Journal of Mechanical Engineering 2011 12

57 (2011) 1 12

Platnica SV-JME 57(2011)12_04.pdf 1 9.12.2011 12:53:53

Since 1955

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Journal of Mechanical Engineering - Strojniški vestnik

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12 year 2011 volume 57 no.

Platnica SV-JME 57(2011)12_04.pdf 2 9.12.2011 12:53:53

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 Co-Editor Borut Buchmeister University of Maribor Faculty of Mechanical Engineering, Slovenia Technical Editor Pika Škraba University of Ljubljana Faculty of Mechanical Engineering, Slovenia

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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 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 Cover: In September 2011 the Taurus G4 electrically powered aircraft flew at an equivalent fuel efficiency of 403 passenger miles per gallon at a speed of 172.1 km/h, winning the Green Flight Challenge 2011. This is both twice as efficient and twice as fast as a Toyota Prius, and is almost six time more efficient than a typical 4-seat general aviation aircraft! The aircraft was designed and built by Pipistrel, a Slovenian manufacturer of self- launching sailplanes and light sport aircraft.

Image courtesy: Pipistrel d.o.o., Slovenia

ISSN 0039-2480 © 2011 Strojniški vestnik - Journal of Mechanical Engineering. All rights reserved. SV-JME is indexed / abstracted in: SCI-Expanded, Compendex, Inspec, ProQuest-CSA, SCOPUS, TEMA. The list of the remaining bases, in which SV-JME is indexed, is available on the website. The journal is subsidized by Slovenian Book Agency.

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, Bay Zoltan Inst. for Mater. Sci. and Techn., Hungary Jože Flašker, UM, Faculty of Mech. Engineering, Slovenia Bernard Franković, Faculty of Engineering Rijeka, Croatia Janez Grum, UL, Faculty of Mech. Engineering, Slovenia Imre Horvath, Delft University of Technology, Netherlands Julius Kaplunov, Brunel University, West London, UK Milan Kljajin, J.J. Strossmayer University of Osijek, Croatia Janez Kopač, UL, Faculty of Mech. Engineering, Slovenia Franc Kosel, UL, Faculty of Mech. Engineering, Slovenia Thomas Lübben, University of Bremen, Germany Janez Možina, UL, Faculty of Mech. Engineering, Slovenia Miroslav Plančak, University of Novi Sad, Serbia Brian Prasad, California Institute of Technology, Pasadena, USA Bernd Sauer, University of Kaiserlautern, Germany Brane Širok, UL, Faculty of Mech. Engineering, Slovenia Leopold Škerget, UM, Faculty of Mech. Engineering, Slovenia George E. Totten, Portland State University, USA Nikos C. Tsourveloudis, Technical University of Crete, Greece Toma Udiljak, University of Zagreb, Croatia Arkady Voloshin, Lehigh University, Bethlehem, USA President of Publishing Council Jože Duhovnik UL, Faculty of Mechanical Engineering, Slovenia Print Tiskarna Present d.o.o., Ljubljana, Slovenia, printed in 480 copies 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 peer-review process.

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.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12 Contents

Contents Strojniški vestnik - Journal of Mechanical Engineering volume 57, (2011), number 12 Ljubljana, December 2011 ISSN 0039-2480 Published monthly Editorial 867 Papers Tine Tomažič, Vid Plevnik, Gregor Veble, Jure Tomažič, Franc Popit, Sašo Kolar, Radivoj Kikelj, Jacob W. Langelaan, Kirk Miles: Pipistrel Taurus G4: on Creation and Evolution of the Winning Aeroplane of NASA Green Flight Challenge 2011 Jesús Meneses, Cristina Castejón, Eduardo Corral, Higinio Rubio, Juan Carlos García-Prada: Kinematics and Dynamics of the Quasi-Passive Biped “PASIBOT” Nikola Korunović, Miroslav Trajanović, Miloš Stojković, Dragan Mišić, Jelena Milovanović: Finite Element Analysis of a Tire Steady Rolling on the Drum and Comparison with Experiment Ivo Pahole, Dejan Studenčnik, Karl Gotlih, Mirko Ficko, Jože Balič: Influence of the Milling Strategy on the Durability of Forging Tools Ljubodrag Tanovic, Pavao Bojanic, Radovan Puzovic, Sergey Klimenko: Polycrystalline Cubic Boron Nitride (PCBN) Tool Life and Wear in Turning of Amorphous-Crystalline IronBased Coatings Andrej Lotrič, Mihael Sekavčnik, Christian Kunze, Hartmut Spliethoff: Simulation of WaterGas Shift Membrane Reactor for IGCC Plant with CO2 Capture Ayyannan Devaraju, Ayyasamy Elayaperumal, Srinivasan Venugopal, Satish V. Kailas, Joseph Alphonsa: Characteristics of Surface Treated Nuclear Grade Stainless Steel Type AISI 316 LN at 25 to 500 °C Aleksandar Sicovic, Momčilo Milinovic, Olivera Jeremic: Experimental Equipment Research for Cryogenic Joule-Thompson Cryocoolers Comparison in IR Technology Sensors Instructions for Authors

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Editorial Let’s Observe Our Achievements and Harness the Power Of Our Diversity “It is not our differences that divide us.

It is our inability to recognize, accept and celebrate those differences.”

(Audre Lorde, 1934-1992).

Although Audre Lorde did not have technical scientists (engineers) in mind when she penned those words, it behoves us to recognize, accept, and celebrate our differences. The power of our technical society lies in the fact that we are not monolithic. We are engineers active in different fields: material science, mechanics, kinematics, thermodynamics, energy and environment, mechanics and robotics, fluid mechanics, tribology, cybernetics, industrial engineering, policy making, and more. We benefit from multidisciplinary collaboration so that our whole is greater than the sum of our parts. We do research grounded in the fundamental scientific principles. We identify different problems with different underlying fundamental principles. We know how to assess the physical, mechanical, chemical, medical and other root causes of those problems, find engineering solutions, and help to derive policies that positively affect our lives.

Our scientific diversity is reflected in the wealth of topics associated with the 92 articles that were published in Volume 57 of the SV-JME. But, are these articles really reflecting the new trends and proven practices in mechanical engineering and the closely related sciences? Are the special issues of the SV-JME with selected articles covering the best international conferences? The real answer lies in the number of citations an article receives in internationally acknowledged journals. But as the success and quality of magazines, SV-JME included, is a reflection of the excellence of its reviewers, we are going to publish in the first 2012 issued of SV-JME a list of reviewers who collaborated in reviewing articles for the 2011 Volume of SV-JME. We are deeply grateful to each and every reviewer for

their excellent cooperation and for their time. We would also like to take this opportunity and invite the scientists and other top and prominent experts from around the world to contact us if they are willing to do review work, as we have deployed an electronic review system which facilitates the work significantly. Further information and the history of SV-JME journal are available at our web site.

The mankind enters the nature and the global world with the products of its hands and intellect. To our grandchildren and the world to come, we are leaving a diverse technical heritage, which we celebrated this year in Slovenia with a festive academy and international conference of mechanical engineers to honour the 200. anniversary of higher education in mechanical engineering and the 50. anniversary of the Association of Mechanical Engineers of Slovenia.

Together we are looking forward to any new or advanced technical product or achievement that introduces state-of-the-art technology, contributes to environmental protection and reduced energy consumption, or anything else bringing a safe and healthy life for the generations to come, and at the same time advances the technical culture.

One of such remarkable achievements of technical research which attracted attention worldwide is the success of Slovenian researchers in the field of aviation.

In September 2011, the Taurus G4 electrically powered aircraft flew at an equivalent fuel efficiency of 403 passenger miles per gallon at a speed of 107 miles per hour (172.1 km/h), winning the Green Flight Challenge 2011 (GFC).

This is both twice as efficient and twice as fast as a Toyota Prius, and almost six times as efficient as a typical 4-seat general aviation aircraft. The aircraft was designed and built by a Slovenian manufacturer of self-launching sailplanes and light sport aircraft.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 867-868

The festive academy at the 200. anniversary of higher education in mechanical engineering and the 50. anniversary of the Association of Mechanical Engineers of Slovenia

a mere two years ago the thought of flying 200 miles (321.8 km) using battery power alone was pure science fiction, and now two aircrafts have shown that this is possible.

Our ability to address these and many other future challenges will benefit from continuing advances in analytical methods, material science, computational fluid dynamics and more. But perhaps our most valuable asset is the diversity, creativity, and youthfulness of our technical society. Let’s observe what we are capable of and seize the great opportunities our technical society has to address the grand challenges of the future, and to contribute to the progress and to the health of people around the world.

Two teams from the 14 enrolled at the GFC have shown that electricity is a viable (indeed, a beautiful) way to power aircrafts. The Taurus G4 and the eGenius are both virtually inaudible as they fly overhead, even at low altitudes.

Electricity is comparatively cheap: at 12 cents per kilowatt-hour it costs under €7 to fly the Taurus G4 for nearly 3 hours. At the GFC, the power for the batteries came from a geothermal power plant, so the flights were truly emissions-free. We are witnessing, and participating in, the birth of affordable and truly green personal air travel.

Startup companies, small businesses and aircraft homebuilders all over the world are working on electrically powered aircrafts. Many electrically powered aircrafts are expected to take the skies over the next few years, as well as continuous performance improvements as the battery technology continues to improve.

The Slovenian company believes that progress comes from bold vision and goals. If NASA accepts their challenge, the company will contribute $100,000 from its Green Flight Challenge 2011 prize to the first electrically powered aircraft that can fly faster than the speed of sound. They fully expect someone to claim this prize within the next five years. Remember that 868

Good luck!

Vincenc Butala References: Butala, V. (2010). Editorial – Anniversary: 55 Years of Strojniški vestnik – Journal of Mechanical Engineering. Strojniški vestnik – Journal of Mechanical Engineering, vol. 56, no.

12, p. 789-790.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 869-878; DOI:10.5545/sv-jme.2011.212

INVITED PAPER

Pipistrel Taurus G4: on Creation and Evolution of the Winning Aeroplane of NASA Green Flight Challenge 2011

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

Tine Tomažič1,* – Vid Plevnik1 – Gregor Veble1 – Jure Tomažič1 – Franc Popit2 – Sašo Kolar3 – Radivoj Kikelj4 – Jacob W. Langelaan5 – Kirk Miles5 1 Pipistrel d.o.o. Ajdovščina, Advanced Light Aircraft, Slovenia 2 Aeroidea s.p., Slovenia; 3 Aerotech s.p., Slovenia; 4 Lupina s.p., Slovenia 5 The Pennsylvania State University, Aerospace Engineering Department, United States The paper presents the journey of the Pipistrel’s Taurus G4, the World’s first four-seat electric aeroplane and winner of the NASA Green Flight Challenge 2011, sponsored by Google, from idea to completion of competition. At the beginning, the race event and qualification requirements are presented.

The selection of the aeroplane’s configuration and sizing is discussed, and emphasis given to the unique twin-fuselage configuration. Next, the architecture of the electrical systems on board the aeroplane, including the large battery pack and propulsion elements is presented. In the second part of the paper, authors give an insight to the 40-hour flight test, the challenges encountered and the preparation to the race event itself. A non-linear mathematical performance model was developed to predict performance of the Taurus G4 during the two race events and used for flight planning and in-flight online performance optimisation. Instead of the conclusion, the results of the NASA Green Flight Challenge 2011 are presented and discussed, together with a comparison to other participating teams’ results. © 2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: electric aeroplane, structural efficiency, NASA Green Flight Challenge 2011, large battery pack, performance modelling, flight test 0 INTRODUCTION NASA organises Centennial Challenges [1] as part of their commitment to on-going innovation and spreading-out engineering efforts to small businesses to tackle large problems.

The Green Flight Challenge 2011 [2], sponsored

by Google (GFC 2011) was the first challenge organised after the Personal Air Vehicle 2007 [3] challenge, where Pipistrel won with the Prototype of Virus SW 80 aeroplane and General Aviation Technology 2008 [4] challenge, where Pipistrel also won with the Prototype of Virus SW 100 aeroplane.

The goal of the GFC 2011 was simple to understand, however the formulation of the rules of engagement [5] and the pre-requirements to even qualify for the race event itself made it difficult to process for some and impossible to meet for others. The winner was to complete two main event sorties i.e. the Economy Flight and the Speed Flight, each 200 miles (321.8 km) long, in less than 2 hours and with an enegy efficiency of the equivalent of 200 passenger-Miles-Per-Gallon *Corr. Author’s Address: Pipistrel d.o.o. Ajdovščina, Goriška cesta 50a, SI-5270 Ajdovščina, Slovenia, tine@pipistrel.si

(pMPG), which is eqivalent to approx 170 km per litre of fuel per passenger. Implicite to this formulation is that the contenders had to fly the course at an average speed of at least 160.9 km/h, which included the take-off.

There were, however, many requirements, which the competing aeroplanes and teams had to meet. Most important were the stall speed of 52 mph (83.5 km/h) or below, good stability and manoeuvrability, based on the CooperHarper scale, maximum noise level on take-off of 78 dB(A), FAR 25 [6] cockpit visibility, cabin comfort and the ability of carrying 200 pounds (91 kg) per declared seat, which also needed to be evacuatable in one minute without outside assistance. The number of the seats was not limited, however the Maximum Take-off Mass of the airplane was not to exceed 5,700 pounds (2,591 kg). The rules also did not impose a certain means of propulsion, but as the overall energy consumed, recalculated to the caloric equivalent of 1 US gallon (3.78 litres) of automotive fuel via the British Thermal Unit (BTU) principle, mattered for the final result in pMPG, the efficiency of 869

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the powertrain mattered almost as much as the efficiency of the powertrain. This important fact was the turning point for the conception of the Pipistrel’s all-electric competition aeroplane, the later named the Taurus G4.

1 CONCEPTION PHASE At this point, it needs to be stressed that the conception, development and construction of the Taurus G4 took only 5 months (20 December, 2010, through 18 May, 2011), so the design team needed to think in a most creative way and envision an efficient design, which would become an all-electric aeroplane, thus maximising the pMPG score by incorporating the most efficient means of propulsion. First calculations showed that an 85+% powertrain efficiency could be achieved with an all electric (battery, power controller, electric motor) type of powertrain, whereas hydrogen-fuel cells, hybrid or internalcombustion-engine-based powertrains could come no-where close to such figures.

The other part of the pMPG score is the number of passengers carried. The relation between the pMPG score and the number of passengers carried is directly linear; therefore the concept of Pipistrel’s GFC 2011 racer was to have as many seats as possible, but economically feasible. This built was not sponsored by any entity outside the Pipistrel d.o.o. Ajdovscina company itself. Given the tight time frame, the team decided to use certain parts, which were available at the workshop from the existing production airframes.

Efforts were immediately given on Taurus Electro G2 airframe, Pipistrel’s two-seat electric powered self-launching glider, which is powered by a 40 kW electric motor. While the motor’s power would be sufficient to produce the required 100 mph cruise with two pilots on board, it would have been necessary to extend the wingspan by at least 2.5 meters, and the fuselage by at least 0.7 meters to achieve adequate stability and enough physical space for all the batteries on board.

Parasite and trim drag associated by having the propeller mounted on top of the retractable mast present a performance penalty versus more conventional solutions. This idea was scrapped also because it was believed that most of the 14 870

teams, which had enrolled for the GFC 2011, would be flying two-seat aeroplanes, and the modified Taurus Electro G2 would have no advantage in the number of passengers carried.

A similar concept with three people aboard a single Taurus fuselage, with greatly extended wingspan (extra 4.5 meters) and fuselage (extra 1.3 meters) was evaluated, but quickly dismissed because of structural inefficiency and construction complexity. The main issue with this design was the necessity of a completely new and different retractable undercarriage system and powertrain installation.

While the advantage of having three passengers aboard the airframe was not great over having only two, it was noted in the process that by carrying four passengers the score becomes notably i.e. in the order of 10% higher than that achievable by two-seat designs. This is, of course, accounting for extra airframe drag of a four-seat design.

1.1 Airframe Configuration and Sizing How does one make a four-seat aeroplane, possibly utilising parts from existing airframes, which are available in-house? Take two airframes and make a new, larger one – this is how the twinfuselage idea was born.

Fig. 1. First configuration and sketches for the Taurus G4 It turns out that span efficiency as well as structural efficiency of a twin-fuselage aeroplane can present obvious advantages, especially with

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 869-878

an all electric aeroplane, where the mass of the batteries accounts for a large portion of the total airframe weight. Such design could carry four passengers and have enough room for all the Lithium Ion-Polymer batteries, then judged at a little less than 500 kg themselves. The undercarriage would not need to be redesigned, and the modification of the pushrod-based control system seemed straightforward. There were challenges in propeller clearance / propulsion nacelle positioning as well as sizing of the central wing. These were also the major components of the to-be Taurus G4, which needed to be designed and built from scratch.

The question persisted in how slender the middle wing could be, to maintain structural (especially torsional) rigidity. Several concepts were evaluated, including such with conjoined tails and different wing platforms. Drag induction devices (one large flap, one smaller flap with airbrakes, just airbrakes) were also discussed, given the fact that the Taurus G4 was showing to become a near 1,500 kg aeroplane with a Lift/ Drag (L/D) ratio of around 30:1. Good handling and landing characteristics needed to be achieved by all means.

characteristics. Two separate horizontal tails were used instead of the single conjoined tail for aero elastic reasons. At this point, design was simultaneously ran in three separate directions, aerodynamic design and optimisation, structural calculations and design/development of the electrical/propulsion system.

1.2 Aerodynamic Design Immediate attention was given to designing a flapped airfoil for the central wing, which had to be built the soonest. A 15% thick airfoil was designed to provide high lift, but to maintain optimal lift distribution in sync with outboard wings, which were stock-geometry Taurus Electro G2 wings, where the airfoil could not be modified. A slotted flap (35 degrees) was selected for weight/lift benefit/drag induction reasons. The result was a system capable of reaching CLmax of 2.8 and enabling the required stall speed of the aeroplane at <83.5 km/h with its five-meter span at a calculated Maximum Take-off Weight of 1,500 kg. Special wing-body-joint fairings were also designed to reduce interference drag from the central wing towards the two fuselages.

Fig. 2. Further concept focusing on central-wing sizing, horizontal tail solution and drag induction devices Before the final design concept was determined, all of the above were modelled and subject to CFD and structural FEM analyses.

The favourite design closely resembled the top sketch on Fig. 2, presenting the best combination of structural and aerodynamic

Fig. 3. CFD optimisation of central nacelle The next challenge was the design of the central nacelle. Although the concept sketches

Pipistrel Taurus G4: on Creation and Evolution of the Winning Aeroplane of NASA Green Flight Challenge 2011

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 869-878

envisioned a much smaller nacelle, it was seen through CFD analyses that a larger and longer nacelle would actually be beneficiary due to parasite and interference drag reasons. This gave the opportunity to utilise the nacelle not only to house the key elements of the propulsion system (electric motor, power controller, cooling system) but also the ballistic parachute rescue system, part of the battery system and the flap actuator mechanism. Laminar flow over the nacelle was ensured with CFD-based shaping and surface optimisation, as seen in Fig. 3.

engine characteristics as well as competition requirements. Electric propulsion can deliver power at flexible RPM, which allows the use of a fixed pitch design. The propeller’s unique shape is a result of optimization using in-house developed computer tools towards three GFC 2011 goals; the thick, inboard section of the blade is responsible for providing thrust during take off, whereas the thinner, outboard part of the blade is designed for climb and especially cruise efficiency.

It must also be noted that the ratio between take-off power and cruise power of the aeroplane is in the order of 5:1!

1.3 Structural Design

Fig. 4. Propeller of Taurus G4

When it was decided that there is enough physical space inside the central nacelle to also house part of the battery system, the advantages of the twin-fuselage concept became apparent from the structural point of view. The mass of the batteries/propulsion system/passengers could be more evenly spaced along the span than with the conventional single-fuselage configuration (Fig. 6).

Fig. 6. Span-wise weight distribution, bending moment for Taurus G4 Fig. 5. Propeller thrust vs. airspeed and RPM The unique propeller was hand crafted from ash wood and reinforced with a thin composite layer. Despite being fixed-pitch, the propeller behaves much like constant-speed propellers on piston engines, thanks to electric motor’s characteristics (Fig. 5).

The propeller of the Taurus G4 is a very special component and was tailored for electric 872

Due to relatively large masses spaced outboard from the centre of gravity, the weight of the structural parts was lighter. In combination with the tailor-developed central wing/airfoil, the following lift distribution was achieved (Fig. 7); enabling the Taurus G4 to use unchanged stock outboard wing panels taken from Taurus Electro G2 two-seat self-launching glider.

The complete structural design, analyses and testing were carried out according to FAR 23c

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 869-878

specifications, rendering the Taurus G4 a proper +3.8 G, 1.52 G aeroplane. The G-load envelope (Vn diagram) is shown in Fig. 8.

Fig. 9. Locations of battery groups

Fig. 7. Span-wise lift distribution, Taurus G4

Within groups, the cells are organised in eight packs where each is equipped with a BMS ‘slave’ electronic board, capable of measuring individual cell voltages, voltage balancing and measuring temperature. A total of 33 eight packs are mounted in the fuselages and in the central wing inside glass fibre cages, which are electrically nonconductive. In case the BMS would detect a cell failure during flight, the relay would disconnect the corresponding battery group and the airplane would continue its flight with the rest of the two chains.

Fig. 8. Vn diagram, Taurus G4 1.4 Electrical Systems All electrical systems for the Taurus G4 were developed in-house, apart from the battery cells themselves, which are a batch of specially commissioned high-energy-density LithiumPolymer cells. There are three battery groups connected in parallel, each group contains 88 LiPo cells, which are connected in series. The array of 264 cells weighs in at 520 kg and has a total energy capacity exceeding 90 kWh at a nominal voltage of 325 volts. To best of authors’ knowledge it is the largest battery array ever used on an aeroplane.

The three battery groups are located one in each fuselage and one in the central nacelle.

Individual groups can be switched on/ off during the flight via the 400 A fuses & 500 A vacuum relays and are monitored by the proprietary Battery Management System (BMS).

Fig.10. Battery group in one of the fuselages Each of the battery groups is controlled by a ‘master’ unit, which handles all operations regarding the given battery group. In addition to controlling the on/off chain power relay, the ‘master’ collects cell voltages and temperature data from the BMS slaves to balance the cell voltages.

The power-controller and electric motor are liquid cooled and designed for 150 kW maximum

Pipistrel Taurus G4: on Creation and Evolution of the Winning Aeroplane of NASA Green Flight Challenge 2011

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power. All elements from the Electrical Systems, including the 12 kW charger, are controlled via a centralised avionics suite located in the centre cockpit console and communicate with oneanother via a hybrid wire/optical fibre CANbus network.

There is a separate avionics/systems batteries, which takes care of instruments and central flap/landing gear actuation. An additional 17 watt solar cell is located on the starboard fuselage as a back up to power instruments in event of total electrical system failure.

was dispatched to the USA in a flight-ready condition. The rules of GFC 2011 called for 40 flight hours be accumulated in the USA prior to the competition itself.

1.5 Rapid Prototyping The Taurus G4 is essentially a prototype, with many components built with advanced robot prototyping techniques. All designs were transferred from 3D CAD/CAM models into physical form using Pipistrel’s 8-axis robot-mill (Fig. 11).

Fig. 12. Central-wing root ribs being fabricated

Fig. 13. Central-wing-spar fitted to wing-box assembly for the first time 2 TEST PHASE Fig. 11. Robot milling a Taurus G4 component Prototype moulds made of either styro foam or Polyurethane block materials were used to fabricate individual composite parts (Fig. 12).

Such a process allowed for the whole composite structure to be completed within 14 weeks from beginning of the project. The remaining 6 weeks were spent integrating electrical systems, controls, avionics, etc.

The completed Taurus G4 airframe was geometrically identical to the computer-designed shape and 4 kg lighter than calculated.

On May 18th, 2011, roughly five months after the beginning of the project, the aircraft 874

The test phase of the Taurus G4 took place at three different locations in the USA. Having already completed all structural tests at Pipistrel’s headquarters in Slovenia, the test phases could be named: integration and environmental testing, initial test flights and performance oriented test flights.

2.1 Integration and Environmental Testing The Taurus G4 arrived to Mifflin County Airport, PA, USA in the first days of June 2011.

There, the team assembled the airframe and began testing the systems, particularly the elements found inside the central nacelle.

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

StrojniĹĄki vestnik - Journal of Mechanical Engineering 57(2011)12, 869-878

controls. The reason behind the delay between the testing in Pennsylvania and the first flight was the postponement of the competition date.

Fig. 14. After arrival to Mifflin Co, PA, USA Pennsylvania is known for its hot and humid weather in early summer, hence the focus of the integration and environmental testing was on cooling, operation and reliability of the electrical systems, including motor, power controller, disconnect relays and the BMS. NASA-provided eTotalizer measurement board was also integrated and its operation verified for safety at that point.

The space inside the nacelle is crowded, so adequate cooling is critical. The motor and power controllers are water cooled by a 50/50 water/ glycol coolant, which is force-circulated inside a radiator-closed system. The cooling opening is the mouth below the propeller in the front of the nacelle; the warm air exhaust is at the rear of the nacelle.

Fig. 15. Central nacelle inner elements: motor left, batteries right, parachute rescue system in the middle A series of charge-discharge cycles also were conducted at Mifflin, simulating the power regimes the Taurus G4 would be subject to in flight during the races. Cooling as well as battery performance and endurance was verified.

2.2 First Flights At Oshkosh, WI, USA, the Taurus G4 flew for the first time on August 11, 2011 with the experienced test pilot Dave Morss at the

Fig. 16. Taurus G4 high above Oshkosh, WI With Morssâ&#x20AC;&#x2122; feedback and data collected with the on-board flight/engine logger, the electrical system was further optimised and responses tweaked. Over the course of one week, the first 11 flight hours were completed and the Taurus G4 cleared FAA imposed Phase 1 of flight testing, which meant it could be flown anywhere in the USA from then onwards.

2.3. Performance Oriented Testing The Taurus G4 was moved again, this time to Hollister, CA, USA, which is within flight distance to Santa Rosa, CA, where the GFC 2011 was to be held. Hollister is also very predictable in terms of weather and allowed flying with the Taurus G4 every single day of the race-preparation phase.

Every day brought new knowledge about the aeroplane and as the data accumulated over time, the team assembled a mathematical performance model [7] of the Taurus G4, which allowed predicting the aeroplaneâ&#x20AC;&#x2122;s behaviour in a race environment with high fidelity.

The model, assembled in Matlab/ Simulink computer environment later served as a basis for pre-flight planning and in-flight online performance optimisation. It includes data on battery system performance and behaviour, aerodynamic data and can determine the influence of a given wind situation aloft.

While it was originally planned to use a netbook computer in the cockpit during the race and run the model during the actual race flights, this idea was abandoned due to reliability issues with the hardware selected. Instead, the model was used on ground where several flight plans for the pilots were prepared, based on different simulation runs. With the wind data known from aviation forecasts, simulations were made for different average travel speeds along the then-

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known course. The pilots received a very simple checklist, which only contained time, at which the aeroplanes must be at a certain turning point, the planned electric power to achieve the desired average travelling speed (taking wind into account) and the cumulative energy consumption in kWh.

Fig. 17. Performance testing at Hollister, CA Careful readers would notice that the flighttesting was never conducted for different weights and centre-of-gravity positions. This is closely linked to the twin-fuselage configuration of the Taurus G4. The aeroplane can actually be flown in only one configuration (always at Maximum Take-off Mass) and one, carefully determined centre of gravity position.

3 TAURUS G4 SPECIFICATIONS Due to the unique configuration of the aeroplane, some aspects of the performance envelope were not known in advance.

While the efficiencies of the propeller and powertrain matched theoretical data within one percent, the L/D ratio proved to be rather different than anticipated. This was linked to a very complex aerodynamic form or the Taurus G4, with three different wing sections and three different fuselage bodies. On the middle wing there was some flow separation at speeds just below 100 mph (160.9 km/h), which was sourced to complex propeller wash influencing the laminar boundary layer locally. Hence, the global maximum of L/D characteristics was shifted to a higher airspeed and the shape of the curve was mutilated to some degree. To designers’ satisfaction, L/D ratio at higher airspeeds, however, proved to be superior to the calculated one.

The specifications and performance data as obtained through flight-testing of the Taurus G4 876

are summarised in Table 1. The 3-view drawing of the Taurus G4 is presented in Fig. 198.

Fig. 18. Taurus G4 3-view drawing Table 1. Pipistrel Taurus G4 basic data Proportions Wing span 21.36 m Length 7.40 m Weights Empty weight (excl. batteries) 632 kg Empty weight (incl. batteries) 1132 kg Maximum take-off weight 1500 kg GFC 2011 competition weight 1496 kg Powertrain Power, RPM 150 kW at 5500 RPM System voltage (nominal) 325 V Battery capacity 3x30 kWh Performance Stall speed 82 km/h Maximum speed 217 km/h Cruise speed 160 to 201 km/h Take-off distance (over 15 m) 600 m Best climb speed Vy 156 km/h Best climb rate 4.5 m/s L/D at 100 mph 28+: 1 Req. PWR for 100 mph cruise 32 kW Endurance 2:45+ h Range 400+ km 4 THE RACE AND RESULTS The NASA Green Flight Challenge 2011, sponsored by Google, was a NASA organised

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

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event, designed to promote efficiency of flight, with particular emphasis on all-electric aircraft.

The actual event consisted of two flights, each oriented toward a different aspect of aircraftâ&#x20AC;&#x2122;s performance.

The Economy Flight was a 200 statute mile flight, in the form of four 50-mile laps, where the competitors had to fly the course as economically as possible. To qualify for the win, the aeroplane had to exhibit a fuel consumption of no less than 200 passenger-miles-per-gallon, while the average flight speed had to be in excess of 100 mph (160.9 km/h). The winner of this flight was the aeroplane that exhibited the highest passenger-miles-pergallon score, where one gallon was converted to an equivalent of 33.7 kWh for electric aeroplanes.

events were flown even more precisely than ever before, finishing one flight 20 seconds ahead of schedule and 0.2 kWh short of energy consumption prediction, and the second flight 6 seconds ahead of schedule and 0.4 kWh ahead of target energy prediction. This degree of model vs. reality precision allowed Pipistrel team to maximise their score by flying as fast as the conditions allowed while maintaining the required 30-minute power reserve exactly. The results together with final score and performance of other teams are presented in the NASA official table of results, in Table 2.

Fig. 20. Taurus G4 at start of the Speed Flight Fig. 19. Behind the scenes, just before take-off for the Economy Flight The Speed Flight was flown over the same 200 statute mile course as the economy flight.

However, the aeroplanes had to fly as fast as possible, but still have enough energy on board after landing to demonstrate a 30 minute reserve in power, based on the average power used during the Economy Flight.

The overall winner was determined by the following scoring Eq.:

Score =

1 , (1) 1 2 + a b

where a is the average flight speed achieved during the Speed Flight and b the passengermiles-per-gallon efficiency achieved during the Economy Flight.

Before the race events, there were multiple occasions where it was proven that the fidelity of the model closely matched what the Taurus G4 and their pilots could do in-flight. Both race

Fig. 21. Ground track from Speed Flight 5 CONCLUSION The Taurus G4 finished the Green Flight Challenge as the winner, demonstrating unprecedented efficiency of flight (403.5 pMPG).

Moreover, it is the largest, most powerful, fastest (enduring) and most economic electric aeroplane to date.

Part of this efficiency came about because of an inspired design. The Taurus G4 used a multi-body concept, which accomplished a 61%

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useful load fraction (or empty weight fraction of 39%). The Taurus G4 also achieved a motor/ controller efficiency of 96%, and utilized the largest battery pack ever assembled for auto or aviation use of 90+ kWh, which included a 30-minute reserve capacity. It is noticeable that it was the combination of configuration, structural efficiency, aerodynamic efficiency and the unique electric powertrain that gave the edge to Pipistrel’s Taurus G4 over the most serious competitor, the eGenius, which flew a more conventionally configured aeroplane.

Table 2. Official table of results in SI units Pipistrel

e-Genius

Fuel used

Phoenix 14.45

14.45

Liters 100 LL

Energy used

65.4

34.7

3.8

kWh

Equivalent fuel used

7.45

3.89

15.00

15.50

Liters auto fuel

Flight time (for speed)

1:47:16

1:48:27

2:25:01

2:00:48

Flight time (for milage)

1:49:37

1:50:23

2:25:43

2:04:07

Distance (for speed)

308.9

307.3

300.4

228.5

km

Distance (for milage)

315.2

311.6

302.2

238.3

km

Milage

171.8

159.9

40.1

30.7

km/l/P

Speed

172.8

170.1

124.4

123.4

km/h

25.0

15.8

Liters 100 LL

Energy used

68.3

37.5

Equivalent fuel used

7.67

4.20

Flight time (for speed)

1:41:55

Flight time (for milage)

1:44:10

Distance (for speed)

Speed Competition Fuel used

3.0

kWh

26.08

7.19

Liters auto fuel

1:47:45

1:22:11

1:42:21

1:50:24

1:22:57

1:4:53

310.5

310.0

303.1

234.9

Distance (for milage)

316.7

315.7

304.9

234.0

km

Milage

165.3

150.0

23.4

27.9

km/l/P

Speed

182.8

172.6

221.2

134.0

km/h

Score

72.7

68.3

35.1

25.2

km

To conclude, the authors would like to quote a letter by an aviation enthusiast, who also witnessed the preparations for the race: “I was there for the last flight of the Taurus G4 at Hollister after their return from the awards ceremony. There was just me – no other visitors – they had just won the prize and there was no one there but me. I was at the last big National Air Race in Cleveland, 1947, maybe. There were several hundred thousand spectators there. What

878

6 ACKNOWLEDGEMENTS

The team, Pipistrel-USA.com

Efficiency Competition EcoEagle

a contrast! I have no idea how disruptive electric propelled flight will be, but surely it interests more than just me. Congratulations to the team!

They’ve shown they can do things none of the other 7 billion people in the world can do!”

Sincerely, Bob Lockhart

Operation part financed by the European Union, European Social Fund (Junior Researchers). Operation part financed by the Slovenian Research Agency (ARRS L2-3644).

7 REFERENCES [1] www.nasa.gov/challenges/, accessed on 2011-11-25 [2] http://cafefoundation.org/v2/gfc_main.php, accessed on 2011-11-25 [3] h t t p : / / c a f e f o u n d a t i o n . o rg / v 2 / p a v _ pavchallenge_2007_results.php, accessed on 2011-11-25.

[4] h t t p : / / c a f e f o u n d a t i o n . o rg / v 2 / p a v _ gatchallenge_2008_results.php, accessed on 2011-11-25.

[5] http://cafefoundation.org/v2/pdf_GFC/GFC. TA.07.28.09.pdf, accessed on 2011-11-25.

[6] http://cafefoundation.org/v2/pdf_GFC/GFC. FOV.122709.pdf, accessed on 2011-11-25.

[7] Tischler, M., Remple, R. (2006). Aircraft and Rotorcraft System Identification: Engineering Methods With Flight-test Examples. American Institute of Aeronautics and Astronautics, Reston.

[8] http://cafefoundation.org/v2/gfc_2011_ results.html, accessed on 2011-11-25.

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 879-887 DOI:10.5545/sv-jme.2010.210

Paper received: 01.10.2010 Paper accepted: 27.10.2011

Kinematics and Dynamics of the Quasi-Passive Biped “PASIBOT”

Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C. Jesús Meneses – Cristina Castejón* – Eduardo Corral – Higinio Rubio – Juan Carlos García-Prada University Carlos III de Madrid, Department of Mechanical Engineering, MAQLAB group, Spain A quasi-passive biped (having only one actuator) developed into a Spanish project called “PASIBOT” [1] is presented in this article. We focus on the PASIBOT’s topology, kinematics and dynamics, and we describe a program designed for carrying out the corresponding calculations. This code provides for all kinematic and dynamic data, as functions of time, along one step: position, velocity and acceleration of all members, as well as all the forces and torques on each of them, motor torque included. This latter information has helped us to choose the required motor, as this choice depends on some parameters of interest that can be modified in the program, like density or link dimensions. Also, we will be able to get strain-stress data in all links in the course of a step, and then optimize those dimensions. To finish, some results are also presented that confirm the interest of the developed code. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: mechanism analysis, quasi-passive biped, walking robots 0 INTRODUCTION Currently, service robotics is one of the main research priority areas. The application of robots for service tasks (personal assistance, education, social tasks, etc.) makes its design very important. Current mobile robots are not adapted to be used in domestic environments due to their large volume and/or weight, and their lack of maneuverability in these complex scenes. The interest in the development of humanoid robotics (which perform the tasks commented before) is increasing and such robots are being developed by a great number of research groups in the entire world [2] and [3].

