Journal of Mechanical Engineering 2011 9

57 (2011) 1 9

Platnica SV-JME 57(2011)9_05.pdf 1 15.9.2011 11:53:32

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Journal of Mechanical Engineering - Strojniški vestnik

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9 year 2011 volume 57 no.

Platnica SV-JME 57(2011)9_05.pdf 2 15.9.2011 11:53:32

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 Above: Conventional cryogenic machining process with flooding of oil-based emulsions to the cutting zone.

Below: Alternative cryogenic machining process, where liquid nitrogen at the delivery cools the cutting zone, evaporates and leaves the cutting product clean and dry.

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., Ižanska cesta 383, Ljubljana, Slovenia General information Strojniški vestnik – Journal of Mechanical Engineering is published in 11 issues per year (July and August is a double issue).

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ISSN 0039-2480

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9 Contents

Contents Strojniški vestnik - Journal of Mechanical Engineering volume 57, (2011), number 9 Ljubljana, September 2011 ISSN 0039-2480 Published monthly Papers Franci Pušavec, Janez Kopač: Sustainability Assessment: Cryogenic Machining of Inconel 718 Blaza Stojanović, Nenad Miloradović, Nenad Marjanović, Mirko Blagojević, Lozica Ivanović: Length Variation of Toothed Belt during Exploitation Van Tuan Do, Ui-Pil Chong: Signal Model-Based Fault Detection and Diagnosis for Induction Motors Using Features of Vibration Signal in Two-Dimension Domain Eduard Niţu, Monica Iordache, Luminiţa Marincei, Isabelle Charpentier, Gaël Le Coz, Gérard Ferron, Ion Ungureanu: FE-Modeling of Cold Rolling by In-Feed Method of Circular Grooves Marko Sedlaček, Luis Miguel Silva Vilhena, Bojan Podgornik, Jože Vižintin: Surface Topography Modelling for Reduced Friction Marek Sadílek, Robert Čep, Igor Budak, Mirko Soković: Aspects of Using Tool Axis Inclination Angle Mohammadreza Shabgard, Mirsadegh Seyedzavvar, Samad Nadimi Bavil Oliaei: Influence of Input Parameters on the Characteristics of the EDM Process Mihael Volk, Blaž Nardin, Bojan Dolšak: Application of Numerical Simulations in the DeepDrawing Process and the Holding System with Segments’ Inserts Instructions for Authors

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647 DOI: 10.5545/sv-jme.2010.249

Paper received: 13.12.2010 Paper accepted: 03.08.2011

Sustainability Assessment: Cryogenic Machining of Inconel 718 Pušavec, F. ‒ Kopač, J. Franci Pušavec* ‒ Janez Kopač University of Ljubljana, Faculty of Mechanical Engineering, Slovenia

This paper disseminates the foreground of sustainable cryogenic machining technology that has a high potential to cut production costs and improve competitiveness by reducing resource consumption and creating less waste. The sustainability issues on a shop floor level are pointed out via a life cycle assessment, concluding that the future of sustainable production is going to entail the use of cryogenic machining to reduce environmental burdens and health risks, while increasing machining performance and profitability. In addition, machining evaluation is covered by an experimental study undertaken to understand the likely impacts of the use of cryogenic technology on production costs. The case study refers to the machining of high-temperature Ni-alloy (Inconel 718). It is shown that tooling costs greatly contribute to the total production cost and that cryogenic machining offers a clean and a cost-effective route to improve sustainability performance in comparison to conventional machining. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: machining, sustainability, cryogenics, LCA, cost evaluation

0 INTRODUCTION This paper presents a case study that highlights the importance of sustainable cryogenic machining in achieving sustainable development objectives. Global environmental problems caused by the consumption of natural resources and the production/life of technical products have led to increased political pressure and stronger regulations applied to both, the producers and users of such products [1] and [2]. The adoption of sustainable development in the production offers the industry a cost effective route [3] to [5] to improve economic, environmental, and social performance [6].

Sustainable development initiatives are established on a political level within the UN, the OECD, the EU, and national levels. These initiatives are well positioned and promoted on the production macro level [7], but there is a lack of implementation practices on the shop floor dealing with machining. With the implementation of sustainability principles on the shop floor level, users have the potential to save money and improve their environmental and social performance even if their production stays in the same range or decreases, as shown in the machining evaluation included in this paper. In this way conducted evaluation of sustainability in

cryogenic machining is more than a method for supporting technology design and an instrument for supporting decision-making. It is also a tool for supporting technology policy and for encouraging its adoption and application in the industry.

The ambitions of implementing cryogenic machining to the practice aims at transitioning towards sustainable production on the machining technology/shop floor level and include the following research and application issues: • Design and development of sustainable cryogenic technologies (machining processes and fluid delivery systems) that can be accommodated over different machining operations.

• Optimization of cryogenic fluid delivery system with a controlled pressure, mass flow and flow phase (liquid or vapor).

• Optimization of nozzles for selectively dispensing liquid phase cryogenic fluid.

• Evaluation of cryogenic machining in various industrial case studies.

• Identification of research/application demands and future industrial opportunities.

1 CRYOGENIC MACHINING It is known that oil-based cooling lubricating fluids (CLFs) are one of the most

*Corr. Author’s Address: University of Ljubljana, Faculty of Mechanical Engineering, Aškerčeva 6, SI-1000 Ljubljana, Slovenia, franci.pusavec@fs.uni-lj.si

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unsustainable elements of machining processes.

Most CLFs are formulated from mineral oils, which are extracted from crude oil, primarily for economic reasons. Although alternative, naturally derived CLFs are available (vegetable oils), there has been limited use of these CLFs. This is partly due to higher costs and partly due to a reduced performance [8].

In cryogenic machining a cryogenic CLF (non-oil-based) is delivered to the cutting region of the cutting tool, shown in Fig. 1, which is exposed to the highest temperature during the machining process, or to the part in order to change the material characteristics and improve machining performance.

Fig. 1. Cryogenic liquid nitrogen delivery The CLF is nitrogen fluid, which is liquefied by cooling to -196 °C (liquid nitrogen – LN). Nitrogen is a safe, non-combustible, and noncorrosive gas. The LN in cryogenic machining systems quickly evaporates and returns to the atmosphere, leaving no residue to contaminate the part, chips, machine tool, or operator, thus eliminating disposal costs. Additionally, cryogenic machining could help to machine parts faster, with higher quality, increased machining performance, and a reduced overall cost [9]. Some potential benefits of cryogenic machining are: • Considerably reduced friction coefficient on the tool-chip interface [10].

• LN applied locally to the cutting edge is superior to emulsion in lowering the cutting temperature [11].

• Increased tool-life due to lower abrasion and chemical wear [12].

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Increased material removal rate with no increase in tool-wear and with reduced cutting tool changeover cost, resulting in higher productivity [13].

• Improved machined part surface quality with the absence of mechanical and chemical degradation of the machined surface [14].

Most cryogenic CLF applications have been examined in the machining of heat resistant super-alloys. However, some studies also included machining of low/high carbon steel and bearing steel [15]. One of those high-temperature alloys are Nickel-based alloys that are normally machined with WC-Co grades, with cutting speeds of about 50 m/min. With the introduction of sialon materials, it is possible to increase the cutting speed by a factor of 5, and more recently even higher cutting speeds are possible with silicon carbide whisker-reinforced aluminia tools. The other alternatives are ceramic cutting tools that show lower chemical affinity with Ni materials.

Unfortunately, the accumulation of cutting heat on the cutting edges of ceramic tools causes many problems and sometime leads to early tool failure [16]. Additionally, their fracture toughness is much lower than that of the other widely used tool materials such as carbides. This is the reason that the geometry of the ceramic tools is mostly neutral or even negative (negative rake angle), while carbide cutting tools are available also with very sharp cutting edges and highly positive rake angles. Both properties lead to the reduction in temperature and significant lower cutting forces.

While in this work the goal was oriented towards finishing process, the carbide cutting tools are used and analyzed.

In the machining of high-temperature alloys, conventional oil-based CLFs are not always effective enough in terms of decreasing the high cutting temperature, increasing toollife, reducing machining costs and improving environmental/social sustainability. The problem is that conventional CLFs do not access the toolpart and tool-chip interfaces, which are under high contact pressure, as they vaporize at a high temperature generated close to the cutting tool edge. Taking this into account, it becomes clear that technologies employing conventional CLFs are ineffective and unsustainable when machining materials with high shear strength and low thermal

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StrojniĹĄki vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647

conductivity. In this case, the avoidance of conventional CLFs, would yield an enormous gain from the sustainability point of view [17].

2 LIFE CYCLE ASSESSMENT (LCA) In order to evaluate sustainability of a product, system, or process, the impact resulting from each stage of its life cycle has to be considered [18]. Although cryogenic machining offers to reduce the negative impacts of conventional CLFs usage on the environment, the additional sustainability characteristics of this technology have yet to be analyzed. Therefore, a quantitative assessment of the CLF’s production and use is evaluated with additional qualitative measures associated with health, safety and performance.

These assessments consider machining (turning or milling) over a one-year production period. At this stage, it is assumed that the production and the quality of machining do not vary depending on the CLF type. In other words, LCA was performed strictly for the production and delivery of the CLF into the cutting zone, while machining performance was assumed to be the same. This assumption is limited to CLF environmental and health influence comparisons.

Conventional machining

Evaporation

(oil-based emulsion) Tools

Water

Workpiece

Base oil

Machining energy

Used tools

MWF carryout

0.74 %

0.60 %

Storage tank

Additives

(Liquid nitrogen)

Industrial pretreatment Machine tool

Disposal

Product Waste heat Recirculation (Large facility)

Nitrogen Used tools

100.00 %

Tools Workpiece

Mist

Pump

1.34 %

Nonionic surfactant

Characterization

Looking at the machining system schemes, shown in Fig. 2, it can be seen that in conventional machining, the CLFs (emulsions) are composed of lubricant oil (petroleum-based) and a lubricant carrier (water). Additionally, water serves also as a coolant. In the case of cryogenic machining, the lubricant and carrier are provided by LN, which is known to be stable, while oil-based emulsions are not and thereof need to contain emulsifiers for their stabilization. With regard to oil-based CLF production, most of the CLFs are used as emulsions containing mineral oil and surfactants based on petroleum. Producing these components requires a distillation and processing of crude oil, which creates several by-products. The components considered in this case are a semi-synthetic CLF system containing oil, two surfactants, and water. Consideration of surfactants is important since they play a dominant role in the overall environmental emissions of water/oil surfactant systems [19].

On the other hand, LN can be used as a CLF, which is a liquefied atmospheric gas produced industrially in large quantities by fractional distillation of air. The input into the process is

97.32 %

Anionic surfactant

Cryo machining

2.1 CLF Quantitative (Production and Use)

N2 exhaust Product

LN

Waste heat

Machining energy Machine tool Material flow (incl. material emissions)

Energy Flow

Infrastructure (not modeled)

Fig. 2. Comparison of cryogenic with conventional machining system Sustainability Assessment: Cryogenic Machining of Inconel 718

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electrical energy (approximately 0.5 kWh/kg) and cooling water (50 l/kg at 15 °C), while the output is LN and the remaining components of the air as a waste. The data for LN production were given by the LN supplier (SIAD – Istrabenz Plini). It must be pointed out that there is no other waste, such as CO2, SO2, etc., when producing LN. However, LN production is an energy-intensive process that can be directed towards sustainable development by powering the cooler in LN production with renewably generated electricity or through direct mechanical work from hydro or wind turbines.

Firstly, considering the LCA, the CLF compositions must be known. Typical oil-based CLFs are composed of water, oil, surfactants, and approximately ten other specific chemicals.

In general, CLFs are sold as concentrates with 10 to 30% mineral oil (semi-synthetic fluids) and are diluted 10 to 20 times with water, thus forming the CLF as an emulsion. In the LCA of environmental impact of CLF production, the considered factors were related to: • water use, • solid waste production, • land use, • energy use, • global warming potential (GWP), • acidification.

The specific data included in the LCA are given in [20].

Knowing the CLF composition and CLF component production environmental impact, the remaining missing data are the CLF usage amounts (conventional machining), CLF consumption rate (cryogenic machining), and machining system usage in a fixed time period. In the following analyses, for ease of comparability, a one-year production period is considered.

A major difference between oil-based systems and cryogenic-based systems is that oil-based CLF systems recirculate CLFs, while cryogenic fluid is delivered only once due to immediate evaporation upon delivery. Therefore, the concept of consumption takes on different meanings in these systems. In cryogenic fluid delivery, the consumption rate is determined by the mass flow rate of LN through the nozzle. In contrast, in conventional flooding systems the consumption rate is determined by the volume 640

of emulsion per machine tool and the disposal interval.

