekočin skupaj s prenosom toplote, je v literature mogoče najti različne matematične pristope. Glavni pristop uporablja turbulentne modele zgorevanja kar omogoča uporabo statističnih lastnosti skalarnega polja. Ti pristopi omogočajo v večini primerov podrobne informacije o tokovnem polju ter temperaturi vendar so pri napovedovanju koncentracij nižjih vrst manj natančni. Za kotlovske naprave je bila že predlagana nova metoda (Faravelli et al., 2001). Ta pristop (SFIRN) temelji na originalni zasnovi hibridne metode. Tokovna ter temperaturna polja niso pod vplivom nastanka NOx. Za opis nastanka toplote v CFD modeliranju je dovolj manjše število hitrih kemijskih reakcij. Spisek oziroma skupina idealnih reaktantov je definirana na osnovi rezultatov pridobljenih z numerično simulacijo. Ti reaktanti skupaj predstavljajo število enakih celic in se rešujejo s pomočjo zelo podrobne kemijske kinetike. Podrobne eksperimentalne meritve 1SF gorilnika predstavljajo pomemben testni primer z namenom potrditve numeričnega postopka, kar je mogoče kasneje razširiti na kompleksnejše primere. Realno kurilno napravo smo preučevali z uporabo računalniške dinamike tekočin; podrobneje s programskim paketom ANSYS CFX. V nalogi smo uporabili eksperimentalno dognane robne pogoje vhodnih parametrov ter nekaj parametrov drugih avtorjev. Za numerične simulacije so bili uporabljeni turbulentni modeli, modeli zgorevanja ter sevanja. Nadalje smo primerjali rezultate različnih izbranih modelov z eksperimentalnimi ter ugotavljali primernost le teh. Cilj modeliranja je bila izbira primernih modelov za uporabo numeričnega preizkušanja novih industrijskih gorilnikov, kar smo v primerjavi z eksperimentom okarakterizirali s tokovnim in temperaturnim poljem, hkrati pa z lokalno stopnjo izgorelosti (CO). Rezultati

numeričnega modeliranja so bili povsod primerjani z eksperimentalnimi meritvami. Prikazan znanstveni pristop omogoča CFD analizo tokovnih lastnosti kot tudi reaktivnega toka že v fazi načrtovanja novih zasnov gorilnikov kar omogoča hitrejši in bolj zanesljiv razvoj novega izdelka. DIPLOMSKE NALOGE Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 23. oktobra 2013: Gorazd KEŠE z naslovom: »Identifikacija dinamskih lastnosti visokonapetosnega lameliranega vodnika« (mentor: prof. dr. Miha Boltežar); Tjaša KODRIČ z naslovom: »Hitra zamenjava orodij« (mentor: prof. dr. Marko Starbek, somentor: izr. prof. dr. Janez Kušar); dne 25. oktobra 2013: Marko BEVC z naslovom: »Ultra tanka toplotna cev« (mentor: prof. dr. Iztok Golobič); Luka CESAR z naslovom: »Eksergijska učinkovitost lamelnega prenosnika toplote« (mentor: prof. dr. Iztok Golobič); Borut GROŠELJ z naslovom: »Vpliv superhidrofobnega površinskega nanosa na proces sreženja« (mentor: prof. dr. Iztok Golobič); Luka JANEŽ z naslovom: »Eksperimentalna raziskava učinkovitosti mešala z asimetrično zapognjenimi lopaticami« (mentor: doc. dr. Andrej Bombač); Marko KOLENC z naslovom: »Fasadni fotonapetostni toplotni sprejemnik sončne energije« (mentor: izr. prof. dr. Andrej Kitanovski, somentor: prof. dr. Alojz Poredoš); Anja RADULOVIĆ z naslovom: »Izraba odpadne toplote izpušnih plinov s pomočjo termoelektričnega generatorja« (mentor: prof. dr. Iztok Golobič). * Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv univerzitetni diplomirani inženir strojništva: dne 3. oktobra 2013: Janže BEBAR z naslovom: »Retro oblikovna zasnova mestnega kolesa« (mentor: izr. prof. Vojmir Pogačar); dne 30. oktobra 2013: Teo MILOŠIČ z naslovom: »Primerjava fizikalnih modelov Adblue tekočine pri SCR sistemih SI 143


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ter vpliv mrež različnih gostot« (mentor: prof. dr. Leopold Škerget, somentor: izr. prof. dr. Jure Ravnik).