Nowadays, humanoid robots are formed by a high number of actuators, used to control the high degrees of freedom (DOF) they have [4]. On the other hand, one of the biggest drawbacks in humanoids is the weight and power consumption.

In the majority of them, around 30% of the total weight is due to the actuators and wires, and more than 25% is due to the reduction systems coupled [5]. For this reason, this work focuses on the design of new mechanisms and kinematic chains which, maintaining the robot capabilities, require a smaller number of actuators. This would reduce the robot mass and hence, its power consumption and the total cost.

During the last few years, different research groups have developed robots based on passive walking techniques [6]. An example is the one meter length Robot Ranger of Cornell University [7] with three joints in each of its long legs. The robot can walk in a similar way to a human, by means of the balance and the dynamic of the natural swinging, in order to consume the minimum energy to walk. Robot Toddlers from MIT University [8] is a small robot that only has a single passive pin joint at the hip, while the 3D movement is achieved by means of the feet surface design. Toddlers is designed only to walk down a shallow slope. Robot Denise from Delft University [9] is a pneumatically powered walking robot with human configuration and five DOF, and the last robot of interest in the passive theory is the one developed in the Nagoya Institute of Technology [10] whose topology is similar to Denise Robot but it only presents two legs, and it includes a stability mechanism of fixed point [11].

In this article, a human size biped, called PASIBOT, with low DOF, which represents a qualitative improvement in the service robotic field, is presented. The innovative design has been carried out with the combination of classical mechanisms (Peaucellier, Watt, pantograph [12], etc.). This prototype is based on the one designed and built at the Laboratory of Robotics and Mechatronics in Cassino (LARM) [13] and

*Corr. Author’s Address: University Carlos III de Madrid, Department of Mechanical Engineering, Avda. De la Universidad, 30, 28911 Leganés, Madrid, Spain, castejon@ing.uc3m.es

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[14], following the philosophy of low cost [15]. A similar leg design has also been used in a four legs walking chair [16].

The proposed mechanism is an arrangement of links in the planar movement that has only one DOF. In this manuscript, the planar kinematics and dynamics analysis of PASIBOT is presented.

The study is performed from a theoretical point of view, and aims at obtaining the linear and angular position coordinates, velocities and accelerations for all links, as well as all the forces and torques between links including motor torque, for any time in the course of one step. The expressions have been implemented in MATLAB® code, and the corresponding results have been used in the design and construction of a real prototype, and they are being used in movement control tasks.

In the section “topological description of PASIBOT”, the biped PASIBOT mechanism is described by defining its subassemblies and parts, and the nomenclature used; next, in the section “Kinematics of PASIBOT”, it is explained how to deduce the expressions for the angular and linear position, velocities and accelerations for all the links of the biped; then, in “Dynamics of PASIBOT”, the method of obtaining all the forces and torques on every link, at every time in the course of one step of PASIBOT is presented; in “Numerical results” a code designed to calculate the kinematical and dynamical values for different set of parameters (motor angular velocity, density, link dimensions, etc.), including the corresponding results is presented; and finally, the conclusions of this work are presented.

two trajectories tracked by the points of interest.

Chebyshev mechanism transfers the motor rotational movement at its crank into a continuous cyclical trajectory, which is composed of a curved section and a quasi-straight one, at the end of its connecting rod (Fig. 1, point C). This point is then linked to a pantograph mechanism in such a way that its free end (Fig. 1, point E) executes a trajectory that is inverted and amplified with respect to that described above. The ratio of magnification of the pantograph depends on the dimension of its bars; for the design presented in this work, this ratio is two.

1 TOPOLOGICAL DESCRIPTION OF PASIBOT

The relative positions of points A, B and D in Fig. 1, are fixed at the member called “hip”, shown in Fig. 2. The hip also carries a slot (see Fig. 2) which is the base of the stabilization system: a set of links arranged in parallelograms with the two longest bars of the pantograph. This stability extension guarantees the parallelism between the supporting foot and the stabilizing bar, whose end slides along the slot at the hip. The first approach is to align the slot with the linear section of the Chebyshev trajectory, in such a way that the supporting foot remains also parallel to the slot.

To provide the opposite leg with the proper movement, the corresponding crank is phased out

The biped presented in this article (see Figs. 2 to 4), is a mechanism that can be divided into three essential subassemblies or “submechanisms”, each of them having a particular function: 1. Quasi-straight line generator mechanism (Chebyshev); 2. Amplifier mechanism (pantograph); 3. Stability extension and foot (parallelogram extensions).

In Fig. 1, the coupling Chebyshevpantograph mechanism is shown, together with 880

Fig. 1. Coupling Chebyshev-Pantograph mechanism with the corresponding trajectories of interest

Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C.

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p rad (see Fig. 2) in the same motor axis. In fact, both cranks take part of the same rigid element.

Table 1. A list of symbols for a general link (number i) is presented li ϑi ϑiland ωi αi mi Ii

Fig. 2. Sub-mechanisms of PASIBOT; a) nomenclature and numeration for the supporting leg, b) angular positions for the links; the members of the opposite leg will be referred to using primes; Chebyshev sub-mechanism: links 7, 8 and 9; Pantograph sub-mechanism: links 2, 3, 4 and 6; Stability extension sub-mechanism: links 1, 5, 10, 11 and 12 As can be seen in Fig. 2, each link has been numerated and named, using prime (x’) for the links belonging to the flying leg, to distinguish from those belonging to the supporting leg. Each leg has 12 links, but since the motor crank (link number 8) is shared with both legs (hence, there is no link number 8’), the number of links for PASIBOT, including the single hip (link number 13), is 24.

rij rijx rijy fij fijx fijy

Length of the link i [mm] Angle between the link i and the hip (direction defined by points A-B in Fig. 1), also called ϑihip, when necessary [rad] Angle between the link I and the land (direction defined by the supporting foot) [rad] Rotational velocity of link [rad/s] Rotational acceleration of link [rad/s2] Mass of the link i [kg] Inertia Moment for the link i with respect to the axis perpendicular to the movement plane containing its centre of mass [kg mm2] Position vector of the ij joint from the link i centre of mass [mm] x-projection of the position vector [mm] y-projection of the position vector [mm] Force exerted by the link i on the link j [N] x-projection of the fij [N] y-projection of the fij [N]

In Fig. 4, a sequence for one step of PASIBOT is presented, as simulated with a mechanical program. Note that one step corresponds to a half rotation (p rad) of the motor crank.

Fig. 4. A track for a PASIBOT gait along one step (from ϑ8= p to 3p/2 rad) 2 KINEMATICS OF PASIBOT

Fig. 3. Some photographs of PASIBOT while walking

The kinematical study presented here is related to one PASIBOT’s step, having one of its feet (the supporting foot) always in contact with a horizontal surface (x axis). No relative motion between the supporting foot and the ground is considered, so this foot could be referred to as the ground element. Taking into this account, the PASIBOT is a planar mechanism with one DOF,

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so the angular positions of any link can be referred to the angular position of the motor crank (θ8):

θi = θi(θ8), i = 1, 2, …, 1’, 2’, …

(

)

6e jϑ6 + 3e jϑ3 + 2.5 e jϑ7 + e jϑ9 − 12 j = 0. (4)

(1)

Then, the x, y coordinates for its centre of mass, can be easily expressed with respect to that angle: xi = xi(θ8); yi = yi(θ8), i = 1, 2, …, 1’, 2’, … (2) Furthermore, if the time dependent function for the motor crank angle is known, those coordinates can also be expressed as time dependent functions. The corresponding angular velocities and accelerations, as well as the centre of mass linear velocities and accelerations are obtained by taking the first and second derivatives of functions in Eqs. (1) and (2).

The biped kinematics is based on three close loop kinematic chains (one for each submechanism described above) which lead to the following three equations systems (the link lengths have been particularized for the designed PASIBOT, and normalized to the crank length, so that l8 = 1, as the resulting angles are independent of the scale): 1. Chebyshev chain (formed by links number 7, 8, 9 and 13) In a Chevyshev mechanism, the distance between motor crank and rocker arm fixed points (A and B in Fig. 1, respectively) is 2 l8, the rocker arm length is 2.5 l8, the connecting rod length is 5 l8, and the rocker arm and connecting rod are joined at the middle point of the latter. Taking into account these lengths, the Chebyshev close loop kinematic chain provides (see Fig. 5):

2.5e

jϑ7

− 2.5e

jϑ9

−e

jϑ8

+ 2 = 0.

(3)

In Eqs. (3) to (5), both projections (vertical and horizontal) for each close loop equation are written in a compact form following the Euler’s formula, where j is the imaginary unit.

2. Pantograph chain (formed by links number 9, 7, 3, 6 and 13) In our model, the tendons length is 6 l8, whereas the distance between the connecting rod-femur and upper tendon-femur joints is 3 l8, and the distance between rocker arm-hip and upper tendon-hip joints (points B and D respectively) is 12 l8, so the pantograph close loop kinematic chain provides (see Fig. 6): 882

Fig. 5. Chevyshev chain (lengths in units of l8)

Fig. 6. Pantograph chain (lengths in units of l8) 3. Stability chain (formed by links number 8, 7, 10 and 13) In our model, the stabilizing link length is 4.2 l8.

Note that, in order to align the slot with the linear section of the Chebyshev trajectory, the vertical distance between the motor crank joint and the slot at the hip must be equal to 4 l8. Calling x the horizontal projection distance between the motor crank joint and the end of the stabilizing link, the stability close loop kinematic chain provides (see Fig. 7):

4.2e jϑ10 − 5e jϑ7 + e jϑ8 − x + 4 j = 0. (5)

As stated below, these equations determine the angles for all the links as functions of that for

Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 879-887

 2 2   ϑ = a cos  −4 ⋅ cos ϑ8 + 13 ⋅ cos ϑ8 − 10 + sin ϑ8 ⋅ −16 ⋅ cos ϑ8 − 60 ⋅ cos ϑ8 + 100   7   25 − 20 ⋅ cos ϑ8    , (6)  2 2    −4 ⋅ cos ϑ8 + 13 ⋅ cos ϑ8 − 10 − sin ϑ8 ⋅ −16 ⋅ cos ϑ8 − 60 ⋅ cos ϑ8 + 100  ϑ9 = a cos    −25 + 20 ⋅ cos ϑ8      2  ϑ = a cos  ( 2.5 ⋅ (cos ϑ7 + cos ϑ9 ) ⋅ ( −27 − A ) − 2.5 ⋅ (sin ϑ7 + sin ϑ9 − 12 ) ⋅ 144 ⋅ A − ( −27 − A )  6 12 ⋅ A      ( 2.5 ⋅ cos ϑ + cos ϑ ) ⋅ ( 27 − A ) + 2.5 ⋅ (sin ϑ + sin ϑ ) − 12 ) ⋅ 36 ⋅ A − ( 27 − A )2   7 9 7 9  ϑ3 = a cos  6⋅ A      

   

, (7)

where: A = ( 2.5 ⋅ (cos ϑ7 + cos ϑ9 ))2 + ( 2.5 ⋅ sin ϑ7 + sin ϑ9 ) − 12 )2 . the motor crank, ϑ8, which is also a function of time.

ϑ1 = ϑ5 = ϑ10 , ϑ12 = ϑ4 = ϑ3 , ϑ2 = ϑ13 = ϑ6 . (9) For the links belonging to the opposite leg, we apply a phase out of π radians on ϑ8:

Fig. 7. Stabilization chain (lengths in units of l8) Solving the Eq. system (3), the expressions (Eq. (6)) for the connecting rod and rocker arm angles are found.

From Eq. (4), the femur and tibia angles are found as functions of the previous ones. (See Eq. (7)).

Finally, the Eq. (5) gives the solution for the stabilizing angle:

 5 ⋅ sin ϑ7 − sin ϑ8 − 4  ϑ10 = asin   . (8) 4.2  

As can be seen in Fig. 2, the rest of the angles involved are identical to one of the given ones in the Eqs. (6) to (8), in particular:

ϑi' ( ϑ8 ) = ϑi ( ϑ8 + π ) . (10)

All these angles have been calculated using a reference system fixed at the hip, the x-axis direction being defined by the points A and B in Fig. 1. In order to apply the second Newton’s law, all the kinematic values must be referenced to an inertial system. An inertial system can be placed at the ground (or at the supporting foot, link number 1 in Fig. 2, as no relative motion between this link and the ground is considered). The corresponding base change is described in Eq. (11):

ϑiground = ϑihip − ϑ1hip , (11)

where ϑiground is the angle of the i-link related to the ground system, and ϑihip is the corresponding one related to the reference system fixed at the hip.

Once the angles are determined in the new reference system, the positions of the center of mass for all the links are easily obtained using trigonometric relations (for example, x2 = L2cosϑ2/2, y2 = L2sinϑ2/2; x3 = L2cosϑ2 + L3cosϑ3/2, y3 = L2sinϑ2 + L3sinϑ3/2; and so on). Then, by time differentiating once and twice, the angular velocity and acceleration respectively for any link,

Kinematics and Dynamics of the Quasi-Passive Biped “PASIBOT”

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as well as the linear velocity and acceleration of its center of mass are found.

In summary, the planar kinematics of PASIBOT is solved. As an application, all the kinematical values for one step of PASIBOT have been calculated, considering a motor crank constant angular velocity, ω8: ϑ8(t) = ω8·t . (12)

3 DYNAMICS OF PASIBOT The inputs for the dynamical problem are the previously calculated kinematic magnitudes for every link, that is to say, its angular acceleration, αi, and its center of mass acceleration, (aix, aiy). The dynamical magnitudes involved in the complete mechanical study of PASIBOT are the weight of every link, mig, the motor torque, T8, and all the forces between links, fji (exerted by link j on link i).

All those kinematical and dynamical magnitudes are presented in Fig. 8, for a general link i.

  f jix − fik x = mi aix  j <i k >i   ⇒ f jiy − fik y = mi g + mi aiy , (13)  j <i k >i  Ti + rijx f jiy − rij y f jix − rik x fik y − rik y fik x = I i αi k >i j <i  i = 2 ,3,...,13,1′, 2′,...7′,9′,...12′

∑

∑

∑(

∑

∑

) ∑(

)

Since there are 23 links (excluding the supporting foot) and there are three equations for each link, the system describing the dynamics of the whole mechanism consists of 69 linear equations. The linear equation system (Eq. (13)) is expressed in a matrix form (Eq. (14)), and then solved via matrix inversion, with a MATLAB® code.  f12 x      m2 a2 x   a11 a12 …   f12 y   a    m2 g + m2 a2 y  a … ,  21 22 ·  =  I2α2     T      8  

…      (14)       [ A( coefficient )][· F ( force )] = [ I ( inertia )] ⇓

[ F ] = [ A]−1·[ I ] .

4 NUMERICAL RESULTS

Fig. 8. Dynamical entities for a general link i For any link, i, the dynamical equations for the motion of the center of mass and for the rotation of the rigid body, using the action-reaction Newton’s law to reduce the number of unknown forces, are exposed in Eq. (13):   Fi = mi ai   i    fij = − f ji  ⇒

 Ton i = I i α   i

∑

∑

884

The kinematical and dynamical equations have been implemented in a MATLAB code in order to obtain solutions (position, velocities, acceleration, as well as forces and torques) depending of a set of parameters (link dimensions, masses, and densities, motor angular velocity) entered by the user. This code is being integrated into the movement control task. In Fig. 9, the flow chart of the developed algorithm is presented.

For every time step, t, the program first finds the corresponding value of ϑ8(t) and using Eqs. (6) to (11), it obtains those of the rest of the angles and the positions of the centres of mass. Then, using the same variables for t-1 and t-2, it calculates the corresponding angular and linear velocities and accelerations. These data that define the kinematic state of the biped at the time t, form the inertia matrix, [I] in Eq (14). Finally, the program inverts the coefficient matrix, [A], by means of a matrix

Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C.

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inversion subroutine and multiplies both matrices to provide the forces and torques between links at this time step. These values are stored to be plotted and the calculations restart for the next time step.

(see Fig. 2).

The code programmed has been validated by comparison with others simulation programs (Working model and ADAMS code). The main advantage of the developed program is that it let us perform fast modifications, making easier the final robot design by selecting new materials, choosing actuators and reduction devices, or even applying any type of optimization process.

As an example of the capacities of the program, some results are presented. Fig. 10 shows the actuator torque in the crank (link number 8) related to time, for different values of the motor angular velocity, ω8.

Fig. 9. Flow chart for the PASIBOT kinematics and dynamics calculus code As the first result, the program implemented in MATLAB® has helped us to choose the suitable actuators and transmission systems, as it provides the motor torque required to perform the prescribed movement. This torque actually depends on many parameters that can be easily changed in the program (the masses and sizes of every link independently, or their densities, the motor angular velocity function, among others). The process of choosing the actuator and/or transmission system is iterative: first, the maximum torque without additional weight is calculated, and a motor that provides higher torque, at the prescribed speed, is proposed. Then, the required torque with the additional definite weight of the proposed engine is recalculated, and if this torque exceeds what the proposed motor can provide, the immediately superior motor is proposed, the torque with the corresponding weight is recalculated, and so on; if not, the process ends and the proposed motor is chosen. This code will allow us to quickly obtain all the dynamic and kinematical parameters of interest for control and optimization tasks.

As an example, during construction of the first prototype, the links “femur”, “fibula” and “tibia” had to be doubled in order to increase their resistance. With the program, the calculation of the new torque required was immediate. In that case, the program helped us to design the reduction phase required to be coupled with the same motor

Fig. 10. Actuator torque for different crank velocities. T is the semi-period for the rotation of the motor crank, that is, the time for one PASIBOT’s step In Fig. 10, the same shape for each case can be appreciated apart from the torque obtained from the highest velocity value (ω8 = 5 rad/s). With this velocity; the dynamical effect of the acceleration (inertia force) becomes important.

Fig. 11. Actuator torque for different hip extra loads

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Another interesting result concerns the motor torque required when the robot load increases (when different actuators, transmissions, batteries, wires, etc. are included in the robot hip).

In Fig. 11 the crank torque is represented again, but for different loads (5, 10 and 15 kg) added to the hip.

The graph in Fig. 11 shows that the required motor torque depends slightly on the added load.

This is because the hip remains at almost the same level in a course of a step. Nevertheless, since the robot begins from the rest position, the differences between the required torques are significant only over the first period of the step. In fact, in the stationary walking state, practically all the motor torque is spent raising alternatively the flying leg, while the supporting leg sustains the rest of the weight -that of the hip, mainly- in such a way that the hip centre of mass moves almost horizontally at constant speed.

5 CONCLUSIONS In this article, a prototype of a quasipassive biped called PASIBOT has been presented.

Kinematical and dynamical basic expressions for a PASIBOT step have been obtained. A program code has been developed to get parametric solutions from these expressions. The program has been validated by comparison of the kinematical results with those provided by other commercial softwares. The developed code has been used to study the PASIBOT behavior before its construction, reducing the complexity in the design process; it is also to be used in control tasks for the real prototype walking. With respect to the presented numerical results, we can highlight the dependency of the load at the hip and the rotational input speed in the actuator torque, which let us to study the most suitable actuator for the movement requirements.

6 ACKNOWLEDGMENTS The authors wish to thank the Spanish Government for financing provided through the MCYT project DPI-2006-15443-C02-02, and the LARM laboratory, particularly professors Ceccarelli and Ottaviano for their suggestions in the design process of PASIBOT. 886

7 REFERENCES [1] MAQLAB, PASIBOT from: http://maqlab. uc3m.es/proyectos/proyectos.htm, accessed on 2010-02-15.

[2] Masato, H., Kennichi, O. (2007). Honda humanoid robots development. Philosophical Transactions of the Royal Society A, Series A, Mathematical, Physical and Engineering Science, vol. 365, no. 1850, p. 11-19.

[3] Akachi, K., Kaneko, K., Kanehira, N. (2005).

Development of humanoid robot HRP-3P. Proceedings 5th IEEE-RAS International Conference on Humanoid Robots, p. 50-55, DOI:10.1109/ICHR.2005.1573544.

[4] Ogura, Y., Aikawa, K., Shimomura, K., Lim, H., Takanishi, A. (2006). Development of a New Humanoid Robot WABIAN-2.

Proceedings IEEE International Conference on Robotics and Automation, p. 76-81.

[5] Hirose, M., Ogawa, K. (2007). Honda humanoid robots development. Philosophical Transactions of the Royal Society A-Mathematical, Physical and Engineering Sciences. vol. 365, no. 1850, p. 11-19.

[6] Collins, S., Ruina, A., Tedrake, R. (2005).

Efficient bipedal robots based on passivedynamic walkers. Science, vol. 307, no. 5712, p. 1082-1085.

[7] Collins, S.H., Ruina, A. (2005). A bipedal walking robot with efficient and humanlike gait. IEEE International Conference on Robotics and Automation (ICRA), p. 19831988.

[8] Tedrake, R., Zhang, T.W., Fong, M.F. (2004).

Actuating a simple 3D passive dynamic walker. IEEE International Conference on Robotics and Automation, p. 4656-4661.

[9] Wisse, M., Feliksdal, G., van Frankenhuyzen, J. (2007). Passive-based walking robot – Denise, a simple, efficient, and lightweight biped. IEEE Robotics & Automation Magazine, vol. 14, no. 2, p. 52-62, DOI:10.1109/MRA.2007.380639.

Ikemata, Y., Yasuhara, K,, Sano, A., [10] Fujimoto, H. (2008). A study of the legswing motion of passive walking. IEEE International Conference on Robotics and Automation (ICRA), p. 1588-1593, DOI:10.1109/ROBOT.2008.4543428.

Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 879-887

[11] Ikemata, Y., Sano, A., Fujimoto, H. (2006).

A physical principle of gait generation and its stabilization derived from mechanism of fixed point. Proceedings of the 2006 IEEE International Conference on Robotics and Automation, p. 836-841, DOI:10.1109/ ROBOT.2006.1641813.

[12] Erdman, A.G., Sandor, G.N. (1997).

Mechanism design: analysis and synthesis.

3rd ed., Prentice-Hall, New Jersey.

[13] Gu, H., Ceccarelli, M., Carbone, G. (2008). Design and operation of 1-DOF anthropomorphic arm for humanoid robots.

Proceedings of the 17th International Workshop on Robotics in Alpe-Adria-Danube Region RAAD08.

[14] Tavolieri, C., Ottaviano, E., Ceccarelli, M., Di Rienzo, A. (2006). Analysis and design

of a 1-DOF leg for walking machines.

Proceedings of RAAD’06, 15th International Workshop on Robotics in Alpe-Adria-Danube Region.

[15] Castejón, C., Carbone, G., García-Prada, J.C., Ceccarelli, M. (2010). A multiobjective optimization of a robotic arm for service tasks. Strojniški vestnik - Journal of Mechanical Engineering, vol. 56, no. 5, p.

316-329.

[16] Hu, Y., Nakamura, H., Takeda, Y., Higuchi, M., Sugimoto, K. (2007). Development of a power assist system of a walking chair based on human arm characteristics. Journal of Advanced Mechanical Design, Systems and Manufacturing, vol. 1, no. 1, p. 141-154, DOI:10.1299/jamdsm.1.141.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 888-897 DOI: 10.5545/sv-jme.2011.124

Paper received: 09.06.2011 Paper accepted: 30.09.2011

Finite Element Analysis of a Tire Steady Rolling on the Drum and Comparison with Experiment Korunović, N. – Trajanović, M. – Stojković, M. – Mišić, D. – Milovanović, J. Nikola Korunović* – Miroslav Trajanović – Miloš Stojković – Dragan Mišić – Jelena Milovanović University of Niš, Faculty of Mechanical Engineering, Serbia

A finite element (FE) model for analysis of tire rolling on the drum, based on a specially developed CAD model, is presented in the paper. All the changes performed on the geometry of CAD model are automatically propagated to FE model. This makes the FE model very suitable for parametric studies, which help tire designer to quickly find the optimal values of tire design parameters. In this way the tire design process is shortened and the quality of resulting tires improved. The results of finite element analyses conducted on the model have directly been compared to experimental ones, confirming model validity. Equipment and methods used for experimental determination of braking and cornering characteristics of the tire as well as for experimental determination of friction coefficient of tire tread have been shown. The difference between experimental and numerical results was smaller after the calibration of friction coefficient had been performed and in such a way a further improvement of the existing model was achieved. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: finite element analysis (FEA), tire design, steady-state rolling (SSR), cornering, braking, parametric 0 INTRODUCTION Due to efficiency in prediction of cornering and braking behavior of tires and moderate requests towards computing resources, steadystate rolling analysis (SSRA) using the finite element method, based on the mixed Eulerian / Lagrangian approach, is often deployed for tire rolling simulations. While purely Lagrangian approach may also be used to simulate the response of a steady rolling tire [1], it is not as efficient and accurate as the mixed one. Explicit FEA, which is a must for dynamic analysis where transients play a significant role, may also be used for tire analysis, but it is neither essential nor efficient for the analysis of a steady rolling tire.

While the results of SSRA and explicit analyses may be very similar and in good correlation with the experimental ones, SSRA generally takes a considerably less time to finish [2]. The reasons for such a difference have been found in the smaller mesh size and fewer analysis increments needed for SSRA. In SSRA, finite element (FE) mesh may be refined only in the vicinity of contact area, thanks to the nature of FE formulation.

SSRA of tires yields various kinds of results that may efficiently be used in tire design. Such results can be directly compared to 888

experimental ones, obtained by laboratory testing or testing on specialized vehicles. Typical results of SSRA include the following functions: driving force in terms of angular velocity, cornering force in terms of slip angle or self-aligning torque in term of slip angle. Tire construction parameters like belt width, belt angle or number of plies and operating parameters like inflation pressure, tire load or camber angle may be varied in order to reach the optimal design. Examples of SSRA of tires and parametric studies in which it has been used may be found in [3] to [5].

Finite element analysis (FEA) is generally performed on tires rolling over flat surface, in order to closely simulate their behavior during service. FEA results may directly be compared with experimental ones, obtained using flat bad or flat surface tire testing machines [6] and [1], or specialized vehicles. However, such equipment is space consuming and expensive and is generally hosted by large tire manufacturers or specialized laboratories. Characteristics of a rolling tire may also be evaluated on drum machines [1], [6] and [7] which require less space and are generally more affordable. When such machines are used, a certain level of approximation in footprint geometry is introduced, which becomes more significant as the drum diameter gets smaller. It is

*Corr. Author’s Address: University of Niš, Faculty of Mechanical Engineering, Aleksandra Medvedeva 14, 18000 Niš, Serbia, nikola.korunovic@masfak.ni.ac.rs

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also difficult to apply a realistic road surface to the drum, because of its curvature. Nevertheless, it is assumed that if tire rolling behavior is successfully modeled and verified on one kind of surface for a range of operating parameters, it may also be used to predict its behavior in different road conditions [1]. In order to compare the results of tire FEA with drum machine testing results, the finite element model has to be modified, to include the definitions of drum surface and drum rotation.

In SSRA of tires special attention should be given to friction modeling, as it has a significant influence on the accuracy of the results. The nature of rubber friction has still not been properly explored. In practice, phenomenological friction models are often used, where friction coefficient of the Coulomb model is defined as a function of sliding velocity, contact pressure and (optionally) of temperature [8]. These models rely on experimental data, obtained using various kinds of equipment intended for sliding of small rubber specimens over artificial road surface [8] to [10].

Previous work of the authors, described in [11] to [14], covers the areas of parametric and knowledge-based tire design, rubber modeling for tire FEA, static analysis and SSRA of tires.

This paper describes the use of the methodology presented in [14] for an analysis of the tire rolling on the drum. The purpose of the analysis was to validate the finite element model by comparing FEA results with experimental ones, which were obtained using test rig described in [7] and [10].

A short description of finite element models is given first, followed by a description of equipment and procedures used for braking and cornering tests as well as for determination of tire tread friction coefficient. Then, the results of SSRA, a comparison with the experiment, calibration of friction coefficient and discussion are given, followed by concluding remarks.

1 FINITE ELEMENT MODELS Finite element tire models described in this paper are highly similar to the ones depicted in [14]. The main feature of new models is that the rolling surface is not flat but cylindrical, so that rolling on a drum may be simulated. Hence, only the main features of FE models will be described

here, especially their differences compared to [14].

Two FE models are used for the analyses, which are performed in FE code ABAQUS according to algorithm given in [14]. An axisymmetric model is used mainly as a starting point for 3D model creation and for tire inflation analysis, while 3D model is used for inflation, static footprint, acceleration-to-braking and cornering analyses. The models are based on the CAD model developed by the authors (Fig. 1). In adition to basic geometry of tire profile, the CAD model contains a parameterized network of lines and points. Following dimensional changes of tire profile and its structural components, lines and points allocate accordingly, to form the basis for mapped FE mesh. Axisymmetric FE model needs to be created in FE preprocessor only once, before the first of the analyses, using geometry entities exported from the CAD model as described in [14]. For any subsequent design change, the following procedure is deployed: design parameters are changed inside the CAD model; the group of newly positioned points is exported to a neutral format and their spatial position is mapped to the position of nodes of FE model.

Mapping and translation of points to nodes is done using a translation program, which is especially written for this purpose. In such a way, it becomes possible to explore a large number of design variations in a very short time, i.e. to conduct parametric studies of tire design. FE model described here may be used for all tire types based on the profile that is constructed in the same or similar manner. If an essentially different type of tire is to be analyzed, a new CAD model needs to be created. This process is a relatively simple one, based on a predefined procedure.

Finite elements of axisymmetric model are grouped accordingly in order to represent the purely rubber components of tire, like sidewall, tread, or bead filler. Composite structural components of the tire, carcass and belts, are created by embedding of surface elements in volume ones. Inside of surface elements rebar layers are defined, which represent steel or rayon cords. The definition of one rebar layer contains a cord area, the distance between the cords, a cord angle and cord material. Bead wire is modeled as isotropic material and the rim is substituted by

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rigid supports. Rubber components are described using Mooney-Rivlin form, where material coefficients are taken from earlier FE models [12].

Steel and rayon have been modeled as linearly elastic materials.

Fig. 1. 2D parametric CAD model of an existing 165/70 R13 tire

21712 nodes and 18538 elements. The numbers of active degrees of freedom are 934 and 63480, respectively.

The drum is modeled as a rigid surface (Fig. 3) and friction between the drum and tread surface is defined using Coulumb model with viscous stick formulation. The variable coefficient of friction is obtained by testing of rubber specimens on Mini–μ–road [10], as described in the next chapter.

Using the 3D FE model, the following analyses have been conducted: inflation analysis, static footprint analysis, straight line rolling under the action of driving or breaking torque, straight line rolling in fine increments to find the angular velocity of free rolling and cornering analysis at free-rolling.

Types of finite elements that have been used include axisymmetric hybrid elements with twist for purely rubber structural components, axisymmetric surface elements with a twist for modeling of carcass and belts and axisymmetric solid elements for modeling of the bead wire.

3D FE tire model (Fig. 2) is obtained by rotation of axisymmetric model around tire axis.

Dense mesh is created only in the vicinity of tire footprint, as described in [14].

Fig. 2. 3D tire model 3D finite elements, which represent the equivalents to the corresponding axisymmetric ones, are created during the rotation of axisymmetric model. In this way, the 3D FE model is composed of: 8 or 6-noded hybrid elements that represent rubber, 4-noded surface elements for modeling of carcass and 8-noded solid elements used to model the bead wire.

Axisymmetric model contains 473 nodes and 403 elements, while the 3D model contains 890

Fig. 3. 3D tire model for FEA of the tire rolling on the drum; diameter of cylindrical surface is equal to diameter of the drum 2 EXPERIMENTAL SET-UP AND PROCEDURES FOR TIRE BRAKING AND CORNERING AND RUBBER FRICTION TESTING 2.1 Tire Testing Rig In order to test tire behavior during braking and cornering, dynamic tire force measuring device [7] (Fig. 4), set up at Helsinki University of Technology, Laboratory of Automotive

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Engineering, has been used. The diameter of machine’s drum, which is equipped with the dynamometer, equals to 1219 mm. Maximal circumferential speed at which reasonable results may be obtained is around 150 km/h. The drum is coated with a layer of safety walk paper. Wheel load can be adjusted up to the maximal value of 5000 N. The slip angle can be adjusted during measuring process, in the range of –10 to +10°.

Camber angle may vary from –4 to +4°. All six forces and moments acting between the tire and the drum, as defined in [6] and [14], may be measured, as well as the rotational speed of the tire. The tire, photographed while one of the cornering tests was performed, is shown in Fig. 5.

1. Braking tests: • Initial speed: 10 km/h, wheel loads: 2845 N, 3580 N and 4287 N. • Initial speed: 50 km/h, wheel loads: 2845 N, 3580 N and 4287 N. 2. Cornering tests, slip angle varying from –10 to +10°. Inflation pressure has in turn been set to 0.12, 0.15, 0.2 and 0.25 MPa, wheel load to 2845, 3580 and 4287 N, drum speed to 10, 50 and 80 km/h and camber angle to 0, 2 and 4 degrees, thus some 20 different combinations of tire operating parameters have been used.

The operating parameters for cornering tests were chosen appropriately, in order to analyze the influence of wheel load, circumferential drum speed, inflation pressure and camber angle to cornering (side) force and self-aligning torque of free-rolling cornering tire.

Quantities that are directly output by testing rig are: time, t [s], slip angle, α [deg], camber angle, γ [deg], dynamic tire radius, rd [mm], temperature T [°C], circumferential drum speed v [km/h], frequency of wheel rotation, N [1/s], longitudinal force, Fx [N], lateral force, Fy [N], wheel load, Fz [N], overturning moment, Mx [Nm], rolling resistance moment, My [Nm] and self-aligning moment, Mz [Nm].

Quantities that were calculated in order to compare test results to FEA results are: • angular velocity of the tire, ω [rad/s]: ω = 2πN , (1)

Fig. 4. Test rig used for tire cornering experiments

Fig. 5. Tire and drum during cornering test Two different kinds of tests, with variable operating parameters, have been performed:

•

effective (rolling) radius, re [mm]:

re = (277.78 · v) / ω , (2) •

slip:

slip = 1 – ( ( 277.78 · v ) / ( rd · ω )) . (3)

Test cycles have been adjusted in order to get optimal accuracy of output quantities. An optimal test cycle of a cornering test is shown in Fig. 6.

Cornering test cycle consists of the following intervals: time needed for stabilization of tire rotation, slip angle change from 0 to +10°, slip angle change from +10 to ‒10°, slip angle change from –10 to 0° and the time needed to stop the rotation. Depending on the ratio of slip angle change, hysteretic effect is more or less visible on the experimental curves that represent the relation

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between slip angle and side force or self-aligning torque. Since experimental results were meant to be compared to the results of quasistatic FEA, the rate of slip angle change has been chosen such as to minimize hysteretic effect (Fig. 7). The vertical force of 3580 N represents nominal load, force of 4287 N is equal to the maximum allowable load, while the force of 2845 N simulates the load less than nominal. [deg] v [km/h] / 10 Fy [kN] Mz [Nm] / 50

10 5

Rubber specimens are usually square-shaped (60 × 60 mm), but tread blocks, cut from the tire, may also be used.

In this study, tread blocks, cut from another tire, were used for determining the friction coefficient. The other tire was exactly the same as the one that was used on tire testing machine, considering size, type and production date. In such a way, the influence of specimen geometry to test results was meant to be diminished. Double specimens were used to enlarge contact surface and enable testing at lower pressures. On the other hand, to enable testing at higher pressures, single specimens were used (Fig. 9).