A comparison of material production impacts broken down by components, given in [20], suggests that surfactants dominate the emissions for three of the six impact categories: energy use, acidification, and solid waste. The need for surfactants in conventional machining means that this technology has a significant environmental impact. However, in the case of cryogenic machining, there is only the need for electrical energy to produce the CLF, and process cooling water, while it has no other impact on the environment. When comparing obtained data, it has been proved that the level of energy needed for the liquidation of nitrogen is lower than for mineral oil production, when comparing the same amount of the two items. However, in the case of cryogenic machining, the delivered LN evaporates and is not reusable in the process, as it is in conventional machining. Therefore, a higher portion of energy is needed for its production.

In addition to production, the CLFs have to be delivered to the cutting zone (by a pump or pressure) from a reservoir. The energy for this depends on the CLF delivery rate. For this reason, the emissions and energy consumption for the delivery are given explicitly as a function of volumetric/mass flow rates and/or pressure.

The production of the CLF delivery equipment is not taken into consideration since the impacts are relatively small compared to the use phase and the equipment has a working life of several years [21]. The conditions used to estimate the power consumption from the delivery phase are the following: • Conventional machining: pump is used with a power of 500 W, providing 0.2 MPa pressure and a volume flow rate in the range of 0 to 8 l/min. Actual energy consumption is related to the CLF volume flow rate.

• Cryogenic machining: LN reservoir is pressurized and therefore pressure itself forces the CLF to the cutting zone. For this, no additional energy is needed.

In conventional machining, energy use is strongly influenced by the annual amount of CLF consumption. The electrical consumption of the delivery pump is small and therefore much less than CLF production energy (822 MJ). In

Pušavec, F. ‒ Kopač, J.

StrojniĹĄki vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647

cryogenic machining, energy use deriving from the LN production (136,858 MJ) is the dominating factor. A detailed analysis of energy consumption related to CLF production and use is given in [20].

Environmental impacts considered in the LCA are summarized in Fig. 3. The observing functional unit is case dependent. For ease of comparability and based on the case study, the following parameters were chosen: • Conventional machining: 1000 l/year of oilbased CLF usage, 2112 h/year machining hours, and 60 l/min emulsion delivery rate. • Cryogenic machining: 2112 h/year machining hours and 0.6 kg/min LN delivery rate.

The presented comparative life cycle burdens reveal that in cryogenic machining, burdens such as GWP, acidification, water use, and solid waste are eliminated at the price of the increased energy use required for nitrogen extraction and liquidation. In short, in cryogenic machining there is a trade-off with regard to higher energy use and a cleaner machining process. As we know that the production of LN requires immense electrical energy consumption, it is possible to talk about cryogenic machining being a sustainable process only when using a renewable energy to produce the LN. Conventional Cryo

1000

Water usage [kg]

GWP [kg CO2eq

1000

750 750 500 500 250 1000

750

500

250

250

Acidification [kg SO2eq]

Energy use [MJ]

38000

250

75000

110000

150000

250 500 500 750 750

1000

Solid waste [g]

1000

Land use [m2]

Fig. 3. Life cycle burdens In conventional machining, the CLF has to be disposed of. It needs to be removed from the part and chips after the machining, and then collected and recycled. All this represents additional processes, costs, and environmental burdens.

The usual procedure for oil-based CLF disposal consists of drying the emulsion and its subsequent combustion. In contrast to oils, emulsions do not have high energy values; therefore the combustion process must take into account the high potential for additional environmental burdens. Although combustion does recover some energy from the waste CLF, it additionally highly impacts GWP and acidification.

2.2 CLF Qualitative Characterization (Health, Safety and Performance) With regard to health and safety of machine tool operators, the effort to find machining alternatives has been driven largely by a desire to make machining processes safer, healthier, and thus more sustainable in view of the society. Chronic inhalation of oil-based mists has been shown to be responsible for serious health risks [17]. Such emulsion mists can harbor bacteria, and contain surfactants, biocides, chlorinated fatty, chelating agents and defoamers, all of which are harmful to health. This is notable since surfactants and biocides have been found to impair lung functioning [22]. In addition to mists, oil-based CLFs can cause dermatitis and other skin irritations. They also tend to result in the accumulation of an oily sludge on and around the production plant over time. Spills can also be a rather regular workplace hazard.

None of those are present in evaporated LN. More importantly, it has been proven that machining mist can be eliminated in cryogenic machining. On the other hand, in cryogenic machining less likely but more serious safety issues are required, related to the extremely low temperature of the pipes delivering the LN, which can cause physical burns in the event of contact. In addition, concerns have been raised in relation to the inhalation of metallic aerosols resulting from cryogenic machining. It has been reported that the development of lung disease due to the presence of micro metallic particles (tungsten carbide, cobalt, titanium) is a rare event and is almost unrelated to the duration and extent of exposure [23].

With regard to machining performance the LCA assumed that both processes employ the same machining time, while having comparable machining performance. In practice this is usually

Sustainability Assessment: Cryogenic Machining of Inconel 718

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not the case, therefore a detailed experimental case study that extends the LCA through a detailed machining evaluation, has been carried out.

3 MACHINING EVALUATION The machining evaluation has been carried out to demonstrate that cryogenic machining offers a cost-effective route to improve sustainability performance in comparison to conventional machining of Inconel 718. To determine the applicability of cryogenic machining, costs need to be calculated.

In the calculation of costs, tool-life is a key factor, therefore machining experiments were performed. Experiments were conducted on turning of Inconel 718 bars with a diameter of 40 mm and length of 100 mm. Machining employed SANDVIK GC 1105 grade carbide tool inserts and the CNMG120408-23 ISO tool geometry designation. Tool-life was assessed according to the ISO 3685, regulating an average tool-life criteria, VBmax = 0.4 mm. Tool-wear was measured with an optical microscope.

Conventional machining employed vegetablebased CLF (6.7% emulsion), with a flow rate of 6 l/min. The cryogenic machining was performed by applying LN under 1 MPa pressure and a flow rate of approximately 0.6 kg/min.

3.1 Hourly Rate of Machining System Usage The machining system usage hourly rate calculations covering operation and labor are presented in Table 1. The calculations consider 80% operating machining system efficiency and do not include shop floor space/rental costs.

The main differences in costs are the higher initial costs in the case of cryogenic machining, due to the additional equipment needed for the CLF delivery system. The results show that when using additional equipment, machining system usage costs are higher, and therefore benefits have to be gained within the process itself. In addition to machining system usage costs, the labor cost has to be added. For operating more sophisticated equipment a higher skill level of labor is required.

Summing machining system usage costs and labor cost determines the overall machining cost, Cmh, i.e. hourly rate. Machining costs for cryogenic 642

machining are 17% higher in comparison to conventional machining.

Table 1. Hourly rate of machining system usage Categories Machine tool costs [€] (a) Tooling (3% of (a))[€] (b) CLF delivery system [€] (c) Machining system costs [€] Annual depreciation [€/year] Maintenance (1.5 % of (a)+(b)+(c)) [€] Insurance (0.4 % of (a)+(b)+(c)) [€] System usage costs [€/h] Direct labor [€/h] (d) Indirect labor (10 % of (d)) [€] Supervision (12 % of (d)) [€] Fringe benefit (33 % of (d)) [€] Cost of labor [€/h] Machining costs, Cmh [€/h]

Conv.

150000 4500 0 157500 22500

Cryo.

150000 4500 10000 167500 23929

2318

2468

618

658

12.53 12

13.29 15

1.20

1.50

1.44

1.80

3.96

4.95

18.60 31.13

23.25 36.54

3.2 CLF Consumption and Costs In order to determine the CLF contribution to the machining cost, the hourly rate has to be calculated. This calculation is given in Table 2.

Table 2. CLF consumption and costs Categories Conv. Cryo.

CLF concentrate [€/l] 10 / CLF disposal [€/l] 0.2 / CLF amount [l] 450 / CLF concentrate needed [l] 28.12 / CLF concentrate costs [€] 281.25 / CLF disposal [€] 90 / CLF maintenance [€] 60 / Overall CLF costs [€] 431.25 / Duration life [h] 2640 / Nonreturnable CLF [kg/min] / 0.6 Nonreturnable CLF costs [€/kg] / 0.21 CLF costs, Cch [€/h] 0.17 7.51 From an environmental perspective, cryogenic machining is preferable due to the complete elimination of oil-based CLFs. For the conventional case, the amount of usage and its

Pušavec, F. ‒ Kopač, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647

lifetime has to be determined in order to calculate its hourly rate. In this case, calculations for an annual period are made.

In cryogenic machining, where LN is used, the CLF is not reusable because it evaporates into the air immediately when it is delivered. Due to this, CLF represents a directly consumed item, which has a relatively high cost but does not need to be recycled. In calculating costs, the LN hourly rate (Cch = 7.51 €/h) is significantly higher than for the CLFs in the conventional method (Cch = 0.17 €/h).

3.3 Costs Associated with Waste Another important sustainability measure that has to be considered refers to the waste produced during machining. Waste products are mostly connected to oil-based CLFs, as discussed above, worn cutting tools, and chips/swarf. In the presented study, the amount of swarf produced is assumed to be equal for both machining cases. This results in equal costs related to swarf compacting (including shredding if needed), which is required to ease transportation. However, conventional machining includes an additional cost for separation of CLF from the swarf. The CLF separation usually includes separation of unemulgated oils through skimming, separation of the hard particles by means of filtration, emulsion separation and treatment of the separated water.

3.4 Total Production Cost per Part The total production cost per part is given in Table 3. The presented partial costs included in the total production cost are valid for machining with the following parameters: • Cutting speed: vc = 90 m/min.

• Feed: f = 0.25 mm/rev.

• Depth of cut: ap = 1.2 mm.

In the total production cost, cp,p, machining cost, cm,p, represents the expenses of the machining system and labor needed for one part.

During machining tool-wear occurs, therefore tool cost per part, ct,p, is added, as well as CLF usage cost per part, cCLF,p. The energy consumption cost, cE,p, is divided into the cost when the machine tool is performing the actual machining and the cost when the machine tool is in a stand-by mode

(e.g. when changing parts and cutting inserts).

The additional costs include the cost for cleaning the CLF on the surface of a machined part, cpcl,p, the cost for separating CLF from the swarf, and the cost for compacting swarf for ease of transportation, csp,p.

Table 3. Production costs Categories Number of parts per tool life time Tool changing time [s/part] Machining cycle time [s/part] Part production rate [part/h] Machining cost, cm,p [€/part] Tool cost, ct,p [€/part] CLF cost, cCLF,p [€/part] Down time [s/part] Electrical usage [kWh/part] Cost of electricity, cE,p [€/part] Part cleaning cost, cpcl,p [€/part] Swarf preparation cost, csp,p [€/part] Total production cost, cp,p [€/part]

Conv.

1 180 243.5 14.8 2.11 2.50 0.002 210 0.076 0.009 0.063

Cryo.

6 30 93.5 38.5 0.95 0.42 0.07 66 0.055 0.007 0

0.015 0.004 4.69

1.45

From Table 3, it can be seen that the total production cost is greatly affected by tool-life, CLF cost, disposal, and waste management.

Considering the tool-life in conventional machining, a huge improvement can be observed in cryogenic machining, as described in [24].

It has been shown that the conventional method yields the shortest tool-life for all cutting speeds in the tested range. If cryogenic machining is used, a significant increase in tool-life was achieved due to highly efficient cooling needed when machining high-temperate alloy (Inconel 718). If this condition is not satisfied, the temperature in the cutting zone rapidly rises, softening the cutting tool material and causing rapid tool-wear, as in the case of conventional machining.

Moving one step further, production time is not correlated only to the machining time, but also to the machining system down time due to changing the worn cutting tool insert. It has been confirmed in our previous experimental study that conventional machining yields the longest production time. By applying cryogenic machining production time can be decreased by up to 63%, depending on which cutting speed

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is used [24]. This reduction can be attributed to the tool changing time, which has a much higher contribution to total production cost in conventional machining. The actual cutting tool changing time is not longer, but the number of changes is higher, due to the increased rate of toolwear in conventional machining.

By combining tool-life and production times, the total production cost per part can be determined for different cutting speeds, as presented in Fig. 4.

4 3.5

Cryogenic

Production costs per part cp,p [€],

Excessive tool-wear

Conventional

5 4.5

3

cm,p ct,p cE,p cCLF,p ccl,p

2.5 2 1.5 1 0.5 0

30

60

90 120 30 60 Cutting speeds vc [m/min]

90

120

Fig. 4. Production costs The total production costs for different cooling conditions are represented by stacked bars, indicating the individual contributions to the total of the production cost at the cutting speed of vc = 75 m/min. The solid line represents the changing trends of the total production cost per part due to the cutting speed variation.