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Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv magister inženir strojništva: dne 25. oktobra 2013: Jure BROVČ z naslovom: »Analiza vpliva parametrov umetnega staranja duroplasta BMC na dimenzijske lastnosti izdelka« (mentor: izr. prof. dr. Tomaž Pepelnjak); Andrej JANEŽIČ z naslovom: »Analiza varivosti nerjavnih dupleks jekel po postopku TIG« (mentor: prof. dr. Janez Tušek); Toni KAMBIČ z naslovom: »Analiza vpliva geometrijskih in materialnih parametrov na postopek krivljenja žice« (mentor: izr. prof. dr. Tomaž Pepelnjak).

Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva: dne 24. oktobra 2013: Lovro ZAJEC z naslovom: »Rekonstrukcija arheoloških najdb s pomočjo 3D tehnologij« (mentor: izr. prof. dr. Bojan Dolšak, somentor: izr. prof. Vojmir Pogačar); dne 29. oktobra 2013: Tomislav OPREŠNIK z naslovom: »Optimiranje rezalnih parametrov optičnega laserja Var-laser HFL015« (mentor: prof. dr. Franci Čuš, somentor: asist. Tomaž Irgolič); dne 30. oktobra 2013: Nejc PERKO z naslovom: »Dimenzioniranje konstrukcije pralnika dimnih plinov Bloka 6 v Termoelektrarni Šoštanj« (mentor: doc. dr. Janez Kramberger, somentor: prof. dr. Srečko Glodež).

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Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva: dne 17. oktobra 2013: Tomaž MARKOVIČ z naslovom: »Jeklena konstrukcija skladišča 36 m x 8,2 m x 5,2 m« (mentor: doc. dr. Boris Jerman); Marin SUBANOVIĆ z naslovom: »Izboljšanje zmogljivosti letalskega Ottovega motorja z uporabo dušikovega oksida« (mentor: izr. prof. dr. Tomaž Katrašnik); dne 18. oktobra 2013: Željko BILIĆ z naslovom: »Projektni informacijski sistem« (mentor: izr. prof. dr. Janez Kušar, somentor: prof. dr. Marko Starbek); Andrej BOŠKOVIĆ z naslovom: »Optimalni razvoj letališča v Portorožu« (mentor: doc. dr. Patrick Vlačič, somentor: izr. prof. dr. Tadej Kosel); Martin LEBAR z naslovom: »Optimalna razmestitev stiskalnic« (mentor: izr. prof. dr. Janez Kušar, prof. dr. Marko Starbek); Urška ŽELEZNIK z naslovom: »Analiza vpeljave Sistema letališkega sodelovalnega odločanja (A-CDM) na letališču Jožeta Pučnika Ljubljana« (mentor: izr. prof. dr. Tadej Kosel, somentor: pred. mag. Andrej Grebenšek); Christian Joël Charles TOSOLINI z naslovom: »Optimizacija proizvodnje valjčkov varnostne ključavnice« (mentor: doc. dr. Davorin Kramar, somentor: prof. dr. Janez Kopač).

Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva (UN): dne 10. oktobra 2013: Andrej SALOBIR z naslovom: »Uporaba alternativnih materialov za prototipna orodja za brizganje plastike« (mentor: izr. prof. dr. Igor Drstvenšek, somentor: asist. mag. Tomaž Brajlih); dne 16. oktobra 2013: Luka PERČIČ z naslovom: »Tehnološke izboljšave pri paletiziranju izdelkov« (mentor: prof. dr. Miran Brezočnik, somentor: asist. dr. Simon Brezovnik ); dne 29. oktobra 2013: Filip MLINŠEK z naslovom: »Tehnična zasnova in razvoj namenskih brusov za ostrenje robnikov smuči« (mentor: doc. dr. Marjan Leber); Samo POSTRUŽNIK z naslovom: »Metode izračunov rezalnih pogojev za postopke odrezavanja« (mentor: prof. dr. Franci Čuš, somentor: asist. Tomaž Irgolič); Denis POTOČNIK z naslovom: »Ergonomsko oblikovanje delovnega mesta v podjetju Henkel d.o.o.« (mentor: doc. dr. Nataša Vujica Herzog, somentor: Igor Čebašek, univ. dipl. inž. str.); Gregor SLEMNIK z naslovom: »Konstruiranje sinhrona za pogonsko prikolico« (mentor: prof. dr. Srečko Glodež, somentor: prof. dr. Miran Brezočnik); dne 30. oktobra 2013: Žan PODBREGAR z naslovom: »Konstrukcija naprave za trajnostni preizkus stikal mikrovalovne pečice« (mentor: doc. dr. Aleš Belšak, somentor: izr. prof. dr. Miran Ulbin);

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SI 144


Strojniški vestnik - Journal of Mechanical Engineering 59(2013)11, SI 141-145

Klemen ŽUŽEK z naslovom: »Rekonstrukcija preoblikovalnega orodja za izdelavo hladilnozamrzovalne naprave« (mentor: izr. prof. dr. Ivan Pahole). * Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv diplomirani inženir mehatronike (UN): dne 30. oktobra 2013: Primož BITENC z naslovom: »Razvoj komunikacije med Siemens simatics S120 in mikrokrmilnikom« (mentorja: izr. prof. dr. Karl Gotlih, izr. prof. dr. Aleš Hace, somentor: asist. dr. Simon Brezovnik, ); Tadej MOTALN z naslovom: »Avtonomno raziskovanje mobilnega robota z meta-operacijskim sistemom ROS (Robotski operacijski sistem)« (mentorja: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Suzana Uran). * Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva (VS): dne 17. oktobra 2013: Primož LUKAČ z naslovom: »Zasnova pnevmatičnega preklopnega piezoventila« (mentor: izr. prof. dr. Niko Herakovič); dne 18. oktobra 2013: Matej AHAČIČ z naslovom: »Analiza ekonomske upravičenosti zamenjave vira toplote v sistemu daljinskega ogrevanja« (mentor: prof. dr. Alojz Poredoš); Domen STRENČAN z naslovom: »Statično testiranje letala« (mentor: izr. prof. dr. Tadej Kosel); Klemen SIMERL z naslovom: »Eksperimentalna zasnova hranilnika toplote z uporabo Na(CH3COO)

·3H2O« (mentor: izr. prof. dr. Andrej Kitanovski, somentor: prof. dr. Alojz Poredoš); Miha VELKAVRH z naslovom: »Varjenje z gnetenjem aluminijeve zlitine 6082« (mentor: prof. dr. Janez Tušek, somentor: doc. dr. Damjan Klobčar); * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva (VS): dne 29. oktobra 2013: Zvonko BEDENIK z naslovom: »Analiza uvajanja horizontalnega obdelovalnega centra v proizvodnjo ulitkov« (mentor: prof. dr. Jože Balič); Aleš KUPLJEN z naslovom: »Načrtovanje aktivnosti pri izdelavi ponudbe v podjetju SŽ Oprema Ravne« (mentor: doc. dr. Marjan Leber); Janez REPA z naslovom: »Izboljšava vpenjalne priprave in vzdrževanja na CNC rezkalnem centru« (mentor: doc. dr. Leber Marjan); dne 30. oktobra 2013: Jožef KEKEC z naslovom: »Eksperimentalna analiza lopatične kaskade v zračnem toku« (mentor: prof. dr. Aleš Hribernik, somentor: dr. Gorazd Bombek); Denis MUHIČ z naslovom: »Koncipiranje trikolesa z udobnim nazaj nagnjenim sedežem« (mentor: izr. prof. dr. Stanislav Pehan); Nejc ROŠER z naslovom: »Proizvodnja žičnih transportnih trakov« (mentor: izr. prof. dr. Ivan Pahole, somentor: doc. dr. Mirko Ficko); Simon ROTOVNIK z naslovom: »Eksperimentalna analiza osamljenega krila v zračnem toku« (mentor: prof. dr. Aleš Hribernik, somentor: asist. dr. Gorazd Bombek); Alen URŠNIK z naslovom: »Geometrijska parametrizacija paha stiskalnice« (mentor: izr. prof. dr. Karl Gotlih, somentor: izr. prof. dr. Ivan Pahole).