0 -5 -10 0

20

40

60

t [s]

Fig. 6. Cornering test cycle of free-rolling cornering test at circumferential drum speed v = 50 km/h and vertical load Fz = 3580 N p = 0.2 MPa, v = 50 km/h Fz = 2845 N Fz = 3580 N Fz = 4287 N

5000 4000 3000

FY [N]

2000

Fig. 8. Mini–μ–road, linear friction tester that enables sliding of rubber specimens over various kinds of surfaces

1000 0

-1000 -2000 -3000 -4000 -5000 -10

-8

-6

-4

-2

0

α [°]

2

4

6

8

10

Fig. 7. Experimentally obtained dependency of cornering force on slip angle, at three different values of vertical load 2.2 Rubber Friction Testing Rig In order to find the values of the friction coefficient of tire tread as a function of sliding speed and contact pressure, rubber specimens have been tested using Mini–μ–road [10] (Fig. 8). The testing rig is situated in the cold chamber, where different kinds of arctic weather conditions and sliding on ice may optionally be simulated.

892

Fig. 9. Single and double specimens have both been used for determination of tire tread friction coefficient At the beginning of each test cycle, rubber specimen is pressed to the surface (Fig. 10).

Pneumatic cylinder that produces normal force as well as force transducer that measures tangential force may be seen in the picture. The required normal force is then calculated in such a way that, when divided by area of specimen contact surface, it produces the desired value of mean contact pressure. While the specimen slides along the surface, at predefined sliding speed, tangential

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force is recorded. Coefficient of friction is then obtained by dividing the mean value of tangential force by the value of normal force.

A measurement has been done in total of 63 cycles, obtained as combinations of 7 different contact pressures and 9 different sliding speeds. Values of friction coefficient obtained by experiment correspond to points shown in Fig. 11.

The figure also contains approximating surface, based on rational Chebyshev polynomials. This surface was fitted to experimental points to represent the friction coefficient in given domain of sliding speeds and contact pressures. Based on the surface, a table that contained the values of friction coefficient at equally spaced intervals of sliding speeds and contact pressures was extracted. This table was used to provide input values of tire tread friction within finite element model of the tire.

Fig. 10. Close photo of rubber specimen being mounted on the machine and pressed against the surface

Fig. 11. Friction coefficient of tread rubber (μ), as a function of sliding speed (v) and contact pressure (p) 3 RESULTS AND DISCUSSION The procedure for SSRA of the tire rolling on flat surface has been described in detail in [14].

Model changes, i.e. introduction of a more realistic representation of friction and the replacement of the flat surface with cylindrical one, have contributed to a considerable increase of resulting deformations, strains, stresses and reaction forces comparing to those shown in [14].

Fig. 12 shows FE tire model during simulation of braking.

Fig. 12. Deformed shape of the tire model and contact pressure distribution at the footprint during braking analysis Experimental results related to braking are of lower quality than those related to cornering, as additional sensors of lower resolution had to be used for this kind of test. However, when the experimental parameters were selected so as to minimize measurement errors, the data that could be used for comparison with numerical results, as shown in Fig. 13 were obtained. Lower resolution and accuracy of the measuring device, as well as insufficient strength of drum brake, cause a significant dissipation of experimental data.

Much more attention has been paid to cornering experiments. In this case, experimental results of a considerably better quality than in the case of braking were obtained, as force and moment sensors were not changed. As a substantial number of operating parameters has been varied during cornering experiments, the complete set of experimental results is quite large.

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Thus, only a chosen subset of those will be shown and compared to results of SSRA.

Fx [N]

1600

800

0

Braking from 10 km/h, Fz = 1367 N Experiment FEA -12

-6

improvement in accuracy of numerical results was achieved by calibration of friction coefficient, as shown at the end of the paper.

The intensity of residual side force (ply steer) obtained numerically is overestimated, which is considered to be the consequence of insufficient density of belt mesh and the absence of detailed tread on finite element model. It is well known that the tread is usually designed to diminish the ply steer effect. The difference between experimental and numerical results also gets larger for higher values of slip angles.

0

Slip [%]

Fig. 13. The comparison of experimentally and numerically obtained correlation of braking force and slip ratio Fig. 14 shows the deformed shape of tire model during cornering analysis. Very large deformations occur during the analysis, especially at larger values of slip angle. For this reason the FE model should be carefully created. This primarily refers to regularity and size of finite element mesh, choice of material model for all structural components, especially tread and carcass, and the definition of friction between the tread and the ground.

Fig. 14. Detail of deformed tire model during cornering analysis, at slip angle of 10° Fig. 15 shows contact pressure distribution at the footprint, at different stages of cornering analysis.

Having in mind all the approximations which were necessarily introduced into the analyses, the correlation between experimental and numerical results is considered to be good (Figs. 16 and 17), especially in the case of the right turn (left side of the graphs). Nevertheless, further 894

Fig. 15. Footprint shape and contact pressure distribution for tire rolling at 50 km/h, under nominal vertical load of 3580 N, inflation pressure of 0.2 MPa, camber of 0° and slip angles of 0, 3, 6 and 10° Fig. 18 shows the comparison of slip angle – side force curves obtained numerically for three examined values of vertical load. Their correlation is very similar to the correlation of experimental ones, given in Fig. 7.

Previous figures show the influence of the vertical force on cornering characteristics of the

Korunović, N. – Trajanović, M. – Stojković, M. – Mišić, D. – Milovanović, J.

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tire. Experimental and numerical procedures have also been used for examination of the influence of rolling speed inflation pressure and inclination angle to cornering characteristics. p = 0.2 MPa, v = 10 km/h Fz = 3580 N Experiment FEA FEA, µ calibrated

4000 3000

FY [N]

2000 1000 0

to experimental friction data were used as an alternative to original one, but no significant change in numerical results was noticed. Then, the experiment used to obtain friction data was simulated using FEA (Fig. 19), in order to determine how the nature of the test itself influences the resulting values of friction coefficient. The data on friction coefficient obtained by surface fitting to experimental points, as described earlier, were used as an input.

-1000 -2000

4000

-3000

3000

-4000 -8

-6

-4

-2

0

α [deg]

2

4

6

8

2000

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Fig. 16. Comparison of experimental and numerical slip angle – side force curves for the tire inflated to 0.2 MPa rolling at 10 km/h at vertical load of 3580 N

FY [N]

-10

p =0.2 MPa, v = 50 km/h FZ = 2845 N FZ = 3580 N FZ = 4287 N

1000 0

-1000 -2000 -3000 -4000 -10

p = 0.2 MPa, v = 50 km/h Fz = 2845 N Experiment FEA

4000 3000

FY [N]

2000 1000 0

-8

-6

-4

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0

α [deg]

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Fig. 18. Comparison of numerical slip angle – side force curves for the tire inflated to 0.2 MPa rolling at 50 km/h at three different values of vertical load

-1000 -2000 -3000 -4000 -10

-8

-6

-4

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0

α [deg]

2

4

6

8

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Fig. 17. Comparison of experimental and numerical slip angle – side force curves for the tire inflated to 0.2 MPa rolling at 50 km/h at vertical load of 2845 N The comparison between numerical and experimental results showed that the intensity of numerically determined forces always had the tendency to be lower, the difference being between 10 and 15%. The difference was also becoming more significant at higher values of the slip angle.

Thus, the room for improvement in numerical results was spotted.

At higher slip angles, a large percentage of tread surface is slipping and thus the friction at the footprint plays the most important role in generation of tire forces and moments. For that reason, the approach to friction modeling was once more reconsidered. At first, different fits

Fig. 19. FE model for simulation of experimental determination of tire tread friction coefficient The simulations showed that there exists a large inhomogenity of contact pressure at specimen surface and that the pressure is much higher at leading edge of the specimen (Fig. 20). This phenomenon is considered to be the cause of noticeable differences in actual and predicted coefficient of friction. Thus, a calibrating function was introduced, by which the originally fitted friction surface was multiplied in order to obtain the corrected one, similar to the approach described in [15]. Fig. 21 illustrates this procedure at mean contact pressure of 0.11 MPa. When the corrected coefficient of friction (curve Fit corrected) was used in the simulation of the experiment, the results were obtained that

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 888-897

correspond to experimental ones much better (curve FEA-corrected).

coefficient has been used in a couple of repeated tire cornering simulations. The results of those are shown in Figs. 22 and 23. When the calibrated results were used, the difference between experimental and numerical results was noticeably smaller, reaching a maximum of 5%. p = 0.2 MPa, v = 50 km/h Fz = 2845 N Experiment FEA FEA, µ calibrated

4000 3000

Fig. 20. Contact pressure distribution at the surface of double tire tread specimen sliding on flat surface obtained by FEA, at mean contact pressure of 0.11 MPa and sliding speed of 10 mm/s

1.65 1.60 1.55 1.50

p = 0.11 MPa Experiment Fit FEA Fit corrected FEA corrected

µ

1.40

0

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α [deg]

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4 CONCLUDING REMARKS

1.35 1.30 1.25 1.20 10

100

1000

v [mm/s]

Fig. 21. Calibration of friction coefficient, shown at contact pressure of 0.11 MPa p = 0.2 MPa, v = 10 km/h Fz = 3580 N Experiment FEA FEA, µ calibrated

4000 3000 2000

FY [N]

1000

Fig. 23. Comparison of experimental and numerical slip angle - side force curves for the tire inflated to 0.2 MPa rolling at 50 km/h at vertical load of 2845 N, calibrated vs. noncalibrated coefficient of friction

1.45

1000 0

-1000 -2000 -3000 -4000 -10

-8

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α [deg]

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Fig. 22. Comparison of experimental and numerical slip angle – side force curves for the tire inflated to 0.2 MPa rolling at 10 km/h at vertical load of 3580 N, calibrated vs. noncalibrated coefficient of friction In order to test the effectiveness of the calibration process, the calibrated friction 896

FY [N]

2000

The finite element model for FEA of tire rolling on the drum has been presented, which demonstrates the flexibility of CAD based meshing approach introduced by the authors. The results of the analyses conducted on the model have successfully been compared to experimental ones, confirming FE model validity. The differences between experimental and numerical results were decreased after the calibration of friction coefficient had been performed. In such a way the further improvement of the existing model was achieved.

The presented FE tire model and associated analyses are used for performing parametric studies within the tire design process, helping the tire designer to quickly find the optimal values of tire design parameters. The tire design process is thus shortened and at the same time greater predictability and improvement of tire performance are achieved.

In order to further improve the existing model, future activities are planned. On the one hand, the work on the improvement in modeling of tire tread friction will be undertaken. On

Korunović, N. – Trajanović, M. – Stojković, M. – Mišić, D. – Milovanović, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 888-897

the other, the FE tire model will be modified to include the representation of detailed tire tread, by enhancing the CAD based meshing approach to 3D geometry.

5 AKNOWLEDGMENT The authors would like to express their gratitude to Mr. Panu Sainio and Research Group for Vehicle Engineering, School of Engineering, AALTO-University (formerly Helsinki University of Technology, TKK), for their invaluable help in tire and friction testing.

6 REFERENCES [1] Tönük, E., Ünlüsoy, S. (2001). Prediction of automobile tire cornering force characteristics by finite element modeling and analysis.

Computers and Structures, vol. 79, no.

13, p. 1219-1232, DOI:10.1016/S00457949(01)00022-0.

[2] Kabe, K., Koishi, M. (2000). Tire cornering simulation using finite element analysis.

Journal of Applied Polymer Science, vol. 78, no. 8, p. 1566-1572, DOI:10.1002/10974628(20001121)78:8<1566::AID-APP140> 3.0.CO;2-I. [3] Ghoreishy, M.H.R. (2006). Steady state rolling analysis of radial tire: comparison with experimental results. Proceedings of the Institution of Mechanical Engineers, Part D: Journal of Automobile Engineering, vol. 220, no. 6, p. 713-721, DOI:10.1243/09544070JAUTO268.

[4] Olatunbosun, O.A., Bolarinwa, O. (2004).

FE simulation of the effect of tire design parameters on lateral forces and moments.

Tire Science and Technology, TSTCA, vol. 32, no. 3, p. 146-163, DOI:10.2346/1.2186779.

[5] Dorsch, V., Becker, A., Vossen, L. (2002).

Enhanced rubber friction model for finite element simulations of rolling tyres. Plastics, Rubber and Composites, vol. 31, no. 10, p. 458464, DOI:10.1179/146580102225006486.

[6] Pottinger, M.G. (2005). Forces and Moments.

Gent, A.N., Walter, J.D. (Eds.), The pneumatic tire. National Highway Traffic

Safety Administration, U.S. department of Transportation, Washington D.C., p. 286-363.

[7] Helsinki University of Technology, Laboratory of Automotive Engineering, Dynamic tyre force measuring device, from http://edp.tkk.fi/en/research/research_ facilities/tkk_dynamic_tyre_force_measuring_ device.pdf, accessed on 2009-09-05.

[8] Guo, K.H., Zhuang, Y., Chen, S.K., Willlam L. (2006). Experimental research on friction of vehicle tire rubber. Frontiers of Mechanical Engineering in China, vol. 1, no. 1, p. 14-20, DOI:10.1007/s11465-005-0001-z.

[9] Božić, A., Petrović, I., Matuško, J. (2009).

Experimental investigations and modeling of the rubber-asphalt sliding pair dynamics.

Strojniški vestnik – Journal of Mechanical Engineering, vol. 55, no. 5, p. 303-318.

[10] Helsinki University of Technology, Laboratory of Automotive Engineering, Mini–μ–road, Rubber Friction Measuring Device, from http://edp.tkk.fi/en/research /research_facilities/tkk_mini-m-road.pdf, accessed on 2009-09-05.

[11] Stojković, M., Manić, M., Trajanović, M. (2005).

Knowledge-embedded template concept. CIRP – Journal of Manufacturing Systems, vol. 34, no. 1, p. 71-79.

[12] Korunović, N., Trajanović, M., Manić, D., Manić, M. (2004). Rubber modeling for tire FEA. World of Polymers, vol. 7, no. 3, p. 8594. (In Serbian) [13] Korunović, N., Trajanović, M., Stojković, M. (2007). FEA of tires subjected to static loading. Journal of Serbian Society for Computational Mechanics, vol. 1, no. 1, p.

87-98.

[14] Korunović, N., Trajanović, M., Stojković, M. (2008). Finite element model for steadystate rolling tire analysis. Journal of Serbian Society for Computational Mechanics, vol. 2, no. 1, p. 63-79.

[15] Huemer, T., Liu, W.N., Eberhardsteiner, J., Mang, H.A. (2001). A 3D finite element formulation describing the frictional behavior of rubber on ice and concrete surfaces.

Engineering Computations, vol. 18, no. 3-4, p.

417-437, DOI:10.1108/02644400110387109.

Finite Element Analysis of a Tire Steady Rolling on the Drum and Comparison with Experiment

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 898-903 DOI:10.5545/sv-jme.2010.078

Paper received: 13.04.2010 Paper accepted: 07.11.2011

Influence of the Milling Strategy on the Durability of Forging Tools Pahole, I. ‒ Studenčnik, D. ‒ Gotlih, K. ‒ Ficko, M. ‒ Balič, J. Ivo Pahole1* ‒ Dejan Studenčnik2 ‒ Karl Gotlih1 ‒ Mirko Ficko1 ‒ Jože Balič1 1Faculty of Mechanical Engineering, University of Maribor, Slovenia 2Unior Nigbo Forging Co. LTD, China

The quality of a tool’s surface has a direct influence on the number of well-produced parts. For the machining of an active tool surface, two technological processes are used: electrical discharge machining and high-speed milling. These two processes are used when machining new tools and for the repairing of used forging tools. In both cases, the material has already been thermally treated, so it has to be used for hard milling. Practical experience shows that the milling strategy has a big influence on the durability of a forging tool. This paper shows the influence of the CNC machining direction during high-speed milling on the durability of the engraving within the forging tool. In some cases the correct milling strategy can increase the durability of the forging tool by about one third. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: forging tools, surface quality, high speed cutting, CNC milling 0 INTRODUCTION Hot or cold-forging using forging tools enables us to approach to the product using one or more forging-operations. The engraving within the forging-tool is approximately a negative shape of the product although it is never completely equal to the negative of the product itself. It depends on the forging technology. The forming technology is directly responsible for the durability of the forging tool when forging.

It is important to manage those factors that most influence the durability of the forming-tool.

When manufacturing active surfaces, one of the more important influences is the roughness of the engraving surface. The requirements for the roughnesses of surfaces when forging are N6. It has been shown, however, that – in addition to the conditionally required measured roughness – it is also necessary also to take into consideration the direction and the type of final milling regarding the engraving with CNC, respectively HSC (highspeed cutting).

This paper shows the two flow-line milling with HSC of a forging tool for forging on a pneumatic hammer. It has been shown the flowline milling is more appropriate for the engraving in comparison to closed/open line milling, and enables a higher duration of engraving from the forging-tool. Flow-line milling also has 898

advantages within the smaller dynamical loads on the forging machine, which influences the smaller roughnesses of the engraving surfaces [1].

The milling strategy of the forging tool active surfaces depends on the forged product.

Certain guidelines are also given, which enable the user to influence the durability of the forging tool using appropriate technological operations at production or renewal of forging tools. An increase in the durability of forging-tools is one of the main research directions within forging.

Fig. 1. Example of forging tool 1 DEPENDENCY OF MILLING STRATEGY AND TECHNOLOGICAL DEMANDS Modern CAM software for the programming of CNC milling-tools provides the possibility of producing various toolpaths for the tool, all leading to similar results. The characteristic shape of the toolpath is described

*Corr. Author’s Address: University of Maribor, Faculty of Mechanical Engineering, Smetanova ulica 17, Maribor, Slovenia, ivo.pahole@uni-mb.si

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 898-903

by the term milling strategy. When working on conventional machines, this strategy did not normally take precedence because, in most cases, there were very simple possibilities for control.

Nowadays, this is completely different on CNC machines because the controls enable various possibilities for tool-control within the workspace [2].

Whilst for rough-milling the success is measured by the amount of removed material within a time unit, for surface-milling the quality of the surface is the most important factor when measuring the effectiveness of the process. In rough-milling there are very few different shapes of toolpaths, because the milling strategy is determined by the cutting-parameters (velocity, feed velocity, cutting depth, step-over), which are manly dependent on the used tools, and the machine tool. During surface-milling, the toolpath is not as strongly determined regarding the capability of the machine, and the cutting tool, as it depends more on the surface quality. In addition to the quality of the surface as measured by the roughness, there are other surface properties that cannot only be described by the measurement of quality or roughness. These surface-properties are a consequence of the milling-strategy. One of these properties is the direction of the roughness, respectively the orientation of the scallop. The direction of the scallop depends on the direction of the trajectories (Fig. 2), while the height and shape of scallop depend on the shape of the milling tool, and the technological parameters.

The scallop arises, not just as a result of the moving tool, but also as a result of the feed-movement, and the low number of cutting edges [3] and [4]. The primary scallop follows

the direction of the tool, the secondary scallop is perpendicular to the direction of the tool’s movement [5]. In the case of a high-velocity of feed movement and low rotational speed the secondary scallop can exceed the height of the primary crests [3]. Special care must be taken concerning these crests, especially during HSC milling. Generally, the surface roughness during CNC milling, depends on [6]: • The depth of cut. The depth of cut influences surface quality in an indirect way. Increasing the depth of cut increases the cutting resistance and the amplitudes of any vibrations. The cutting temperature also rises. Therefore, it is expected that the surface quality will deteriorate.

• The feed-rate per cutter tooth. Experiments show that, as the feed-rate increases the surface roughness also increases [7]. In any case, using feed-rates under a certain limit does not yield any substantial improvement in surface quality.

• The cutting speed. An increase of cuttingspeed generally improves surface quality.

• Stepover of tool.

• The cutting-tool wear. The irregularities of the cutting-edge due to wear are reproduced on the machined surface. Apart from this, as tool-wear increases, other dynamic phenomena such as excessive vibrations will occur, further deteriorating the surface quality.

• The use of cutting-fluid. The cutting-fluid is generally advantageous in regard to surface roughness because it affects the cuttingprocess in three different ways.

Fig. 2. Common used strategies for surface treatment Influence of the Milling Strategy on the Durability of Forging Tools

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 898-903

Fig. 3. Directions of material flow and the direction of feed Based on their experience, engineers can identify those successful strategies for individual shapes or characteristics of the surface, from the point of view of surface quality.

The more frequently used strategies for the treatment of free-surfaces are those strategies regarding surface treatment at z-levels and the strategy of parallel-passing, as in Fig. 2. The strategy of treatment at z-levels has, from the point of view of control and cutting, many advantages because it proceeds along the surface parallel to the x-y working surface. It is best suited for milling surfaces or parts of steep surfaces.

The strategy of the treatment using parallelpassing is the projection of the trajectory on to the x-y surface in the shape of a parallel-passing.

Although this passing is not defined by the inclination of the surface, the used toolpath lies in the direction of the surface’s inclination. The tool is moving “up and down” on the inclination.

In most cases the milling goes in the downward direction of the inclination. This strategy is the best for the treatment of more gentle surfaces.

Although the more often used milling strategies are used as one for the steeper and another for gentler slopes, they can also be used for different shapes of surfaces. The two previously-mentioned strategies can also be used when it is technically and technologically better to mill the surface in different directions, as in the case of the material flowing within the forming tool, as shown in Fig. 5. In addition to these two strategies, there are many different strategies that can be used when milling for special requirements.

Fig. 3 shows the use of directional material flow during a forging process for determining the direction of milling and the forming of a primary scallop.

900

The quality of the treated-surface, correspondingly the height of the scallop, can mostly be managed with cutting parameters, and the parameters of the strategy. The cutting parameters are set by the producers of the tools, the carbide inserts, and from experience. From a technological point of view it is best, if the CAM software supports the programming with respect to the heights of the deviations from the desired surface (the scallop height). The scallop height is defined for round milling tools and plane surfaces by:

h = R−

( 2 R )2 − p 2 , (1) 4

where R is the radius of the milling tool and p is the distance of passing the tool.

Generally, the surface, which is parallel to the desired surface and touches the peaks of the crests, [4], is defined by (Fig. 4):

PS,C(u,v)=P(u,v)+nh ,

(2)

where P(u,v) is the desired shape, n is the normal vector of the surface and h is the height of deviations (crests).

The surface treatment of a forging tool surface can be different according to the type of treatment. It is important to focus on the final treatment of the engraving. With respect to the surface milling when milling, we differentiate between: a) line-milling, b) circular-milling, c) closed-line milling, and d) open-line milling.

In most cases, circular-treatment is used for the milling of round, oval, and spherical shapes. The flat parts of the engraving are most

Pahole, I. ‒ Studenčnik, D. ‒ Gotlih, K. ‒ Ficko, M. ‒ Balič, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 898-903

often treated with line-milling, which can be open or closed [8].

shapes of the cutting tools, analysis of forces at cutting, analysis of the shapes of the cut-off material).

Fig. 4. Description of the surface and the biggest deviations According to the forging direction line milling can be oriented at different angles. During the process of forging, the biggest material flow is perpendicular to the forging direction. It is reasonable that the direction of the milling is oriented in the direction of the main material flow within the tool. The leftovers/scallop of the final treatment does not hinder the materialflow. Consequently, the wear is reduced. The real influence of the final direction treatment’s on the durability of the engraving can be stated: • as the measurement of the engraving’s surface roughness, and • as the tracking of the engraving’s durability during the forging process.

The direction of the material-flow within forging tool (Fig. 5) can be determined on the basis of the process’s technological plan for a particular product. The experiences of the tooldesigners and engineers play an important role [9]. Computer programmes are a big support when designing forging tools on the basis of the finite elements method.

These are usually created computer programmes. The basic efficiencies of the software (DEFORM 3D) are: – simulation of deformations and the heat-transfer for cold and hot forming, – material flow analysis, filling of the engraving, – the forming-forces, stresses on the tool, shape of the bean, diffusion process, – the hardness distribution, – analysis of the sintering, – analysis of the cutting process (wear of tools, optimization of cutting parameters and the

Fig. 5. Direction vectors of material flow in the forging tool

Fig. 6. Flow-chart of forging-tool production It is possible to define the directions of material flow within the forging-tool using simulations of the forging process. The sodefined direction of the material flow is the basis for defining the strategy of milling during the treatment. Fig. 6 shows the course of a basic design, the construction and the production of a forging tool [10].

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2 TECHNOLOGICAL PARAMETERS OF THE FORGING PROCESS The theoretical dependencies are quite well-known for the forging process. Therefore, it is logical to expect the same response for repeated conditions.

The coincidence of gluing the part to the tool can appear along with thickness changes and cracks appear. It can be concluded that noninfluential factors are known. In spite of wellknown accesses to the technology of forging, the experiences of the people who work in this field play an important role.

For forging parameters the interdependency of the parameters is important. The connections may be found within the parameters. Therefore, for any change of thickness ds on the press holds:

There are some estimations that there are at least ten process parameters that must be controlled for mastering the process.

3 RESULTS AND CONCLUSION An experiment was prepared for the forging tool of an equal part. The surface milling of the tool surface was carried out using the following strategies: • line-milling under an angle of 45° (example 1), Fig. 8, and • closed-line milling (example 2), Fig. 9.

ds = 1 / ( k · dF) , (3)

where ds is a change of the thickness, dF change of the forming force and k the stiffness of the press.

The reasons for changing the thickness and the process of changing, can be split (Fig. 7): • direct influences (technological influences of the first order), and • indirect influences (physical process quantities of the second and higher orders).

Fig. 8. The engraving is treated with linetreatment under an angle of 45˚ [1] (example 1) Both tools were used for forging, with the same forging parameters. The results of the experiment are shown in Table 1. The results of this experiment have shown that line-milling gives better results. The change in milling strategy resulted in an approximately 30% longer duration of the forging tool.

Fig. 7. Influence of the forming force on the thickness of the part 902

Pahole, I. ‒ Studenčnik, D. ‒ Gotlih, K. ‒ Ficko, M. ‒ Balič, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 898-903

Fig. 9. The engraving was treated with closed-line treatment [1] (example 2) Table 1. Comparison of the experimental results The tool was milled with lineRoughness treatment at an parameter angle of 45º (example 1) Ra [mm] 0.78 4.78 Rz [mm] 5.67 Rmax [mm] Number of produced 28,784 parts

The tool was milled with closed-line treatment (example 2) 0.86 6.02 7.94 20,148

The results of this experiment clearly show the influence of milling-strategy on its durability when applied to the active surface of a forging tool. It is favourable to use such a milling strategy that has toolpaths following the flow of the material during the forging process. Although the results are very promising, it should be noted that the research is incomplete, as yet. In the future we are planning to research the connections between various milling-parameters and milling-strategies, on forging tool durability.

4 REFERENCES [1] Studencnik, D. (2004). Meaning and influence of different method renovation forging tools to abstinence of forging tools. Specialist thesis. Faculty of Mechanical engineering, University of Maribor.

[2] Kramar, D., Kopač, J. (2009). High pressure cooling in the machining of hard to machine materials. Strojniški vestnik – Journal of Mechanical Engineering, vol. 55, no. 11, p.

685-694.

[3] Chen, J.S., Huang, Y.K., Chen, M.S. (2005).

A study of the surface scallop generating mechanism in the ball-end milling process.

International Journal of Machine Tools and Manufacture, vol. 45, no 9, p. 1077-1084, DOI:10.1016/j.ijmachtools.2004.11.019.

[4] Feng, H.Y., Li, H. (2002). Constant scallopheight tool path generation for three-axis sculptured surface machining. ComputerAided Design, vol. 34, p. 647-654, DOI:10.1016/S0010-4485(01)00136-1.

[5] Chen, J.-S., Huang, Y.-K., Chen, M.-S. (2005). A study of the surface scallop generating mechanism in the ball-end milling process. International Journal of Machine Tools and Manufacture, vol.

45, no. 9, p. 1077-1084, DOI:10.1016/j. ijmachtools.2004.11.019.

[6] Benardos, P.G., Vosniakos, G.C. (2002).

Prediction of surface roughness in CNC face milling using neural networks and Taguchi’s design of experiments. Robotics and Computer-Integrated Manufacturing, vol.

18, no. 5-6, p. 343-354, DOI:10.1016/S07365845(02)00005-4.

[7] Kaczmarek, J. (1983). Principles of machining by cutting, abrasion and erosion.

Peter Peregrinus, London.

[8] Klancnik, S., Senveter, J. (2010). Computer based work piece detection on cnc milling machine tools using optical camera and neural networks. Advances in Production Engineering & Management, vol. 5, no. 1, p.

59-68.

[9] Stupan, J., Potrc, I. (2000). Computer simulation forging processes. Strojniški vestnik – Journal of Mechanical Engineering, vol. 46, no. 9. p. 641-647.

[10] Kecelj, B., Kopač, J., Kampuš, Z., Kuzman, K. (2004). Speciality of HSC in manufacturing of forging dies. Journal of Materials Processing Technology, vol. 157-158, p. 536-542, DOI: 10.1016/j.jmatprotec.2004.07.112.

[11] Dotcheva, M., Milward, H. (2008). A generation of more efficient CNC tool paths using simulation. International Journal of Simulation Modelling, vol. 7, no. 3, p. 135145, DOI:10.2507/IJSIMM07(3)3.108.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 904-910 DOI:10.5545/sv-jme.2011.010

Paper received: 12.01.2011 Paper accepted: 02.08.2011

Polycrystalline Cubic Boron Nitride (PCBN) Tool Life and Wear in Turning of Amorphous-Crystalline Iron-Based Coatings

Tanovic, L. ‒ Bojanic, P. ‒ Puzovic, R. ‒ Klimenko, S. Ljubodrag Tanovic1 ‒ Pavao Bojanic1,* ‒ Radovan Puzovic1 ‒ Sergey Klimenko2 1University of Belgrade, Faculty of Mechanical Engineering, Serbia 2V. Bakul Institute for Superhard Materials of the National Academy of Sciences of Ukraine, Ukraine The paper presents PCBN-Ciborit cutting tools life and wear test results. The effects of the machining regime when turning amorphous-crystalline Fe80B20 and Fe79Cr16B5 coating systems applied to conventional workpiece materials were assessed. It has been shown that the observed tool wear mechanisms are complex in their character and are dominated by abrasive-mechanical, adhesive and chemical effects in the cutting zone. Under changing turning conditions tool life is affected by the structural-phase composition and by the non-homogeneous structure of the coating. Specifically, when turning gas-flame coatings deposited with a Fe80B20 electrode and electro-arc coatings with a Fe79Cr16B5 electrode the lowest wear and the highest tool life was achieved at cutting speeds of v = 1.1 to 1.2 m/s and a back rake angle of γ = –10º. It has been demonstrated that a change of the back rake angle from γ = 0 to –10º does not have a great effect on tool life contrary to the case with γ = -20º. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: cutting, tool wear, tool life, amorphous-crystalline coating

0 INTRODUCTION Recently, materials possessing amorphous and amorphous-crystalline structures have been used to manufacture workpieces with improved physical-chemical and mechanical properties.

Extraordinary enhancements in performance can be gained when, for surface layer reinforcement, amorphous alloys in the form of coatings are deposited on workpieces made of conventional design materials.

A relatively new trend is the manufacture of Fe-B, Fe-Si-B, Fe-Cr-B alloy coatings, which form amorphous structures. These coatings are relatively inexpensive and possess a uniform complex of properties, however, the quality of the machined surface frequently lacks desirable characteristics. The available technical literature discriminates between two directions of research in this area. The first deals with the different types of amorphous coatings and their application technologies, while the second deals with the tools for their machining.

Surfaces of the amorphous metal-metalloid alloy Fe80B20 prepared by laser annealing were 904

investigated [1] and [2]. Their results show that laser annealing leads to large enhancements in boron concentration in the first 0.5 to 1.0 nm of the surface, as a consequence of the surface oxidation of boron and iron. The supercooled liquid region, the tensile fracture strength and the Vickers hardness of the Fe-B-Zr amorphous alloys were measured [3] and [4]. The addition of Nb to the ternary Fe-B-Zr system was found to lead to significant changes in glass-forming ability (GFA), strength, and hardness. Their results are discussed in the context of the three empirical rules for the achievement of high GFA metallic glass and the change in packing density of the amorphous phases. A novel Gas Metal Arc Welding (GMAW) process, referred to as double-electrode GMAW or DE-GMAW, has been developed to make it possible to control melting at a desired level [5].

By means of differential scanning calorimeter (DSC) measurements, the thermal stability of an amorphous Fe80B20 alloy after various periods of low-energy ball milling has also been studied [6] and [7]. The results indicate that the thermal stability of the amorphous

*Corr. Author’s Address: University of Belgrade, Faculty of Mechanical Engineering, Department for Production Engineering, Kraljice Marije 16, 11120 Belgrade 35, Serbia, pbojanic@mas.bg.ac.rs

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Fe80B20 ribbons can be enhanced upon mechanical deformation with a low milling intensity.

The second group of papers is related to PCBN tools and their use in hard turning. Many parameters, such as the composition and hardness of the workpiece and of the tool material, the environment and the machining parameters influence the different degradation mechanisms, which will eventually affect the workpiece surface finish and tool life [8] and [9]. Degradation mechanisms in PCBN tools are particularly dependent on the temperature and stress during cutting and on the chemical composition of both the tool and workpiece materials. According to [10], three main mechanisms are involved in the wear of PCBN tools: (1) chemical wear caused by the interaction with the environment (atmospheric oxidation and interaction with the workpiece), (2) formation of a protective layer on the surface of the tool at high temperatures, and (3) removal of this layer at very low temperatures or lower cutting speed, leading to abrasive wear and further chemical wear. The protective layer has also been considered [11].

Different classifications of tool wear in metal cutting can be found in the literature.

In typical classifications the most prominent mechanisms of tool wear are abrasion, adhesion and diffusion. CBN tool flank wear is considered to be the main wear pattern and an important tool life metric in hard turning, and it has been extensively studied [12] to [15]. While flank wear is generally used as an indicator of tool wear, it does not tell the whole story. The brittle materials used for cutting tools, such as CBN, require large wedge angles and negative rake angles. However, as crater wear progresses, the effective rake angle becomes more positive, leading to changes in the tool’s cutting geometry [16] and [17].

The application of amorphous-crystalline coatings is limited to a considerable extent in practice due to the lack of scientifically based data on their mechanical machining properties.

In the engineering literature not many works are available related to mechanical machining of coatings that possess an amorphous-crystalline structure. There is virtually no information about the effects of structure properties and amorphouscrystalline coatings on cutting tool life and wear.

Coatings are considerably more difficult to

machine by cutting as compared to monolithic materials of identical chemical composition and hardness due to the changing mechanical properties and machining allovance size as well as due to the presence of solid particles and pores in the structure.

The answer to the question about the effects of the coating material’s non-homogeneity on tool life and wear is crucial in the choice of machining regime parameters. A substantial improvement in techno-economic parameters of the cutting process in coated components can be achieved by using tools made of super-hard materials such as PCBN. Therefore, the principal aim of this paper is to experimentally assess the influence of key machining partameters on PCBN tool life and wear.

1 EXPERIMENTAL SETUP AND PROCEDURES The experiments were carried out using a PH 42-CNC lathe with a 16 kW spindle with a speed range from 16 to 5200 rpm, with PCBNCiborit cutter bits (Institute of superhard materials, Kiev, UA), a scanning electron microscope (JSPM 5200), and workpiece coatings with the following structures: Fe80B20 and Fe79Cr16B5 formed by various technologies (Fig. 1). The base workpiece material on which the coatings were deposited was steel 40Ch-GOST (1040 SAE).

Fig. 1. Experimental set-up Table 1 lists the mechanical characteristics of the coating materials while Table 2 gives the hardness values of the coatings. The tests were performed under the following conditions: cutting speed 0.6 to 2.7 m/s; feed 0.05 to 0.18 mm/rev; cutting tool, insert PCBN-Ciborit, RNMN 0703-T, back rake angle 0 to ‒20°.