From the presented results, it is possible to assert that conventional machining is significantly more expensive than cryogenic machining. This trend is even more dominant if high cutting speeds are employed. What is interesting from these plots is that at lower cutting speeds conventional machining can be the cheapest. However, this production rate is not optimal. Therefore, cryogenic machining should be used when high efficiency and high productivity are required.

Regarding the contribution to the overall production cost, it can be seen that machining costs and costs arising from changing the cutting tool are the highest in conventional machining.

Energy consumption costs are almost negligible in both cases, while coolant costs are almost 644

negligible in conventional machining. The exception is cryogenic machining, where the price of the LN is much higher. While conventional machining has low CLF cost, the costs of cleaning are higher, while they are negligible in the case of cryogenic machining.

4 LCA UPGRADE - ADDING TOOL INSERT PRODUCTION ENERGY CONSUMPTION Since two comparing processes (cryogenic and conventional machining) differ in cutting tool life and thus in the overall cost for consumed cutting inserts, also the environmental impacts associated with their production should be considered. In this way, the energy spent for their production is going to be determined and compared with the additional electrical energy spent in cryogenic machining for the production/ extraction of nitrogen from the air.

In this work, carbide tools were used in experiments and have been analyzed in this work. In practice, the starting materials for the manufacture of carbide tools are hard refractory carbides (i.e. tungsten carbide) and metal binder (i.e. Cobalt), both in the form of powder. To achieve good toughness and hardness of WC– Co tool, fabrication of composite powder of nanometer size is preferred. For this process ball milling comminution method in carbide industry is used. The WC–Co composite powder will then go through three sequential procedures, i.e. pressing, sintering and machining to convert shape to the final product i.e. cutting tool. All those processes, especially the milling processes are energy intensive processes. Therefore, in this paper, life cycle inventory related to energy consumption, for producing cutting insert is evaluated based on the process data extracted from [25]. The boundary of energy consumption is limited to include ball milling, pressing, sintering and grinding processes. The energy/material consumption inventory for producing WC–Co cutting insert is listed in Table 4. Since the tool insert used in this paper has CNMG shape with a total weight of 9 g, the inventory data are scaled based on weight, except for the grinding time, which is scaled based on total area of surface requiring grinding.

Pušavec, F. ‒ Kopač, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647

Categories Quantity Cutting tool weight [g] 9 Surface to grind [cm2] 5.3 Milling specific energy [kWh/g] 0.52 Milling energy consumption [kWh] 4.68 Pressing energy consumption [kWh] 0.008 Sintering energy consumption [kWh] 0.5 Grinding energy consumption [kWh] 0.34 5.53 Total energy consumption [kWh] =19.9MJ * In addition to energy consumption there are also consumables like inert argon gas, water, grinding CLF, etc. that also affect life cycle impact. However, this is beyond the scope of this work.

From the results in Table 4 it can be seen that 5.53 kWh of electrical energy is spent for the production of a cutting tool. This means that lowering the tool-wear or a prolongation of toollife in machining production is going to result in lower energy consumption in production of cutting tools. Additionally, this can be used to justify the difference of cryogenic and conventional machining process overall energy consumption that was initially on the side of conventional machining on account of the liquid nitrogen production (extraction of air).

To correlate these results, the total energy consumption is calculated for conventional and cryogenic machining/production of product specified in section 3 at different batch sizes. The analysis includes: • energy consumption for production of CLF, • energy consumption for production of cutting tools, • energy consumption for machining process.

The results, correlating the batch size, productivity via cutting speed and total energy consumption are shown in Fig. 4. It can be seen that the batch size does not significantly affect the total energy consumption process, while cutting speed is a significant parameter. With increasing cutting speed, the tool-wear is nonlinearly increasing, while the machining time (for the whole batch size) is nonlinearly shortening. The first results in lower energy consumption for cutting tool production, while the latter means lower energy consumption in using CLF and the

energy spent for the machine tool usage (both were scaled from the annually calculations to the exact batch size production time). It is possible to see that at small cutting speeds, the LN extraction is the dominant energy consumption source, while at higher cutting speeds, cutting tools become dominant energy consumption sources (high toolwear). The threshold, where cryogenic machining becomes more energy efficient than conventional machining in Inconel 718, is at approximately vc = 90 m/min.

Energy consumption [MJ]

Table 4. Input inventory of WC–Co cutting insert manufacturing

Conventional Cryogenic Treshold

Fig. 5. CLF, cutting tools production and machining process total energy consumption Based on these results it can be concluded tha cryogenic machining, even due to the fact that LN extraction is a very intensive process in terms of energy, can be more energy efficient that conventional machining on the account of the prolongation of tool-life and thus decreased cutting tool consumption.

5 CONCLUSIONS Challenges in production with regard to the economy, society, and the environment in view of machining technology, are discussed in the first part of this paper. More specifically, LCA aims to convince the industry of the merits of sustainable machining technologies, taking into account the overall life cycle of the CLFs. In this respect, cryogenic machining is presented as a viable and sustainable machining technology in comparison to conventional machining. The LCA demonstrated that transitioning from oil-

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 637-647

based CLFs to LN used in cryogenic machining is a move towards more sustainable machining, which results in a significant reduction in solid waste, water usage, global warming potential, acidification, and in an increased energy use for CLF production. However, it has been found that in addition to the nitrogen extraction, the production of cutting tools is also a very energy consuming process is also. Based on the comparative analysis and calculation of total production energy consumption, it has been proved that cryogenic machining can be more energy efficient than conventional machining. This goes on account of drastic reduction on cutting tools consumption.

In the second part of this paper an experimental machining study of Inconel 718 is discussed. The tool-wear was measured and used for the determination of tool-life. Once cutting tool-life was known, the cryogenic CLF was evaluated in terms of the total production cost per part, covering all sustainability measures. It has been shown that the elimination of conventional CLFs, the reduction of costs associated with waste and higher tool-life in cryogenic machining drastically reduces total production cost per part in comparison to conventional machining (by up to 30%). This confirms that even though the initial cost and effort involved with the cryogenic machining system is higher, it can obviously offer significant sustainability benefits through shorter production cycles and a lower cost needed to machine a part as well as enhanced productivity due to higher output.

An additional evaluation of Inconel 718 machining, not reported here, proved that cryogenic CLF increases machining reliability while maintaining dimensional tolerances and improving machined surface integrity. However, the reliability of the cryogenic CLF delivery system itself is not yet clear. For this, the industrial implementation of the system that reduces consumption rates (costs), environmental burdens, and health risks, while simultaneously increasing machining performance and profitability, is required.

6 ACKNOWLEDGEMENTS The work has been performed in collaboration with the industrial partner 646

ISTRABENZ Plini (Slovenia, EU) that provided the equipment and helped with the technical solutions. Gratitude also goes to the EUREKA platform project foundation that is financially supporting the SusCryMac research project.

7 REFERENCES [1] Ana, Q.L., Fub, Y.C., Xub, J.H. (2011).

Experimental study on turning of TC9 titanium alloy with cold water mist jet cooling. International Journal of Machine Tools & Manufacture, vol. 51, p. 549-555.

[2] Adamczak, S., Čuš, F., Miko, E. (2009).

A model of surface roughness constitution in the metal cutting process applying tools with defined stereometry. Strojniški vestnik Journal of Mechanical Engineering, vol. 55, no. 1, p. 45-54.

[3] Zuperl, U., Cus, F. (2004). A determination of the characteristic technological and economic parameters during metal cutting. Strojniški vestnik - Journal of Mechanical Engineering, vol. 50, no. 5, p. 252-266.

[4] Zuperl, U., Cus, F., Gecevska, V. (2007).

Optimization of the characteristic parameters in milling using the PSO evaluation technique.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 53, no. 6, p. 354-368.

[5] Cus, F., Zuperl, U., Kiker, E. (2007). A modelbased system for the dynamic adjustment of cutting parameters during a milling process.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 53, no. 9, p. 524-540.

[6] Jovane, F., Westkamper, E., Williams, D. (2009). The Manufuture Road. SpringerVerlag, Berlin, Heidelberg.

[7] Jovane, F., Yoshikawa, H., Alting, L., Boer, C.R., Westkamper, E., Williams, D., Tseng, M., Seliger, G., Paci, A.M. (2008). The incoming global technological and industrial revolution towards competitive sustainable manufacturing. CIRP Annals - Manufacturing Technology, vol. 57, no. 2, p. 641-659.

[8] Herrmann, C., Hesselbach, J., Bock, R., Zein, A., Ohlschlager, G., Dettmer, T. (2007).

Ecologically benign lubricants - evaluation from a life cycle perspective. Clean Soil Air Water, vol. 35, no. 5, p. 427-432.

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[9] Yildiz, Y., Nalbant, M. (2008). A review of cryogenic cooling in machining processes.

International Journal of Machine Tools & Manufacture, vol. 48, no. 9, p. 947-964.

[10] Hong, S.Y., Ding, Y., Jeong, W. (2001).

Friction and cutting forces in cryogenic machining of Ti-6Al-4V. International Journal of Machine Tools & Manufacture, vol. 41, p. 2271-2285.

[11] Hong, S.Y., Ding, Y. (2001). Cooling Approaches and Cutting Temperatures in Cryogenic Machining of 6Al-4V. International Journal of Machine Tools & Manufacture, vol. 41, p. 1417-1437.

[12] DaSilva, F.J., Franco, S.D., Machado, A.R., Ezugwu, E.O., Souza, A.M.J. (2006). The performance of cryogenically treated HSS tools. Wear, vol. 261, p. 674-685.

[13] Pusavec, F., Kramar, D., Krajnik, P., Kopac, J. (2010). Transition to sustainable production - Part II: Evaluation of sustainable machining technologies. Journal of Cleaner Production, vol. 18, no. 12, p. 1211-1221.

[14] Pusavec, F., Deshpande, A., M’Saoubi, R., Kopac, J., Dillon, O.W.J., Jawahir, I.S. (2008). The modeling and optimization of machining of high temperature nickel alloy for improved machining performance and enhanced sustainability. Proceedings of the 11th CIRP conference on Modeling of Machining Operations, p. 21-28.

[15] Zhao, Z., Hong, S.Y. (1992). Cooling strategies for cryogenic machining from a materials viewpoint. Journal of Materials Engineering and Performance, vol. 1, no. 5, p. 669-678.

[16] Wang, Z.Y., Rajurkar, K.P., Fan, J., Lei, S., Shin, Y.C., Petrescu, G. (2003). Hybrid machining of inconel 718. International Journal of Machine Tools & Manufacture, vol. 43, no. 13, p. 1391-1396.

[17] Skerlos, S.J., Hayes, K.F., Clarens, A.F., Zhao, F. (2008). Current advances in sustainable metalworking fluids research. International Journal of Sustainable Manufacturing, vol. 1, no. 1-2, p. 180-202.

[18] Ostlin, J., Sundin, E., Bjorkman, M. (2009). Product life-cycle implications for remanufacturing strategies. Journal of Cleaner Production, vol. 17, no. 11, p. 9991009.

[19] McManus, M.C., Hammond, P.H., Burrows, C.R. (2004). Life-cycle assessment of mineral and rapeseed oil in mobile hydraulic system.

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[20] Pusavec, F., Krajnik, P., Kopac, J. (2010).

Transition to sustainable production – Part I: Application on machining technologies.

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[21] Clarens, A.F., Zimmerman, J.B., Keoleian, G.A., Hayes, K.F., Skerlos, S.J. (2008).

Comparison of life cycle emissions and energy consumption for environmentally adapted metalworking fluid system.

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[22] Cohen, H., White, E.M. (2006). Metalworking fluid mist occupational exposure limits discussion of alternative methods. Journal of Occupational and Environmental Hygiene, vol. 3, no. 9, p. 501-507.

[23] Ruediger, H.W. (2000). Hard metal particles and lung disease: coincidence or causality?

Respiration, vol. 67, p. 137-138.

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Proceedings of the 42nd CIRP Conference on Manufacturing Systems, Grenoble, France.