SI 145


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 440 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 Chamber of Commerce and Industry of Slovenia, Metal Processing Industry Association President of Publishing Council Branko Širok, UL, Faculty of Mech. Engineering, Slovenia Vice-President of Publishing Council Jože Balič, UM, Faculty of Mech. Engineering, Slovenia Cover: Microstructure of injection molded high impact polystyrene structural foams. This is a multilayer material consisting of two solid skins and a foamed core. The cell distribution at the core is dependent of the type of mold (steel or hybrid) and the processing conditions. The more efficient temperature control in the steel mold leads to better uniformity of the cell size which is much smaller than in hybrid moldings. Image Courtesy: Carla Martins, IPC/I3N – Institute for Polymer and Composites, University of Minho.

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 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.

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http://www.sv-jme.eu

59 (2013) 11

Since 1955

Papers

637

Fantina Rosa Esteves, Tiago Alexandre Carvalho, António Sérgio Pouzada, Carla Isabel Martins: The Influence of Processing on the Aesthetic, Morphological and Mechanical Properties of Structural Foam Mouldings of High-Impact Polystyrene

646

Jan Deckers, Jean-Pierre Kruth, Ludwig Cardon, Khuram Shahzad, Jef Vleugels: Densification and Geometrical Assessments of Alumina Parts Produced Through Indirect Selective Laser Sintering of Alumina-Polystyrene Composite Powder

662

Markus Gottfried Battisti, Walter Friesenbichler: Injection-Moulding Compounding of PP Polymer Nanocomposites

669

Kim Ragaert, Filip De Somer, Stieven Van de Velde, Joris Degrieck, Ludwig Cardon: Methods for Improved Flexural Mechanical Properties of 3D-Plotted PCL-Based Scaffolds for Heart Valve Tissue Engineering

677

Vito Speranza, Umberto Vietri, Roberto Pantani: Monitoring of Injection Moulding of Thermoplastics: Adopting Pressure Transducers to Estimate the Solidification History and the Shrinkage of Moulded Parts

683

Giovanni Lucchetta, Marco Fiorotto: Influence of Rapid Mould Temperature Variation on Appearance of Injection-Moulded Parts

689

Frederik Vogeler, Wesley Verheecke, André Voet, Hans Valkenaers: An Initial Study of Aerosol Jet® Printed Interconnections on Extrusion-Based 3D-Printed Substrates

697

Alejandra Costantino, Valeria Pettarin, Julio Viana, Antonio Pontes, Antonio Pouzada, Patricia Frontini: Polypropylene/Clay Nanocomposites Produced by Shear Controlled Orientation in Injection Moulding: Deformation and Fracture Properties

Journal of Mechanical Engineering - Strojniški vestnik

Contents

11 year 2013 volume 59 no.

Strojniški vestnik Journal of Mechanical Engineering

Journal of Mechanical Engineering 2013 11  

The Strojniški vestnik – Journal of Mechanical Engineering publishes theoretical and practice oriented papaers, dealing with problems of mod...

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