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Table 1. Mechanical characteristics of the coating materials Material

Fe80B20 Fe79Cr16B5

Vickers hardness, HV [GPa] Stress fracture s [GPa] Young’s Modul E [GPa]

10.8 10.3

3.5 3.06

Table 2. Hardness values of the coatings deposited by different methods Coatings Gas-flame coating with Fe80B20 electrode Gas-flame coating with Fe80B20 powder Electro-arc coating with Fe80B20 electrode Gas-flame coating with Fe79Cr16B5 powder Electro-arc coating with Fe79Cr16B5 electrode

Hardness, HRc 56 to 58 52 to 56 56 to 58 54 to 58 56 to 58

2 TOOL WEAR AND TOOL LIFE RESULTS The turning of workpieces with amorphouscrystalline coatings Fe80B20 and Fe79Cr16B5 with PCBN tools was performed at relatively high cutting speeds, lower feeds and depths of cut due to constraints imposed by the coating thickness of a = 0.1 to 0.6 mm. As the depth of cut was small, the contact area between the tool and the chip was also small, resulting in a very high specific load in the contact area. The experiments were replicated five times under identical machining conditions. Tool flank wear was measured and

170 -

σ/E

HV / σ

0.02 -

3.1 3.3

the mean measured value was taken for the result, with deviations of up to 5%. For every coating and corresponding cutting regime a new Ciborit cutter bit was used. The investigations were based on observations and measurements of flank wear (VB) in turning of the corresponding coatings and the determination of the tool’s life. Fig. 2 shows a photograph of the back rake and the flank surface wear patterns of the Ciborit cutting tools, while Fig. 3 presents the nature of the changes in tool flank wear when turning Fe80B20 and Fe79Cr16B5 coatings. In Fig. 3 the curves 1, 2 represent the curve of Fe80B20 coating wear, while the curves 3 and 4 represent the curve of Fe79Cr16B5 coating wear. Due to a large range of values along the T axis, the wear curves are represented in two segments such as: the initial segments of the curves on the T axis are denoted with T2,4, while the continuation of the wear curves is denoted with T1,3.

Experimental tool life data when machining two types of coatings formed by various technologies, assuming a VB = 0.25 mm tool wear criterion, are shown in Fig. 4. Investigations on the effects of the cutting regime and of the tool’s back rake angle on tool life were conducted on Fe80B20 based coatings and the results obtained are presented in Fig. 5 and in the 3D coordinate system in Fig. 6.

Fig. 2. PCBN-Ciborit wear (1-back rake surface; 2-cutting edge; 3-flank; f = 0.05 mm/rev; a = 0.2 mm; γ = -10º) when machining of coatings; a) gas flame with electrode Fe80B20,v = 1.7 m/s, × 150; b) electro arc with electrode Fe79Cr16B5, v = 2.7 m/s, × 150) 906

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Fig. 3. Change in flank wear VB during cutting (1, 2: gas flame coating with electrode Fe80B20; 3, 4: electro arc coating with electrode Fe79Cr16B5): v = 2.7 m/s, f = 0.05 mm/rev; A = 0.2 mm; γ = -10º

Fig. 4. Change in tool life (VB = 0.25 mm) as a function of cutting speed (f = 0.05 mm/rev, α = 0.2 mm, γ = -10º): 1: gas flame coating with electrode Fe80B20, 2: electro arc coating with electrode Fe80B20, 3: gas flame coating with powder Fe80B20, 4: electro arc coating with electrode Fe79Cr16B5, 5: gas flame coating with powder Fe79Cr16B5

Fig. 5. Change in tool life (VB = 0.25 mm) in turning of gas flame coatings with electrode Fe80B20 as a the function of: a and c) cutting speed and feed (a = 0.1 mm, γ = -10º), b and d) cutting speed and back rake angle (f = 0.05 mm/rev, a = 0.2 mm) 3 DISCUSSION When turning of coatings whose structure is over 70% amorphous (Fe80B20) a continuous/ ribbon chip is being formed. This chip is contact with tool back rake surface over a long period of time, resulting in tool wear across the back rake surface as depicted in Fig. 2a. Due to the low

thickness of the chips, the tool contact surface approaches the cutting edge and results in its rapid wear and destruction. When turning coatings whose amorphous structure accounts for 50% (arc coatings formed with electrodes Fe79Cr16B5) the tool predominantly wears across the flank (Fig. 2b). This is the consequence of the fact that the structure of these coatings is formed of separate

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Fig. 6. Change in tool life in the 3D coordinate system T, v, f and T, v, γ, (machining conditions, Fig. 5) fragments which are, unlike highly amorphous coatings, insufficiently deformed in the layering process and weakly inter-connected, primarily due to diffusion processes and mechanical interaction.

In machining of both types of coatings initial tool wear is observed, which then turns into uniform wear, as noticeable in Fig. 3.

Simultaneously with tool back rake and flank surface wear, when turning the respective coatings, the occurrence of deposits was observed on the tool’s back rake surface. Based on the tests performed, it was observed that when turning coatings of identical chemical composition but formed by different technologies deposits with various characteristics occur on tool back rake surface. Namely, after turning of electro arc coating Fe79Cr16B5 deposits of relatively high density are formed on the tool’s working surfaces, while when turning of heterogeneous gas flame coatings of the same chemical composition loose unstable deposits are observable and are periodically removed in the machining process leading to more intensive wear and lower tool life. Different deposit forms and structures are created during machining of coatings of the same chemical composition, but formed by different technologies, indicating that the characteristics of the contact inteactions in the cutting zone are different and determined by the specific method by which each coating was formed.

It is well-known that coatings on workpieces are of relatively low thickness but high price, therefore tool life determination is very costly and, in some cases, tests are difficult 908

to carry out. Hence, when determining tool life, it is assumed that [18] if in the process of machining tool wear intensity under machining conditions (v, f, γ) is higher or lower compared to tool wear intensity under a different set of machining conditions (v‫׳‬, f‫׳‬, γ‫)׳‬, then the trend in the changes will be retained along the entire tool wear curve.

This makes it possible to conduct tests to find out a sharp tool’s life in turning of various types of coatings, where VB = 0.1 mm is taken as the tool life criterion (T0.1).

Experimental data presented in Fig. 3 indicate that tool life relationships can be established for a VB = 0.25 mm wear (T0.25) value relative to the tool life for a VB = 0.1 mm wear. Namely, when turning gas flame coatings with electrode Fe80B20 for VB = 0.1 mm, the tool life is T0.1 = 0.88 min. The continuation of the machining process until VB = 0.25 mm is reached results in a tool life of T0.25 = 57 min. According to the above, tool life relationships for machining of coatings can be established. Specifically, for the cases considerd here: a) gas-flame coating formed with an electrode Fe80B20 → T0.25 = 64.7 T0.1, b) electro-arc coating formed with an electrode Fe79Cr16B5 → T0.25 = 45 T0.1.

It was observed that on the tool’s working surfaces after turning of electric-arc deposited coatings, Fe79Cr16B5 deposits of relatively higher density were formed, while when turning heterogeneous gas-flame coatings of the same chemical composition loose unstable deposits were created. The latter deposits were periodically

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removed during the machining process, leading to more intensive tool wear and decrease in tool life.

Based on the results for tool life measurements under different machining conditions of the respective coatings, tool life measures were established, as presented in Fig. 4.

A general look of the deposits and their position relative to the tool’s cutting edges changes, depending on machining conditions.

When it comes to machining of Fe80B20 based coatings, the maximal tool life is achieved during machining of the gas-flame deposited coating with electrodes, while minimal life was achieved with powder coating. It should be pointed out that a lower influence of coating heterogeneity on tool life is observed with an increase in cutting speed.

Thus, cutting edge life in turning at a cutting speed of v = 2 m/s (f = 0.05 mm/rev, a = 0.2 mm) of the arc coating Fe79Cr16B5 amounts to 18 minutes, while that of the gas coating Fe79Cr16B5 is 16 minutes. However, when the cutting speed is reduced to v = 1.2 m/s under analogous conditions the tool life is at the 27 and 23 minutes level at the VB = 0.25 mm benchmark for wear. The cutting speed shows a great influence on the contact conditions and on the physical characteristics of the phenomena that accompany tool wear.

Cutting speed largely affects tool wear, and maximal tool life is achieved at a cutting speed of v = 1.1 to 1.2 m/s.

Under this machining regime, contact conditions are created for which the intensity of the abrasive interaction in the cutting zone is significantly decreased and the tool is subjected to optimal dynamic stresses, adhesion contact and chemical interactions. All the aforementioned conditions lead to a minimum intensity of wear.

With an increase in the cutting speed, tool life gradually decreases as a consequence of the elevated intensity of wear and changes in both the temperature and dynamic conditions of cutting, as well as of the development of highly unfavorable impact loads on the Ciborit cutter bit.

The influence of feed is essentially coupled with that of the cutting speed. It is characteristic of the investigated cutting speed range that tool life decreases with an increase in the feed. In turning at speed of v > 1.2 m/s the feed influences tool life more significantly than at lower cutting speeds. Namely, an increase in the feed leads to

a temperature rise in the cutting zone leading to considerably intensified chemical reactions that, in turn, lead to a decrease in tool life. In turning of the coatings being considered in the present work, the tool back rake angle also appreciably influences tool life. Thus, a change in γ = 0 to ‒10º does not show a high influence on tool life change at cutting speeds up to v = 1.2 m/s, which is not the case when γ = ‒20º (Figs. 5a and c, 6). This is explained by the increase in the chip and tool back rake contact surface area. So, an optimum value for the back rake is γ = ‒10º, because at this value the tool possesses the highest tool life, which is needed to achieve the lowest values of the machined surface’s roughness parameters.

4 CONCLUSION The paper presents tests results for PCBN Ciborit tool wear and the determination of its tool life when turning coatings that possess a heterogeneous amorphous-crystalline structure and that are formed by various technologies.

On the basis of what has been said, the mechanism of PCBN cutting edge wear in turning of amorphous-crystalline coatings is complex in its characteristics and is influenced by abrasive, mechanical, adhesive and chemical interactions in the cutting zone. The wear characteristics are dependent upon the structure of the machined material, cutting regime and tool geometry.

The investigations carried out in turning of gas-flame coatings Fe80B20 with Ciborit tools show that the highest tool life was achieved at a cutting speed of 1.2 m/s, feed 0.06 mm/rev and rake angle γ = ‒10º (Fig. 5).

The results indicate that in the tested cutting speed range, an increase in the feed leads to a tool life decrease. In turning at v > 1.2 m/s, the influence of the feed on tool life decrease is greater than at lower cutting speeds. It has been demonstrated that a change in the back rake angle γ = ‒10 to ‒20º at the applied cutting speeds and feed leads to tool life decrease, and that as a consequence of tool wear increase. It should be pointed out that as the cutting speed is increased the influence of heterogeneity of the coatings Fe80B20 and Fe79Cr16B5 on tool life is decreased, as indicated by Fig. 3.

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5 REFERENCES [1] Gimzewski, J.K., Moller, R., Myhra, S., Schlittler, R.R., Stoll, E. (1990).

Investigations of the surface of the amorphous alloy Fe80B20 by STM, XPS and AES. Journal of Non-Crystalline Solids, vol. 116, p. 253261, DOI:10.1016/0022-3093(90)90699-M. [2] Fan, G.J., Quan, M.X., Hu. Z.Q. (1996).

Deformation enhanced thermal stability of an amorphous Fe80B20 alloy. Journal of Applied Physics, vol. 80, no. 10, p. 6055-6057, DOI:10.1063/1.363563.

[3] Ma, L., Wang, L., Zhang, T., Inoue, A. (1999).

Effect of Nb addition on glass-forming ability, strength, and hardness off Fe-B-Zr amorphous alloys. Materials Research Bulletin, vol.

34, no. 6, p. 915-920, DOI:10.1016/S00255408(99)00089-6.

[4] Hirata, A., Hirotsu, Y., Amiya, K., Nishiyama, N., Inoue, A. (2008). Nanocrystallization of complex FeB-type structure in glassy Fe-Co-B-Si-Nb alloy. Intermetallics, vol.

16, no. 4, p. 491-497, DOI:10.1016/j. intermet.2007.11.006.

[5] Li, K., Zhang, Y. (2007). Metal transfer in double-electrode gas metal arc welding.

ASME, Journal of Manufacturing Science and Engineering, vol. 129, no. 6, p. 9911000, DOI:10.1115/1.2769729.

[6] Fan, G.J., Quan, M.X., Hu, Z.Q. (2008).

Deformation-enhanced thermal stability of an amorphous Fe80B20 alloy. Intermetallics, vol.

16, no. 4, p. 491-497.

[7] Alaeddine, M., Ando, T.R., Doumanidis, C.C. (2005). Modeling the melting and dissolution stages during thermal processing of intermetallic coatings from layered precursors. ASME, Journal of Manufacturing Science and Engineering, vol. 127, p. 148156.

[8] Angseryd, J., Coronel, E., Elfwing, M., Olsson, E., Andren, H.O. (2009). The microstructure of the affected zone of a worn PCBN cutting tool characterised with SEM and TEM. Wear, vol. 267, no. 5-8, p. 10311040, DOI:10.1016/j.wear.2008.12.075.

[9] Velkavrh, I., Kalin, M., Vižintin, J. (2008).

The performance and mechanisms of DLCcoated surface in contact with steel in 910

boundary-lubrication conditions - a review.

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[10] Chou, K., Evans, Y., Barash, C.J. (2002).

Experimental investigation on CBN turning of hardened AISI 52100 steel. Journal of Materials Processing Technology, vol.

124, p. 274-283, DOI:10.1016/S09240136(02)00180-2.

[11] Klimenko, S.A., Mukovoz, Yu.A., Lyashko, V.A., Vashchenko, A.N., Ogorodnik, V.V. (1992). On the wear mechanism of cubic boron Nitride base cutting tools. Wear, vol. 157, no. 1, p. 1-7, DOI:10.1016/00431648(92)90183-9.

[12] Poulachon, G., Moisan, A., Jawahir, I.S. (2001). Tool-wear mechanisms in hard turning with polycrystalline cubic boron nitride tools. Wear, vol. 250, p. 576-586, DOI:10.1016/S0043-1648(01)00609-3.

[13] Poulachon, G., Bandyopadhyay, B.P., Jawahir, I.S., Pheulpin, S., Seguin, E. (2004).

Wear behavior of CBN tools while turning various hardened steels. Wear, vol. 256, no. 3-4, p. 302-310, DOI:10.1016/S00431648(03)00414-9.

[14] Barry, J., Byrne, G. (2001). Cutting tool-wear in the machining of hardened steels. Wear, vol. 247, p. 152-160, DOI:10.1016/S00431648(00)00531-7.

[15] Huang, Y., Liang, S.Y. (2004). Modeling of CBN tool flank wear progression in finish hard turning. ASME Journal of Manufacturing Science and Engineering, vol. 126, no. 1, p.

98-106, DOI:10.1115/1.1644543.

[16] Huang, Y., Liang, S.Y. (2004). Modeling of CBN tool crater wear in finish hard turning. International Journal of Advanced Manufacturing Technology, vol. 24, no. 9-10, p. 632-639, DOI:10.1007/s00170-003-17445.

[17] Huang, Y., Dawson, Ty.G. (2005). Tool crater wear depth modeling in CBN hard turning. Wear, vol. 258, no. 9, p. 1455-1461, DOI:10.1016/j.wear.2004.08.010.

[18] Ber, A., Kaldor, S. (1982). The first seconds off cutting. Annals of the CIRP, vol. 31, no. 1, p. 13-17, DOI:10.1016/S00078506(07)63260-0.

Tanovic, L. ‒ Bojanic, P. ‒ Puzovic, R. ‒ Klimenko, S.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 911-926 DOI: 10.5545/sv-jme.2011.100

Paper received: 15.03.2011 Paper accepted: 28.07.2011

Simulation of Water-Gas Shift Membrane Reactor for Integrated Gasification Combined Cycle Plant with CO2 Capture Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H. Andrej Lotrič1,* ‒ Mihael Sekavčnik1 ‒ Christian Kunze2 ‒ Hartmut Spliethoff2 1University of Ljubljana, Faculty of Mechanical Engineering, Slovenia 2Technische Universität München, Institute for Energy Systems, Germany

The effectiveness of energy conversion and carbon dioxide sequestration in Integrated Gasification Combined Cycle (IGCC) is highly dependent on the syngas composition and its further processing. Water gas shift membrane reactor (WGSMR) enables a promising way of syngas-to-hydrogen conversion with favourable carbon dioxide sequestration capabilities. This paper deals with a numerical approach to the modelling of a water gas shift reaction (WGSR) in a membrane reactor which promotes a reaction process by selectively removing hydrogen from the reaction zone through the membrane, making the reaction equilibrium shifting to the product side. Modelling of the WGSR kinetics was based on Bradford mechanism which was used to develop a code within Mathematica programming language to simulate the chemical reactions. The results were implemented as initial and boundary conditions for the tubular WGSMR model designed with Aspen Plus software to analyze the broader system behaviour. On the basis of selected boundary conditions the designed base case model predicts that 89.1% CO conversion can be achieved. Calculations show that more than 70% of carbon monoxide conversion into hydrogen appears along the first 40% of reactor length scale. For isothermal conditions more than two thirds of the heat released by WGSR should be extracted from the first 20% of the reactor length. Sensitivity analysis of the WGSMR was also performed by changing the membrane’s permeance and surface area. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: IGCC, water-gas shift reaction, membrane reactor 0 INTRODUCTION Gasification is an exothermal chemical process at high temperatures and pressures between a hydro-carbonaceous material and an oxidizer (air, oxygen and/or steam). In general, the feedstock used for gasification consists of hydrocarbons and can be in different forms like coal, oil, heavy refinery residuals, coke, biomass or even municipal waste. As the oxygen supply is limited (generally 20 to 70% of the oxygen needed for complete combustion) only partial combustion of the feedstock occurs. The released heat from the chemical reaction drives the secondary reaction that further converts organic material to syngas.

Syngas is mostly a mixture of CO, H2, CO2 and some CH4, H2O and N2. The ratio between CO, H2 and CO2 can be quite different as it depends on the type of a gasifier (movingbed, fluid-bed or entrained-flow gasifier) and the gasifying feedstock. To achieve the desired composition of syngas the correct operating temperature – hence the correct amount of

oxidizer for partial combustion – and pressure of gasification must be selected.

The operating temperature can range between 800 and 1800 °C and the pressure between 10 and 100 bar.

Gasification is one of the main processes in integrated gasification combined cycle (IGCC) energy systems. In general, all existing IGCC power plants follow the chain of events presented in Fig. 1 with an exception of carbon dioxide sequestration (CCS). This chain can be broken down into ten key processes [1]: 1. Air separation unit (ASU) separates air into oxygen to supply the gasifier and nitrogen, which can be used as a carrier, sweep or dilution gas.

2. Coal particles are transferred – using pneumatic conveying or in a form of watercoal slurry – into a gasifier.

3. By-products captured in the gasifier (ash and slag) can have commercial value, depending on local market conditions.

*Corr. Author’s Address: University of Ljubljana, Faculty of Mechanical Engineering, Laboratory for Heat and Power, Aškerčeva 6, 1000 Ljubljana, Slovenia, andrej.lotric@gmail.com

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4. The syngas also has small amounts of other impurities (e.g. H2S, COS, NH3 and also chlorides, fluorides, mercury etc.) which are removed during the gas clean-up. Before the clean-up process syngas is cooled, usually in a syngas cooler, where part (or all) of the steam used in the gasification process is produced.

5. Around 99% of H2S is separated from the syngas and converted to elemental sulphur or possibly sulphuric acid.

6. Nowadays, the syngas is passed through a two stage WGS reactor and reacted over a catalyst with added steam to convert the majority of the CO into CO2 and additional H2. The H2 is separated from the gas mixture usually by using the pressure swing adsorption (PSA) process.

7. The future potential of IGCC systems is the CO2 removal, leaving H2-rich syngas behind.

8. Syngas can also be used in variety of processes; like production of synthetic fuels

or chemicals and pure H2 can be used in fuel cells or combined cycle to produce electricity.

9. The main cause for IGCC plants being more efficient than conventional power plants is the combined use of a gas and a steam turbine to produce electricity.

10. Much of the water used in this process is recycled in the plant, while some is evaporated in a cooling tower.

1 WATER-GAS SHIFT REACTION PROCESS 1.1 Membrane Reactor In general, the membrane reactor (MR) is a device that combines a membrane separation process with a chemical reactor in one unit. The MR is capable of promoting a reaction process by selectively removing at least one of the products from the reaction zone through the membrane, making the equilibrium reaction shifting to the product side.

Fig. 1. Schematic presentation of the syngas production, hydrogen separation, CCS and hydrogen utilisation in the IGCC 912

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Fig. 2. WGSR process in the MR

The membrane used is highly selective to the product of interest therefore, the product can be directly recovered during the reaction, eliminating the need for additional product purification steps.

The rate of the product flux through the membrane is proportional to the product’s partial pressure difference between the process (feed) side and the permeate side; therefore, the extracted product is recovered at a lower pressure than the process stream pressure. To increase the pressure difference, the inert diluting (or carrier) gas is sometimes used. The diluting gas is carrying away the permeated product therefore, reducing the concentration on the permeate side and thus, increasing the driving force for permeation. The pressure drop through the membrane is material dependent, and it obeys different laws for different membrane materials.

With MR the yield can be increased (even beyond the equilibrium value for equilibrium reactions) and/or the selectivity can be improved by suppressing other undesired side reactions or the secondary reaction of products. Due to the integration of reaction and separation, chemical processes become simpler leading to a much lower processing cost.

According to Eq. (1) changing the pressure of the WGSR should not have any considerable effect on changing the equilibrium concentrations because the equation is equimolar. However, the experimental results, obtained from [2], show that very high pressures slightly favour the CO conversion which can be seen in the Table 1.

Table 1. Effect of pressure on equilibrium CO concentrations (inlet dry gas: 13.2% CO, 10.3% CO2, 35.3% H2, 41.2% N2, steam-to-dry gas = 0.5) Temperature [°C] 200 300 400 500 600 700

p= 3.04 bar [% CO] 0.12 0.68 1.98 3.93 6.15 8.38

p= 30.39 bar [% CO] 0.12 0.65 1.94 3.88 6.10 8.34

p= 303.9 bar [% CO] 0.07 0.48 1.61 3.46 5.68 7.95

1.2 Water-Gas Shift Reaction The equilibrium of reversible reactions can be shifted toward more product formation by changing reaction conditions such as pressure and temperature or concentrations of reactants or products.

The main focus of this paper is the Water‑Gas Shift Reaction (WGSR), which is also an equilibrium reaction and it is represented with the following equation:

CO + H2O ↔ CO2 + H2 ‒ 41 MJ/kmol .

(1)

Fig. 3. Reactant conversion for the WGSR where H2 and CO2 are not present at the beginning [3] On the other hand, it is well known that increasing the temperature has a decreasing effect on equilibrium CO conversion. Therefore, H2 yield can be considerably enhanced at high

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reaction temperatures, at which the equilibrium CO conversion would otherwise be low, by extracting either CO2 or H2 from the reaction mixture.

1.3.1 General Definition of Rate Law Chemical kinetics investigates how different experimental conditions can influence the chemical reaction rate. The main factors that can speed up the reaction rate include increase in the concentrations of the reactants, increase in temperature or pressure at which the reaction occurs and whether or not any catalysts are present in the reaction.

Usually, what appears to be a single step conversion is in fact a multistep reaction and thus, the reaction mechanism is a step by step sequence of elementary reactions. Each step has its own rate law and molecularity (the number of species taking part in that step); therefore the slowest step is the one that determines the overall reaction rate – the rate limiting step.

Only for simple (elementary) reactions a partial order of reaction is the same as the stoichiometric number of the reactant concerned.

For stepwise reactions there is no general connection between stoichiometric numbers and partial orders. Such reactions may have more complex rate laws and the orders of reaction are in principle always assigned to the elementary steps.

By conducting experiments involving reactants A and B, one would find that the rate of the reaction r is related to the concentrations [A] and [B] in a rate law as: r = k [A]a [B]b , (2)

where k is the rate constant and the powers a and b are called the partial orders of the reaction. The overall order of the reaction is found by adding up the partial orders. The rate constant is not a true constant because it is dependent on the reaction’s temperature T and the activation energy Ea and is defined via Arrhenius equation as:

k = k0 ⋅ e

−

Ea R⋅T

Assuming that above reactions are elementary, the reaction rate can be expressed as:

r = k1·[A]a·[B]b ‒ k2·[C]c·[D]d , (5)

where k1 is the rate coefficient for the forward reaction which consumes A and B; k2 is the rate coefficient for the backward reaction, which consumes C and D. In a reversible reaction, like WGSR, chemical equilibrium is reached when the rates of the forward and backward reactions are equal and the concentrations of the reactants and products no longer change (r = 0 in balance).

1.3.2 Bradford Mechanism a) Backward WGSR Bradford [4] assumed that the reaction mechanism would follow the simple gas-phase chain-reaction mechanism given below (M indicates any gas-phase collision partner). The chain is initiated by the gas-phase dissociation of hydrogen (Eq. (6)). Propagation steps are represented by reactions in Eqs. (7) and (8). The termination step corresponds to the gas-phase reassociation of H2 in Eq. (9), consuming the chain carriers. k

1 H 2 + M  → 2H + M , (6)

2 H + CO 2 ←   → CO + OH , (7) k

3 OH + H 2 ←   → H 2 O + H , (8) k

−1 2H + M  → H 2 + M . (9)

k

−2

k

−3

k

With the assumption of low conversion and a stationary state for the concentrations of the intermediates (H and OH concentrations do not change significantly with respect to time), the following rate equation rb is obtained,

, (3)

where k0 is the pre-exponential factor.

914

a A + b B ⇔ c C + d D . (4)

1.3 Water-Gas Shift Kinetics

A pair of forward and backward reactions may define an equilibrium process where A and B react into C and D and vice versa (a, b, c and d are the stoichiometric coefficients):

rb =

d [ CO ] dt

1/ 2

 k  = 1   k−1 

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

⋅ k2 ⋅ [ H 2 ]

1/ 2

⋅ [ CO 2 ] . (10)

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 911 -926

Therefore, the rate constant for the backward WGSR kb may be expressed as: 1/ 2

 k  kb =  1  ⋅ k2 , (11)  k−1  where the temperature dependence of the rate constant is described with the Arrhenius Eq. (3).

Consequently, the expression for the rate of reaction in terms of kb becomes: d [ CO ] 1/ 2 rb = = kb ⋅ [ H 2 ] ⋅ [ CO 2 ]. (12) ,

dt

b) Forward WGSR The gas-phase, chain-reaction mechanism proposed by Bradford [4] can also be used to describe the forward WGSR. Reaction in Eq. (13) provides the chain initiation by the reaction of H2O with any gas-phase molecule (designated

by M). Reactions in Eqs. (14) and (15) are the propagation steps, while the reaction in Eq. (16) is the termination step. k4

H 2 O + M → H + OH + M , (13)

5 CO + OH ←   → H + CO 2 , (14) k

6 H 2 O + H ←   → OH + H 2 , (15) k

−4 H + OH + M  → H 2 O + M. (16)

k

−5

k

−6

k

The stationary-state approximation for the concentration of the chain-carriers (H and OH) under the conditions of low conversions leads to the following expression for the forward WGSR rate of reaction rf :

rf =

d [ CO 2 ] dt

= 1/ 2

 k  =  4 ⋅ k5 ⋅ k6  k  −4 

(17) ⋅ [ CO ]

1/ 2

⋅ [ H 2 O].

The rate constant for the forward WGSR kf is defined as: 1/ 2

 k  k f =  4 ⋅ k5 ⋅ k6  k  −4 

, (18)

and the rate can be expressed as:

rf =

d [ CO 2 ] dt

= k f ⋅ [ CO ]

1/ 2

⋅ [ H 2 O ] . (19)

2 MODEL OF WATER-GAS SHIFT MEMBRANE REACTOR At the beginning it was assumed that the use of different unit models integrated in simulation software Aspen Plus will be sufficient to model the MR. However, it turned out that the model could not correctly predict the process taking place inside of the reactor. This is why the decision was made to include the WGSR kinetics, which gave a better overview of what occurs inside the reactor.

The modelling of the WGSMR was conducted by using both the Mathematica programming language and Aspen Plus. The calculations made in Mathematica were based on WGSR kinetics and used to predict the reaction

Fig. 4. Boundary conditions for the modelled system and link between Mathematica and Aspen Plus environment Simulation of Water-Gas Shift Membrane Reactor for Integrated Gasification Combined Cycle Plant with CO2 Capture

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rate and the permeation through the membrane.

The calculated data were then imported in Aspen Plus as input (or initial) values for simulations.

Aspen Plus was used to predict the behaviour of a WGSMR by using basic engineering relationships, such as mass and energy balances, and phase and chemical equilibrium. According to the calculated data from Mathematica, predicted operating conditions, and different design models, the behaviour of MR was simulated.

around 1.1 can be reached with quench water at 260 °C.

2.1 Initial and Boundary Conditions In IGCC system the MR will be situated after the gasifier and before the gas turbine combustor, replacing the WGS reactor and PSA system (see Fig. 1).

The boundaries for the modelled system are marked with the dashed line in Fig. 4. In this figure n represents the molar flow, x the composition, T the temperature and p the pressure of gas stream. Heat duty is represented with Q and CO conversion with XCO. Index GAS stands for syngas, R for retentate side and P for permeate side of the membrane.

Based on the position in the chain of processes that take part in the IGCC, the initial and boundary conditions for “base case” studies were defined: a) Based on gasifier: • The modelling is based on a dry-feed entrained-flow gasifier with water quench.

Therefore, the maximum outlet pressure from the gasifier is presumed to be around 50 bar.

• The outlet temperature from the gasifier is set to 1500 °C. • Composition of syngas has a major role in defining the steam-to-CO ratio and especially in defining the H2 partial pressure. b) Based on quench: • The quench temperature is set to 800 °C. In order to reach this temperature the syngas has to be quenched with water and not steam.

• With pressure of 50 bar, equal to syngas, quench water can only reach temperatures up to 264 °C, otherwise it would begin to vaporise.

• Steam-to-CO ratio is also dependent on the temperature of quench water. Maximum ratio 916

Fig. 5. Schematic presentation of the reactor model c) Based on MR: • For modelling the tubular MR with gauge measurements L = 40 m and a = 10 m was selected (see Fig. 5). The outer tubular membrane diameter was set to the value 2r = 7.5 cm.

• The operational temperature range of the MR is presumed to be between 700 and 900 °C. • The pressure on the retentate side is defined by the outlet from the gasifier. In this study it is presumed to be 50 bar.

• Low pressure on the permeate side increases the driving force through the membrane.

However, from the point of energy loss needed for compression the minimum pressure should not be less than 1 bar.

• In order to improve the driving force the N2 sweep gas is introduced. In view of syngas the flow of N2 is in counter-current direction where H2-to-N2 molar ratio at the exit is 1:1.

• Since the WGSR is an exothermal reaction, the reactor is water cooled. In the present study the MR will operate at isothermal conditions.

• The desired CO conversion is set to be around 95%.

• Based on the temperature range, pressure conditions, hydrogen purity and permeability, the selected membrane material is Palladium based with permeance: k′ = 3·10-4 mol m-2 s-1 Pa-0.5.

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 911 -926

Fig. 6. Presentation of the numerical model for calculations of WGSR kinetics • • • •

Due to the high temperatures no catalyst is used for the WGSR. In this model it is presumed that the membrane material can withstand the acid environment therefore no pre-cleaning of syngas is needed.

The model of the reactor is based on a plug flow reactor. This means that there is no change in concentration in radial direction.

In the kinetic model it is presumed that there are no pressure losses on the retentate or on the permeate side.

where permeance was held at constant value k′ = 3·10-4 mol m-2 s-1 Pa-0.5.. This value was selected based on the experimental study made

by Ciocco et al. [5] where the permeance of pure Pd membrane k′ = 3·10-4 mol m-2 s-1 Pa-0.5. was achieved.

Based on Fig. 6 the Law of Conservation of Mass must apply for each section therefore, the molar flow of each specie was calculated with the following set of Eqs.: n RH 2 ,i +1 = n RH 2 ,i + nreaction ,i +1 − n H 2 ,i +1 , (22)

2.2 Numerical Model

nCO 2 ,i +1 = nCO 2 ,i + nreaction ,i +1 , (23)

The numerical model in Mathematica was created in such a way that each reactor tube was divided in NO = 10,000 infinitesimal sections with length Dx. The calculation started with values obtained from initial and boundary conditions, which gave the values for the first section. In general, the calculations were conducted in such a way that the results from section marked with index i were used to calculate the values indexed as i+1 until reaching the last NOth section. a) Retentate side: Combining Eqs. (5), (12) and (19) the reaction rate for section i+1 was calculated as:

nCO ,i +1 = nCO ,i − nreaction ,i +1 , (24)

, n H 2O ,i +1 = n H 2O ,i − nreaction ,i +1. (25)

ri +1 = k f ⋅ [CO ]i ⋅ [ H 2 O ] − kb ⋅ [ H 2 ]i ⋅ [CO 2 ]i . (20) 1/ 2

1/ 2

The flux in section i+1 was calculated with the following Eq.:

J i +1 = k ′ ⋅ ( pR H 2 ,i − pP H 2 ,i ), (21)

In the above set of Eqs. nreaction,i+1 and nH2,i+1 were defined as:

nreaction ,i +1 = ri +1 ⋅ NTUBES ⋅ ∆V , (26)

n H 2 ,i +1 = J i +1 ⋅ NTUBES ⋅ ∆ A. (27)

The molar flow of the retentate gas stream was a sum of all calculated molar flows and the inert species in this study (Ar, H2S, N2, COS, HCl, CH4 and others): n R ,i +1 = n RH 2 ,i +1 + nCO 2 ,i +1 + (28)

+ nCO ,i +1 + nH 2O ,i +1 + ninert .

The molar fraction of each species was calculated using the equation:

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Table 2. Activation energies and pre-exponential factors for forward and backward WGSR obtained from literature; the rate constants for WGSR were calculated at operating temperature 1073 K Eaf

Source Culbertson [8] Graven & Long [9] Bustamante [10]

[J/mol]

k0f

[(l/mol)0.5 s-1]

334.72 7.7·1045·T-10 (i) 281.58 5.0·1012 253.55 1.52·1010

Eab

[J/mol]

/ 238.49 218.40

k0b

kf

kb

/ 9.5·1010 6.65·108

6.10 3.10 0.22

/ 7.37 0.49

[(l/mol)0.5 s-1] [(cm3/mol)0.5 s-1] [(cm3/mol)0.5 s-1]

(i)The pre-exponential factor is also temperature dependant therefore the correct unit is K10 (l/mol)0.5 s-1.

x j ,i +1 =

n j ,i +1 n R ,i +1

, (29)

where j represents individual specie (H2, CO2, CO or H2O).

Similarly, the concentration of each species was calculated: n j ,i +1 pR p c j ,i +1 = ⋅ = x j ,i +1 ⋅ R , (30)

n R ,i +1 R ⋅ T R ⋅T where the pressure on the retentate side was held at constant value pR = 50 bar. The partial H2 pressure on the retentate side was calculated as:

pR H 2 ,i +1 = xR H 2 ,i +1 ⋅ pR . (31) In order to calculate heat duty the heat of reaction DHR needs to be calculated. This was done by using the van’t Hoff Relation based on source [6]: ∂ ∆H R = − R ⋅ T 2 ⋅ ln K eq ( T , p ), (32)

∂T where the logarithm of equilibrium constant Keq was obtained with the equation reported in Bustamante [10] which was based on the study from Singh and Saraf [7]:  9998.22  − 10.213 +   T   1  +2.7456 ⋅10−3 ⋅ T −  ln K eq = ⋅ . (33)  1.987  −6 2  −0.453 ⋅10 ⋅ T −   −0.201 ⋅ ln T    And the CO conversion XCO was deduced as follows: n − n X CO ,i +1 = CO ,in CO ,i +1 . (34)

nCO ,in b) Permeate side: Similarly as on the retentate side the Law of Conservation of Mass must also apply for each 918

section on the permeate side. The molar flow of H2 and N2 mixture was calculated as: n P ,i +1 = n P ,i − nH 2 ,i +1 . (35)

Because the molar flow of the N2 sweeping gas is constant throughout the reactor, the molar flow of H2 on the permeated side can be written as:

n P H 2 ,i +1 = n P ,i +1 − n N 2 . (36) Samilarly as on the retentate side, the molar fraction of each species was calculated using the Eq.: n j ,i +1 x j ,i +1 = , (37)

n R ,i +1 where j represents individual specie (H2 or N2).