[25] Zhao, F., Bernstein, W.Z., Naik, G., Cheng, G.J. (2010). Environmental assessment of laser assisted manufacturing: case studies on laser shock peening and laser assisted turning.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 648-654 DOI:10.5545/sv-jme.2010.062

Paper received: 19.03.2010 Paper accepted: 13.06.2011

Length Variation of Toothed Belt during Exploitation

Stojanović, B. ‒ Miloradović, N. ‒ Marjanović, N. ‒ Blagojević, M. ‒ Ivanović, L. Blaza Stojanović* ‒ Nenad Miloradović ‒ Nenad Marjanović ‒ Mirko Blagojević ‒ Lozica Ivanović University of Kragujevac, Faculty of Mechanical Engineering, Kragujevac, Serbia Timing belt drives are a relatively new concept in power transmission, accepted nowadays in all areas of industry. Teeth, equally spaced on the inner side of timing belts, come into contact with the belt pulley’s teeth through their grooves. By this meshing, a connection between the belt and the belt pulley is achieved and torque is transmitted.

This paper reviews basic tribomechanical systems in timing belt drives, focusing on the analysis of the tribomechanical system “belt’s teeth - belt pulley’s teeth”. In addition, analysis of pitch variation and belt’s length variation during testing is conducted. Friction is the main cause of wear of the flanks of the belt’s teeth, which leads to an increase of pitch. Testing of the timing belt was conducted on a specially designed test bench and showed that the belt’s elongation is especially distinct in a period of running in. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: timing belt drives, belt pitch, friction, wear, tribology, tribomechanical systems 0 INTRODUCTION A timing belt drive is a relatively young drive designed by Case in 1946 [1]. It was a rubber belt with trapezoidal teeth profile used as a transmitter on a sewing machine. Despite the advantages in operation, timing belt drives have just recently gained large application. It was not until the application of timing belts as IC engine’s camshaft drive that usefulness of their application became obvious. The intensification of design demands in terms of an increased service life and a decreased construction mass have initiated the appearance of a large number of tests of timing belt drives [2] and [3].

Gerbert et al. [4] formed the first model of a timing belt and conducted a detailed analysis of forces acting on the belt teeth. They introduced friction force into the analysis of load distribution.

Further steps in analysis of these transmitters were achieved by testing of their tractive characteristics [5], by analysing load distribution and by pretension of the timing belt [6] and [7]. Kido et al.

[8] presented the first analysis of load distribution using the finite element method. Soon afterwards, the finite element method was beginning to be increasingly more introduced in the analysis of timing belt drives. Karolev and Gold [9] presented a new, changed form of the timing belt model. The analysis of load distribution takes into account the friction force and belt deformation and tests are conducted under variable torque.

648

Considering an increasingly more frequent use of timing belt drives, but also their limited service life, the analysis of friction and wear became more and more significant in the late 1990’s. Dalgarno, Childs et al. [10] to [15] introduced new models of timing belts, taking into account the friction force between belt’s teeth and belt pulley, together with friction force analysed until then only between belt pulley’s teeth apexes and the belt’s groove. Dalgarno et al. gave detailed analyses of meshing and tribological processes on contact surfaces.

Introduction of friction force into a new model of belt [16] and [17] and belt’s model with mass [18] and [19] showed that the results obtained by application of numerical methods completely coincide with experimental results, regarding load distribution. Paper [21] presents the analysis of deformation state through a model developed by application of finite element method.

Papers [22] to [23] demonstrate the modelling of friction induced noise of timing belt drives.

The retrospective of the existing tests of timing belt drives is mostly related to analyses of meshing and load distribution, both analytical and numerical. The aim of this review is to identify basic tribomechanical systems in timing belt drive and analyze them. Wear that leads to an increase of pitch and elongation of the belt appears to be a direct consequence of friction on the flanks of the belt and of the belt pulley. Elongation of the belt induces changes in kinematics of coupling and

*Corr. Author’s Address: University of Kragujevac, Faculty of Mechanical Engineering, Sestre Janjić 6, 34000 Kragujevac, Serbia, blaza@kg.ac.rs

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 648-654

increased friction that brings about larger power losses, that is reduction of the efficiency of the drive.

Table 1. Tribomechanical systems and types of motion in timing belt drives Tribomechanical system

1 TRIBOMECHANICAL SYSTEM IN TIMING BELT DRIVES

belt’s tooth – belt pulley’s tooth

The largest amount of motion and power is transferred by shape, while only a small amount is transferred by friction. The influence of friction should not be neglected. Appearance of friction in timing belt drives and its consequences have not been thoroughly explained. In contrast to other transmissions of power and motion (gears, chain drives, cardanic transmissions, etc.) in which friction mostly occurs in the contact of the two metal surfaces, in timing belt drives, there are one metal and one non-metal surface. The timing belt pulley of the tested timing belt drive is made of C45 steel (WNr 1.0503) with the initial teeth roughness of Ra = 1.6 µm. The belt’s material is rubber whose initial teeth roughness depends on measurement point (apex, flank or groove), i.e.

Ra = 9 to 12.5 µm.

The basic tribomechanical systems in the timing belt drives are (Fig. 1) [24]: 1. belt’s tooth – belt pulley’s tooth, 2. belt’s face – flange, 3. the belt groove – apex of the belt pulley’s tooth.

belt’s face - flange

Fig. 1. Timing belt drive and basic tribomechanical systems Types of motion that occur in these tribomechanical systems are given in Table 1.

The flank of the belt’s teeth makes contact with the flank of the belt pulley’s teeth, after entering the meshing. In addition, the inner surface of the belt groove and the outer surface of the belt pulley and, from time to time, the front surface of the belt pulley with the flange ring, are in contact.

the belt groove – apex of the belt pulley’s tooth

Type of motion - impact - sliding - rolling - impact - sliding - sliding - rolling

The belt’s tooth enters the meshing with the drive belt pulley, maximally strained due to previous tension. While entering the meshing, the belt tooth’s apex contacts the flank of the belt pulley’s tooth. At that moment, a line contact occurs, i.e. pure roll occurs. Due to the interference, the belt’s tooth cuts into the flank of the belt pulley’s tooth. Due to elastic properties of the belt and the large stiffness of the belt pulley, deformation of the belt’s tooth occurs (Fig. 2, position 4). The usual initial assumption is that the stiffness of the belt pulley’s is indefinitely large (rigid body) compared to the belt’s stiffness [4], [9], [10], [15], [17]. Deformation of the belt’s tooth grows, while, at the same time, the contact surface between the belt and the belt pulley increases. The contact point between the belt’s tooth and the belt pulley tooth moves from the belt pulley’s tooth apex towards its root.

Fig. 2. Layout of belt’s teeth entering the meshing with the teeth of the drive belt pulley Maximal tooth deformation takes place in position 2 (Fig. 2). The reduction of deformations occurs due to action of internal stresses and turning of the belt and the belt pulley. Full coincidence of the flanks of the belt’s teeth and the belt pulley’s teeth occurs in position 1 (Fig. 2).

Now, contact over surface occurs. Relative sliding

Length Variation of Toothed Belt during Exploitation

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of their flanks, with appearance of the friction force, follows the process of belt’s teeth entering the meshing with the belt pulley. The value of normal force varies according to parabolic law, which leads to variation of the friction force. The greatest values of normal force and friction force are at the teeth’s roots.

Value of the friction force increases with the increase of the length of the sliding path and achieves its greatest value at the root of the belt’s tooth (Fig. 3). At the same time, the action point of the resultant component of normal force moves from the tooth’s apex towards its root. The normal force changes according to parabolic law [25]: Ni = −

N max lt2

2

⋅ ( l − lt ) + N max ,

(1)

where Nmax is maximal value of normal force (Nmax ≈ 1.5 Fo / z01) and lt is the length of friction path.

The friction force occurs at the flank of the belt’s tooth and its value is determined according to the following Eq. (2):

Fti = Ni × µ =

Foi × µ , cos( β / 2)

(2)

where Ni is normal force acting on the belt’s tooth, µ is the friction coefficient, Foi is circumferential 650

2 TESTING OF TIMING BELT DRIVE Testing of timing belt drive was conducted on a specially designed test bench, made at the Laboratory for Mechanical Constructions and Mechanization of the Faculty of Mechanical Engineering in Kragujevac. Test bench operates on a principle of opened loop power [26].

Fig. 4. Test bench for testing of timing belt

Fig. 3. Friction force at the flanks of the belt’s tooth

force acting on the belt’s tooth and β is the angle of the belt’s profile.

are: 1.

2.

3.

4.

5.

6.

7.

8.

9.

10.

Basic elements of the test bench, Fig. 4,

drive unit (electric motor), cardanic drive, measuring (input) shaft, input shaft’s rotational speed transducer, input shaft’s torque transducer, tested drive (timing belt drive), output shaft, mechanical brake, tension mechanism and amplifier bridge.

A toothed disc with 30 teeth on circumference is mounted on a measuring shaft.

Rotational speed is read on the amplifier bridge that acquires a signal from inductive sensor HBM M1 and rotational speed pulse receiver HBM DV2556. The torque of the shaft is measured with contactless torque sensor that consists of a strain gauge, a signal transmitter HBM MT2555A and a signal receiver HBM EV2510A. Values of rotational speed and torque were displayed on a display of a digital amplifier DA24.

Stojanović, B. ‒ Miloradović, N. ‒ Marjanović, N. ‒ Blagojević, M. ‒ Ivanović, L.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 648-654

In order to obtain a true picture of tribological characteristics of the timing belt, measurement of roughness parameters and determination of geometrical values are conducted. Measurement of these values is conducted according to previously determined dynamics.

Prior to testing, the state of the contact surfaces and initial values of the belt’s geometrical values were established. Further measurements were conducted after a certain operation time and are shown in Table 2.

conducted on optical microscope ZEISS ZKM01250C. Belt’s pitch variation (Δt) may be shown with the help of the following Eq.:

Table 2. Time intervals of measurement of roughness parameters and the belt’s geometrical values

Table 3. Belt’s pitch variation Δt = t - to [µm]

Number of 1 measurement Operation 0 time [h]

2

3

4

5

6

7

8

9

10

5

10 20 50 100 150 200 250 300

Fig. 5. Measured geometrical values of the belt 3 VARIATION OF BELT’S PITCH Measurement of geometrical values of timing belts was conducted on eight belt’s teeth.

The following values were measured (Fig. 4) [27]: 1. belt’s pitch (t), 2. belt’s width (b), 3. belt groove’s thickness (hb = hs ‒ ht) and 4. belt’s total height (hs).

Belt’s pitch is a distance between centres of two consecutive teeth and is measured on the so-called pitch line. Considering the design realization of the belt and the available apparatus, the belt’s pitch is measured at the tooth’s root.

Belt’s pitch is determined with the rollers.

Namely, the rollers having corresponding radii are set at the roots of the two successive belt’s teeth (Fig. 5). Then, the belt with the rollers is set on the auxiliary tool. The tension force exerted on the belt is the same for all measurements. The total length of the belt is determined based on the belt’s pitch and the number of teeth. Measurement is

Δt = t ‒ to , (3)

where: t is the measured value of belt’s pitch during testing and to is the initial value of belt’s pitch.

Measurement results of belt’s pitch variation during the operation for all eight teeth are given in Table 3 and shown in Fig. 6.

Time of operation [h] 5 10 20 50 100 150 200 250 300

1 123 123 135 152 170 175 185 204 205

2 121 122 133 152 169 174 184 197 202

3 121 121 133 151 167 173 184 197 207

Δt Belt’s tooth 4 5 118 121 119 120 129 131 148 150 163 167 170 172 181 182 196 205 208 210

6 118 118 129 147 166 170 178 196 198

7 121 121 131 151 168 174 183 201 211

8 119 119 130 150 166 173 182 198 200

In addition, the belt’s length is obtained with the help of centre distance, by mounting of the same belt on the two belt pulleys having transmission ratio of 1, with no previous tension.

Values of the belt’s length determined in both ways were approximately equal, before the belt was tested.

4 ANALYSIS OF TRIBOLOGICAL PROCESSES In the period of running in, there is a sudden increase of belt’s pitch. This increase originates from plastic deformation of the belt’s tractive element and from the wear of teeth’s flanks. Wear by roll-formation (a special form of elastomeric wear) is especially emphasized there and the consequences are a removal of material from the belt’s teeth and an increase of pitch [28].

In the period of normal wear, which appears after 20 hours of operation, variation of geometrical values is still strong. After 20 hours of operation, the belt’s pitch is still increasing.

Variation of the belt’s pitch is more pronounced

Length Variation of Toothed Belt during Exploitation

651

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in the period from 20 to 50 hours of operation, after which it becomes approximately linear. The results obtained by measurement on all the eight teeth almost do not deviate from each other.