The partial H2 pressure on the permeate side was calculated as: pP H 2 ,i +1 = xP H 2 ,i +1 ⋅ pP . (38)

where the pressure on the permeate side was held at the constant value pP = 1 bar.

2.3 Model Design in Aspen Plus The simulation was performed using the Aspen Plus reactor model unit where chemical reactions of WGS and side reactions (H2S and COS formation) were modelled. The input data were temperature, pressure, molar flow, composition of syngas and formulas of chemical reactions that take place inside the reactor. The shortcoming of the model was that the program had no information about the kinetics of the WGSR and hence no knowledge of the extent of the CO conversion and the H2 permeation through the membrane. These two parameters were calculated based on the kinetics of the WGSR and imported from Mathematica (see Fig. 4). That gave us more accurate predictions of the molar flow of H2 in the permeate stream and the achieved CO conversion

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 911 -926

in the WGSMR. Based on these two parameters and the initial and boundary conditions (see subchapter 2.1) the simulation of the WGSMR was performed.

The model of the MR was designed to achieve the isothermal temperature distribution throughout the entire reactor. This is done in such a way that indirect water cooling is applied in a co-current direction to syngas flow. To attain efficient heat removal, the water evaporates almost throughout the entire reactor and gets slightly superheated at the end of the reactor. It is presumed that the WGSR will be most intensive at the beginning of the reactor thus, most of the heat will be released in the first part of the reactor. The sweeping N2 gas enters the reactor in a countercurrent way. The reason for that is to keep the H2 partial pressure difference reasonably high in all parts of the reactor in order to achieve better flux through the reactor’s membrane.

3 RESULTS AND DISCUSSION 3.1 Calculations in Mathematica In the subchapters bellow the WGSR kinetics are discussed and calculations for the base case model are presented. A sensitivity analysis of the model is made by varying different parameters like the membrane’s permeance, length (indirectly the surface area) of the reactor and the molar fraction of H2 in the permeate stream.

3.1.1 WGSR Kinetics There is some discrepancy between the uncatalysed WGSR kinetic data available from

literature. For this purpose, three different studies were compared all using the Bradford mechanism [4] to describe the WGSR. Table 2 presents activation energies and pre-exponential factors obtained from these studies.

The rate coefficients for forward and backward reaction were calculated according to Eq. (3). As shown in Table 2, the rate coefficients calculated with values obtained from Bustamante [10] are one order smaller than those obtained with values from Graven and Long [9]. Even though Bustamante’s research [10] was conducted at elevated pressures (16.21 bar), and Graven and Long’s [9] only at atmospheric pressure, the results from Graven and Long’s study [9] are in better agreement with Culbertson [8] who conducted his study at even higher pressures and temperatures (196.5 to 496.4 bar and 1200 to 2100 K).

The calculations showed that the reaction rate according to Bustamante [10] is one order smaller than the one predicted from Graven and Long [9] (see Fig. 7). This has also been confirmed in the study from Culbertson [8], therefore the decision was made that the coefficients deduced from Graven and Long [9] will be used in the WGSR kinetic study.

3.1.2 Base Case Studies One of the purposes of this study was to determine if the permeation through the membrane is sufficient or if there is a build-up of H2 on the retentate side. First, the comparison was made by using the molar flows of both processes calculated

by Eqs. (26) and (27).

Only in the first two sections the reaction rate is slightly faster than the permeation (on Fig.

a) Bustamante’s model b) Graven & Long’s model Fig. 7. Rate of reaction based on different models Simulation of Water-Gas Shift Membrane Reactor for Integrated Gasification Combined Cycle Plant with CO2 Capture

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8 the difference is negative) after which the rate of permeation becomes faster than the production of H2. As seen from Fig. 8, the maximum difference is reached around the 400th section and the permeation rate becomes almost equal to the reaction rate around the 4000th section. The rate of reaction rises slightly at the end because at that point the concentration of H2 is close to zero and the effect of the backward WGSR almost disappears. The increase in the reaction rate also causes the rate of permeation to increase, because more H2 is produced, which immediately permeates through the membrane because the process is not permeation limited.

in the retentate stream at the beginning and H2 converted from CO until that point permeate through the membrane. From that point on, only the H2 that is converted from CO is permeating through the membrane.

Rapid permeation of H2 is responsible for molar fractions of H2O and CO to experience the first inflection point around the 400th section.

The fractions reach the second inflection point around 2000th section, where majority of H2 on the retentate side already permeated through the membrane. From that point on both fractions gradually reduce towards the end of reactor where the consumption of H2O and CO again slightly increases on behalf of increase in the reaction rate.

The molar fraction of CO2 is increasing fast in the first 2000 sections because of H2 permeation and because of CO conversion. This can be seen from a graph where after 2000 sections the rate of CO2 formation starts to decrease since virtually all of H2 that was present in the stream permeates through the membrane. At the end, the rate of CO2 formation slightly increases on behalf of the increase in the reaction rate.

Fig. 8. Difference between the rate of H2 production and the rate of H2 permeation in each section of the reactor As it can be seen from Fig. 9, the molar fraction of H2 actually rises at the beginning even though the rate of permeation is greater as the rate of formation. By constantly extracting H2 from the retentate stream, the joint molar flow also decreases. Since at the beginning the difference between the permeation and the production rates of H2 is not big enough, the H2 molar flow in retentate stream is decreasing but not as fast as the molar flows of H2O and CO. As a consequence, the molar fraction of H2 actually rises slightly until the 64th section where, on behalf of a sharp increase in the permeation rate, H2 starts to rapidly permeate through the membrane.

After reaching the peak at the 64th section the molar fraction of H2 starts to decrease fast until the 1000th section. Around the 2000th section the permeation already slows down considerably and around the 4000th section most of H2 present 920

Fig. 9. Change in molar fractions of species and the CO conversion throughout the reactor Based on the input values, the calculated molar composition of the retentate stream is presented in Table 3. Below, these values are also compared to the simulation results from Aspen Plus.

In order to keep the MR operating at isothermal conditions, approximately two thirds of all heat, produced from the WGSR, would need to be removed from the reactor in the first 2000 sections (see Fig. 10). The cooling tubes should be distributed in such a way that they would insure

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 911 -926

the reactor to operate at isothermal conditions and to prevent the formation of any hotspots in the reactor.

Table 3. Input and output values from Mathematica program species xi species xi

Input values (SYNGAS) H2O CO H2 CO2 0.410 0.375 0.154 0.052 Output values (GAS-2) CO2 H2O CO inert 0.754 0.148 0.080 0.018

inert 0.009

600 sections. Beyond that point permeation is faster but it remains fairly low. At the beginning permeation slightly rises because of the H2 buildup which increases the H2 partial pressure and enhances the flux through the membrane. The WGSR proceeds fast in the first 1000 sections and then it remains low through the entire reactor. The blue, dashed line shows the difference between both processes.

H2 3.0∙10-4

It should also be noted that by introducing the cooling tubes into the reactor the CO conversion may suffer minor penalties because the reaction volume for the WGSR will be reduced.

The shape of the curve resembles the curve of the CO conversion, because with more CO converted more heat is released. The calculated cumulative heat duty released from the WGSR is Q = 63.4 MW.

Fig. 10. Cumulative distribution of heat released in the WGSMR

Fig. 11. Difference between the rate of H2 production and the rate of H2 permeation with one order smaller permeance The graph in Fig. 12 shows the build-up of H2 molar fraction in the retentate stream. After the reaction rate becomes steady, the H2 fraction starts to gradually decrease on behalf of slow permeation. As the permeation is very limited, all of H2 does not succeed in permeating through the membrane, therefore in the end there is still around 4% of H2 left in the retentate stream.

3.1.3 Influence of Membrane’s Permeance on CO Conversion The influence of membrane’s permeance on the CO conversion was studied using the MR designed for base case studies. The only parameter that was changed was the membrane’s permeance.

In this analysis the permeance was set to the value k′ = 3·10-5 mol m-2 s-1 Pa-0.5. The calculations show that in the MR with such material permeance the processes would become permeation limited.

The graph in Fig. 11 shows that the reaction rate is faster than the permeation rate in the first

Fig. 12. Change in molar fractions of species and the CO conversion with one order smaller permeance

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The reactor achieves only 50% CO conversion because the extraction of H2 is too slow and it does not shift the equilibrium as far to the product side as in the base case scenario.

the CO conversion. By enlarging the membrane surface area, more H2 can permeate through, which shifts the WGSR more to the product side, and helps improve the CO conversion.

To further determine the influence of surface area, the reactor with 70 m in length and H2 molar fraction xH2 = 0.45 was simulated. The reactor achieved only 92.3% CO conversion. This shows that extending a reactor further would help achieve better conversions, but from economic point of view investment costs would be enormous to build such a big reactor. As discussed in earlier chapters, it would be better to use the catalyst because the surface area is already large enough and in the latter part of the MR the limiting process is the WGSR rate.

3.1.4 Influence of Reactor’s Surface Area and H2 Molar Fraction in Permeated Stream on CO Conversion The influence of the reactor’s surface area on CO conversion was studied by varying the length of the reactor. Additionally, the influence of different molar compositions in the permeate stream was studied. By changing both parameters Table 4 was obtained.

The table shows that reducing the molar fraction of H2 in the permeate stream would favour the CO conversion. This is due to the fact that the partial pressure of H2 is lower and thus, enhances the flux through the membrane which indirectly affects the CO conversion.

Increasing the reactor’s surface, by extending the length of the reactor, also favours

3.2 Simulation Results in Aspen Plus The model simulated on the basis of the calculations obtained from the Mathematica base case model is presented in the previous subchapter.

Table 4. CO conversion at different H2 molar fractions and lengths of reactor L = 30 m XCO

xH2

nN2

L = 40 m nP,H2

[mol/s]

[mol/s]

XCO

nN2

L = 50 m nP,H2

[mol/s]

[mol/s]

XCO

nN2

nP,H2

[mol/s]

[mol/s]

0.45

88.0

3281.43

2684.73

90.1

3335.05

2728.57

91.2

3363.35

2751.82

0.5

87.0

2664.53

2664.49

89.1

2707.63

2707.57

90.2

2730.35

2730.3

0.55

86.0

2162.35

2642.9

88.0

2196.87

2685.16

89.1

2215.03

2707.32

Table 5. Molar compositions of gas streams(see Fig. 13) species xi

CO 0.560

H2 0.230

species xi

H2O 0.410

CO 0.375

species xi

CO2 0.769

H2O 0.140

species xi

CO2 0.835

H2O 0.146

species xi

H2 0.5

N2 0.5

922

Rawgas (on the exit from the gasifier) H2O CO2 Ar H2S N2 0.120 0.077 4.8∙10-3 4.5∙10-3 4.0∙10-3 Syngas (after quench) H2 CO2 Ar H2S N2 0.154 0.052 3.2∙10-3 3.0∙10-3 2.7∙10-3 Gas-2 (retentate stream on the exit of the MR) CO Ar H 2S N2 H2 0.072 6.3∙10-3 5.9∙10-3 5.3∙10-3 5.3∙10-4 Gas-3 (after catalytic burner) SO2 Ar N2 O2 CO 6.2∙10-3 5.2∙10-3 2.0∙10-3 0 6.4∙10-3 GT-gas (permeate stream on the exit of the MR) CO H 2O CO2 Ar H2S 0 0 0 0 0 Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

COS 3.9∙10-4

HCl 6.7∙10-5

COS 2.6∙10-4

HCl 4.5∙10-5

COS 5.2∙10-4

HCl 8.8∙10-5

H2 0

H2S 0

COS 0

HCl 0

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This chapter deals with the influence of absolute pressure on the permeate side as regards to the power demand for compression of H2/N2 mixture.

3.2.1 Base Case Studies The input variables for the simulation in Aspen Plus were the same as the values for model in Mathematica. The only additional data that the model in Aspen needed were the values for CO conversion and the molar flow of permeated H2. On the basis of the simulation, the molar compositions of gas streams were obtained (see Table 5). The other calculated data are presented in the flowsheet diagram presented in Fig. 13.

The molar fraction of CH4 was left out of this table because the gasifier model in Aspen Plus predicts that only a small fraction of methane xCH4 =3.8·10-5 is formed.

The values predicted from the simulation in Aspen Plus and the calculations from Mathematica are in good agreement. In general, the model in Aspen predicts that more CO should be converted for a given CO conversion, because the fractions of CO2 and H2 are slightly higher and the fractions of CO and H2O are slightly smaller as predicted in Mathematica. The calculated heat duty is in very good agreement between both models.

At the exit of the catalytic burner, the temperature of CO2 rich stream reaches 1200 °C; therefore, there is a great amount of heat available for other processes. A considerable amount of SO2 present in the stream can be cleaned of sulphur using the conventional Flue gas desulphurisation (FGD) technologies. At this point it should be mentioned that the MR can also be used for direct thermal decomposition of H2S. Studies on this topic have already been conducted by several authors [12] to [14] and this could be one of the future features integrated in the WGSMR. 3.2.2 Influence of Absolute Pressure on Power Demand for Compression The influence of the absolute pressure at the permeate side on the power demand for compression of H2 and N2 mixture was studied at three different pressures and at constant H2 molar fraction xH2 = 0.5.

The change in the absolute pressure on the permeate side effects the H2 permeation through the membrane because the partial pressure is also changed. As a consequence different H2 molar flows on the exit of the MR are obtained. To adequately compare the power demand at different operating conditions, the specific work wcomp was introduced and defined as:

Fig. 13. Flowsheet diagram used for WGSMR simulations Simulation of Water-Gas Shift Membrane Reactor for Integrated Gasification Combined Cycle Plant with CO2 Capture

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wcomp =

Pcomp n P

,

(39)

where Pcomp is the joint power demand of all stages and nP the molar flow of the permeate stream. The calculated data are presented in Table 6.

Table 6. Comparison of specific work for compression of H2 to 26 bar p [bar] Pcomp [MW] nP [mol/s] wcomp [kJ/mol]

1

2

3

68.497

51.127

45.475

5415.2

5222.2

5075.6

12.65

9.79

8.96

The calculated data show that specific power demand for the base case (1 bar) is relatively high compared to other two cases. In cases of 2 and 3 bar the power demand is not that different; therefore, operating at 2 bar would be an inviting option. In the future, it would be reasonable to study the trade-off between the higher H2 production and the lower power consumption for compression.

4 CONCLUSIONS The WGSR kinetics was described with Bradford mechanism [4] together with the kinetic data obtained from Graven and Long’s [9] experimental results. The calculations performed in Mathematica were used to calculate the CO conversion and the H2 flux through the membrane.

The calculated data were then imported in Aspen Plus as input (or initial) values for simulations.

The model in Aspen Plus was designed as a black box model, simulating the processes based on the initial and boundary conditions and the calculations obtained from Mathematica. The results, derived from both models, are in good agreement with each other and support the correctness of the designed model.

Based on the operational conditions of the designed reactor the membrane material based on Pd was selected. The calculations show that the membrane material with permeance k′ = 3·10-4 mol m-2 s-1 Pa-0.5. would be sufficient for the CO conversion because the process would be reaction and not permeation limited. The CO conversion could be enhanced by increasing the 924

reaction rate with the use of high temperature catalyst or with higher process temperatures.

Other studies emphasized that Pd also has some catalytic properties for the WGSR but this was not included in the designed model. However, in the future it would be reasonable to determine the catalytic properties of Pd to accurately predict the WGSR rate, as well. It was established that one order smaller membrane permeance would reduce the CO conversion to 50% because the process would be permeation limited.

Based on the selected boundary conditions the designed model predicts that 89.1% CO conversion can be achieved. With length increase of 75% and H2 molar fraction of xH2 = 0.45 in the permeate stream 92.3% conversion was achieved.

But from economic point of view the gain in CO conversion would not justify the investment cost for such a big reactor.

Changing the absolute pressure on the permeate side shows that the trade-off between the CO conversion and the power demand for compression of H2/N2 gas stream needs to be approached by further research. The specific power demand shows that it would be reasonable to set the absolute pressure to 2 bar.

The retentate stream is fed to the catalytic burner, where the unconverted CO, unpermeated H2 and H2S are burned to produce additional heat.

Afterwards, the stream contains the SO2 which can be removed with conventional FGD technologies.

Another possibility for the MR in the future may be the thermal decomposition of H2S. However, in order to incorporate the process into the MR, it needs to be studied further.

In general, the study shows that the WGSMR is feasible, but there are still many issues that need to be addressed. First of all, the membrane material must retain its structural stability and permeability during the operating conditions that are stated. If the material cannot withstand the acid environment, pre-cleaning of syngas would be needed. Proper cooling of the reactor also needs to be designed to avoid hotspots and to keep the reactor operating at isothermal conditions. By introducing cooling tubes into the reactor, the CO conversion may suffer minor penalties because less reaction volume will be available for the WGSR to take place.

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

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5 ACKNOWLEDGEMENT The authors would like to thank the reviewers for their constructive criticism, which helped to improve the quality and understanding of this article.

6 NOMENCLATURE General acronyms ASU Air Separation Unit CCS Carbon Capture and Storage FGD Flue Gas Desulphurisation IGCC Integrated Gasification Combined Cycle MR Membrane Reactor PSA Pressure Swing Adsorption WGS Water-Gas Shift WGSR Water-Gas Shift Reaction WGSMR Water-Gas Shift Membrane Reactor Physical quantities Outer diameter of tubular membrane, [m] 2r [A] Concentration, [mol/m3] a Partial order of reaction a Width and height of MR, [m] c Molar concentration, [mol/m3] Membrane’s surface area of infinitesimal ∆A

section, [m2] ∆HR Heat of reaction, [J/mol] ∆V Volume surrounding infinitesimal section

of the membrane, [m3] ∆x Length of infinitesimal section, m Ea Energy of activation, [J/mol] Energy of activation for backward Eab

WGSR, [J/mol] Energy of activation for forward WGSR, Eaf [J/mol] index representing the ith section in i

numerical calculations index representing individual gas specie j J Flux through membrane, [mol/(m2 s)] k Rate constant, [(cm3/mol)0.5 s-1] kb Rate constant for backward WGSR, [(cm3/mol)0.5 s-1] kf Rate constant for forward WGSR, [(cm3/mol)0.5 s-1] k0 Pre-exponential factor, [(l/mol)0.5 s-1] k0b Pre-exponential factor for backward

WGSR, [(l/mol)0.5 s-1] k0f Pre-exponential factor for forward

WGSR, [(l/mol)0.5 s-1]

k’ Membrane’s permeance,

[mol m-2 s-1 Pa-0.5.]

Keq Equilibrium constant L Length of reactor, [m] n Molar flow, [mol/s] Molar flow of retentate gas stream, nR [mol/s] Molar flow of permeate gas stream, nP [mol/s] Molar flow of CO, [mol/s] nCO nCO2 Molar flow of CO2, [mol/s] Molar flow of H2 permeating through nH2

membrane, [mol/s] nH2O Molar flow of H2O, [mol/s] ninert Molar flow of inert gases, [mol/s] nN2 Molar flow of N2 (carrier gas) on

permeate side of the membrane, [mol/s] nreaction Molar flow of species produced or

consumed during the WGSR, [mol/s] nPH2 Molar flow of H2 on permeate side of the

membrane, [mol/s] nRH2 Molar flow of H2 on the retentate side of

the membrane, [mol/s] NTUBES Number of tubes in MR p Pressure, [bar] pP Pressure on permeate side of MR, [bar] pPH2 Partial pressure of H2 on permeate side

of the membrane, [bar] pR Pressure on retentate side of MR, [bar] pRH2 Partial pressure of H2 on retentate side of

the membrane, [bar] Pcomp Power demand for H2 compression, [W] Heat duty, [W] Q r Rate of reaction, [mol/(cm3 s)] rb Rate of reaction for backward WGSR, [mol/(cm3 s)] rf Rate of reaction for forward WGSR, [mol/(cm3 s)] R Universal gas constant, [J/(mol K)] T Temperature, [K] wcomp Specific work for H2 compression, [J/mol] x Molar composition or molar fraction xRH2 Molar fraction of H2 on retentate side of

the membrane xPH2 Molar fraction of H2 on permeate side of

the membrane XCO CO conversion

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7 REFERENCES [1] Donaldson, A.M., Mukherjee, K.K. (2006).

Gas turbine “refueling” via IGCC POWER Magazine, from: http://www.powermag.com/ gas/Gas-turbine-&quotrefueling%22-viaIGCC_540_p3.html, accesed on 2009-06-14.

[2] Ladebeck, J.R., Wagner, J.P. (2003). Catalyst development for water-gas shift. Handbook of Fuel Cells - Fundamentals, Technology and Applications, vol. 3, part 2, p. 190-201.

[3] Killmeyer, R., Rothenberger, K., Howard, B., Ciocco, M., Morreale, B., Enick, R., Bustamante, F. (2003). Water-gas shift membrane reactor studies. US DOE, NETL, Berkeley.

[4] Bradford, B.W. (1933). The water-gas reaction in low-pressure explosions. Journal of the Chemical Society, p. 1557-1563, DOI 10.1039/jr9330001557.

[5] Ciocco, M.V., Iyoha, O., Enick, R.M., Killmeyer, R.P. (2007). High-temperature water-gas shift membrane study. NETL, South Park.

[6] Lebon, G., Jou, D., Casas-Vazquez, J. (2008). Understanding Non-equilibrium Thermodynamics: Foundations, Applications, Frontiers. Springer-Verlag, Berlin, DOI: 10.1007/978-3-540-74252-4.

[7] Singh, C.P., Saraf, D.N. (1977). Simulation of high-temperature water-gas shift reactors.

Industrial and Engineering Process Design and Development, vol. 16, no. 3, p. 313-319, DOI:10.1021/i260063a012.

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[8] Culbertson, B., Sivaramakrishnan, R., Brezinsky, K. (2008). Elevated pressure thermal experiments and modeling studies on the water-gas shift reaction. Journal of Propulsion and Power, vol. 24, no. 5, p.

1085-1092, DOI:10.2514/1.31897.

[9] Graven, W.M., Long, J.F. (1954). Kinetics and mechanisms of the two opposing reactions of equilibrium CO + H2O = CO2 + H2. Journal of the American Chemical Society, vol. 76, p.

2602-2607, DOI:10.1021/ja01639a002.

[10] Bustamante-Londono, F. (2004). The hightemperature, high-pressure homogeneous water-gas shift reaction in a membrane reactor. Ph.D. Thesis, University of Pittsburgh, Pittsburgh.

[11] Enick, R., Bustamante, F., Iyoha, O., Howard, B., Killmeyer, R., Morreale, B., Ciocco, M. (2004). Conducting the homogeneous watergas shift reaction in a palladium/copper alloy membrane reactor at high temperature and pressure. NETL, Pittsburgh.

[12] Edlund, D. (1996). A membrane reactor for H2S decomposition. Bend Research, Inc., Morgantown.

[13] Ma, Y. H., Moser, W.R., Pien, S., Shelekhin, A.B. (1994). Development of hollow fiber catalytic membrane reactors for high temperature gas cleanup. Worcester Polytechnic Institute, Worcester.

[14] US 7,163,670 B2 - Patent, Agarwal, P.K., Ackerman, J. (2007). Membrane for Hydrogen Recovery from Streams Containing Hydrogen Sulfide, University of Wyoming, Wyoming.

Lotrič, A. ‒ Sekavčnik, M. ‒ Kunze, C. ‒ Spliethoff, H.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 927-935 DOI: 10.5545/sv-jme.2010.260

Paper received: 24.12.2010 Paper accepted: 02.09.2011

Investigation on the High Vacuum Tribological Characteristics of Surface Treated Nuclear Grade Stainless Steel Type AISI 316 LN at 25 to 500 °C Devaraju, A. – Elayaperumal, A. – Venugopal, S. – Kailas, S.V. – Alphonsa, J. Ayyannan Devaraju1,* – Ayyasamy Elayaperumal1 – Srinivasan Venugopal2 – Satish V. Kailas3 – Joseph Alphonsa4 1Department of Mechanical Engineering, Engineering Design Division, College of Engineering, Anna University, India 2Metallurgy and Materials Group, Indira Gandhi Centre for Atomic Research, India 3Department of Mechanical Engineering, Indian Institute of Science, India 4Facilitation Centre for Industrial Plasma Technologies, Institute for Plasma Research, India

Although some researchers have published friction and wear data of Plasma Nitride (PN) coatings, the tribological behavior of PN/PN Pairs in high vacuum environment has not been published so far. In order to bridge this knowledge gap, tribological tests under dry conditions have been conducted on PN/PN Pairs for varying temperatures of 25, 200, 400 and 500 °C in high vacuum (1.6×10-4 bar) environment. The PN coatings showed good wear resistance layer on the ring surface. The PN coatings were removed only from the pin surface for all the tests since it contacts at a point. The friction and wear were low at lower temperatures and it eliminated adhesion between the contact surfaces until the coating was completely removed from the pin surface. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: nuclear materials, surface treatment, PN layer, elevated temperature, high vacuum, tribological characteristics 0 INTRODUCTION Austenitic stainless steels exhibit good results in corrosion, vacuum, CO2 and argon environments [1]. However, this material is restricted to be used directly due to its low surface hardness, high friction and strong adhesion when it is sliding against itself and other metals. Hence, it is to be coated to reduce or eliminate these problems. Among many surface modification techniques, plasma nitriding is the most successfully and widely employed surface treatment to coat the austenitic stainless steels at different temperatures, durations and gas mixtures.

It is well known that plasma nitriding offers many advantages over traditional gas nitriding and salt bath nitriding, particularly in terms of reduced gas consumption, reduced energy consumption and the removal of environmental hazards [2] to [7].

In plasma nitriding process, the sputter cleaning effectively removes the oxide film (Cr2O3) from the stainless steel surface and thus accelerate the nitrogen mass transfer [4]. Hence, plasma nitriding has been selected for the current study.

In general, when the metals are in sliding contact under high vacuum environment, they exhibit tribological problems of high friction, stick -slip motion and high wear due to strong adhesion between the contact surfaces. In high vacuum environment, metals will become atomically clean due to the absence of oxygen or oxide layer and moisture which usually separates the sliding surfaces and act as lubricants thereby reducing friction and wear [8] and [9]. Only few researchers have done the dry wear tests at elevated temperatures on plasma nitride (PN) coatings [10] and [11]. They have reported that the wear rate was reduced at elevated temperature due to the formation of oxide layer. The ceramic coating looses its hardness and shear strength at elevated temperature [12] and [13]. These studies are useful to know the tribological behaviors of nitride coatings in air environment. However, the wear behavior of plasma nitrided stainless steel in high vacuum environment has not been studied so far. Furthermore, the Prototype Fast Breeder Reactor (PFBR) which under construction at Kalpakkam, India, austenitic Stainless Steel type

*Corr. Author’s Address: Department of Mechanical Engineering, Engineering Design Division, College of Engineering, Anna University, Chennai-600 025, India, adevaa2008@rediffmail.com

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Fig. 1. Schematic view of the vacuum based high temperature Pin-On-Disc machine AISI 316 LN (316LN SS) is the main structural material with liquid sodium is being used as the coolant for transferring the heat from the reactor to the steam generators [14]. In PFBR, the many important components are sliding at the temperatures of 25 to 500 °C in a high vacuum environment. To simulate its condition, the wear tests were conducted in an equally matching environment. Therefore, in this work 316LN SS was plasma nitrided and its high vacuum tribological characteristics have been evaluated.

1 VACUUM BASED PIN-ON-DISC (VPOD) TRIBOMETER The schematic view of VPOD tribometer is shown in Fig. 1. The important parts of VPOD tribometer are the testing chamber, the sensor chamber, the disc rotating mechanism, the heating system, loading mechanism, the vacuum built up circuit, the control board and the data measurement system. The testing chamber is a hollow cylindrical shell of diameter 0.45 m and height of 0.35 m. Its base has been closed with a flat stainless steel circular plate. The top has been covered by hemi-spherical shell called sensor chamber. The cylindrical and hemispherical shells are double walled structure. There is a gap in the double walled structure and through this gap; the water is circulated to maintain the structure at room temperature. The testing chamber has been vacuum sealed for all the openings in which 928

high temperatures and high vacuum experiments are conducted. The rectangular hinged door (not shown in the diagram) has been fitted in front of the testing chamber for loading and unloading of specimens. An ‘eye piece’ has been fixed in the hinged door to observe the sliding interface between the pin and the disc during the experiments. The disc has been mounted on a horizontal shaft, which was connected to the a d.c servomotor through a timer-belt which avoids the slipping between the belt and pulley. The servomotor provides constant speed with plus or minus one revolution.

Fig. 2. Schematic view of the heater An electric resistance heater is used to obtain the different temperatures upto 600 °C. It has been placed in a box cover, so that the heat generated is completely directed to the area where the pin and disc are positioned (Fig. 2). The Proportional Integral Derivative (PID)

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controller has been used to maintain the steady temperatures with ±5 °C. The stationary pin is loaded horizontally against a vertically rotating disc as shown in Fig. 3, which is different from the conventional pin on disc tribometer where the rotating disc was mounted in the horizontal position [15]. The main feature of the VPOD is that it eliminates or reduces the trapping wear debris in the wear track (i.e., third body effect), which alters the wear rate and mechanism. The rotary and diffusion pumps have been connected in a series to the left side of the testing chamber. The initial vacuum is obtained using a rotary pump and the high vacuum (up to 10-5 bar) is maintained by the diffusion pump. A penny gauge is used to measure the pressure up to 0.001 bar and a pirani gauge is used to measure the pressure greater than 0.001 bar. The sensor chamber consists of load cell, Linear Voltage Displacement Transducer(LVDT) and the loading lever mechanism. When the dead weight is added in the normal load arm, the loading lever mechanism magnifies one and a half times, approximately. The load cell is capable of measuring loads from 1 to 2000 N. The linear movement of the pin towards the disc can be measured using LVDT with ±1 micron accuracy.

The control board contains the pressure indicator, the temperature indicator, the eater switch and Vacuum pump switch. All the data from the VPOD is acquired using a personal computer. The machine has been fixed on an anti-vibration stand.

Fig. 3. Schematic diagram of loading configuration of VPOD 2 EXPERIMENTAL METHODS 2.1 Specimen The 316LN SS was used as the substrate material in this study and its composition was in

wt. % C: 0.02, Ni: 12.1, Cr: 17.9, Si: 0.3, N: 0.07, Mn: 1.8, Mo: 2.4 and Fe: Balance. In order to reduce the material cost and difficulties involved during the post examination works, the discs have been prepared in the form of rings. The 316LN SS was cut from a plate and then machined to the ring geometry of 92 mm outer diameter, 72 mm inner with 1.05 mm radius at one end were tested. The samples were ground using various ranges of the Silicon carbide emery paper and then a mirror polished using diamond paste to obtain the roughness (Ra) of 0.04 µm. The pins and rings were plasma nitrided at 570 °C for 24 hr in a gas mixture of 20% N2 to 80% H2 to obtain thicker PN layer under a working pressure of 5 mbar.

2.2 The Microstructure, Microhardness and XRD Test The cross sectional microstructure of the PN coated surface is examined in a Scanning Electron Microscope (SEM) and its phases were identified by the X - ray Diffraction (XRD) test.

It is to be noted that the XRD spectrum was taken in the powder mode with Cu K alpha target and Ni filter. The hardness of the PN coatings was measured by using a Vickers microhardness tester at 25 g load.

2.3 Dry Wear Test The dry sliding wear tests were conducted on the PN coated ring against PN coated pin at 25, 200, 400 and 500 °C to evaluate the tribological characteristics of PN coatings using VPOD tribometer. All the tests were conducted at the constant sliding speed (0.0576 ms-1), sliding distance (103 m) and normal load (4.1 N) in high vacuum (1.6×10-4 bar) environment. The initial Hertzian contact pressure was found to be 2 GPa for the load of 4.1 N. Two or three experiments have been conducted for each condition for the reasonable reproducibility of the results. Prior to wear testing, the rings and pins were cleaned using acetone and then dried for 15 minutes. The electronic balancing machine with the accuracy of 0.00001 g was used to record the mass of material loss during wear tests. The morphology of worn surfaces of pins and rings were examined by SEM.

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The 2D optical profilometer was used to measure the depth of the wear track.

3.2 The Tribological Test on PN/PN Pair at 25 °C

3 RESULTS AND DISCUSSION

The PN coated ring– PN coated pin wear test was conducted for one hour for the sliding speed of 0.0576 ms-1, load of 4.1 N and at 25 °C. The variation of coefficient of friction (COF) as a function of the sliding distance is presented in Fig. 7. At 25 °C, the COF was 0.6 up to the sliding distance of 25 m and then, it increased continuously till 82 m sliding distance and it maintains steady at 1.8 (Fig. 7). The PN coated pin slides on the PN coated ring without the coating failure up to 25 m sliding distance. The pin coating started to fail at 25 m. This could be the reason for lower friction to the sliding distance of 25 m.

3.1 The Microstructure, Microhardness and XRD Results A cross sectional SEM micrograph of plasma nitrided 316LN SS sample is shown in Fig. 4. It reveals that the average thickness of the PN coating layer was 70 μm. The maximum coating thickness was 90 μm near to the edge.

The microhardness was measured along the depth on the cross sectional plasma nitrided 316LN SS sample and presented in Fig. 5. The surface hardness was 1040 HV25g which is five times higher than the hardness of substrate material (210 HV25g). In a certain depth (65 to 75 µm) below the nitrided surface, microhardness is reduced gradually and reached to substrate hardness at 300 µm. It is due to the nitrogen diffusion layer. The XRD analysis on the PN coating was performed over the 2θ range from 35 to 90°. The results of the XRD analysis carried on the PN surface (i.e., after the plasma nitriding process) are presented in Fig. 6. They reveal that the PN layer consists of CrN, Fe4N and Fe3N phases, in addition to the observation of the austenite reflections from the substrate material.

Fig. 5. Microhardness profiles of Plasma nitrided 316LN SS sample

Fig. 6. The XRD patterns of Plasma nitrided 316LN SS sample Fig. 4. The cross sectional SEM micrograph of plasma nitrided 316LN SS sample

930

When the sliding distance increased beyond 25 m, the real area of contact increased and the coating started to get removed from the pin surface. Therefore, instead of PN/PN contact, it became partly 316LN/PN in addition to PN/PN between the sliding distance 25 m and 82 m. Thus, the friction increased continuously from 0.6 to 1.8.

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The majority of the PN coatings were removed from the pin surface between the sliding distance 25 and 82 m. Beyond 82 m sliding distance, the contact almost became 316LN/PN and the friction coefficient was steady (1.8). However, the ring coatings were not removed. It was confirmed through the 2D optical profilometer measurement on the wear track which revealed that the worn depth is considerably low (-3.12 µm) compared to PN coating thickness (70 µm). The SEM micrograph on the wear track which appeared as smooth surface (Fig. 8a) whereas the pin surface clearly revealed that the PN coatings were removed and then the substrate material (316LN SS) was directly contacted with the hard and rough PN surface of the ring. Therefore, the substrate material of the pin was plastically deformed and came out of the contact region (Fig. 8b).

However, it produced low mean COF and metal loss comparatively (Figs. 9 and 10). Furthermore, the PN coating layer on the ring surface has been shown to be wear resistant whereas majority of the wear occurred on the pin surface since it contacts at a point during the sliding wear test. Thus, PN coatings have removed from the contact surface of the pin. The PN/PN Pairs produced a similar result in the air environment also [16].

the mean COF was the lowest (0.924) of all the tests (Fig. 9). The ring metal loss recorded was also low (Fig. 10) compared to higher temperature experiments. The SEM micrograph on the wear track was smooth with mild abrasive marks (Fig. 12a). The reason for the mild abrasive marks is that the small spherical or elliptical shape debris slide on the wear track. The wear debris collected during the experiment is shown in Fig. 12b. It contains much micron level spherical or elliptical shape debris. Therefore, the mild abrasion wear mechanism has been exhibited during this test.