Belt pitch variation [μm]

250 200 150 100 50

tooth 8

Operation time [h]

250

150

50

10

0

0

tooth 1

Fig. 6. Belt’s pitch variation during exploitation

Variation [μm]

250

205.13

200 150 76.375

100 50 0

54.875

26.25

1

2

3

4

Geometric value

Fig. 7. Average values of variations of geometric values; 1 - Belt’ pitch, 2 - Belt’s width, 3 - Belt’s total height, 4 - Belt groove’s thickness Absolute average values of variation of geometrical values are presented in Fig. 7.

The histogram shows that the belt’s pitch changes the most. The belt’s pitch increases for approximately 0.2 mm, which leads to an increase of the belt’s length. Total elongation of the belt is approximately 23 mm. An increase of the belt’s pitch induces variations in kinematics of coupling, reduction of contact surface, increase of friction and the need for additional tightening.

Variation of the belt’s pitch arising in the period of running in is a direct consequence of 652

elongation of the tractive element, i. e. elongation of the belt. Thereat, the tractive element was permanently elongated. Elongation of the belt’s pitch is the largest in the period of running (Table 2) and amounts 60% of total elongation. After the period of running in, during the period of normal wear, the belt’s pitch was still increasing.

The variation of the belt’s pitch has almost linear tooth 1 form and it changes continuously until the end oftooth the 2test. At the same time, the total length of thetooth belt3 changes very little, i.e. the total variation oftooth the 4belt’s pitch is considerably larger than the variation tooth 5 of the belt’s length. A newly created variation of pitch arises, in the first place, due to tooth 6 wear of the flanks of the teeth. tooth 7 tooth 8

5 CONCLUSIONS

An increase of the belt’s pitch arises due to plastic deformations of the belt’s tractive element and wear by roll-formation of teeth’s flanks. An increase of length (elongation) of the tractive element gradually occurs during exploitation and stays permanently, even after the belt is unloaded.

A large portion of this increase, approximately 70%, occurs due to plastic deformation of the belt, while the rest of it is due to wear by rollformation of the belt teeth’s flanks. Participation of wear by roll-formation in total elongation of the belt increases with the increase of timing belt’s operation time as plastic deformation is the greatest in the period of running in. Variation of the belt’s pitch leads to disturbance in operation of timing belt as there are changes in load distribution, reduction of carrying capacity and unevenness in operation. There is a need for additional tensioning of the belt, which directly affects the service life of the drive.

6 REFERENCES [1] Case, Y.R. (1954). Timing belt drive. McGraw Hill Book Company, Inc., New York.

[2] Stojanovic, B. (2007). Characteristics of tribological processes in timing belts.

Master’s thesis, Faculty of Mechanical Engineering from Kragujevac, Kragujevac. (in Serbian) [3] Stojanovic, B., Miloradovic, N. (2009).

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1780-1787.

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[8] Kido, R., Kusano, T., Fujii, T. (1995). Finite element analysis for distribution of toothed belts in two axes drive system. Transactions of the Japan Society of Mechanical Engineers, vol. 61, p. 130-136. (in Japanese) [9] Karolev, N.A., Gold, P.W. (1995). Load distribution of timing belt drives transmitting variable torque. Mechanism and Machine Theory, vol. 30, no. 4, p. 553-567.

[10] Dalgarno, K.W., Day, A.J., Childs, T.H.C. (1993). Finite element analysis of synchronous belt tooth failure. Proceedings of the Institution of Mechanical Engineers, Part D, Journal of Automobile Engineering, vol. 207, p. 145-153.

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[14] Childs, T.H.C., Dalgarno, K.W., Day, A.J., Moore R.B. (1998). Automotive timing belt life laws and a user design guide. Proceedings of the Institution of Mechanical Engineers, Part D, Journal of Automobile Engineering, vol. 212, p. 409-419.

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[21] Zupancic, B., Nikonov, A., Florjancic, U., Emri, I. (2007). Time-dependent behavior of drive belts under periodic mechanical loading – an analysis of the location of a single line spectrum. Strojniški vestnik - Journal of mechanical Engineering, vol. 53, no. 10, p.

696-705.

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at timing belt’s tooth flank. Tribology in Industry, vol. 31, no. 3-4, p. 53-58.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 655-666 DOI:10.5545/sv-jme.2010.162

Paper received: 20.07.2010 Paper accepted: 27.07.2011

Signal Model-Based Fault Detection and Diagnosis for Induction Motors Using Features of Vibration Signal in TwoDimension Domain Do, V.T. – Chong, U.-P. Van Tuan Do1,* – Ui-Pil Chong2 1 VTT Technical Research Centre of Finland, Finland 2 University of Ulsan, Department of Computer Engineering and Information Technology, South Korea

In this paper, we propose an approach for vibration signal-based fault detection and diagnosis system applying for induction motors. The approach consists of two consecutive processes: fault detection process and fault diagnosis process. In the fault detection process, significant features from vibration signals are extracted through the scale invariant feature transform (SIFT) algorithm to generate the faulty symptoms. Consequently, the pattern classification technique using the faulty symptoms is applied to the fault diagnosis process. Hence, instead of analyzing the vibration signal to determine the induction motor faults, the vibration signal can be classified to the corresponding faulty category, which presents the induction motor fault. We also provide a framework for the pattern classification technique that is applicable to SIFT patterns. Moreover, a comparison with two other approaches in our previous work is also carried out. The testing results show that our proposed approach provides significantly high fault classification accuracy and a better performance than previous approaches. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: fault detection and diagnosis, SIFT, feature vector, texton dictionary, two-dimension domain, classification accuracy

0 INTRODUCTION Systems for detection and diagnosis of malfunctioning machines play an important role in industrial fields. They are critical in the manufacturing industry, since a bad manufacturing machine may produce many defective products dangerous to consumers. An investigation for the earliest possible detection for a machine before it becomes faulty is therefore compulsory.

The approaches for fault detection and diagnosis consist of two processes: fault detection process and fault diagnosis process as shown in Fig. 1. The fault detection process analyzes the measured signals such as vibration, noise, acoustic sound, pressure or bases on the analytical parameters to generate the faulty symptoms, which can be analytical symptoms or heuristic symptoms [1] to [2]. The faulty symptoms are the input of the fault diagnosis process that determines the size, type and location of the system fault [3].

Depending on a specific application, simple or complex techniques for fault detection process and fault diagnosis process can be applied.

fault Process Measured signals or observed variables

Signal or process model Feature extraction features Feature database

FAULT DETECTION

Symptoms Fault diagnosis: Fault size Fault location Fault types Etc.

FAULT DIAGNOSIS

Fig. 1. General structure of model-based fault detection and diagnosis For example, limit checking techniques are widely used for simple fault detection systems.

However, they have two disadvantages: it is impossible to predict the fault in advance since the fault has already occurred when detected and the methods do not provide the type, size and location of faults, which can be provided by applying

*Corr. Author’s Address: VTT Technical Research Centre of Finland, FI-02044, Espoo, Finland, dtuan@ualberta.ca

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model-based fault detection technique such as approaches in [4] to [6].

The Induction motor is a three phase AC motor and is the most widely used machine. Its characteristic features are: simple and rugged construction, low cost and minimum maintenance, high reliability and sufficiently high efficiency, it needs no extra starting motor and need not be synchronized. During the operation; however, there are several types of faults frequently happening such as: bearing faults, gear faults, rotor bar eccentricity and stator winding failures and misalignment. The fault detection and diagnosis methods for induction motors are wide such as current spectrum analysis, vibration analysis and acoustic analysis for different types of motor fault identification [7] to [10]. Since previous research showed that more than 40 percent of faults in induction motors are related to bearing faults [11], a number of research works have been done with bearing fault detection using wavelet technique for vibration data [12] to [13]. Gear faults are also common in the induction motor. In [14], the authors used an adaptive method for vibration to detect the gear tooth faults. Neural networks and classification techniques also widely used in fault diagnosis [15] to [17].

In this paper, a new approach for fault detection and diagnosis system for induction motors in which the fault detection process is based on vibration signals and the fault diagnosis process is based on the pattern classification technique [18], is proposed. Since the vibration of the induction motor is the root cause of motor faults [19], the vibration signal can be analyzed to indicate the state of the motor. The characteristics of vibration signals from the motor in a normal condition are different from those of in a faulty condition. The most common fault detection techniques based on vibration signal focus on the vibration signal’s frequencies and magnitude, which are dealing with one dimension domain [7] to [16], [20] to [21]. In this paper, however, other features of the vibration signals that are local features by translating the vibration signal into an image (two dimensions), are explored. The local features from the image are extracted using the SIFT algorithm. The SIFT features, then, are used for the pattern classification process.

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The rest of the paper is organized as follows: In Section 1, a short explanation of SIFT algorithm is briefly introduced. Section 2 covers the vibration signal to image translation, feature extraction, framework for pattern classification, and discussion of SIFT feature advantages. The experimental setup and vibration signal database for the fault detection and diagnosis systems are discussed in Section 3. Results and discussions are provided in Section 4, and conclusions in Section 5.

1 SCALE INVARIANT FEATURE TRANSFORM (SIFT) The SIFT algorithm consists of four main filtering stages, which are: scale-space extreme detection, keypoint localization, orientation assignment and keypoint descriptor. In this section, the algorithm containing these four steps is briefly introduced. The details of the SIFT algorithm can be found in [22].

1.1 Scale-Space Extrema Detection The scale space called L(x, y, σ) is defined by the following function: L(x, y, σ) = G(x, y, σ) * I(x, y) ,

(1)

where, ‘*’ notation is the convolution operator, G(x, y, σ) is a variable-scale Gaussian kernel and I(x, y) is the intensity of the pixel, which its coordinates are x and y. The SIFT is one such technique which locates scale-space extrema from Gaussian image differences called D(x, y, σ) given by: D(x, y, σ) = L(x, y, kσ) ‒ L(x, y, σ) ,

(2)

where k = 1, 2, 3, ... is used to present the different scale space. To detect the local maxima or minima of D(x, y, σ), each point is compared with its eight neighbors on the same scale, and its nine neighbors on the up and down scale. If this value is larger than all 26 neighbors it is a maxima, if it is smaller then it is minima.

1.2 Keypoint Localization

Do, V.T. – Chong, U.-P.

The location of extrema, z, is given by:

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z=

−∂ 2 D −1 ∂D . (3) ∂x 2 ∂x

If the function value at z is below a threshold value then this point is discarded. This removes extrema that has a low contrast. Edge extrema that have large principle curvatures but small curvatures in the perpendicular direction are eliminated. Using 2×2 Hessian matrix H computed at the location and scale of the keypoint, principle curvatures which are proportional to eigenvalue of H can be computed.  Dxx H =  Dyx

Dxy  .

Dyy 

(4)

The elimination criteria can be constructed as follows:

Tr ( H ) 2 (r + 1) 2 α < , r = , Det ( H ) r β

(5)

where σ is Eigenvalue with larger magnitude, and β is Eigenvalue with smaller magnitude. If this inequality is true, the keypoint is rejected.

This stage aims to assign a consistent orientation to the keypoints based on local image properties. The keypoint descriptor is represented relative to this orientation because it is invariant to rotational movements of the keypoints. The approach taken to find an orientation has five steps described below: Step 1: Use the keypoint scale to select the Gaussian smoothed image L. Compute gradient magnitude, m(x,y) and orientation, θ(x,y) by two following Eqs.: m ( x, y ) =

( L( x + 1, y ) − L( x − 1, y )) 2 + ( L( x, y + 1) − L( x, y − 1)) 2

θ ( x, y ) = tan −1 ( L( x, y + 1) −

1.4 Keypoint Descriptor The local gradient data, used above, is also used to create keypoint descriptors. The gradient information is rotated to line up with the orientation of the keypoint and then weighted by a Gaussian kernel with a variance of keypoint scale multiplied by 1.5. This data is then used to create a set of histograms over a window centered on the keypoint. Keypoint descriptors typically use a set of 16 histograms, which are aligned in a 4×4 grid, each with eight orientation bins, one for each of the main compass directions and one for each of the mid-points of these directions. These results in a feature vector contain 128 elements.

2 METHODOLOGY 2.1 Vibration Signal to Image Translation

1.3 Orientation Assignment

create a keypoint with that orientation. Some points will be assigned multiple orientations.

Step 5: Fit a parabola to the three histogram values closest to each peak to interpolate the peaks position.

,

(6)

L( x, y − 1) ). (7) L( x + 1, y ) − L( x − 1, y )

Step 2: Form an orientation histogram from gradient orientations of sample points.

Step 3: Locate the highest peak in the histogram.

Step 4: Use this peak and any other local peak within 80% of the height of this peak to

The proposed approach deals with extracting the features of a vibration signal in twodimension domain. By translating the vibration signal into an image, the local features are extracted using the SIFT algorithm. To translate the vibration signal into an image, the amplitude of each sample of vibration signal is first normalized ranging from 0 to 255, which is the significant pixel intensity range for a gray image.