However, the surface damage on the ring surface is insignificant. It confirmed that PN coatings maintain its strength at lower temperature.

Fig. 7. The variation of COF as a function of sliding distance at 25 °C for the PN coated ring PN coated pin Pair 3.3 The Tribological Test on PN/PN Pair at 200° C The PN coated ring vs. PN coated pin wear test was conducted for one hour at 200 °C. The variation of COF as a function of the sliding distance is presented in Fig. 11. The friction almost remained steady from the initial state. At 200 °C,

Fig. 8. SEM micrograph taken after tribological test for PN coated ring-PN coated pin Pair at 25 °C; a) wear track generated on PN ring; b) surface of the PN coated pin Alternatively, there could be a possibility that friction and wear remained low at lower temperatures because of the presence of

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contaminants between the contact surfaces, which could be the reason for low wear and friction at 200 °C in a vacuum environment. Similar results have been absorbed by Kazuhisa Miyoshi [8].

Therefore, PN coating eliminates the adhesion between the surfaces even in a high vacuum environment at 200 °C.

Fig. 9. The Error bar for the mean COF at 25, 200, 400 and 500 °C

Fig. 12. SEM micrograph taken after tribological test for PN coated ring – PN coated pin Pair at 200 °C; a) wear track generated on PN ring; b) wear debris 3.4 The Tribological Test on PN/PN Pair at 400 °C Fig. 10. The Error bar for the mass of ring metal loss at 25, 200, 400 and 500 °C

Fig. 11. The variation of COF as a function of sliding distance at 200 °C for the PN coated ring - PN coated pin Pair 932

The PN coated ring – PN coated pin wear test was conducted for one hour at a speed of 0.0576 ms1, load of 4.1 N and at 400 °C temperature. The variation of COF as a function of sliding distance is presented in Fig. 13. The mean friction and ring metal loss was very high at 400 °C (Figs. 9 and 10) and the worn surface generated showed more grooves (Fig. 14a). The pin PN layer was removed from the pin contact region during the beginning stage of the experiment itself. Therefore, the pin surface plastically deformed in which the harder and sharper worn particles have embedded (Fig. 14b). At 400 °C the base metal (316LN SS) of pin was softened and held the worn particles. Then, it made grooves on the ring surface. Therefore, the wear mechanism involved was abrasive. The 2D

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 927-935

optical profilometer measurement on the wear track reveals that the worn depth was shallow (-5.2 µm) compared to PN layer thickness (70 µm). Therefore, PN coatings were not removed from the ring surface and PN layer on the ring surface was proved as an excellent wear resistant layer at 400 °C in a high vacuum environment, whereas the coatings were removed only from the pin surface since it contacts at a point during the sliding wear test.

than 400 °C. Furthermore, wear track shows the adhesive marks like pits (Fig. 16a). The weight loss scatter for 400 and 500 °C tests (Fig. 10) are bigger than lower temperature tests. This scatter is due to the PN layer of the removed pin with a small difference in the interval because of coating softening at elevated temperature. This is the reason why weight loss is also relatively high.

In air, the PN coatings exhibit high friction and wear at lower temperatures and low at elevated temperatures [9]. Whereas in high vacuum environment, it was just reverse. Although friction and wear has decreased slightly at 500 °C, it remained relatively higher than 200 °C.

Fig. 13. The variation of COF as a function of sliding distance at 400 °C for the PN coated ring - PN coated pin Pair 3.5 The Tribological Test on PN/PN Pair at 500 °C The variation of COF as a function of the sliding distance for the tribological pair of PN vs. PN pair at 500 °C is presented in Fig. 15.

Generally, all materials lose their hardness and mechanical properties at elevated temperatures.

The CrN coatings have softened at 500 °C [13].

In the current work, in addition to the elevated temperature, the surfaces are atomically clean because of a high vacuum environment. Therefore, the hardness and shear strength of PN coatings as well as substrate should have reduced for both, pin and ring surfaces. Therefore, adhesion occurs between the contact surfaces. The worn surface generated at 500 °C had more adhesive marks which can also be seen on the surface crack (Fig. 16a). The pin coatings were immediately removed.

Therefore, the substrate material (316LN SS) of the pin was plastically deformed and some worn particles embedded on the pin surface (Fig. 16b) and it made the grooves on the ring surface similar to 400 °C. However, depth of grooves are lower

Fig. 14. SEM micrograph taken after tribological test at 400 °C for PN coated ring – PN coated pin Pair; a) Wear track generated on PN ring; b) surface of the PN coated pin The PN/PN Pairs have produced low COF and low wear at lower temperatures since pin coatings were removed after the considerable

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sliding distance. As long as the PN coating layer was with the pin surface, it avoids the strong adhesion between contact surfaces even in high vacuum environment and maintains lower friction and wear. At higher temperatures and a high vacuum environment, the contact surfaces are free from contaminants.

Fig. 15. The variation of COF as a function of sliding distance at 500 °C for the PN coated ring - PN coated pin Pair

Therefore, direct metal to metal contact occurs and it results in pin coatings being removed immediately. Thereafter, the soft pin substrate material (316LN SS) directly contacts the hard PN coating surface of the ring and exhibits very high friction and high wear. Since the pin coatings were removed in all tests, the author could not conduct the experiments for PN/PN Pairs for longer sliding distance. Hence, pin (hemispherical end) on ring configuration was not recommended to investigate the PN/PN Pairs as the pin is contacted at a point which coatings removed shortly. However, PN coatings were not removed from the ring surface whereas coating failure and subsequently plastic deformation was observed only on the pin surface.

Hence, the PN coatings layer on the ring surfaces have been proved as an excellent wear resistance layer.

4 CONCLUSIONS The COF was the highest at 400°C and lowest was at 200 °C. The friction and wear was low at lower temperatures and high at higher temperatures and it was the reverse result of air environment.

The PN layers were removed from the pin contact surface for all the PN / PN tests due to high initial contact pressure (2 GPa) and pin substrate material was plastically deformed. Hence, it is concluded that the pin (hemispherical end) on disc configuration is not recommended to investigate the wear resistance of PN/PN Pairs.

However, PN/PN Pairs maintain lower COF as long as the PN layer was attached to the pin surface since it avoids the strong adhesion between the contact surfaces. Although PN coatings exhibit very high friction at elevated temperatures in a high vacuum environment, it was proved as an excellent wear resistant layer on the ring surface.

5 REFERENCES

Fig.16. SEM micrograph taken after tribological test at 500 °C for PN coated ring – PN coated pin pair; a) wear track generated on PN ring; b) surface of the PN coated pin 934

[1] Smith, A.F. (1986). Influence of environment on the unlubricated wear of 31 6 stainless steel at room temperature. Tribology International, vol. 19, p. 1-10, DOI:10.1016/0301679X(86)90088-5.

Devaraju, A. – Elayaperumal, A. – Venugopal, S. – Kailas, S.V. – Alphonsa, J.

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[2] Larisch, B., Brusky, U., Spies, H.J. (1999).

Plasma nitriding of stainless steels at low temperatures. Surface and Coatings Technology, vol. 116-119, p. 205-211, DOI: 10.1016/S0257-8972(99)00084-5.

[3] Musil, J., Vlček, J., Růžička, M. (2000).

Recent progress in plasma nitriding. Vacuum, vol. 59, p. 940-951, DOI:10.1016/S0042207X(00)00404-8.

[4] Borges, C.F.M., Hennecke, S., Pfender, E. (2000). Decreasing chromium precipitation in AISI 304 stainless steel during the plasma- nitriding process. Surface and Coatings Technology, vol. 123, p. 112-121, DOI:10.1016/S0257-8972(99)00506-X. [5] Collins, G.A., Hutchings, R., Short, K.T., Tendys, J., Li, X., Samandi, M. (1995).

Nitriding of austenitic stainless steel by plasma immersion ion implantation. Surface and Coatings Technology, vol. 74-75, p. 417424, DOI:10.1016/0257-8972(95)08370-7.

[6] Hannula, S.P., Nenonen, P., Hirvonen, J.P. (1989). Surface structure and properties of ion nitride austenitic stainless steel.

Thin Solid Films, vol. 181, p. 343-350, DOI:10.1016/0040-6090(89)90502-6.

[7] Marchev, K., Cooper, C.V., Blucher, J.T., Giessen, B.C. (1998). Conditions for the formation of a martensite single-phase compound layer in ion-nitrided 316L austenitic stainless steel. Surface and Coatings Technology, vol. 99, p. 225-228, DOI:10.1016/S0257-8972(97)00532-X. [8] Miyoshi, K. (1999). Considerations in vacuum tribology (adhesion, friction, wear, and solid lubrication in vacuum.

Tribology International, vol. 32, p. 605-616.

DOI:10.1016/S0301-679X(99)00093-6.

[9] Buckley, D.H. (1981). Surface effects in adhesion, friction, wear and lubrication.

Elsevier Scientific Publishing Company, New York.

[10] Karamis, M.B., Gercekcioğlu, E. (2000).

Wear behaviour of plasma nitride steels at ambient and elevated temperatures. Wear, vol. 243, p. 76-84, DOI:10.1016/S00431648(00)00426-9.

[11] Taktak, S., Ulker, S., Gunes, I. (2008). High temperature wear and friction properties of duplex surface treated bearing steels. Surface and Coatings Technology, vol. 202, p. 33673377.

[12] Staia, M.H., PérezDelgado, Y., Sanchez, C., Castro, A., Bourhis, E.Le., Puchi-Cabrera, E.S. (2009). Hardness properties and hightemperature wear behavior of nitrided AISI D2 tool steel, prior and after PAPVD coating.

Wear, vol. 267, p. 1452-1461, DOI:10.1016/j. wear.2009.03.045.

[13] Polcar, T., Parreira, N.M.G., Novák, R. (2007). Friction and wear behaviour of CrN coating at temperatures up to 500 °C. Surface and Coatings Technology, vol. 201, p. 52285235, DOI:10.1016/j.surfcoat.2006.07.121.

[14] Bhaduri, A.K., Indira, R., Albert, S.K., Rao, B.P.S., Jain, S.C., Asokkumar, S. (2004). Selection of hardfacing material for components of the Indian Prototype Fast Breeder Reactor. Journal of Nuclear Material, vol. 334, p. 109-114, DOI:10.1016/j. jnucmat.2004.05.005.

[15] Konca, E., Cheng, Y.T., Weiner, A.M., Dasch, J.M., Alpas, A.T. (2005). Vacuum tribological behavior of the non-hydrogenated diamondlike carbon coatings against aluminum: Effect of running-in in ambient air. Wear, vol. 259, p. 795-799, DOI:10.1016/j.wear.2005.02.034.

[16] Devaraju, A., Elayaperumal, A. (2010).

Tribological behaviors of plasma nitrided AISI 316 LN type stainless steel in air and high vacuum atmosphere at room temperature.

International Journal of Engineering Science and Technology, vol. 2, no. 9, p. 4137-4146.

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Paper received: 24.12.2010 Paper accepted: 18.10.2011

Experimental Equipment Research for Cryogenic JouleThompson Cryocoolers Comparison in IR Technology Sensors Sicovic, A ‒ Milinovic, M. ‒ Jeremic, O. Aleksandar Sicovic1,* ‒ Momčilo Milinovic2 ‒ Olivera Jeremic2 1Le-tehnika, d.o.o., Slovenia 2University of Belgrade, Faculty of Mechanical Engineering, Serbia

This paper shows experimental setup for research of Joule-Thompson cryocoolers performances integrated in Dewar vessels with temperature detector simulator of IR sensor. Experimental solutions of transient temperature regimes of detector cooling on cryogenic cooldown temperature by using nitrogen as the coolant have been analyzed. In experiments of transient temperature influences, three principal methods of the fluid flow regulation, were compared. Results of experiments have shown important differences in quality and rate of cooling for IR detector integrated in cooling system with or without coolant flow regulation. The chosen thermodynamic regime parameters in experiments were suitable for the real necessities of detector cooling in IR sensor application. Scheme of the experimental equipment, its components, method of research and monitoring of desired transient function of temperature versus time in cooling (cooldown regime) of detector, and thermodynamic performances of the overall micro-cryogenic cooler are shown in this paper. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: cool-down regimes, Joule-Thompson micro-cryogenic coolers, sensitive element, cryogenic gasses, temperature measurements, experimentation of transient 0 INTRODUCTION The basic purpose of cryogenic coolers is sustaining required temperatures of IR sensitive elements of different IR sensors like experimental sensitive camera, of the high resolution used in Infrared Thermographs method represented in paper [1], and other IR equipment, used in military and civil applications, [2] to [5]. Joule Thompson cryocooler is the representative type of this device and the most often used for these purposes. It is designed by dimensions of storage spaces which are usually, not more than several centimeters, and also to meet thermo-technical and general mechanical properties, necessary for rigid operational conditions. The device has to be operable inside the time interval, in different environmental conditions and under different external loads ranging between few Newton’s up to several hundreds of Newton’s. Furthermore, the Joule-Thompson cryocooler must provide the regulated temperature of cooling element inside the desired tolerance in the given time interval on the very low cryogenic temperature level, usually less than 100 K, [6] to [11]. These 936

requirements are further aggravated by the requirement of the desired temperature of IR sensors sensitive detectors achieved in a short time of less than a few seconds. The subject of this paper is a comparative analysis of tested transient temperature results, on the integrated laboratory assembly, presented in Fig. 1, the real designed Joule Thompson cryocooler, the real designed Dewar vessel, and an appropriate, experimental and designed simulation sample of temperaturesensitive element, as the model, of the real designed, expensive, mock-up detector.

The experimental set up of JouleThompson cryocooler assembly also includes the regulator and heat exchanger cooler packup, mounted in Dewar vessel, [12]. The image of the experimental, closed to real, design of the subassembly, used in this research is given in Fig. 1.

The regulator is a special subassembly that is separately integrated in the mockup Dewar vessel assembly with the heat exchanger module, [13], and joined with the detector simulator setup, which is all mounted in the internal space of cooler assembly. The aim of the test was to study the *Corr. Author’s Address: Le-tehnika, d.o.o., Šuceva 27, Kranj, Slovenia, aleksandarsicovic@yahoo.com

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optimum conditions of regulated or unregulated cooldown transient temperature regimes on the cooling detector sample as the model with properties that the real sensitive element requires in operation.

Fig. 1. Design of Dewar model vessel The difference between the mock-up experiment assembly, Dewar vessels with coolers and the original assembly used on IR sensors is in the sample taken as the sensitive element instead of real sensors special elements determined by the same requirements of temperature conductivity.

Emulation of detectors cooling rate on the mockup assembly is achieved by the specific heat

capacity of sensitive element sample, rearranged by the calculation with the sample thickness.

This corresponded to the effect of temperature conductivity rate, which was same as on the special element of real detector. Additional variations in experimental heat exchanges on the detector in the simulation process of Joule heat consumption rates, is also designed in experiments, (marked with D3 subassembly, Fig. 1). This was controlled by electric powered source, of 50 mW DC power supply, connected on the sample of sensitive element, made of prepared copper. Dewar vessel model, with JouleThompson cryocooler and nitrogen as coolant is universally used as an assembly for research of variable regulated or unregulated test conditions, required by sensor temperature sensitivity performances. The conditions are emulated by sensitive element through the loop D3.

Experimental testing of regulated and unregulated coolers was implemented by modifying the subassembly of cooler regulator, in the same unique, and universally designed, Dewar vessel.

Experimental research of the cooling performance test on the assembly using two representative regulator types inserted in the unregulated cooler assembly to compare the cooling and coolant efficiency designed in the same manner with modified regulation functions. The representative

Fig. 2. Measurement flow chart Experimental Equipment Research for Cryogenic Joule-Thompson Cryocoolers Comparison in IR Technology Sensors

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regulator control subassemblies used here were the continuously designed self-regulation type, and pulsing coolant flow regulation type. All tested cooler types used nitrogen as the coolant.

1 EXPERIMENTAL ASSEMBLY The setting method of experimental research of the universal Dewar vessel mock-up assembly, with cooler and emulation detector, in terms of different exchangeable coolant flow regulators, is given in Fig. 2.

A similar attempt has been done in the papers, [14] to [15], by Alexeev et al. and Luo at al., regarding mixed coolant and mixed JouleThomson cycle for different micro-coolers types.

The equipment consists of: • Controlling block equipment and following data; • Coolant supply control block equipment with capability to control; • Temperature sensor for detector temperature measurements; • Emulation detector with DC current supply; • Dewar vessel assembly with Joule-Thompson cryocooler; • Coolant regulator block inserted in a JouleThompson cryocooler (exchangeable regulator). Block data between positions 1 to 7, Fig. 2 represents controlled functions and measured data. The control functions are set by equipment and measured data, the desired environmental conditions and the desired controlled temperature closed behind the detector. The desired detector temperature and its time development as the transient cooldown performance was the focus of testing in these experiments. The evaluation of the quality of the experimental assembly with and without the use of regulators, was based on the criteria of the measured temperature and their time derivation. The controlling block can be used with different numbers of input functions versus on the type of exchangeable regulators requirements.

The controlling block can also be used for managing of power supplies when the model of the Dewar vessel with cooler is used without the regulator in an unregulated cooling manner. The coolant control and energy resources is achieved 938

on the equipment with the cryogenic gas, which is delivered to the Dewar vessel assembly through the cooler to the detector over the nozzle with fixed or variable cross-section, controlled by continuous or pulsing regulator or uncontrolled (in unregulated regimes). All cases of control functions and measured data integrated on the cryogenic gas equipment and mock-up Dewar vessel assembly with or without regulated JouleThompson cryocooler, are shown in Fig. 2.

Three types of measurements were performed: • without regulators, • with continuous flow regulation of coolant, and • with the pulsing flow regulated coolant.

The source of coolant in all three cases, were pressure vessels with nitrogen under pressure (360 bar), volume between 150 to 700 cm3, and the system of supply valves that provides pressure decreasing vs time [16] to [17]. An additional source of electric power 50 mW was turned on to simulate the constant Joule heat flux on the detector. In the case of unregulated cooler, feedback measurement branch through the A/D converter and the PC is used only for collecting data regarding the behavior of temperature detectors in real time. The same flow chart of data was used also with the continuous regulator.

Controls of flow function were analogous, in this case, from the value of differential pressure to the direct motion of flow rate executive body in the regulator, without using the PC in the loop. The integrated control of temperature measurement and coolant flow was realized with additional PID loop in which an analog signal of the measured temperature is returned back to the controller A/D converter over the PC and D/A converter in to the operating body of pulse regulator. PID then implements pulse movement of the executive body, which controls the flow, thereby PC was controlling the current value of flow. The methodological flow chart settings of the management experiment shown in Fig. 3 is similar to the equipment in paper [18].

The central block C, consists of the control loops for measuring the pressure and the surrounding temperature around the sensor (the Dewar vessel assembly with a cooler) and the sensor detector temperature Td. Other measuring

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Fig. 3. Block diagram of instalation and regulating loops, (secondary) are denoted as blocks A to D3. Their function is to control the coolant supply parameters. The integrated loops enabled the use of the module regulator R as a variable construction in the same experimental scheme. In this way it was able to examine the improvement of detector time and temperature characteristics, (sensitive elements), in a controlled or uncontrolled (unregulated) coolant supply. This experimental block diagram allows, for any modular, dimension, the adjusted design of Joule-Thompson cryocooler and the regulator R assembly to be used under the same conditions on the same equipment, with the same detector in the same Dewar vessel assembly.

It allows mutual comparison of the executive quality regulation of coolant flow.

This equipment permits changing a JouleThompson cryocooler assembly that meets the required dimensional demands of Dewar vessel and particularly research of performances on the developing Joule-Thompson cryocooler using the same detector in the Dewar vessel. The regulation of power heat capacity of the emulated detector can also be achieved by changing the source of electric power in the electric circuit loop in certain limits.

By setting environmental conditions in Block C (appropriate exit temperature and pressure), it was possible to simulate the behavior of the detector in the required environment in which the real sensors could operate. The regulation of different coolant types is, also, possible on this equipment.

The control of the operation mode in JouleThompson cryocooler and the sources of heat fluxes on real detector D3 can be regulated by a variation of pressure and flow of block A, B by the commands D1 and D2. Hence, this equipment simulates all the real parameters of the IR sensor in the desired range of temperatures, either in the transient regime (cooldown), in the operating regime (run-time) or integrated in both modes.

This universal equipment enables modular device exchange that emulates the specific conditions in which the Dewar vessel assembly, sensor and Joule-Thompson cryocooler operating as the key functional elements of cooled IR sensors.

2 CONTROLS OF EXPERIMENTS AND THERMODYNAMIC FLOW CHART The thermodynamic cycle of the experiment consists of the process that changes

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Fig. 4. Thermal cycle of cooled gas the values of the coolant states. These values change from the input to the output values of the Dewar vessel assembly by the extreme gradients in overall thermodynamic state values (Fig. 2 and cycles in Fig. 4).

The most rapid change in the cooling cycle is the density of coolant, ρ, which extremely changes the gradient sign just in expected regulating point, T3 in Fig. 4, of Joule-Thompson cryocooler, making the system extremely sensitive by state values in this cycle point. The mass flow rate of coolant, as the regulation cooler performance, is mainly determined by density ρ; cycle overall, is control by pressure p, while, temperatures T, are only desired guided values of states in the cycle points. It was useful to present relationships between guidance and control major values of state T and p, as the function of the main regulating thermodynamic value ρ, as graphical expressions of their thermodynamic changes in the cooling operations. It was also possible to make the process in the experimental testing set up visible by consequence on the mass flow disturbances intakes by the regulations over ρ, in the first proximity evaluations, in the rearranging guided and control T and p cycle values of state.

The quality of cooling and the procedure of thermodynamic control applied to the JouleThompson cryocooler in order to achieve the desired value of the cooling temperature on the detector, is shown in Fig. 4. The representation is based on the values of state equation for the cryogenic coolant taken from the input point 940

of coolant state 2 to the output point coolant state 5 (Fig. 3), in the open coolant cycle. This means that the process of cryogenic coolant 1-2, corresponds to the state in the pressure vessel (state 1) and expansion up to the entrance of the Dewar vessel (state 2), and have not relevancing influences in cycle, if the supply pipeline is short.

The next process 2-3, corresponds to the state of expansion of the cryogenic coolant in the cooler heat exchanger, while the expansion in the nozzle, which is regulated or unregulated, (fixed throat area), corresponds to the process 3-4. State 4 is the input state of coolant in the free space of the detector, while the detector temperature Td is more or less equal to the temperature in state 4. State 5 is the output state from the Dewar vessel to the environment (ambient conditions C, Figs. 3 and 4). Testing was realized for the three mentioned developing cases of integrated design sensor assemblies with or without the regulated coolant flow rate. An appropriately designed regulator was inserted in the same Dewar vessel and tested on the described universal equipment.

The experimental results of temperature decrease on the simulated detector Td (T4), are compared by rates in, the so-called, transient cooldown regime, as the main quality criteria for developing of sensor assembly. This criterion determines sensor readiness capabilities to start operating in the IR mode. In all the design cases mentioned above is the essential quality performance of the sensor as a feature for establishing rapid desired detector temperature

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of the sensor after the initiation of coolant flow.

The rate of the detector temperature decrease, beyond process 3-4 in the thermodynamic cycle is presented in Fig. 4. The key parameter in this process is cryogenic coolant state magnitude, before and after this process. The requirements of each specified type of regulators is to provide the best behavior of process 3-4 in and after the transient temperature regime related to the requirement to achieve and to maintain the temperature point T4 in the operating regime. The regime of maintaining temperature is well known as the ‘’run-time’’ [7], and depends on the above mentioned initial state of coolant magnitude. The Input flow in Joule Thompson cryocooler placed in Dewar vessel starts by the coolant state 2, Fig. 4, and is finished by the output state 5. Therefore, the flow which changes with the environment is directed by the conditions in the nozzle, process 3-4, by the input and output state magnitudes of coolant. Changes of these magnitudes, which have different disturbances, in the case of regulated and unregulated coolant flow, controls quality of run time processes stabilization after cooldown transient regime also provide regulation possibilities to their diminishing, in run time temperature stabilization, after these unsteady state transients. This is essence item for any developing type subassembly, consists of Dewar vessel with Joule-Thompson cryocooler aimed for detector cooling purposes.

In the detector temperature transient cooldown decreasing phase, all the three developed assembly types studied in this research, had the same values of nozzle throat area and were only maximally opened in this regime. In all three cases of design subassemblies the coolant flow rate achieved maximum. Consequently, the coolant state magnitudes are also equal in all three cases, so the regulated and unregulated flow rate reached the critical regime in the nozzle throat.

This means that the condition for the coolant is the same at any moment in time on the exit of the Joule-Thompson cryocooler and on the entrance to the detector during transient temperature decrease with a strong gradient. After reaching the desired temperature quickly, their further temperature regulation and stabilization, for the run time regime, expressed differences, depend of type of regulation, or consequences of unregulated mass

flow processes. Achieving the critical regime in the nozzle, the mass flow becomes possible for regulation uses as in the critical flow with variable output magnitude states by variation nozzle throat cross section. However, the mass flow rate exchanges with the environment from detector volume in an aggravated form because of the mass accumulating in the free volume of the detector, and the mass flow regulation became sensible on the value of states in this free volume. For the optimization of the detector temperature rates of changes after the cooldown transient regime, and its temperature tolerance field necessary for running time process stabilization and control initiation, the coolant mass flow regulation is welcomed. The disturbances of temperature on the detector are aggravated by the fact that pressure and temperature in a detector free volume increases by the flow accumulation that causes weak detector cooling. The temperature difference of coolant and detector, therefore, decreases, and the heat flux is removed from a detector decreases.

3 RESULTS The first impression is that there is an increase of the mass flow conveyed to the detector and that it can provide faster cooling and establishment of the desired temperature. The provided considerations joined with experimental experiences, have shown that the latter statement is only partially true, whereas the free volume of the cooler, where the detector and coolant are present, exchange the mass flow with the environment under thermodynamic leakage conditions. The conclusion is that any accumulation of the coolant mass, which is not regulated with the incoming flow, causes a temperature and pressure increase, in the part of volume space, where the detector is placed. Regardless of the increased mass in that space, temperature control by the heat exchange from the detector shows weaker performance.

This was visible in case of the unregulated JouleThompson cryocooler, which required longer time for the stabilization of the desired temperature, and their further regulation for the run time regime (Fig. 5; curve 1).

Consequently, the exchange of heat with less consumption of coolant by flow rate, can achieved more rapid in run time temperature

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stabilization and their maintaining in the process.

For the regulated coolers (curves 2 and 3) their time response sensitivities, controls accumulation of coolant in detector space, and by this feature, also control temperature differences between coolant and detector. The best control capabilities has pulsing regulator because of good short reaction time (curve 2, Fig. 5). Continuous regulated cooler have longer time response, because of analogous control loop. Consequently, its regulation capabilities allow gas accumulation in detector space, which makes transient detector temperature, toward run time stabilization regime, more unsteady (curve 3). All regulations are achieved by reducing the mass flow and avoiding the accumulation of gas in its free volume of the detector. Measurement error is considered summary for each unit assembly. Measurement results are presented with curves 1, 2 or 3, Fig. 5. The measurements were carried out under the same conditions in ten independent experiments.

Diagrams of temperatures changes as a function of time and their represent mean values, of measured magnitudes, for one typical model sensor assembly, are given in Fig. 5. Using the same design cryocooler with mass flow regulator, (diagrams 2 and 3), the standard deviation of temperature measurements, in the cooldown mode, has not changed in relation to the diagram

1, for unregulated cooler, and is approximately of STDEV = 0.07. It was determined by the end point of temperature, achieved in cooling versus time assumed as the end of cooldown mode taken as T = 100 K, for all three cooler types.

The relative error of a cooldown time, measured in relation to its average rate for all three types of model assemblies was less than 2.5%. Values of error that appeared larger could be the result of structural dimensional tolerances of individual subassemblies which resulted in thermodynamic parameters error, and cannot be attributed to neither cumulative nor constructive errors. The represented cooldown regimes for advanced design solutions of cryocoolers often use two principals of coolant mass flow rate improvement to accelerate achievement of run time temperature in transient regime. This design solutions are known as double action flow [17] and demand flow, [17], [7]. The double action flow design type is not considered in this paper. The demand flow type corresponds to the designed case of regulated continuous type by bellows and is considered in this paper. The performances of all the mentioned design types depend on the coolant, on the initial run total pressure, on the designed nozzle, on the simulated heat silk on the detector, on the desired run time temperature and general overall thermal mass of the Dewar sensor vessel. These data

Fig. 5. Comparative measured temperature profiles versus time 942

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are not completely known in referenced papers.

The nozzles with fixed, doubled and optimized demand flow, [17], Fig. 6, show more deviation of transient cooldown time caused by different mass flow rates operating in this regime.

Fig. 6. Cool-down performance of a fixed-orifice cryocooler (dash dotted line), a demand-flow cryocooler (continuous line), and a double-action cryocooler (dotted line), [17]

argon. Paper [7] uses nitrogen and also argon (not presented here) and [17] uses only argon as the coolant. As it is much cheaper in this paper, Fig. 5, only nitrogen coolant is used, for the achievement, of approximately the same run time temperature of about 100 K. Cooldown pressures are much different in papers [7] and [17], and also vary from the cooldown pressure in this paper. The regulation of cooldown in the paper [7] is for the multipurpose cooldown initial starting of principally different sensor assemblies. The design presented in [17] is a single used cooldown system as in this paper. The above mentioned differences between this paper and [17] include novel designed pulsed regulation system of flow control. The heat sink is not comparable in these papers. This also points at differences in transient cooldown times for all design types, represented in Figs. 5 to 7. JouleThompson cryocooler systems, which choose cooldown regimes through advance designed nozzles, are able to control a steady state run time regime by a regulation system, independent of transient cooldown, for the required temperature, (Fig. 5). The relative measurement errors are not mentioned in the papers [7] and [17].

4 DISCUSSION

Fig. 7. Cooldown process of the Joule Thompson refrigerator with nitrogen at various pressures, [7] The cooldown in Fig. 7 [7] shows variation dependant on pressure. These facts indicates more sensitivity of cooldown on the mass flow rate values. The cooldown temperature curves profiles in this paper, Fig. 5, are by form very similar to the experiments in papers [17], Fig. 6 and [7], Fig. 7, for both types of coolants, nitrogen and

The measured performances of all the three Joule-Thompson cryocoolers design cases are the essential developing base for upgrading the unregulated sensors by regulated solutions of both regulation cases, continuous or pulsing types of regulators. Scientifically precise methodologies of transient performance simulation understand the comparison of transient regimes of the cooling to the same final temperature. The design of the regulator and Dewar vessel with a cooler assembly and an inserted detector simulator determines the equipment for the so-called cooldown regime testing, which is same but different in transient temperature stabilization for each of the cooler design variations. The gradient of the detector temperature on the decrease in the cooldown regime was the same in all tested cases, but cooldown time was different depending on the transient run time temperature stabilization.

This is different for all three types corresponding with the following explanations.

The unregulated cooler continuously cools the

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detector in run - time (curve 1; Fig. 5) until the end of the full operation, and is not able to maintain a constant temperature. Tolerances of temperature changes in this transient run time temperature stabilization regime are acceptable for this type of sensor. The temperature instability in the run time regime was the consequence of the constant changes of exhaust conditions from the detector free volume during its cooling.

In the design case with continuous flow regulated cooler, transient regime is established on the desired temperature corresponding to the pressure difference between the environment and the free volume of the detector using an analog membrane pressure regulator. The operation in this regulator is based on the pressure differences between the environment condition and conditions in the free volume of the detector. Its reaction time is not rapid, and corresponds to the characteristic transient time (Fig. 5, curve 3).

Pulsing mode regulation is provided by using the direct detector temperature measurement loop. The loop is integrated over PC digital controller to change the amount of coolant flow through the nozzle and, control amount of the coolant mass in to the free volume space. This is achieved by using pulses of flow portions of the coolant, by opening and closing the throat nozzle in accordance with the continual leakage from the free volume in the environment, and to the cooling temperature Td, measured on the detector.

The method of achieving this temperature is accomplished by closed loop pulse flow regulation (Fig. 2, loop D/A-R) of PID controller type. Therefore, the transient regime measured to the same final temperature was not possible since the design of the process was conditioned by the assessment, considered temperature transient regime and their reaction time of detector. This is caused by detector free volume, and environmental relations, which were not independent of the regulation methodology.

Universal equipment provided the same sensor assembly with the Dewar vessel condition to be tested in terms of regulated or unregulated design solutions. In this sense, it was able to vary the thermodynamic cycle by changing the design and input or environmental parameters related to the cryocooler, or to the overall sensors. In further work on this equipment, experimental 944

testing and theoretical simulations, the new features of coolants and new types of detectors for the extended requirements of modern sensors that operate in different environmental civil and military functional conditions would be expected.

5 CONCLUSIONS This procedure provided unique experimental equipment and methods for measuring and checking the same functional parameters of the different model assemblies with the Dewar vessel joint with the Joule-Thompson cryocooler, regulator and IR detector. The testing equipment presented in this paper has the following appropriated unique performances: 1. Experimental integrated accessories are able to simulate environmental heat changes, reflected through the sensor focal plane array, as changeble heat sinks appeared in the real exploatation, Fig. 3, ambient control loop, and D3 loop. This capability is not expressed in equipment used [7] and [17].

2. Experimental integrated accessories, also have special properties to control mass flow rate of cryogenic coolant on the different type of the Joule-Thompson cryocoolers in the required critical flow point, Fig. 3, Tiristor PC control. This provides the possibility of testing for any of the designed cryocoolers type represented [7], [17] and [12]. The represented features provide for the testing of regulation effects for the mass flow rate control in the following: fixed orifices, demand flow types, double action types of cryocoolers, [17], [7], [12] but also provides the testing of special, regulated and digital controlled pulse types, novel designed and tested in this paper.

The method is applicable to both steady and unsteady transient temperature measurements of main subassembly functions for successful employment in the IR sensors.

The key question of the Joule-Thompson cryocooler efficiency is the heat transfer of two phase flow transformations of run time cooling and expansion through system inside of the Dewar vessel. Achieving temperature of detector, as the heat sink thermal mass through corresponding regimes, is a similar process to the two phase

Sicovic, A â&#x20AC;&#x2019; Milinovic, M. â&#x20AC;&#x2019; Jeremic, O.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 936-946

flow, the cooling method explained in [19]. The designed method is for the big PCM systems in environmental conditions. Considerations and model in paper [19], set up the question of how much intermediate obstacles influenced on the heat transfer efficiency, when the heat of phase transformations appears, and are they melting or does evaporation latent heat. For the cryocoolers, the latent heat is the main influenced performance linked with phase transformations of liquid to gas phase of coolant state. Heat transfer with appropriate obstacles of run two phase flows, rigidly influencing the phase changes exposing heat losses or heat sources depend on the coolant state. Paper [19] shows fin influences in the preserved flow system, on heat transfer cooling efficiency, in two phase flow condition. In the equipment presented in the cryocooler design, similar obstacles are avoided, to keep independant heat transfer in two phase flow processes, and to sustain unchanged state of liquid phase. The cooling duration of the emulated detector is the question of the fast two phase flow evaporation in the controlled gas environment as the heat transfer process from the coolant to the detector.

6 ACKNOWLEDGEMENTS This paper is a part of research on the project Interdisciplinary Integral Research 47029 supported by the Ministry of Science and Technology of the Republic of Serbia in 2011.

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[2] Semenov, V.I., Kuznecov, N.S., Bidenko, M.F., Lapunov, S.I., Komarov, N.B. (2004).

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[7] Hong, Y.J., Park, S.J., Kim, H.B. (2004). The Performance of Joule Thomson Refrigerator.

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[8] Hong, Y.J., Park, S.J., Choi, Y.D. (2009).