The intuitive explanation for the translation is indicated in Fig. 2. In this figure, the vibration signal has M multiplying by N samples where the M×N term is the size of the image (M and N values are the row and column of the image, respectively). The M and N values are dependent on the length of the vibration signal. However, the computational complexity of the proposed approach will be directly proportional to those values. Therefore, if the complexity is the matter, M and N values should be chosen as small as possible but they should be large enough in order to retain the most significant features from the origin. A recommendation for M and N values is: M = 128; 256 or 512 and N = 128; 256 or 512.

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The normalized amplitude of each sample point of the vibration signal becomes the intensity of the corresponding pixel in the image. The mapping from normalized amplitude to an equivalent pixel is clearly presented in Fig. 2. The coordinate of the corresponding pixel for the ith sample in the vibration signal is pixel ( j, k) where j = floor (i/N) and k = modulo (i/N).

An example of a vibration signal from a normal motor translated to a gray image is given in Fig. 3. In this example, the vibration signal has 16384 samples (equivalent to 2.048 s) with the sampling rate of 8000 Hz. The translated image size is 128×128 (i.e. 16384 pixels).

2.2. Local Feature Extraction and Texton Dictionary The SIFT algorithm is used to extract a number of local features from a gray image. Each local feature is a 128-dimension vector, which contain information of location of vector, weight

and orientation of each dimension. The details of the SIFT algorithm to extract local feature can be found in [22].

In the proposed fault detection and diagnosis for induction, it has been assumed that all kinds of faulty vibration signals of the induction are known and available in the database (faulty categories). To determine the status of the motor, the testing vibration signal of the motor should be classified into the equivalent faulty category. For classifying, a texton dictionary for each faulty category is used. Each texton dictionary contains the most significant features for that faulty category that are distinctive among the faulty categories. The features from the testing signal will be compared with features in texton dictionaries for classifying based on a pattern classification framework, which will be detailed in Section 2.3.

In the fault detection process, there are two important steps that are feature extraction and texton dictionary creation based on the

Fig. 2. Vibration signal to image translation scheme

Fig. 3. An example of a vibration signal translated into the 128×128 gray image (M = 128 and N = 128) 658

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Fig. 4. The model for creating the faulty database

images translated from the vibration signals. As mentioned above, for the feature extraction, the SIFT algorithm is applied for images to generate the 128-dimension feature vectors. The feature vectors are considered as the features of interest.

For the texton dictionary creation, a number of images translated from vibration signals for each faulty category are used to create the texton dictionary for that category. The adaptive shift clustering algorithm, which is presented in [23] to [25] is utilized to generate C centroid feature vectors (referred to textons); therefore, the C centroid feature vectors are representative of the category. This collection of C centroid feature vectors is; therefore, called “texton dictionary” for the faulty category. The value C for each category is proportionate to the number of vector feature generated (e.g. each subset of 100 random feature vectors in one category is clustered into 10 textons). Hence, value C for each category is different. The texton dictionary creation procedure is illustrated in Fig. 4 (assuming that there are P faulty categories considered).

2.3 Pattern Classification Framework in Fault Diagnosis In the fault diagnosis process, the feature vectors extracted from the SIFT algorithm as the faulty symptoms are used. Using a pattern classification technique, the current vibration signal should be classified to the corresponding faulty category through a pattern classification framework. The pattern classification framework

utilizing the feature vectors and the texton dictionary is proposed as follows: Step 1: Translate the vibration signal to the image.

Step 2: Apply the SIFT algorithm for this image to generate the 128-dimension feature vectors.

Step 3: Calculate the Euclidean distance between each feature vector and each centroid feature vector in the texton dictionary for each faulty category. With two 128-dimension vectors V1 and V2 with the coordinates {V11 , V12 ,..., V1128 } and {V21 , V22 ,..., V2128 } , respectively, the Euclidean distance can be calculated as follows:

D12 =

128

∑ (V

i 1

− V2i ) 2 .

(8)

i =1

Step 4: Find the category containing the centroid feature vector that makes the distance minimal. The centroid feature vector is called the “match” vector.

Step 5: Apply for entire feature vectors of the vibration signal; therefore, the histogram for the “match” vector is created for each faulty category.

Step 6: Classify the vibration signal to the equivalent faulty category based on the created histogram. The vibration signal belongs to the category with the highest number of “match” vector.

The simulation of the detail algorithm is clearly presented in Fig. 5.

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Fig. 5. Framework for classification by model 2.4 Advantages of SIFT Features As mentioned in [22], the SIFT features have some strong characteristics comparing to other local features such as: invariant to image scaling and rotation, partially invariant to change in illumination and 3D camera viewpoint, and highly distinctive.

Machine operating in industrial plants work in noisy environments, and as a result, the useless noise added in the recorded signals is unpreventable. As such, it may be an obstacle for analyzing the vibration signals. However, when the vibration signals are translated into images, the added noise is considered as the illumination of the light to the image. Hence, the effect of noise to the signal is removed when using SIFT features. When using pattern classification technique to classify signals, the highly distinctive characteristic of SIFT feature provides an efficient classification with high accuracy. Those advantages of the SIFT algorithm motivated us exploiting its output features for induction motor fault detection using the classification technique.

data needed under full load conditions. One of the motors operates under normal as a benchmark for comparison with faulty motors. The others are faulty motors: bowed rotor, broken rotor bar, bearing outer race fault, rotor unbalance, adjustable eccentricity motor (misalignment), and phase unbalance as shown in Fig. 7. Therefore, there are eight kinds of vibration signal categories called “faulty categories”; they are: angular misalignment, bowed rotor shaft, broken rotor bar, faulty bearing (out race), rotor unbalance, normal motor, parallel misalignment and phase unbalance.

3 EXPERIMENT SETUP AND TRAINING AND TESTING DATABASE 3.1 Experiment Setup The experiment was setup under a selfdesigned test rig. The experiment consists of motor, pulleys, belt, shaft and fan with changeable blade pitch angle as shown in Fig. 6 (Yang et al, 2006). In the experiment, six 0.5 kW, 60 Hz, 4-pole induction motors are used to create the 660

Fig. 6. Experiment setup The conditions of faulty induction faulty motors are described in Table 1. The motor’s load can be changed by adjusting the blade pitch angle or the number of blades. An accelerometer was used to measure the vibration signals of vertical

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Fig. 7. Faults on induction motor Table 1. Description of faulty induction motor Fault condition Broken rotor bar Bowed rotor Faulty bearing Rotor unbalance Eccentricity Phase unbalance

Fault description Number of broken bar: 12 Maximum bowed shaft deflection: 0.075 mm A spalling on outer raceway Unbalance mass on the rotor (8.4 g) Parallel and angular misalignments Add resistance on one phase

direction. In this paper, the vibration signals are used to evaluate the proposed fault detection and diagnosis algorithm. The sampling frequency of the experiment is 8 Khz; therefore, the maximum frequency obtained is 4 Khz. The vibration signal contains low frequency components; hence, vibration signal with 8000 Hz sampling rate is fully reasonable to consider. Each vibration signal consists of 16384 data points.

3.2 Training and Testing Database For an evaluation of the proposed fault detection and diagnosis, two vibration signal databases are required: training and testing databases. The training database is used for creating the texton dictionary, while the testing database is for testing the efficiency of the proposed approach based on the accuracy of the classification. In this experiment, eight vibration signals were collected from each faulty category

Others Total number of 34 bar Air-gap: 0.25 mm #6203 Adjusting the bearing pedestal 8.4%

from which two random signals for training database and six the other for testing database.

So the testing database consists of 48 vibration signals with six signals for each category, while the training database has 16 signals with 2 signals for each category. An example of waveforms of eight vibration signals from eight faulty categories and their translated images is illustrated in Fig. 8.

As mentioned above, the training database to construct the texton dictionary is used to represent the main characteristics of each category by a collection of centroid feature vectors after using the clustering algorithm. A trained vibration signal from the testing database is analyzed to classify in the corresponding faulty category using our proposed pattern classification framework.

Each faulty category name is denoted as follows.

Am: Angular misalignment, Br: bowed rotor shaft, Brb: broken rotor bar, Fb: faulty bearing, Mu: rotor unbalance, No: normal motor, Pm: parallel misalignment and Pu: phase unbalance.

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Fig. 8. Waveforms of eight vibration signal samples from eight faulty categories and their equivalent translating images 4 RESULTS AND DISCUSSIONS In the experiment, a classification for faulty vibration signals from machines is implemented.

In our test, 10 trials were carried out. For each trial, two randomly selected signals are for training database and six others for testing, therefore, the testing database consisting of 48 vibration signals of eight faulty signal categories (Am, Br, Brb, Fb, Mu, No, Pm, and Pu) with six signals for each, are used for training. Each vibration signal in the testing database is used to extract the SIFT features. The classification framework proposed in section 2.3 is applied with the features and texton dictionary in order to classify the vibration signal to the corresponding faulty category.

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Fig. 9 depicts an example of a classification for a vibration signal in the testing database, which is supposed to classify in the broken rotor bar category. In this example, 610 feature vectors are created after using the SIFT algorithm. The figure indicates that in the 610 feature vectors, there are 30 “match” vectors for Am, 4 for Br, 207 for Brb, 20 for Fb, 68 for Mu, 123 for No, 33 for Pm, and 125 for Pu categories discovered in the texton dictionaries, respectively. The histogram indicates that the trained signal is classified in the broken rotor bar category because of the highest “match” vectors number.

Table 2 provides a local feature distribution of each signal for each faulty vibration signal classification when a representative of each category is trained. The percentage of local

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Fig. 9. Histogram for feature vectors of a vibration signal from the testing database; on the X axis, 1: Am, 2: Br, 3: Brb, 4: Fb, 5: Mu, 6: No, 7: Pm and 8: Pu faulty categories Table 2. Percentage of local feature distribution Signal (Belonged category) 1 (Am) 2 (Br) 3 (Brb) 4 (Fb) 5 (Mu) 6 (Nor) 7 (Pm) 8 (Pu)

Am 44.66 0.78 4.34 0.81 2.37 4.67 27.12 3.83

Br 2.16 91.47 0.16 4.40 0.00 0.81 1.84 0.12

Brb 6.47 0.00 37.21 2.77 10.53 13.69 3.80 15.21

feature distribution represents the probability of the number of features of training vibration signal falling into one faulty category. The highest percentage number indicates the corresponding faulty categories that the vibration signal belongs to (the highest number in a row in Table 2). In Table 2, eight arbitrary vibration signals from eight faulty categories from testing database are selected for training and they are all correctly recognized.

Table 3 shows the final classification results for 10 trials. The average of total classification accuracy after 10 trials reaches 97.9%, which is considerably high. There are four among ten trials even giving 100% of accuracy. In the worse trial (trial #7), the classification accuracy of

Faulty categories Fb Mu 4.75 8.20 6.98 0.00 2.79 13.80 75.37 8.32 4.01 42.73 7.24 18.04 6.75 4.05 1.97 20.79

Nor 5.39 0.00 14.11 3.26 14.54 34.46 3.19 11.61

Pm 16.40 0.77 5.89 2.60 3.41 4.51 50.80 1.74

Pu 11.97 0.00 21.70 2.45 22.40 16.59 2.45 44.71

93.7% is even sufficient enough to be accepted in fault detection and diagnosis with classification technique.

To compare with other approaches in our previous reseach, we used two approaches introduced in [26], which are one-dimension domain and are wavelet-variance based and wavelet-crosscorelation based approaches. For experimental data in this paper, the result is shown in Table 4. The higher classification accuracy of the proposed approach clearly shows a big advantange of exploying two-dimension domain features over that of one-dimension domain.