A Numerical Study of the Performance of a Heat Exchanger for a Miniature JouleThomson Refrigerator. Cryocoolers, vol. 15, p. 379-386.

[9] Chua, H.T., Wang, X., Teo, H.Y. (2006).

A Numerical study of the Hampson-type miniature Joule-Thomson cryocooler.

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[10] Xue, H., Ng, K.C., Wang, J.B. (2001).

Performance evaluation of the recuperative heat exchanger in a miniature JouleThomson cooler.

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[11] Ng, K.C., Xue, H., Wang, J.B. (2002).

Experimental and numerical study on a miniature Joule-Thomson cooler for steadystate characteristics. International Journal of Heat and Mass Transfer, vol. 45, p. 609-618, DOI:10.1016/S0017-9310(01)00165-X. [12] Bradley, P.E., Radebaugh, R., Huber, M., Lin, M.H., Lee, Y.C. (2009). Development of a Mixed-Refrigerant Joule-Thomson Micro-cryocooler. Cryocoolers, vol. 15, p.

425-432.

[13] ter Brake, H.J.M., Lerou, P.P.P.M., Burger, J.F., Holland, H.J., Derking, J.H., Rogalla, H. (2008). Micromachined Joule-Thomson coolers for cooling low-temperature detectors and electronics. IEEE Sensors Conferences, p. 1352-1355.

[14] Alexeev, A., Haberstroh, Ch., Quack, H. (1999). Mixed gas J-T Cryocooler with Precooling stage. Cryocoolers, vol. 10, p.

475-479.

[15] Luo, E.C., Gong, M.Q., Zhou, Y., Liang, J.T. (1999). Experimental Comparison

Experimental Equipment Research for Cryogenic Joule-Thompson Cryocoolers Comparison in IR Technology Sensors

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of Mixed-Refrigerant Joule- Thomson Cryocoolers with two types of counterflow Heat Exchanger. Cryocoolers, vol. 10, p.

481-486.

[16] Maytal, B.Z. (2000). Flow rate pressure dependence of a fixed orifice JouleThomson cryocooler. Advances in cryogenic engineering, vol. 45, no. I, p. 323-328.

[17] Tzabar, N., Lifshiz, I., Kaplansky, A. (2008).

Fast cool-down J-T cryocooler to 88 K.

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Advances in cryogenic engineering, vol.

53B, p. 1025-1032.

[18] Arkhipov, V.T., Yakuba, V.V., Lobko, M.P., Yevdokimova, O.V. (1999). Multicomponent Gas Mixture for J-T Cryocooler.

Cryocoolers, vol. 10, p. 487-495.

[19] Stritih, U., Butala, V. (2011). Energy savings in Building with a PCM Free cooling system.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 2, p. 125-134, DOI:10.5545/sv-jme.2010.066.

Sicovic, A ‒ Milinovic, M. ‒ Jeremic, O.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, 947

Corrigendum Structural Analysis of an Articulated Urban Bus Chassis via FEM: a Methodology Applied to a Case Study Dario Croccolo* ‒ Massimiliano De Agostinis ‒ Nicolò Vincenzi

The following is correction for typographical error in the paper published in Strojniški vestnik Journal of Mechanical Engineering, vol. 57, no 11, p. 799-809.

Page 799, DOI code DOI:10.554.5/sv-jme.2011.077 should read: DOI:10.5545/sv-jme.2011.077 where 10.5545 represents a prefix of a DOI code assigned to SV-JME.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12 Vsebina

Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 57, (2011), številka 12 Ljubljana, december 2011 ISSN 0039-2480 Izhaja mesečno

Uvodnik Povzetki člankov Tine Tomažič, Vid Plevnik, Gregor Veble, Jure Tomažič, Franc Popit, Sašo Kolar, Radivoj Kikelj, Jacob W. Langelaan, Kirk Miles: Pipistrel Taurus G4 – o nastanku in evoluciji zmagovalnega letala na tekmovanju NASA Green Flight Challenge 2011 Jesús Meneses, Cristina Castejón, Eduardo Corral, Higinio Rubio, Juan Carlos García-Prada: Kinematika in dinamika kvazipasivnega dvonožnega robota “PASIBOT” Nikola Korunović, Miroslav Trajanović, Miloš Stojković, Dragan Mišić, Jelena Milovanović: Analiza stacionarnega kotaljenja pnevmatike po bobnu z metodo končnih elementov in primerjava z eksperimentom Ivo Pahole, Dejan Studenčnik, Karl Gotlih, Mirko Ficko, Jože Balič: Vpliv rezkalnih strategij na vzdržljivost orodij za kovanje Ljubodrag Tanovic, Pavao Bojanic, Radovan Puzovic, Sergey Klimenko: Doba uporabnosti in obraba orodij iz polikristaliničnega kubično kristaliziranega borovega nitrida (PCBN) pri struženju amorfno-kristaliničnih prevlek na osnovi železa Andrej Lotrič, Mihael Sekavčnik, Christian Kunze, Hartmut Spliethoff: Simulacija membranskega reaktorja za konverzijo CO z vodno paro za kombinirano plinsko-parno elektrarno z uplinjevalnikom in zajemom CO2 Ayyannan Devaraju, Ayyasamy Elayaperumal, Srinivasan Venugopal, Satish V. Kailas, Joseph Alphonsa: Raziskava triboloških lastnosti površinsko obdelanega nerjavnega jekla za jedrsko uporabo AISI 316 LN v visokem vakuumu pri temperaturah od 25 °C do 500 °C Aleksandar Sicovic, Momčilo Milinovic, Olivera Jeremic: Eksperimentalna primerjava kriogenih Joule-Thompsonovih hladilnikov pri IR-zaznavalih

SI 181

SI 183 SI 184 SI 185 SI 186 SI 187 SI 188 SI 189 SI 190

Navodila avtorjem

SI 191

Osebne vesti Doktorske disertacije, magistrska dela, specialistična dela in diplome

SI 193

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 181-182

Uvodnik Poglejmo svoje dosežke in izkoristimo našo raznolikost “Niso razlike tiste, ki nas ločujejo. To je naša nezmožnost prepoznati, sprejeti in slaviti te razlike.” (Audre Lorde, 1934-1992). Čeprav Audre Lorde ni imela v mislih tehničnih znanstvenikov (inženirjev), ko je zapisala te besede, je primerno, da prepoznamo, sprejmemo in slavimo naše razlike. Moč naše tehnične družbe je v tem, da ni monolitna. Kot inženirji smo aktivni na najrazličnejših področjih: ukvarjamo se z gradivi, mehaniko, kinematiko, termodinamiko, energetiko in okoljem, mehaniko in robotiko, mehaniko fluidov, tribologijo, kibernetiko, industrijskim inženiringom, ustvarjamo politike in še več. Rezultat našega multidisciplinarnega sodelovanja je ta, da celota presega vsoto vseh naših delov. Opravljamo raziskave, ki temeljijo na osnovnih znanstvenih načelih. Opredeljujemo različne probleme in se jih lotevamo z različnimi osnovnimi principi. Mi znamo ovrednotiti fizikalne, mehanske, kemične, medicinske in druge osnovne prvine problemov, poiščemo inženirske rešitve in pomagamo soustvarjati politike, ki pozitivno vplivajo na naše življenje.

Naša znanstvena raznolikost se odraža v bogastvu tem, ki jih obravnava 92 člankov, objavljenih v 57. letniku revije SV-JME. Vprašati pa se moramo, ali ti članki resnično odražajo nove trende in uveljavljene prakse v strojništvu in z njim povezanih znanostih? Ali tematske številke revije SV-JME z izbranimi članki pokrivajo najboljše mednarodne konference? Pravi odgovor je samo v številu citatov, ki jih doseže posamezni članek v mednarodno priznanih revijah. Ker pa je uspešnost in kakovost revij, tudi SV-JME, odsev vrhunskosti recenzentov, bomo v prvi številki SV-JME 2012 objavili seznam recenzentov, ki so sodelovali pri čistih recenzijah člankov SV-JME v letu 2011. Vsakemu recenzentu smo posebej hvaležni za odlično sodelovanje in porabljeni čas. Ob tej priložnosti vabimo znanstvenike ter druge vrhunske in eminentne strokovnjake od

vsepovsod, da nam sporočijo pripravljenost za opravljanje recenzijskega dela, saj elektronski sistem recenzijskega postopka olajša delo. Vse informacije kakor tudi zgodovina revije SV-JME so dosegljive na spletni strani naše revije. Človek vstopa v naravo, v globalni svet, z izdelki svojih rok in razuma. Našim vnukom in našim zanamcem zapuščamo raznoliko tehniško dediščino, kar smo letos v Sloveniji proslavili s slavnostno akademijo in mednarodno konferenco inženirjev strojništva v počastitev dvestoletnice visokošolskega študija strojništva ter petdesetletnice Zveze strojnih inženirjev Slovenije. Skupaj se veselimo vsakega novega ali naprednega tehniškega izdelka oziroma dosežka, ki uvaja naprednejšo tehnologijo, prispeva k varovanju okolja, zmanjšanju rabe energije ali drugemu, kar naj bi prihodnim rodovom omogočalo varno in zdravo življenje z dodano stopnjo tehniške kulture. In med take svetovno odmevne tehniške raziskovalne dosežke lahko uvrstimo letošnji izjemno odmeven uspeh slovenskih raziskovalcev na področju letalstva.

Električno letalo Taurus G4 je septembra 2011 letelo z ekvivalentno porabo goriva 403 potniške milje na galono s hitrostjo 107 milj na uro in si priletelo nagrado Green Flight Challenge 2011 (GFC). Taurus G4 je tako letel z dvakrat manjšo porabo goriva in dvakrat hitreje kot Toyota Prius, oziroma skoraj šestkrat učinkoviteje kot tipično štirisedežno potniško letalo. Letalo je razvil in zgradil slovenski proizvajalec jadralnih letal s pomožnim motorjem in lahkih športnih letal.

Dve ekipi med 14 prijavljenimi na tekmovanju GFC sta pokazali, da je pogon letal z elektriko dejansko izvedljiv (in čudovit). Letali Taurus G4 in eGenius sta med preletom praktično neslišni, tudi na majhnih višinah. Elektrika je razmeroma poceni: pri ceni 12 centov na kilovatno uro lahko Taurus G4 za manj kot 7 evrov leti skoraj tri ure. Električna energija za akumulatorje na tekmovanju GFC je bila proizvedena v SI 181

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 181-182

Slavnostno akademija in mednarodna konferenca inženirjev strojništva v počastitev dvestoletnice visokošolskega študija strojništva ter petdesetletnice Zveze strojnih inženirjev Slovenije geotermalni elektrarni, zato so bili leti resnično brez emisij. Smo priče in soudeleženci pri rojstvu dostopnega in resnično zelenega potniškega zračnega prometa.

Električna letala gradijo novoustanovljena podjetja, mala podjetja in domači mojstri po vsem svetu. Na nebu v prihodnjih letih pričakujemo veliko električnih letal, kakor tudi nenehno izboljševanje zmogljivosti tehnologije akumulatorjev.

V slovenskem podjetju verjamejo, da do napredka vodijo le pogumne vizije in cilji. Če bo NASA sprejela njihov izziv, bo podjetje prispevalo 100.000 dolarjev od svoje nagrade Green Flight Challenge 2011 za prvo električno letalo, ki bo letelo hitreje od zvoka. Pričakujejo, da bodo nagrado lahko oddali v naslednjih petih letih. Ne pozabimo, da je bila komaj dve leti nazaj zamisel o 200 milj dolgem letu samo z električno energijo v akumulatorjih čista znanstvena fantastika, danes pa to zmoreta dve letali.

SI 182

Teh in številnih drugih izzivov prihodnosti se bomo lahko lotevali tudi zaradi nenehnega razvoja analitičnih metod, materialov, računalniške dinamike fluidov in drugih področij. Morda pa je naše največje bogastvo prav raznolikost, kreativnost in mladostnost naše tehnične družbe.

Poglejmo kaj zmoremo in kot tehnična družba izkoristimo velike priložnosti, da se lotimo veličastnih izzivov prihodnosti ter prispevamo k napredku in zdravju ljudi po vsem svetu.

Srečno!

Vincenc Butala Reference: Butala, V. (2010). Editorial – Anniversary: 55 Years of Strojniški vestnik – Journal of Mechanical Engineering. Strojniški vestnik – Journal of Mechanical Engineering, vol. 56, no.

12, p. 789-790.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 183

VABLJENI ČLANEK

Pipistrel Taurus G4 – o nastanku in evoluciji zmagovalnega letala na tekmovanju NASA Green Flight Challenge 2011

Tomažič, T. – Plevnik, V. – Veble, G. – Tomažič, J. – Popit, F. – Kolar, S. – Kikelj, R. – Langelaan, J.W. – Miles, K.

Tine Tomažič1,* – Vid Plevnik1 – Gregor Veble1 – Jure Tomažič1 – Franc Popit2 – Sašo Kolar3 – Radivoj Kikelj4 – Jacob W. Langelaan5 – Kirk Miles5 1 Pipistrel d.o.o. Ajdovščina, Advanced Light Aircraft; Slovenija 2 Aeroidea s.p., Slovenija; 3 Aerotech s.p., Slovenija; 4 Lupina s.p., Slovenija 5 The Pennsylvania State University, oddelek za letalsko in vesoljsko tehniko, Združene države Amerike Taurus G4 je koncept, namenjen zgolj za Nasino tekmovanje, hkrati pa poskusni zajček za mnoge nove tehnologije, ki bodo uporabljene pri nedavno predstavljenem štirisedežnem poslovnem letalu Panthera – letalu, ki bo na voljo v treh različicah bodisi z navadnim bencinskim motorjem, hibridnim motorjem ali v popolnoma električni različici. Prazna masa letala je 1065 kg, od tega je masa akumulatorjev 470 kg.

Maksimalna predpoletna masa pa znaša 1500 kg, kar pomeni, da ima Taurus Electro G4 435 kg uporabne nosilnosti in je torej pravo štirisedežno letalo. Edinstvena zasnova letala G4 je nastala tako, da so v podjetju zaradi zmanjšanja stroškov razvoja povezali dva trupa obstoječega letala Taurus Electro G2, ki so ju spojili s centralno, 5 metrov dolgo sekcijo, na kateri je nameščen električni oziroma hibridni pogon.

Energijo med letom črpa bodisi iz litij-polimernih akumulatorjev, ki so nameščeni v obeh trupih ter v sredinskem delu, po potrebi pa tudi iz generatorja, ki proizvaja elektriko. Letalo poganja brezkrtačni motor z močjo 145 kW, ki je nameščen na sredini centralne gredi med obema trupoma, in poganja posebej za to letalo izdelani propeler s premerom 2 metra. Skupni razpon kril znaša 21,4 metra. Podvozje je uvlačljivo in letalo ima izjemno malo upora, torej ima dobre jadralne lastnosti, kar je tudi pogoj za uspeh na tekmovanju. Čeprav ni bilo načrtovano kot jadralno letalo, izkorišča dobre jadralne lastnosti za to, da kar najbolje uporabi svoj električni pogon. To pomeni, da je energetsko varčno, kar je najpomembnejši kriterij tekmovanja. Namen razvoja in izdelave tega letala je uvrstitev in ubranitev naslova iz let 2007-2008 na tekmovanju CAFE/NASA Green Flight Challenge, ki se je odvijalo septembra 2011 na letališču Santa Rosa v severni Kaliforniji. Po koncu tekmovalnega obdobja bo letalo G4 uporabljeno za raziskovanje prihodnosti alternativnih letalskih pogonov.

Letalo Pipistrel Taurus G4 je zmagalo na tekmovanju NASA Green Flight Challenge 2011, ki ga je sponzoriral Google. Je hkrati največje, najtežje in najhitrejše letalo na električni pogon, trenutno edini električni štirised na svetu. Prispevek predstavi pot letala od ideje, preko načrtovanja do izvedbe, testiranj in udeležbe na tekmovanju. Objavljena so dejstva iz ozadja, tehnični diagrami in ideologija v ozadju koncepta letala. Avtorji prav tako razkrijejo tekmovalno taktiko in znanje, pridobljeno v postopkih testiranja. Hitra izdelava prototipa, sočasni razvoj na več področjih hkrati. Projekt načrtovanja in izdelave Taurusa G4 je trajal pičlih 5 mesecev, zato so bili pri izvedbi potrebni inovativni pristopi. Teoretična ozadja so multidisciplinarna in obsegajo vse od aeronavtičnih, strojniških, elektrotehničnih znanj, pa do statistike pri analizi in obdelavi podatkov s testiranj. Zmogljivosti letala Taurus G4 se z natančnostjo nekaj procentov (1 do 3 %) ujemajo s predvidevanji. Dokazano je bilo, da lahko z uporabo teoretičnih znanj v zelo kratkem času razvijemo izdelek, ki se obnaša predvidljivo in v skladu s predvidevanji, čeprav po svoji vsebini ni konvencionalen. Taurus G4 je po konstrukciji dvotrupec in spada v kategorijo t.i. multibody letal. Prednosti konstrukcije se kažejo v odlični porazdelitvi mase vzolž krila, pri tem pa se lahko doseže izdatno manjša konstrukcijska masa letala, s tem pa tudi manjša raba energije za določeni namen. Konstrukcijski vidik je mogoče uporabiti pri letalih različnih izvedb, kjer je sposobnost nošenja tovora velikega pomena. Pomen in uporabnost hitre izdelave prototipov, inovativne konstrukcijske rešitve, edinstvene aerodinamične rešitve in CFD-optimizacija. Znanja o električnem pogonskem sistemu, znanja za prihodnost načrtovanja letal z majhno specifično energijsko rabo. Izsledki so uporabni tako inženirjem kot raziskovalcem. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: električno letalo, učinkovitost konstrukcije, NASA Green Flight Challenge 2011, velik komplet akumulatorjev, modeliranje zmogljivosti, preskus letenja *Naslov avtorja za dopisovanje: Pipistrel d.o.o. Ajdovščina, Goriška cesta 50a, SI-5270 Ajdovščina, Slovenija, tine@pipistrel.si

SI 183

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 184

Prejeto: 01.10.2010 Sprejeto: 27.10.2011

Kinematika in dinamika kvazipasivnega dvonožnega robota “PASIBOT” Meneses, J. – Castejón, C. – Corral, E. – Rubio, H. – García-Prada, J.C. Jesús Meneses – Cristina Castejón* – Eduardo Corral – Higinio Rubio – Juan Carlos García-Prada Univerza Carlosa III. v Madridu. Oddelek za strojništvo, Skupina MAQLAB, Španija

Humanoidni roboti so dandanes sestavljeni iz velikega števila servomotorjev, ki izvršujejo gibanja z velikim številom prostostnih stopenj. Ena glavnih slabosti humanoidnih robotov je njihova teža in poraba energije – pri večini tovrstnih robotov odpade približno 30 % celotne teže na servomotorje in kablovje, več kot četrtina pa na reduktorje. Naše delo je zato usmerjeno v snovanje novih mehanizmov in kinematičnih verig, ki bi zahtevale manjše število servomotorjev ob neokrnjenih zmogljivostih robota. Na ta način bi se zmanjšala masa robota, s tem pa tudi poraba energije in celotni stroški.

V članku je predstavljen dvonožni robot PASIBOT z eno prostostno stopnjo in še posebej program za izračun njegove kinematike in dinamike, oziroma sil med členi in momentov za doseganje želenih gibanj.

PASIBOT je zasnovan na mehanizmu, ki ga je razvil Laboratorij za robotiko in mehatroniko v Cassinu (LARM), ob tem pa je zasledoval filozofijo doseganja čim nižjih stroškov.

Glavni problem je dimenzioniranje členov in motorjev, ki so potrebni za hojo robota. V prispevku je predstavljena kinematična in dinamična analiza PASIBOT-a s teoretičnega vidika, ki daje linearne in kotne položajne koordinate, hitrosti in pospeške za vse člene, kakor tudi vse sile in momente med členi za katerikoli trenutek znotraj enega koraka.

Prikazana je topologija, kinematika in dinamika robota PASIBOT, kakor tudi program, ki je bil razvit za pripadajoče izračune. Koda MATLAB® vključuje vso kinematiko in dinamiko kot funkcije časa, nekatere parametre, kot so velikosti in mase členov ter zahtevena hitrost, obremenitve itd., pa je mogoče tudi spreminjati. Rezultati so bili uporabljeni pri snovanju in konstruiranju realnega prototipa in so podlaga za krmiljenje gibanj.

Program je bil validiran s primerjavo rezultatov kinematike in rezultatov drugih komercialnih paketov. V članku je podanih nekaj računskih rezultatov, vključno s potrebnim navorom motorja za hojo pri različnih hitrostih in dodatnih obremenitvah.

Izpeljani so bili kinematični in inverzni dinamični izrazi za stanje, ko ima robot eno nogo v zraku, brez upoštevanja zdrsavanja nosilne noge na tleh. Stanje, ko sta obe nogi na tleh, ni bilo preučeno, potrebno pa bi bilo za izračun celega koraka. V prihodnjih raziskavah bi bilo treba upoštevati tudi dinamiko zdrsavanja noge na tleh.

Razvita koda je bila uporabljena za preučitev vedenja robota PASIBOT še pred procesom konstruiranja, s čimer se je zmanjšala zahtevnost procesa snovanja. Koda bo uporabljena tudi za krmiljenje hoje realnega prototipa. V predstavljeni numerični analizi je možno izpostaviti odvisnost obremenitve v boku od vhodne vrtilne hitrosti servomotorja. Ta podatek je bil uporaben pri izbiri motorja, saj je ta odvisna od nekaterih parametrov, ki jih je mogoče spreminjati v programu, kot so gostota, dimenzije členov, hitrost itd. Pridobiti pa bo mogoče tudi podatke o napetostih in deformacijah v vseh členih med hojo, ki so osnova za optimizacijo dimenzij členov. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: analiza mehanizmov, kvazipasivni dvonožni robot, hodeči roboti

SI 184

*Naslov avtorja za dopisovanje: Univerza Carlosa III. v Madridu, Oddelek za strojništvo, Skupina MAQLAB, Avda. De la Universidad, 30. 28911 Leganés, Madrid, Španija, castejon@ing.uc3m.es

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 185

Prejeto: 09.06.2011 Sprejeto: 30.09.2011

Analiza stacionarnega kotaljenja pnevmatike po bobnu z metodo končnih elementov in primerjava z eksperimentom Korunović, N. – Trajanović, M. – Stojković, M. – Mišić, D. – Milovanović, J. Nikola Korunović* – Miroslav Trajanović – Miloš Stojković – Dragan Mišić – Jelena Milovanović Univerza v Nišu, Fakulteta za strojništvo, Srbija

V članku je predstavljena primerjava rezultatov eksperimentov in rezultatov analize kotaljenja pnevmatike po bobnu, opravljene po metodi končnih elementov (MKE), z namenom validacije in izboljšanja modela pnevmatike na osnovi končnih elementov (KE). V članku so opisane oprema in metode, ki so bile uporabljene za eksperimentalno določitev lastnosti pnevmatike pri zaviranju in vožnji v ovinek ter koeficienta trenja dezena plašča.

Za podrobno simulacijo vedenja pnevmatik pri kotaljenju po ravni površini se običajno izvajajo analize po MKE. Rezultate analiz po MKE lahko v tem primeru neposredno primerjamo z rezultati eksperimentov, dobljenimi s stroji za preizkušanje pnevmatik z ravno površino ali s posebnimi vozili.

Takšna oprema pa zahteva veliko prostora in je tudi draga. Značilnosti kotaleče se pnevmatike je mogoče vrednotiti tudi na strojih z bobni, ki so manjši in cenovno dostopnejši, zahtevajo pa določene aproksimacije geometrije naležne površine. Za primerjavo rezultatov analize pnevmatik po MKE in rezultatov preizkusov na strojih z bobni je treba izvirni model pnevmatike s KE spremeniti tako, da bosta upoštevana tudi površina bobna in vrtenje bobna.

Zaradi učinkovitosti napovedovanja obnašanja pnevmatik pri vožnji v ovinek in zaviranju ter zmerne porabe računskih zmogljivosti je bila uporabljena stacionarna analiza kotaljenja (SSRA) na osnovi mešanega Eulerjevega/Lagrangeovega pristopa. Za preizkušanje vedenja pnevmatike med vožnjo v ovinek in zaviranjem je bil uporabljen stroj z bobni. Vrednost koeficienta trenja kot funkcija drsne hitrosti in kontaktnega tlaka je bila določena s pomočjo linearnega stroja za preizkušanje trenja na vzorcih dezena, odrezanih s pnevmatike.

Začetna korelacija med rezultati eksperimentov in numeričnih analiz je dobra, ob upoštevanju vseh aproksimacij, ki so bile vpeljane pri analizi. Natančnost numeričnih rezultatov je bila še dodatno izboljšana z umerjanjem koeficienta trenja. Preizkus, ki je bil opravljen za določitev podatkov o trenju, je bil simuliran z analizo MKE. Simulacije kažejo, da je kontaktni tlak na preizkušancu zelo nehomogen. Ta pojav je verjetno tudi razlog za opazne razlike med dejanskim in napovedanim koeficientom trenja. Zato je bila uvedena kalibracijska funkcija, s katero je bila pomnožena izvirna torna površina. Ko je bil pri ponovljenih simulacijah vožnje v ovinek upoštevan umerjeni koeficient trenja, se je razlika med rezultati eksperimenta in numerične analize občutno zmanjšala, do največ 5 %.

Uporaba metoda s KE je omejena na kotaljenje po suhi in ravni površini. Načrtujemo tudi izboljšanje obstoječega modela: po eni strani bo izboljšan model trenja dezena plašča, po drugi strani pa bo model KE ustrezno prilagojen podrobni 3D-geometriji dezena pnevmatike z ustreznim mreženjem CADmodela.

Modeli pnevmatike s KE, ki so bili uporabljeni pri tej raziskavi, so izdelani na osnovi namenskega CAD-modela, ki so ga razvili avtorji. Vse spremembe geometrije CAD-modela se samodejno prenesejo na model s KE. Model s KE je zato zelo fleksibilen in primeren za parametrične študije, konstruktorji pnevmatik pa lahko z njegovo pomočjo hitro poiščejo optimalne vrednosti konstrukcijskih parametrov pnevmatike. Rezultati analize po metodi končnih elementov so bili neposredno primerjani z rezultati eksperimentov in potrjujejo veljavnost modela s KE. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: analiza po metodi končnih elementov (FEA), konstrukcija pnevmatike, stacionarno kotaljenje (SSR), vožnja v ovinek, zaviranje, parametrično

*Naslov avtorja za dopisovanje: Univerza v Nišu, Fakulteta za strojništvo, Aleksandra Medvedeva 14, 18000 Niš, Srbija, nikola.korunovic@masfak.ni.ac.rs

SI 185

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 186

Prejeto: 13.04.2010 Sprejeto: 07.11.2011

Vpliv rezkalnih strategij na vzdržljivost orodij za kovanje Ivo Pahole1* ‒ Dejan Studenčnik2 ‒ Karl Gotlih1 ‒ Mirko Ficko1 ‒ Jože Balič1 1Univerza

v Mariboru, Fakulteta za strojništvo, Slovenija Nigbo Forging Co. LTD, Kitajska Namen prispevka je znanstveno raziskati in potrditi smernice za rezkanje gravur orodij za kovanje, ki omogočajo izboljšanje tehnoloških lastnosti. V članku je prikazan vpliv smeri obdelave pri visokohitrostnem rezkanju na obstojnost gravure preoblikovalnega orodja za kovanje. Kovanje v preoblikovalnih orodjih v vročem ali v hladnem stanju nam omogoča, da pridemo do končnega izdelka z eno ali več preoblikovalnimi operacijami. Orodje za utopno kovanje imenujemo tudi utop. Gravura utopnega orodja ima približno negativno obliko izdelka, čeprav nikoli ni popolnoma enaka negativu izdelka. Oblika gravure mora biti prilagojena tehnologiji kovanja. Pri pripravi raziskovalnega dela smo izhajali iz v praksi poznanih dejstev, ki smo jih analitično utemeljili in dokazali s pomočjo preizkusov. Pri obdelavi gravur preoblikovalnih orodij se v glavnem uporabljata dva tehnološka postopka: elektroerozija in visokohitrostno rezkanje. Običajno je material že termično obdelan, zato govorimo o obdelavi »v trdo«. Tehnologi imajo pri tem precej možnosti in najboljšo rešitev izbirajo predvsem glede na izkušnje, v mnogo primerih pa izberejo kar standardno strategijo rezkanja. Če pri določitvi strategije rezkanja upoštevamo smer toka materiala med preoblikovalnim postopkom, lahko dosežemo bistveno boljše lastnosti orodja. Obvladovati je treba predvsem tiste lastnosti orodja, ki bistveno vplivajo na povečanje obstojnosti preoblikovalnega orodja. Eden najpomembnejših vplivov pri izdelavi gravure utopnega orodja je hrapavost površine gravure. Hrapavost je skalarno podana vrednost, ki ne zajema smeri in načina končne obdelave gravure pri visokohitrostnem rezkanju. V praksi se je izkazalo, da je primernejša končna obdelava gravure, izvedena v smeri toka materiala, kar se v povprečju izkazuje z večjo obstojnostjo gravure utopnega orodja. Značilno obliko poti orodja opisujemo s pojmom obdelovalna strategija. Strategije obdelave pri strojih brez računalniškega krmiljenja običajno niso bile predmet posebne obravnave, saj so bile zaradi omejenih možnosti vodenja v večini primerov zelo enostavne. Povsem drugače pa je na računalniško krmiljenih rezkalnih strojih, ki imajo velike možnosti vodenja orodja po obdelovalnem prostoru. Medtem ko se pri grobem rezkanju uspešnost procesa meri s količino odrezanega materiala v časovni enoti in ekonomsko učinkovitostjo procesa, je pri fini obdelavi kakovost površine najpomembnejši dejavnik, s katerim merimo učinkovitost procesa. Razen hrapavosti površine se pojavlja še nekaj lastnosti površine, ki niso opisljive samo s tem merilom kakovosti. Te lastnosti površine so največkrat posledica oblike poti orodja pri fini obdelavi. Ena izmed takšnih lastnosti je tudi smer hrapavosti oziroma usmerjenost grebenov, ki so posledica posameznih prehodov orodja. Smer grebenov je odvisna od smeri poti, medtem ko je višina in oblika grebenov odvisna od oblike orodja in od tehnoloških parametrov. Ugotovitve raziskave kažejo, da rezkalne strategije bistveno vplivajo na tok materiala pri preoblikovanju. Tehnologi na podlagi izkušenj določijo uspešne strategije za posamezne oblike, oziroma karakteristike površine z vidika kakovosti površine. Največkrat uporabljeni strategiji za obdelavo prostih površin sta strategija obdelave površine po z-nivojih in strategija vzporednih prehodov. Med procesom kovanja je največji tok preoblikovanega materiala pravokotno na smer kovanja. Tako je smiselno, da je smer končne linijske obdelave v smeri največjega toka preoblikovanega materiala v utop.

Sledi končne obdelave ne zavirajo pretoka materiala v orodje, posledično pa se zmanjša obraba gravure utopa.

Smer toka materiala v kovaškem utopu je možno določiti na osnovi tehnološkega načrta procesa kovanja za določen izdelek. Pri tem imajo največkrat veliko vlogo izkušnje orodjarjev in konstruktorjev orodij. Pomemben pripomoček pri zasnovi in načrtovanju kovaških orodij predstavljajo programska orodja, ki uporabljajo metodo končnih elementov. Na osnovi simulacij poteka kovanja je torej možno določiti tudi smer toka materiala v utopnem orodju. Smer toka materiala je osnova za določanje strategije postopkov rezkanja pri končni obdelavi.

Omejitve raziskave so v številu izvedenih preizkusov, saj različne kombinacije materiala, rezkalnega orodja, rezkalne strategije in rezalnih parametrov dajejo zelo različne rezultate za trajnost orodja. V okviru raziskave je bilo izvedeno tudi preizkušanje kovaškega orodja. Ob enakih pogojih kovanja se je obstojnost orodja bistveno povečala, če je bila gravura orodja obdelana v smeri tečenja materiala med procesom kovanja.

Rezultati kažejo, da je z analizo tečenja materiala ob uporabi ustrezne programske opreme za simuliranje postopka kovanja in programske opreme za programiranje rezkalnega stroja z ustreznim inženirskim znanjem mogoče doseči večjo kakovost dela. Pregled baz znanstvenih člankov je pokazal, da vpliv rezkalnih strategij pri rezkanju gravure orodja za kovanje na trajnost orodja doslej še ni bil raziskan. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: orodja za kovanje, kakovost površine, visokohitrostno rezkanje, CNC-rezkanje 2Unior

SI 186

*Naslov avtorja za dopisovanje: Univerza v Mariboru, Fakulteta za strojništvo, , Smetanova ulica 17, Maribor, Slovenija, ivo.pahole@uni-mb.si

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 187

Prejeto: 12.01.2011 Sprejeto: 02.08.2011

Doba uporabnosti in obraba orodij iz polikristaliničnega kubično kristaliziranega borovega nitrida (PCBN) pri struženju amorfno-kristaliničnih prevlek na osnovi železa Tanovic, L. ‒ Bojanic, P. ‒ Puzovic, R. ‒ Klimenko, S. Ljubodrag Tanovic1 ‒ Pavao Bojanic1,* ‒ Radovan Puzovic1 ‒ Sergey Klimenko2 1Univerza v Beogradu, Fakulteta za strojništvo, Srbija 2Institut V. Bakula za supertrde materiale pri Ukrajinski nacionalni akademiji znanosti, Ukrajina

Uporaba amorfno-kristaliničnih prevlek je v praksi precej omejena zaradi pomanjkanja znanstvenih podatkov o njihovih lastnostih pri mehanski obdelavi. V tehnični literaturi ni veliko del, ki obravnavajo mehansko obdelavo prevlek z amorfno-kristalinično strukturo. O vplivih značilnosti strukture amorfnokristaliničnih prevlek na dobo uporabnosti in obrabo orodij ni praktično nobenih informacij. Namen predstavljene raziskave je preučitev vpliva pogojev obdelave in lastnosti strukture prevlek Fe80 B20 in Fe79Cr16B5 na življenjsko dobo in obrabo rezalnih orodij iz materiala PCBN-Ciborit.

Obdelava prevlek z odrezavanjem je v primerjavi z obdelavo monolitnih materialov enake kemične sestave in trdote bistveno zahtevnejša zaradi spreminjajočih se mehanskih lastnosti in velikosti dodatka za obdelavo, kakor tudi zaradi prisotnosti trdih delcev in por v strukturi materiala.

Odgovor na vprašanje o vplivih nehomogenosti materiala prevleke na dobo uporabnosti in obrabo orodja je ključnega pomena pri izbiri obratovalnih parametrov obdelovalnega stroja.

Izkazalo se je, da so opazovani mehanizmi obrabe orodja kompleksnega značaja, prevladujejo pa abrazivni mehanski, adhezivni in kemični procesi v območju odrezavanja. Na življenjsko dobo orodja pri spremenljivih pogojih struženja vplivata fazna sestava in nehomogena struktura prevleke. Zlasti pri struženju prevlek, izdelanih po postopku plamenskega nabrizgavanja z elektrodo Fe80B20, in prevlek, izdelanih po postopku obločnega nabrizgavanja z elektrodo Fe79Cr16B5, je bila najmanjša obraba in najdaljša doba uporabnosti orodja dosežena pri rezalnih hitrostih v = 1,1 do 1,2 m/s in pri cepilnem kotu γ = –10º. Izkazalo se je, da sprememba cepilnega kota iz γ = 0 na –10° nima večjega vpliva na dobo uporabnosti orodja, v nasprotju s primerom, ko je kot γ = –20°.

Pri struženju z rezalnimi hitrostmi v > 1,2 m/s je vpliv podajanja na dobo uporabnosti orodja večji.

Povečanje podajanja namreč povzroči zvišanje temperature v območju rezanja, s tem pa intenzivnejše kemične reakcije in odstranjevanje enostavno taljivih produktov iz območja rezanja in z orodij.

Prihodnje raziskave bodo usmerjene v dve področji. Prvo področje je razvoj tehnologij nanašanja prevlek z zahtevanimi mehanskimi in fizikalno-kemijskimi lastnostmi na strojne elemente brez naknadne obdelave, drugo področje pa obsega razvoj novih materialov in orodij za primere, ko se obdelavi ni mogoče izogniti.

Predstavljeni rezultati raziskave pojasnjujejo pojave pri interakciji med orodjem in prevleko na primeru dobe uporabnosti in obrabe orodja PCBN-Ciborit, ki so uporabni tudi v praksi. ©2011 Strojniški vestnik. Vse pravice pridržane.

Keywords: PCBN, struženje, območje rezanja, obraba orodja, doba uporabnosti orodja, amorfnokristalinična prevleka

*Naslov avtorja za dopisovanje: Univerza v Beogradu, Fakulteta za strojništvo, Katedra za proizvodni inženiring, Kraljice Marije 16, 11120 Beograd 35, Srbija, pbojanic@mas.bg.ac.rs

SI 187

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 188

Prejeto: 15.03.2011 Sprejeto: 28.07.2011

Simulacija membranskega reaktorja za konverzijo CO z vodno paro za kombinirano plinsko-parno elektrarno z uplinjevalnikom in zajemom CO2 Andrej Lotrič1,* ‒ Mihael Sekavčnik1 ‒ Christian Kunze2 ‒ Hartmut Spliethoff2 1Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija 2Tehnična univerza München, Fakulteta za strojništvo, Nemčija Obravnavan je model membranskega reaktorja za konverzijo CO z vodno paro (WGSMR – WaterGas Shift Membrane Reactor) v kombinirani plinsko-parni elektrarni z uplinjevalnikom (IGCC – Integrated Gasification Combined Cycle) in zajemom CO2. Z uporabo uplinjanja premoga je prikazana tehnologija pridobivanja sinteznega plina za pogon plinsko-parnih postrojenj. Namen integracije membranskega reaktorja v IGCC postrojenje je pridobivanje H2 iz sinteznega plina z uporabo WGS reakcije. Sintezni plin v večini sestavljajo CO, H2, H2O in CO2, z WGS reakcijo pa lahko CO pretvorimo v H2 in CO2. Proces konverzije CO poteka tako, da sintezni plin vstopa v reaktor na zaporno stran membrane, ki prepušča samo H2. V sinteznem plinu vsebovan H2 in H2, ki nastaja pri konverziji, prehaja na prepustno stran membrane, kjer ga iz reaktorja potiska odnašalni plin (navadno N2). Tehnologija pridobivanja H2 iz sinteznega plina z membranskim reaktorjem ima v primerjavi s klasičnim WGS reaktorjem več prednosti, kot so višja temperatura in tlak procesa, boljša konverzija CO, spontano ločevanje H2 znotraj samega reaktorja in manjši investicijski stroški reaktorja, ker združuje več funkcij v eni sami enoti. Kinetiko kemičnih reakcij v reaktorju, pri procesu konverzije CO v H2 in v CO2, smo popisali z uporabo »Bradfordovega« mehanizma.

Ta mehanizem predpostavlja, da WGS reakcija poteka preko štirih elementarnih reakcij. Najpočasnejša od teh reakcij določa hitrost, s katero poteka celotna reakcija. Na osnovi eksperimentalnih podatkov, pridobljenih iz literature, in začetnih ter robnih pogojev smo model reaktorja popisali v programskem jeziku Mathematica. Tako smo izračunali konverzijo CO in prepustnost vodika skozi membrano. Skupaj z začetnimi in robnimi pogoji sta ta parametra tvorila osnovo za izdelavo modela membranskega reaktorja v programskem okolju Aspen Plus. Ta model je izdelan tako, da bo lahko kot element integriran v simulacijo celotnega IGCC sistema. Izračuni so pokazali, da bi bila membrana s prepustnostjo vsaj k′ = 3·10-4 mol m-2 s-1 Pa-0,5 zadostna, ker bi bil proces konverzije CO omejen s hitrostjo reakcije in ne s prepustnostjo H2 skozi membrano. Ugotovljeno je, da bi membrana s prepustnostjo slabšo za en red dosegla le 50 % konverzijo CO, ker bi bil proces omejen s prepustnostjo skozi membrano. Na podlagi izbranih robnih pogojev model membranskega reaktorja napove, da je možno doseči 89,1 % konverzijo CO. S povečanjem dolžine reaktorja za 75% in povečanjem pretoka odnašalnega plina, ter s tem zmanjšanje molskega deleža H2 na prepustni strani membrane na xH2 = 0,45, je lahko dosežena 92,3% konverzija CO. Z ekonomskega staliča povečanje konverzije CO ne bi opravičevalo tako velikega reaktorja. Analiza spreminjanja absolutnega tlaka na prepustni strani membrane je pokazala, da je kompromis med konverzijo CO in rabo energije za kompresijo zmesi H2/N2, pred vstopom v gorilnik plinske turbine, potrebno še podrobneje raziskati. Specifična raba energije nakazuje, da bi bilo smiselno povečati absolutni tlak na prepustni strani membrane na 2 bar, kar pa bi imelo posledice na druge parametre v sistemu. To možnost je potrebno podrobneje raziskati, da se preuči vpliv na celoten sistem. V izdelanem modelu se plin na zaporni strani membrane vodi na katalitični zgorevalnik, kjer se CO, H2S in H2, ki ni uspel preiti skozi membrano, zgorijo in ustvarijo toploto, ki se lahko uporabi v drugih procesih. Po zgorevanju dimni plini od žveplenih spojin vsebujejo samo še SO2, ki se lahko odstrani s konvencionalnimi tehnologijami razžveplevanja. WGSMR je tehnološko izvedljiv vendar je še mnogo vprašanj, ki jih je potrebno razrešiti. Membrana mora obdržati svoje strukturne lastnosti in prepustnost pri izbranih obratovalnih pogojih. Če pri teh pogojih ne bo prenesla kislega okolja, bo potrebno čiščenje sinteznega plina. Zagotovljeno mora biti ustrezno hlajanje reaktorja, da se prepreči pregrevanje reaktorja in da se zagotovi izotermno obratovanje reaktorja. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: uplinjanje premoga, kombinirana plinsko-parna elektrarna z uplinjevalnikom, reakcija vodnega plina, membranski reaktor, Aspen Plus, zajem CO2 SI 188

*Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za strojništvo, Laboratorij za termoenergetiko, Aškerčeva 6, 1000 Ljubljana, Slovenija, andrej.lotric@gmail.com

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 189

Prejeto: 24.12.2010 Sprejeto: 02.09.2011

Raziskava triboloških lastnosti površinsko obdelanega nerjavnega jekla za jedrsko uporabo AISI 316 LN v visokem vakuumu pri temperaturah od 25 °C do 500 °C Devaraju, A. – Elayaperumal, A. – Venugopal, S. – Kailas, S.V. – Alphonsa, J. Ayyannan Devaraju1,* – Ayyasamy Elayaperumal1 – Srinivasan Venugopal2 – Satish V. Kailas3 – Joseph Alphonsa4 1Tehniška fakulteta, Univerza Anna, Indija 2Skupina za metalurgijo in materiale, center Indire Gandhi za jedrske raziskave, Indija 3Oddelek za strojništvo, Indijski institut za znanost, Indija 4Center za spodbujanje rabe industrijskih plazemskih tehnologij, Institut za raziskave plazme, India

V literaturi doslej še niso bile objavljene raziskave tribološkega vedenja v vročem vakuumu za avstenitna nerjavna jekla AISI 316LN (316LN SS), obdelana s plazemskim nitriranjem (PN). Za premostitev te vrzeli v znanju je bilo opravljeno vrednotenje triboloških lastnosti jekla 316LN SS, obdelanega s PN, pri temperaturah 25, 200, 400 in 500 °C v visokem vakuumu (1,6×10-4 bar).

Material 316 LN SS je bil pripravljen v obliki igel in obročev. Igle in obroči so bili 24 ur plazemsko nitrirani pri temperaturi 570 °C in pri delovnem tlaku 5 mbar v plinski mešanici 20% N2 in 80% H2, s čimer je bil ustvarjen debelejši PN-sloj. Faze v PN-sloju so bile določene s preizkusom rentgenske difrakcije (XRD). Trdota PN-prevleke je bila izmerjena z napravo za preizkušanje mikrotrdote po Vickersu pri obtežitvi 25 g. Tribološke značilnosti PN-prevlek v vročem vakuumu so bile ovrednotene z vakuumskim tribometrom vrste igla na disku.

Pri številnih tehničnih aplikacijah (zlasti v letalski in vesoljski industriji ter v jedrski industriji) so pomembne komponente med sabo v drsnem stiku v okolju vakuuma. Tribološko vedenje kovin v okolju vakuuma je drugačno kot na zraku, kljub temu pa je na to temo dostopne le malo literature. To delo je zato namenjeno vrednotenju tribološkega vedenja v vakuumu za nerjavno jeklo 316LN SS, obdelano s plazemskim nitriranjem, z namenom kvalifikacije za jedrske aplikacije.

Povprečna debelina PN-prevleke je bila 70 μm. Trdota površine plazemsko nitriranega jekla 316LN SS je bila 1040 HV25g, oziroma petkrat večja od trdote neobdelanega materiala (210 HV25g). Rezultati preizkusov XRD so pokazali, da je PN-sloj sestavljen iz faz CrN, Fe4N in Fe3N, ugotovljen pa je bil tudi odsev avstenita iz materiala substrata. Pri parih PN/PN je bil koeficient trenja manjši, vse dokler je bil na površini igle PN-sloj, ki preprečuje močnejšo adhezijo med kontaktnimi površinami. Čeprav imajo PN-prevleke pri povišanih temperaturah v okolju visokega vakuuma zelo veliko trenje, so se na površini obroča izkazale z odlično protiobrabno obstojnostjo.

PN-sloji so bili pri vseh preizkusih PN/PN zaradi visokega začetnega kontaktnega tlaka odstranjeni s kontaktne površine igle. Konfiguracija igle (s polkrožno konico) na disku zato ni priporočljiva za raziskave protiobrabne obstojnosti parov PN/PN. Raziskave tribološkega vedenja površinsko obdelanega jekla 316LN SS v vakuumu so redke.

Pregled razpoložljive literature je pokazal, da tribološke lastnosti plazemsko nitriranega jekla 316LN SS v okolju visokega vakuuma doslej še niso bile raziskane. V tem projektu so bile zato ovrednotene tribološke lastnosti plazemsko nitriranega jekla 316LN SS v visokem vakuumu pri temperaturah 25, 200, 400 in 500 °C. ©2011 Strojniški vestnik. Vse pravice pridržane.

Keywords: materiali za jedrsko uporabo, površinska obdelava, PN-sloj, povišana temperatura, visok vakuum, tribološke lastnosti

*Naslov avtorja za dopisovanje: Katedra za strojništvo, oddelek za konstruiranje, Tehniška fakulteta, Univerza Anna, Chennai-600 025, Indija, adevaa2008@rediffmail.com

SI 189

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 190

Prejeto: 24.12.2010 Sprejeto: 18.10.2011

Eksperimentalna primerjava kriogenih Joule-Thompsonovih hladilnikov pri IR-zaznavalih Sicovic, A ‒ Milinovic, M. ‒ Jeremic, O. Aleksandar Sicovic1,* ‒ Momčilo Milinovic2 ‒ Olivera Jeremic3 1 Le-tehnika, d.o.o., Slovenija 2Univerza v Beogradu, Fakulteta za strojništvo, Srbija

V članku je predstavljena postavitev eksperimenta za raziskavo zmogljivosti Joule-Thompsonovih kriohladilnikov, integriranih v Dewarjevih posodah z modeli IR-zaznaval. Eksperimentalno so bili analizirani prehodni temperaturni režimi pri kriogenskem hlajenju detektorja, pri čemer je bil kot hladivo uporabljen dušik. Pri eksperimentalni raziskavi vplivov temperaturnih prehodov so bile primerjane tri glavne metode regulacije pretoka fluida. Rezultati eksperimentov kažejo pomembne razlike v kakovosti in hitrosti hlajenja IR-detektorja, integriranega v hladilnem sistemu z in brez regulacije pretoka hladiva.

Metoda je uporabna za meritve stacionarnih in nestacionarnih temperaturnih prehodov v glavnih podsestavih IR-zaznaval.

Osnovni namen kriogenih hladilnikov je vzdrževanje zahtevane temperature elementov zaznaval, občutljivih na IR-sevanje, ki so vgrajena v visokoločljivostne kamere za infrardečo termografijo in drugo IR-opremo, ki se uporablja v vojaških in civilnih aplikacijah. Za te namene se najpogosteje uporablja Joule-Thompsonov kriohladilnik. S svojo kompaktno izvedbo je primeren za aplikacije, kjer je običajno na voljo le nekaj centimetrov prostora, ima pa tudi toplotne in splošne mehanske lastnosti, ki so potrebne za dane obratovalne razmere. Joule-Thompsonov kriohladilnik mora zagotoviti nadzorovano temperaturo hladilnega elementa, ki je običajno globoko v kriogenem območju pod 100 K, v predpisanih tolerancah in v zelo kratkem času.

Predmet tega članka je primerjalna analiza rezultatov preizkušanja temperaturnega prehoda na integriranem laboratorijskem sistemu, ki je sestavljen iz Joule-Thompsonovih kriohladilnikov različnih izvedb, Dewarjevih posod in ustreznega eksperimentalnega modela temperaturno občutljivega elementa, ki nadomešča drage realne detektorje. Pomen tega dela je v eksperimentalnih podatkih o znižanju temperature na modelnem detektorju, ki omogočajo primerjavo hitrosti in prehodnega režima hlajenja, kot glavnih kriterijev kakovosti pri razvoju senzorskih sestavov. Ta kriterij določa sposobnost zaznavala za delovanje v IR-načinu. Rezultati so primerjani s skromnimi podatki o sodobnih rešitvah kriohladilnikov iz dostopne literature. Pri vseh zgoraj naštetih izvedbah je ključni kriterij kakovosti senzorja sposobnost hitre vzpostavitve zahtevane temperature detektorja po začetku pretoka hladilnega sredstva. V članku so predstavljeni tudi podatki o meritvah, ki so jih opravili avtorji.

Izmerjena zmogljivost vseh obravnavanih izvedb Joule-Thompsonovih kriohladilnikov je osnova za nadgradnjo nekrmiljenih senzorjev v krmiljene rešitve z zveznimi ali impulznimi regulatorji. Raziskave omogočajo edinstvena eksperimentalna zasnova in metode za merjenje in kontrolo parametrov delovanja različnih modelnih sistemov z Dewarjevo posodo, Joule-Thompsonovim kriohladilnikom, regulatorjem in IR-detektorjem. Univerzalna eksperimentalna oprema omogoča preizkušanje istega senzorskega sistema z Dewarjevo posodo v izvedbi s krmiljenjem in brez njega. Postavitev eksperimenta tako omogoča spreminjanje termodinamičnega cikla ob različnih parametrih okolice kriohladilnika in zaznaval. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: režimi hlajenja, Joule-Thompsonovi kriogeni hladilniki, zaznavalni element, kriogeni plini, meritve temperature, eksperimentalna določitev prehodnih režimov

SI 190

*Naslov avtorja za dopisovanje: Le-tehnika, d.o.o., Šuceva 27, 4000 Kranj, Slovenija, aleksandarsicovic@yahoo.com

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 191-192 Navodila avtorjem

Navodila avtorjem Napovednik

Avtorje vabimo, da svoje članke oddajo z uporabo elektronskega sistema na naši spletni strani: http://ojs.sv-jme.eu.

Avtorji lahko enostavno dostopajo do elektronskega sistema z registracijo na spletni strani. Avtorje prosimo, da na spletni strani naložijo članek in s sledenjem sistemu zagotovijo želene podatke, potrebne za indeksiranje in lažje iskanje po reviji. Avtorji lahko naložijo dopolnilne vsebine kot npr. spremno pismo, podatkovne baze, raziskovalne instrumente, programsko kodo in podobno. Članki morajo biti napisani v angleškem jeziku. Strani morajo biti zaporedno označene.

Prispevki so lahko dolgi največ 10 strani. Daljši članki so lahko v objavo sprejeti iz posebnih razlogov, katere morate navesti v spremnem dopisu.

Kratki članki naj ne bodo daljši od štirih strani.

Navodila so v celoti na voljo v rubriki “Informacija za avtorje” na spletni strani revije: http://en.sv-jme.eu/. Prosimo vas, da članku priložite spremno pismo, ki naj vsebuje: 1. naslov članka, seznam avtorjev ter podatke avtorjev; 2. opredelitev članka v eno izmed tipologij; izvirni znanstveni (1.01), pregledni znanstveni (1.02) ali kratki znanstveni članek (1.03); 3. izjavo, da članek ni objavljen oziroma poslan v presojo za objavo drugam; 4. zaželeno je, da avtorji v spremnem pismu opredelijo ključni doprinos članka; 5. predlog dveh potencialnih recenzentov, ter kontaktne podatke recenzentov. Navedete lahko tudi razloge, zaradi katerih ne želite, da bi določen recenzent recenziral vaš članek.

OBLIKA ČLANKA Članek naj bo napisan v naslednji obliki: Naslov, ki primerno opisuje vsebino članka.

Povzetek, ki naj bo skrajšana oblika članka in naj ne presega 250 besed. Povzetek mora vsebovati osnove, jedro in cilje raziskave, uporabljeno metodologijo dela, povzetek rezultatov in osnovne sklepe.

- Uvod, v katerem naj bo pregled novejšega stanja in zadostne informacije za razumevanje ter pregled rezultatov dela, predstavljenih v članku.

- -

- Teorija.

- Eksperimentalni del, ki naj vsebuje podatke o postavitvi preskusa in metode, uporabljene pri pridobitvi rezultatov.

- Rezultati, ki naj bodo jasno prikazani, po potrebi v obliki slik in preglednic.

- Razprava, v kateri naj bodo prikazane povezave in posplošitve, uporabljene za pridobitev rezultatov. Prikazana naj bo tudi pomembnost rezultatov in primerjava s poprej objavljenimi deli. (Zaradi narave posameznih raziskav so lahko rezultati in razprava, za jasnost in preprostejše bralčevo razumevanje, združeni v eno poglavje.)

- Sklepi, v katerih naj bo prikazan en ali več sklepov, ki izhajajo iz rezultatov in razprave.

- Literatura, ki mora biti v besedilu oštevilčena zaporedno in označena z oglatimi oklepaji [1] ter na koncu članka zbrana v seznamu literature.

Enote - uporabljajte standardne SI simbole in okrajšave. Simboli za fizične veličine naj bodo v ležečem tisku (npr. v, T, n itd.). Simboli za enote, ki vsebujejo črke, naj bodo v navadnem tisku (npr. ms1, K, min, mm itd.)

Okrajšave naj bodo, ko se prvič pojavijo v besedilu, izpisane v celoti, npr. časovno spremenljiva geometrija (ČSG).

Pomen simbolov in pripadajočih enot mora biti vedno razložen ali naveden v posebni tabeli na koncu članka pred referencami.

Slike morajo biti zaporedno oštevilčene in označene, v besedilu in podnaslovu, kot sl. 1, sl. 2 itn. Posnete naj bodo v ločljivosti, primerni za tisk, v kateremkoli od razširjenih formatov, npr. BMP, JPG, GIF. Diagrami in risbe morajo biti pripravljeni v vektorskem formatu, npr. CDR, AI. Vse slike morajo biti pripravljene v črnobeli tehniki, brez obrob okoli slik in na beli podlagi.

Ločeno pošljite vse slike v izvirni obliki Pri označevanju osi v diagramih, kadar je le mogoče, uporabite označbe veličin (npr. t, v, m itn.). V diagramih z več krivuljami, mora biti vsaka krivulja označena. Pomen oznake mora biti pojasnjen v podnapisu slike.

Tabele naj imajo svoj naslov in naj bodo zaporedno oštevilčene in tudi v besedilu poimenovane kot Tabela 1, Tabela 2 itd.. Poleg fizikalne veličine, npr t (v ležečem tisku), mora biti v

SI 191

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 191-192

oglatih oklepajih navedena tudi enota. V tabelah naj se ne podvajajo podatki, ki se nahajajo v besedilu.

Potrditev sodelovanja ali pomoči pri pripravi članka je lahko navedena pred referencami. Navedite vir finančne podpore za raziskavo.

REFERENCE Seznam referenc MORA biti vključen v članek, oblikovan pa mora biti v skladu s sledečimi navodili. Navedene reference morajo biti citirane v besedilu. Vsaka navedena referenca je v besedilu oštevilčena s številko v oglatem oklepaju (npr. [3] ali [2] do [6] za več referenc). Sklicevanje na avtorja ni potrebno.

Reference morajo biti oštevilčene in razvrščene glede na to, kdaj se prvič pojavijo v članku in ne po abecednem vrstnem redu. Reference morajo biti popolne in točne. Vse neangleške oz. nenemške naslove je potrebno prevesti v angleški jezik z dodano opombo (in Slovene) na koncu Navajamo primere: Članki iz revij: Priimek 1, začetnica imena, priimek 2, začetnica imena (leto). Naslov. Ime revije, letnik, številka, strani, DOI oznaka.

[1] Hackenschmidt, R., Alber-Laukant, B., Rieg, F. (2010). Simulating nonlinear materials under centrifugal forces by using intelligent cross-linked simulations. Strojniški vestnik - Journal of Mechanical Engineering, vol.

57, no. 7-8, p. 531-538, DOI:10.5545/svjme.2011.013.

Ime revije ne sme biti okrajšano. Ime revije je zapisano v ležečem tisku.

Knjige: Priimek 1, začetnica imena, priimek 2, začetnica imena (leto). Naslov. Izdajatelj, kraj izdaje [2] Groover, M. P. (2007). Fundamentals of Modern Manufacturing. John Wiley & Sons, Hoboken.

Ime knjige je zapisano v ležečem tisku.

Poglavja iz knjig: Priimek 1, začetnica imena, priimek 2, začetnica imena (leto). Naslov poglavja. Urednik(i) knjige, naslov knjige. Izdajatelj, kraj izdaje, strani.

[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. Članki s konferenc: Priimek 1, začetnica imena, priimek 2, začetnica imena (leto). Naslov. Naziv konference, strani.

[4] Štefanić, N., Martinčević-Mikić, S., Tošanović, N. (2009). Applied Lean System in Process

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Industry. MOTSP 2009 Conference Proceedings, p. 422-427.

Standardi: Standard (leto). Naslov. Ustanova. Kraj.

[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.

Spletne strani: Priimek, Začetnice imena podjetja. Naslov, z naslova http://naslov, datum dostopa.

[6] Rockwell Automation. Arena, from http://www. arenasimulation.com, accessed on 2009-09-27.

RAZŠIRJENI POVZETEK Ko je članek sprejet v objavo, avtorji pošljejo razširjeni povzetek na eni strani A4 (približno 3.500 - 4.000 znakov). Navodila za pripravo razširjenega povzetka so objavljeni na spletni strani http://sl.svjme.eu/informacije-za-avtorje/.

AVTORSKE PRAVICE Avtorji v uredništvo predložijo članek ob predpostavki, da članek prej ni bil nikjer objavljen, ni v postopku sprejema v objavo drugje in je bil prebran in potrjen s strani vseh avtorjev. Predložitev članka pomeni, da se avtorji avtomatično strinjajo s prenosom avtorskih pravic SV-JME, ko je članek sprejet v objavo. Vsem sprejetim člankom mora biti priloženo soglasje za prenos avtorskih pravic, katerega avtorji pošljejo uredniku. Članek mora biti izvirno delo avtorjev in brez pisnega dovoljenja izdajatelja ne sme biti v katerem koli jeziku objavljeno drugje.

Avtorju bo v potrditev poslana zadnja verzija članka. Morebitni popravki morajo biti minimalni in poslani v kratkem času. Zato je pomembno, da so članki že ob predložitvi napisani natančno.

Avtorji lahko stanje svojih sprejetih člankov spremljajo na http://en.sv-jme.eu/.

PLAČILO OBJAVE Domači avtorji vseh sprejetih prispevkov morajo za objavo plačati prispevek, le v primeru, da članek presega dovoljenih 10 strani oziroma za objavo barvnih strani v članku, in sicer za vsako dodatno stran 20 EUR ter dodatni strošek za barvni tisk, ki znaša 90,00 EUR na stran.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 193-195 Osebne objave

Doktorske disertacije, magistrska dela, specialistična dela in diplome DOKTORSKE DISERTACIJE Na Fakulteti za strojništvo Univerze v Ljubljani sta z uspehom obranila svojo doktorsko disertacijo: dne 11. novembra 2011 Gašper BENEDIK z naslovom: »Tokovne razmere na rotirajočem disku iz odprtoceličnega poroznega materiala« (mentor: prof. dr. Branko Širok); Doktorsko delo obravnava tokovne razmere na rotirajočem disku, pri katerem se energija iz rotorja na fluid prenaša preko strukture poroznega materiala. Podane so teoretične osnove za popis toka fluida skozi rotirajoč porozen material in pregled obstoječih znanstvenih in strokovnih del. Predstavljene so modelne izvedbe diskov z različnimi konstrukcijskimi parametri. Teoretično in eksperimentalno so določeni padci tlaka in lokalne hitrosti v odvisnosti od volumskega pretoka za različne modelne izvedbe nerotirajočih diskov. Analizirano je hitrostno polje v odvisnosti od lokalne strukture materiala in volumskega pretoka zraka. Eksperimentalna dinamika fluidov za rotirajoče diske vključuje meritve integralnih aerodinamskih in akustičnih karakteristik, meritve lokalnih hitrosti zračnega toka ter vizualizacijo tokovnega polja. Na osnovi eksperimentalne primerjave s klasičnim lopatičnim rotorjem so ugotovljene boljše akustične lastnosti poroznega diska, primerljiva tlačna števila in nižji integralni aerodinamski izkoristki.

Predstavljena je numerična simulacija toka zraka skozi homogen rotirajoč porozen disk. Ugotovljeno je relativno dobro ujemanje numeričnih in eksperimentalnih rezultatov. Predstavljene ugotovitve so primerne za številne aplikacije z rotirajočimi poroznimi materiali; dne 23. novembra 2011 Janez KUNAVAR z naslovom: »Geometrijska optimizacija stabilnosti enoosnega elementa iz gradiva z oblikovnim spominom« (mentor: prof. dr. Franc Kosel); Gradiva z oblikovnim spominom se uporabljajo na različnih področjih človekove ustvarjalnosti. V praksi se uporabljata dva njihova fenomena, oblikovni spomin in superelastičnost. Transformacija iz martenzitnega (nizko-temperaturna faza) v avstenitno (visokotemperaturna faza) stanje ob sočasnem mehanskem obremenjevanju se imenuje ovirana povračljivost,

in ta pojav je za primer sočasnega nateznega obremenjevanja precej dobro raziskan. Ovirana povračljivost je precej manj raziskana za primere, ko se v gradivu z oblikovnim spominom generira tlačna napetost. Pojav osne tlačne napetosti v vitkem enoosnem elementu lahko povzroči uklon tega elementa. V doktorski nalogi so določene vse potrebne termomehanske lastnosti, s katerimi lahko na osnovi predvidene temperature uklona izračunamo dolžino idealno ravne palice okroglega prereza iz gradiva z oblikovnim spominom.

Izdelana je eksperimentalna verifikacija računskih rezultatov. Teoretično in računsko je izdelan tudi postopek geometrijske optimizacije uklona zaradi ovirane povračljivosti za palice okroglega prereza iz gradiva z oblikovnim spominom. Poleg tega je za več različic geometrijske optimizacije izvedena eksperimentalna verifikacija rezultatov. * MAGISTRSKA DELA Na Fakulteti za strojništvo Univerze v Ljubljani je z uspehom zagovarjala svoje magistrsko delo: dne 29. novembra 2011 Galina ANATOLYEVNA KUBYSHKINA z naslovom: »Vpliv različnih sterilizacijskih tehnik na časovno odvisno vedenje poliamidov (The influence of different sterilization techniques on the timedependent behavior of polyamides)« (mentor: prof. dr. Igor Emri). * Na Fakulteti za strojništvo Univerze v Mariboru sta z uspehom zagovarjala svoje magistrsko delo: dne 17. novembra 2011 Nataša CAMLEK z naslovom: »Svetloba v delovnem okolju in njen vpliv na vrednotenje obremenitve vida« (mentor: prof. dr. Andrej Polajnar); dne 17. novembra 2011 Beno ARBITER z naslovom: »Ravnotežni model uplinjanja odpadkov« (mentor: prof. dr. Niko Samec).

SI 193

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 193-195

* SPECIALISTIČNA DELA Na Fakulteti za strojništvo Univerze v Ljubljani sta z uspehom zagovarjala svoje specialistično delo: dne 4. novembra 2011 Dušan NOVAK z naslovom: »Priprava procesne vode za farmacijo« (mentor: prof. dr. Iztok Golobič); dne 11. novembra 2011 Miha VOJIR z naslovom: »Forenzične preiskave strelnega orožja« (mentor: prof. dr. Iztok Golobič, somentor: izr. prof. dr. Ivan Bajsić).

prof. dr. Karl Gotlih, somentor: doc. dr. Janez Kramberger); David KALJUN z naslovom: »Inženirsko oblikovanje energetsko varčne pisarniške svetilke« (mentor: izr. prof. Vojmir Pogačar). *

*

Na Fakulteti za strojništvo Univerze v Mariboru je pridobil naziv univerzitetni diplomirani gospodarski inženir: dne 25. novembra 2011: Borut MATKO z naslovom: »Uporaba vrednostne analize na primeru razvoja novega izdelka Eko-koš« (mentor: doc. dr. Iztok Palčič, somentor: prof. dr. Anton Hauc).

DIPLOMIRALI SO

*

Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 25. novembra 2011: Gašper JEREB z naslovom: »Optimizacija hidravličnih površin v hidrodinamičnem prenosniku moči« (mentor: izr. prof. dr. Mihael Sekavčnik); Tomaž OSTERC z naslovom: »Uporaba parnega krožnega procesa z organsko delovno snovjo za izrabo geotermalne energije pri proizvodnji električne energije« (mentor: izr. prof. dr. Mihael Sekavčnik); Luka SKRINJAR z naslovom: »Nezvezna dinamika kombiniranega električnega zaščitnega stikala« (mentor: prof. dr. Miha Boltežar, somentor: doc. dr. Janko Slavič); dne 28. novembra 2011: Tilen BURŠIČ z naslovom: »Vpliv strukture visokoelastičnega materiala na njegove dušilne lastnosti« (mentor: prof. dr. Igor Emri); Jože MIKOLAVČIČ z naslovom: »Priprava tehnologije navarjanja nikljeve zlitine na konstrukcijsko jeklo« (mentor: prof. dr. Janez Tušek).

Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva: dne 9. novembra 2011: Drago BORŠTNAR z naslovom: »Najnovejši trendi pri preiskavah letalskih nesreč« (mentor: viš. pred. mag. Aleksander Čičerov, somentor: izr. prof. dr. Tadej Kosel); Janez KREVZEL z naslovom: »Projektno vodeno sočasno osvajanje izdelka« (mentor: izr. prof. dr. Janez Kušar, somentor: prof. dr. Marko Starbek); Jernej MARKELJ z naslovom: »Izdelava nihajnega protipovratnega ventila« (prof. dr. Janez Kopač, somentor: doc. dr. Davorin Kramar); Damjan MEDVED z naslovom: »Načrtovanje preizkuševališča za krmilne ventile« (mentor: izr. prof. dr. Ivan Bajsić); Franci ROJC z naslovom: »Napovedovanje materialnih potreb« (mentor: izr. prof. dr. Janez Kušar, somentor: prof. dr. Marko Starbek).

* Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv univerzitetni diplomirani inženir strojništva: dne 24. novembra 2011: Vesna JURJEC z naslovom: »Zasnova prikolice za prevoz jadralnih letal« (mentor: izr.

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* Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva (UN), univerzitetni študijski program I. stopnje STROJNIŠTVO (Razvojno raziskovalni program): dne 17.junija 2011: Ervin STRMČNIK; dne 20. junija 2011: Ivana NANUT; dne 22. junija 2011: Matej MESOJEDEC in Rok KOCEN; dne 13. julija 2011: Nejc ŠKOBERNE;

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)12, SI 193-195

dne 27. julija 2011: Martin BRICL in Matej KRANJEC; dne 2. avgusta 2011: Andrej OŠLAK; dne 4. avgusta 2011: Anton PRESKAR; dne 22. avgusta 2011: Uroš MAVEC in Blaž ZUPANČIČ; dne 23. avgusta 2011: Klemen KRIŽMAN, Andrej JANEŽIČ, Mitja PETACI, Damijan ZORKO, Matic AMBROŽIČ in Hugo ZUPAN; dne 24. avgusta 2011: Matej KOVAČ, Matic VIRANT in Borut ČERNE; dne 25. avgusta 2011: Jure SALOBIR in Jurij KRANJC; dne 29. avgusta 2011: Nejc VOLK, Damir DEBOGOVIĆ, Marko BRUS, Urša LOKAR in Andrej KRAGELJ; dne 30. avgust 2011: Damjan LUKANČIČ; dne 1. septembra 2011: Anica KOKELJ, Matic ZUPANC, Tjaša KODRIČ, Miha TRAMPUŽ, Tomaž STARMAN, Miha BLATNIK, Klemen COTIČ in Jure FILAK; dne 2. september 2011: Benjamin ČERNOŠA, Gašper DUCMAN in Marko MRAK; dne 5. septembra 2011: Rok KAPLER, Jure ŠAVLI in Staš URBANČIČ; dne 6. septembra 2011: Miha ZUPIN in Blaž ŽUGELJ; dne 7. september 2011: Domen LEVIČAR, Dejan PLOS, Peter ARKO in Blaž ŽUN; dne 8. septembra 2011: Matej RAZPOTNIK, Žiga SIMŠIČ, Gašper ČEBAVS, Luka PAVIČ in Blaž STARC; dne 9. septembra 2011: Žiga TANKO dne 12. septembra 2011: Margerita FLORJANČIČ, Matjaž NERED, Simon GODEC, Polona SREBRNJAK MIHALIČ, Manca SREBRNJAK MIHALIČ, Jan OROŽ, Matej JARM in Jernej RIJAVEC;

dne 13. septembra 2011: Tom KUNAVER, Matevž ZUPANČIČ, Božidar BREČKO in Nejc STRAVNIK; dne 14. septembra 2011: Jurij ŠVEGELJ; dne 15. septembra 2011: Miha BERGINC, Matej KOS, Gašper JARC, Mario MARINOVIĆ, Pierre Robert STARE, Kristian NAGLLIC in Aleš MRAK; dne 16. septembra 2011: Simon TROŠT, Aleš CVELBAR in Nejc JEGLIČ; dne 20. septembra 2011: Tine LAZAR; dne 21. septembra 2011: Jure ZADRAVEC; dne 22. septembra 2011: Nejc JERŠIN; dne 23. septembra 2011: David PEJAŠINOVIĆ, Samo JERANKO in Simon FINK; dne 26. septembra 2011: Vita ŠMAJDEK, Nejc ZUPAN in Matic RESNIK; dne 27. septembra 2011: Miran MESERKO, Matej NOVAK in Uroš ROŽIČ; dne 28. septembra 2011: Rok LOVŠIN, Rok HAMLER, Enej RESINOVIČ, Jaka HIRŠMAN in Žiga GREGOR Turšič; dne 29. septembra 2011: Jan LUKAČ; dne 30. septembra 2011: Jure PLEŠKO. * Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv diplomirani inženir strojništva: dne 24. novembra 2011: Simon KRAMBERGER z naslovom: »Dimenzioniranje jeklene skladiščne hale« (mentor: doc. dr. Janez Kramberger); Gregor PERNER z naslovom: »Ultrazvočna preiskava zvarov in vrednotenje napak na parovodu sveže pare« (mentor: izr. prof. dr. Vladimir Gliha).

SI 195

57 (2011) 1 12

Platnica SV-JME 57(2011)12_04.pdf 1 9.12.2011 12:53:53

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