Another comparison of these three approaches was also caried out. The experimental data were taken from [26]. The data are in

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Table 3. Classification accuracy of proposed approach Trials #

Am(6)

Faulty category name (number of testing signals) Br(6) Brb(6) Fb(6) Mu(6) Nor(6) Pm(6)

Total classification accuracy [%]

Pu(6)

Number of sucessful testing signals classified to corresponding categories 1 2 3 4 5 6 7 8 9 10

6 6 6 6 6 6 6 5 6 6

6 6 6 6 5 6 6 6 6 6 6 6 6 6 4 6 6 6 6 6 6 6 6 6 5 6 6 6 6 6 6 5 6 6 4 6 6 6 6 5 6 6 6 6 6 6 6 6 6 5 Total average of classification accuracy for 10 trials

6 6 6 6 6 6 6 6 6 6

6 6 6 6 6 6 6 6 6 6

97.9 100 95.8 100 97.9 100 93.7 95.8 100 97.9 97.9

Table 4. Comparison of proposed approach and two approaches proposed in [26] in term of classification accuracy Proposed approach Average classification accuracy for 10 trials [%]

97.9

Approaches Wavelet-Variance based approach

Wavelet-Crosscorrelation based approach

89.3

78.6

Table 5. Comparison of proposed approach and two approaches proposed in [26] in term of classification accuracy using experiment data in [26] Proposed approach Average classification accuracy for 10 trials [%]

98.1

95.6

microphone format (sound data taken by microphone). In these data, five sound data types from five different faulty categories of induction motors are: fault bearing, loose bearing, unbalance bearing, misalignment bearing and normal bearing. The details of experiment setup and data can be found in [26]. Three approaches are applied with this data. Again, 10 trials of testing were applied. Then, final classification accuracy is shown in Table 5. The avarage classification accuracy from 10 trials for proposed approach gains 98.1% compared to 95.6 and 78.7% for two other ones, respectively. Once again, with this different experiment, the proposed approach provides a better performance.

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Approaches Wavelet-Variance based Wavelet-Crosscorrelation approach based approach 78.7 5 CONCLUSIONS The results of high classification accuracies achieved by the proposed approach clearly demonstrates the potential of exploiting the features from two-dimension domain data of vibration signal by considering the SIFT algorithm for fault detection and diagnosis system.

Different from the previous approaches for a fault detection and diagnosis using features of vibration signal in one-dimension domain, we propose a vibration signal to image translation technique and a classification framework that are applied in a novel fault detection and diagnosis system for induction motors. For a real application,

Do, V.T. â&#x20AC;&#x201C; Chong, U.-P.

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knowledge about the induction motor faults should be fulfilled; therefore, the most common faults needs to be examined and can be represented in the texton dictionary. The results indicate that some features of the testing signal belong to other faulty categories; however, most of features are classified to the corresponding ones so the testing signal is still successfully classified into the correct faulty category. The high classification accuracy in the final results indicates that our proposed fault detection and diagnosis using signal model-based exploiting vibration signal in two-dimension domain is guaranteed.

6 REFERENCES [1] Isermann, R. (1994). Integration of fault detection and diagnosis methods. IFAC Symposium on fault detection, Supervision and Safety for Technical Processes, Espoo, p.

587-609.

[2] Isermann, R. (1997). Supervision, fault detection and diagnosis methods - an introduction. Control Engineering Practice – CEP, vol. 5, no. 5, p. 639-652.

[3] Isermann, R. (2006). Fault diagnosis system: An introduction from fault detection to fault tolerance. Springer, Berlin.

[4] Kimmich, F., Schwarte, A., Isermann, R. (2005). Fault detection for modern diesel engines using signal-and process modelbased methods. Control Engineering Practice, vol. 13, p. 189-203.

[5] Kimmich, F., Schwarte, A., Isermann, R. (2001). Model based fault detection for diesel engines. Aachen Colloquium, Aachen.

[6] Isermann, R. (1984). Process fault detection on modeling and estimation methods – a survey. Automatica, p. 387-404.

[7] Virtic, M., Abersek, B., Zuperl, U. (2008).

Using of acoustic models in mechanical diagnostics. Strojniški vestnik - Journal of Mechanical Engineering, vol. 54, no. 12, p.

874-882 [8] Žumer, J., Biček, A., Boltežar, M. (2008).

Characterization of the vibrations and structural noise of a suction unit’s cover.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 54, no. 12, p. 883-891.

[9] Furlan, M., Rebec, R., Cernigoj A., Celic, D., Cermelj, P., Boltezar, M. (2006).

Vibro-acoustic modelling of an alternator.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 52, no. 2, p. 112-125.

[10] Boyle, C. (1994). Online current monitoring to detect misalignment and dynamic eccentricity in three-phase induction motor drives. 29th universities power engineering conference 1, p. 5-8.

[11] McInerny, S.A., Dai, Y. (2003). Basic vibration signal processing for bearing fault detection. IEEE Transaction on Education, vol. 46, no. 1, p. 149-156.

[12] Kahaei, M., Torbatian, M., Poshtan, J. (2007). Bearing-fault detection using the Meyer-wavelet-packets algorithm. Strojniški vestnik - Journal of Mechanical Engineering, vol. 53, no. 3, p. 186-192.

[13] Arenas, J.P. (2005). Enhancing the vibration signal from rolling contact bearing using an adaptive closed-loop feedback control for wavelet de-noising. Strojniški vestnik Journal of Mechanical Engineering, vol. 51, no. 4, p. 184-192.

[14] Belsak, A., Flasker, J. (2008). Vibration analysis to determine the condition of gear units. Strojniški vestnik - Journal of Mechanical Engineering, vol. 54, no. 1, p.

11-24.

[15] Yang, B.S., Han, T., Yin, Z.T. (2006). Fault diagnosis system of induction motor using feature extraction, feature selection and classification algorithm. JSME International Journal, vol. 49, no. 3, p. 734-741.

[16] Goncalves, V.D., Almeida, L.F., Mathias, M.H. (2010). Wear Particle Classifier System Based on an Artificial Neural Network.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 50, no. 4, p. 284-288.

[17] Patan, K. (2008). Artificial neural networks for the modeling and fault diagnosis of technical processes. Springer Verlag, Berlin.

[18] Earl, G., Richard, J., Steve, J. (1996).

Pattern recognition and image analysis.

Prentice Hall, New York.

[19] Micheal, N., Denis, K. (2003). Fundamentals of noise and vibration analysis for engineers.

Cambridge University Press, Cambridge.

Signal Model-Based Fault Detection and Diagnosis for Induction Motors Using Features of Vibration Signal in Two-Dimension Domain

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[20] Benbouzid, M.E.H. (2000). A review of induction motors signature analysis as a medium for faults detection. IEEE Transactions on Power Electronics, vol. 14, no. 5, p. 984-993.

[21] Ocak, H., Loparo, K.A. (2004). Estimation of the running speed and bearing defect frequencies of an induction motor from vibration data. Mechanical Systems and Signal Processing, vol. 18, no. 3, p. 515-533.

[22] Lowe, D. (2004). Distinctive image features from scale-invariant keypoints. International Journal of Computer Vision, vol. 60, no. 2, p. 91-110.

[23] Krystian, M., Cordelia, S. (2005). A performance evaluation of local descriptors.

IEEE Trans. Pattern Analysis and Machine Intelligence, vol. 27, no. 10, p. 1615-1629.

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[24] Leung, T., Malik, J. (2001). Representing and recognizing the visual appearance of materials using three-dimensional textons.

International Journal of Computer Vision, vol. 43, no. 1, p. 29-44.

[25] Georgescu, B., Shimshoni. I., Meer, P. (2003). Mean shift based clustering in high dimensions: a texture classification example.

9th IEEE International Conference on Computer Vision 1, p. 456-463.

[26] Do V.T., Cho S.J., Chong U.P. (2009). Fault detection and diagnosis for induction motors using variance, crosscorrelation and wavelets.

Transactions of Korean Society for Noise and Vibration Engineering, vol. 19, no. 7, p. 726735.

Do, V.T. â&#x20AC;&#x201C; Chong, U.-P.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 667-673 DOI:10.5545/sv-jme.2010.244

Paper received: 03.12.2010 Paper accepted: 27.07.2011

FE-Modeling of Cold Rolling by In-Feed Method of Circular Grooves

Niţu, E. – Iordache, M. – Marincei, L. – Charpentier, I. – Le Coz, G. – Ferron, G. – Ungureanu, I. Eduard Niţu1,* – Monica Iordache1 – Luminiţa Marincei1 – Isabelle Charpentier2 – Gaël Le Coz3 – Gérard Ferron2– Ion Ungureanu1 1 University of Piteşti, Romania 2 LEM3, University Paul Verlaine-Metz, France 3 LEM3, ParisTech-Metz, France The methods of cold rolling of rods are widely used in manufacturing industries to obtain pieces with complex profiles. In this study, complex profiles with grooves have been formed by in-feed methods using two rolls. An experimental system was constructed to record the process parameters. The microhardness has been measured by the Vickers method in an axial section of the rolled piece. The process has also been simulated by means of finite element calculations using the Abaqus/Explicit code. The material behavior is described by using a 5-parameter strain-hardening law and by accounting for thermal effects at high strain-rates. Finally, a comparison is made between experimental and simulated results. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: groove profile, cold rolling, micro-hardness, finite element modeling

0 INTRODUCTION The advantages of cold rolling, including high productivity, substantial improvement in mechanical properties and low roughness [1] and [2], are clearly apparent in the case of profiled surfaces, such as: threads, grooves, teeth, parts that can be found in various products of the automotive industry, aeronautics, appliances, etc.

An important objective of the deformation processing of metals and alloys is the production of defect-free parts, with the desired microstructure and properties. This goal can be achieved by improving the design, calculation methods and control of process parameters.

In recent years, finite elements (FE) models have been widely used to analyze a number of metal-forming processes [3] to [5].

The accumulated knowledge enabled the forming industry to improve product performance, service life and process competitiveness [6].

The FE modeling of cold rolling uses numerical models of the elements involved in the working process (blank material and tools), with the aim of computing the evolution of different quantities during the process: stresses and strains, material flow paths, and the final profile of the product. The FE modeling of the cold rolling process started in 1990 [1] and [7], but the *Corr. Author’s Address: University of Piteşti, Str. Târgul din Vale, nr. 1, Piteşti, Romania, eduard.nitu@upit.ro

high volume of calculations and the computers incapacity to simulate the process within a reasonable time, restricted these studies to the understanding of the deformation process [6] to [8], by analyzing the state of stresses and strains of circular profiles at different levels of deformation.

The research intensified in 2000, together with the development of efficient FE software and with the growing computational capacity of computers [8] to [10]. The main elements of interest in these research works are the material of the work-piece, piece profile and rolling process, with particular attention to the plastic behavior of the material, the meshing elements and the software used for simulations (MARC, ABAQUS, DEFORM, MSC Super Form ...).

The strain-hardening laws most frequently used for the analysis and simulation of large plastic deformations at room temperature are Hollomon, Ludwik, Ludwik-Hartley and Voce [11]. However, cold rolling processes are affected by the effects of high-speed processing and associated temperature rise, because heat generated by plastic deformation does not have enough time to be evacuated by convection through the surface and by conduction to the connecting parts. These strain-rate and temperature effects are often described using the Johnson-Cook’s law [12].

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In this study, a complex profile with five grooves has been formed by in-feed method using two rolls. This paper focuses on the development of three-dimensional FE models using the stressstrain law characterized in compression tests and an optimal mesh of the work-piece in order to obtain accurate results with a reasonable number of elements and an acceptable computation time.

The validation results are based on forces and micro-hardness measurements. The measured force is the radial force. The experimental microhardness is measured on the axial section of the tooth by Vickers method and the process is simulated using the Abaqus/Explicit FE code.

The piece was cut by electric discharge machining and then finely polished for the measurement of the micro-hardness in the axial section of the tooth. The micro-hardness indentations were performed on several lines along the axial and radial directions, Fig. 2. For each direction the distance between indentations was 0.125 mm and the minimum distance from the surface of the piece was 0.1 mm.

1 EXPERIMENTAL PROCEDURE Threads are formed by the progressive penetration of a set of parallel wedge-shaped indentors into the blank surface during a fixed number of blank revolutions. The predominant loading modes are plane-strain compression and shear in the external part of the work-piece. The profile generated by radial cold rolling using two rolls and in-feed method was a concentric channels surface (five grooves similar in axial section to metric thread M20×2, Fig. 1).

Fig. 2. Schema of micro-hardness indentations An experimental system, Fig. 3, was used to record the process parameters: in-feed of the rolls, force along the radial direction and tools rotation.

2 STRESS-STRAIN BEHAVIOR

Fig. 1. Form of the cold rolling profile The material used in this investigation was AISI 1015 steel. Its chemical composition and initial micro-hardness are given in Table 1. The blank was obtained from a hot extruded bar by turning and grinding.

Micro-hardness measurements were made

by Vickers method, which allows the use small loads and a comparison of the results with other mechanical quantities. The load was taken equal to 300 g, in order to take account of estimated micro-hardness and of grain size.

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The compression test was adopted to characterize the stress-strain behavior of the material, since (1) it allows us to reach high strain levels (up to an effective strain ε ≈ 0.9 in our tests) and (2) the stress state generated during the cold rolling process is mostly compressive.

Considering the high speed of the rolling process, it is interesting to characterize material behavior at high strain-rates. The compression tests were performed at the speeds of 1.8 mm/ min (low speed test, LST) and 180 mm/min (high speed test, HST). They correspond to nominal strain-rates of 10-3 and 10-1 s-1, respectively.

The stress-strain curve for the HST is at first slightly higher, and then it progressively becomes lower than in the LST. This can be interpreted by considering that (1) the effect of positive strain-rate sensitivity predominates at small strains and (2) the temperature rise

Niţu, E. – Iordache, M. – Marincei, L. – Charpentier, I. – Le Coz, G. – Ferron, G. – Ungureanu, I.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 667-673

Table 1. Chemical composition and initial micro-hardness of the steel AISI 1015 C 0.15

Mn 0.65

Chemical composition [wt %] Cr Si Ni 0.11 0.27 0.08

Mo 0.01

P, S < 0.035

Micro-hardness (average) Vickers 300 g 1360 MPa

Fig. 3. Scheme of the experimental system for the radial cold-rolling becomes significantly higher in the HST, which is contributed to a decrease in the flow stress.

The LST is assumed to be isothermal, while strain-rate and temperature sensitivity effects should be taken into account for the HST. Strain-hardening laws involving only three parameters are not able to give a good account of the stress-strain curve in the LST over the whole range up to ε ≈ 0.9 . A five-parameter law which combines Hollomon and Voce’s laws was chosen and the hardening law in the form was expressed:

σ LS =  K ε n + S (1 − A exp(− Bε ))  , (1)

where K, n, S, A and B are material parameters.

The stress-strain curve in the HST is described by introducing strain-rate and temperature sensitivity terms in agreement with the Johnson-Cook’s law, i.e.: σ HS = σ LS

  . ε  ⋅ 1 + C ln  .    ε0

m     ⋅ 1 −  T − T0   , (2)     Tm − T0        

where C and m are strain-rate and temperature sensitivity coefficients, respectively, ε0 = 10-3 s-1 is the reference strain-rate of the LST, ε = 10-1 s-1 is the strain-rate of the HST, T0 = 300 K is the reference (room) temperature, Tm= 1810 K is the melting temperature and T is the current temperature.

Under adiabatic conditions, the increase in temperature corresponds to a fraction β of the

plastic work that converts into heat. Under these conditions, the rate of temperature rise T is obtained with the Eq.:

βσε = c ρT , (3)

where β is the Taylor-Quinney coefficient, generally taken equal to 0.9; C = 0.460 J/gK is the specific heat capacity and ρ = 7.8 × 106 g/m3 is the specific mass.

The identification of the five strainhardening parameters, Eq. (1), is first performed by fitting the LST data with a gradient method implemented in a FORTRAN program. The following is obtained: K = 542.5 MPa; n = 0.135; S = 217.6 MPa; A = 0.99 and B = 9.91. Then, the HST data are analyzed by considering that the actual temperature rise (T – T0) is a fraction η of the value calculated under adiabatic conditions with Eq. (3). Again, the FORTRAN program is used to determine the remaining parameters. The following has been found: η = 0.7; C = 0.0125 and m = 0.78.

The experimental stress-strain curves in compression for the LST and HST are shown in Fig. 4, together with the fit obtained with Eqs. (1) and (2), respectively. The fitted curves are extrapolated up to the value ε ≈ 4 , which corresponds to the strain levels attained in the cold rolling process.

As a result of the assumed continuous increase in temperature, the extrapolated HST curve exhibits a marked decrease at high strains.

FE-Modeling of Cold Rolling by In-Feed Method of Circular Grooves

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a)

Fig. 4. Experimental and extrapolated real stress-strain curves in compression

b)

3 METALLOGRAPHIC CHARACTERIZATION A metallographic examination was performed on a sample piece from the bar stock (undeformed material, Fig. 5a) and on a cold rolled piece, Fig. 5b. The pieces were cut along the axial direction by electric discharge machining, finely polished and 2% Nital etched. The microharness in the centre of cut specimens was found to be very close to that measured on the surface of the undeformed material (mean value 1360 MPa).

Micrographs show that deformation is concentrated in the superficial layers, especially on the root and on the flanks of the tooth, with somewhat less stretching in the tooth interior (Fig. 5b).

4 NUMERICAL PROCEDURE The numerical calculations were performed with the dynamic explicit FE code Abaqus/ Explicit.

The elastic behavior of the work-piece is modeled by assuming isotropic elasticity, with the values of Young’s modulus, E = 200,000 MPa and Poisson’s ratio, ν = 0.3. Strain-hardening is described using the von Mises criterion with the assumption of isotropic hardening. Accordingly, the yield function is given by: f =

3 sij sij − σ , 2

(4)

where sij are the deviatoric stress components,

( 3 / 2 ) sij sij

is the von Mises equivalent stress

and σ is the current yield stress.

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Fig. 5. Microstructure; a) of the undeformed material, and b) of the cold-rolled profile The laws obtained in the LST and HST, section 2, are used as two possibilities to describe the effective stress-strain law in numerical simulations, assuming that their extrapolation to the entire strain range that develops during the rolling process is valid.

Based on the microstructural observations, section 3, a dense mesh must be used in the deformation zone near the blank surface and a much coarser mesh can be applied in the blank interior. The best compromise that could be found between the number of elements and accuracy of results is presented in Fig. 6: • along the axial direction, three areas are defined, Fig. 6a: • area A, with very small deformations, where the size of elements can be very large; • area B1, corresponding to the root of the profile with the largest deformations is very finely meshed; • area B2, corresponding to the flank of the profile with large deformations is finely meshed; • area C, corresponding to the crest of the profile, has moderate deformations and an average size of elements.

Niţu, E. – Iordache, M. – Marincei, L. – Charpentier, I. – Le Coz, G. – Ferron, G. – Ungureanu, I.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 667-673

Table 2. Dimensions of the finite elements Area A 2 × 0.4 p × p/5

Dimension of the element (length × width), [mm] Area B Area C B1 B2 0.4 × 0.02 0.4 × 0.04 0.4 × 0.08 p/5 × p/100 p/5 × p/50 p/5 × p/25

a)

Area D 0.4 × 0.2 p/5 × p/10

b)

Fig. 6. The optimized mesh of the blank; a) along the axial direction, b) in a cross-section •

along the radial direction, two areas are defined, Fig. 6b: • area D, associated to the superficially deformed layer, where the size of elements has to be small; • area E, corresponding to the core, has small deformations and the size of elements can be very large.

In areas A, B, C and D we used C3D8R, 8-nodes solid hexahedral elements with reduced integration, while area E was meshed with C3D4R, 4-nodes solid tetrahedral elements with reduced integration. The dimensions of the elements in the five above-defined areas were established in relation with the main characteristic of the profile, step p between adjacent grooves. The dimensions of elements in the different areas are indicated in Table 2.

The tools are modeled by analytical rigid surfaces. The strain-rate dependency is not handled in the simulations.

5 RESULTS AND DISCUSSION A comparison between experiments and calculations has been made by considering the evolution of radial force during cold rolling and

the distributions of micro-hardness HV in an axial section of the piece.

The experimental and simulated evolutions of the radial force are presented in Fig. 7.

Fig. 7. Experimental and simulated radial forces The two simulated radial forces have very similar evolutions and the value of the maximum force is very close to the experimental one.

However, the evolution of the simulated radial force is somewhat different from the experimental evolution: the increase in radial force is more rapid, and the maximum is predicted earlier in the simulations.

The similarity between radial forces calculated with the LST and HST stress-strain curves is fairly surprising. In fact, these two curves intercept at ε ≈ 0.3 , and the difference between

FE-Modeling of Cold Rolling by In-Feed Method of Circular Grooves

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the two curves is less than 2% up to ε ≈ 0.7 (Fig. 4). Actually, only 10% of the volume of the work piece is deformed up to strain values larger than ε ≈ 0.7 at the end of the rolling process. In other words, only a thin superficial layer is concerned by very high strains, and most of the internal energy expended during rolling corresponds to regions where the strain levels remain moderate.

As a result, the evolution of the force is mainly controlled by the stress-strain law at low and moderate strains. This analysis is corroborated by similar experiments and calculations performed on other materials, where the radial forces are observed to be dependent on the stress level obtained in compression, but independent of the extrapolation adopted in the simulations.

Micro-hardness measurements were also compared with yield stress values obtained by numerical simulations. For blank materials, the Vickers micro-hardness HV is found to be proportional to the initial yield stress σy, i.e. [13]:

HV = α × σy ,

(5)

where the proportionality factor α is close to 3.

For plastically deformed materials, σy in Eq. (5)

should be replaced by the current yield stress σ in order to take account of strain-hardening.

The distributions of micro-hardness HV measured in an axial section of the piece are presented in Fig. 8a. The equivalent strains (PEEQ) obtained in numerical simulations are presented on Figs. 8b and 8c with the hardening laws obtained in the LST and HST, respectively.

The values of micro-hardness HV are also indicated in Fig. 8b, using the calculated σ - values and α = 3 in Eq. (5). Indeed, this conversion would be unrealistic using the stress-values obtained in the HST since these stresses correspond to the high temperatures attained during the rolling process, while micro-hardness measurements are made at room temperature.

In spite of a fairly large scatter in microhardness measurements, the comparison tends to shown a very good correlation, with the higher levels of HV and ε at the root of the profile, and lower values in the central part of the tooth.

The proportionality factor α between HV and σ is very close to 3 when σ is estimated at room temperature with the isothermal stress-strain

b)

a)

c)

Fig. 8. Experimental micro-hardness in [MPa]; a) and equivalent strain distribution for the two numerical models, using the stress-strain laws; b) σLST , and c) σHST 672

Niţu, E. – Iordache, M. – Marincei, L. – Charpentier, I. – Le Coz, G. – Ferron, G. – Ungureanu, I.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 667-673

law σLST. This α-value is in agreement with the usually reported in the literature.

6 CONCLUSIONS FE simulations of the cold rolling process have been performed using an optimized FE mesh and an extrapolation of the stress-strain law identified in compression tests. The results demonstrate that valuable information can be obtained from FE simulations: • the calculated radial force is in very good agreement with its experimental value, • the calculated distribution of yield stress σ in the superficial layers of the piece is in good correspondence with measurements of Vickers micro-hardness HV, considering the usually-accepted value of the proportionality factor (α = 3) between these two quantities.

7 ACKNOWLEDGEMENTS This work was supported by CNCSIS – UEFISCDI, project number PN II - IDEI 711/2008, ANCS project number PN II – CAPACITATI - Bilateral project “Brâncuşi” 211/2009 and the project “Supporting young PhD students with frequency by providing doctoral fellowships”, co-financed from the EUROPEAN SOCIAL FUND through the Sectoral Operational Program Development of Human Resources.

8 REFERENCES [1] Neagu, C., Vlase, A., Marinescu, N.I. (1994).

Volume cold pressing of the parts with thread and teeth. Technical Publishing House, Bucharest. (In Romanian) [2] Ungureanu, I., Iacomi, D., Nitu, E.L. (2004).

Processing technologies of the profiles by cold forming, Design guide. University of Piteşti Publishing House, Piteşti. (In Romanian) [3] Leon, J., Luis, C.J., Luri, R., Reyero, J. (2007). Determination of the neutral point in flat rolling processes. Strojniški vestnik Journal of Mechanical Engineering, vol. 53, no. 11, p. 747-754.

[4] Kamouneh, A.A., Ni, J., Stephenson. D., Vriesen. R. (2007). Investigation of work hardening of flat-rolled helical-involute

gears through grain-flow analysis, FEmodelling, and strain signature. International Journal of Machine Tools & Manufacture, vol. 47, p. 1285-1291.

[5] Gantar, G., Sterzing, A. (2008). Robust design of forming processes. Strojniški vestnik Journal of Mechanical Engineering, vol. 53, no. 4, p. 249-257.

[6] Domblesky, J.P., Feng, F. (2002). Twodimensional and three-dimensional finite element models of external thread rolling.

Proceedings of the Institution of Mechanical Engineers, Part B: Journal of Engineering Manufacture, vol. 216, no. 4, p. 507-517.

[7] Martin, J.A. (1998). Fundamental finite element evaluation of a three dimensional rolled thread form: modeling and experimental results. Fatigue, Fracture and Residual Stresses ASME, vol. 373, p. 457467.

[8] Domblesky, J.P., Feng, F. (2002). A parametric study of process parameters in external thread rolling. Journal of Materials Processing Technology, vol. 121, p. 341349.

[9] Kamouneh, A.A., Ni, J., Stephenson. D., Vriesen, R., DeGrace, G. (2007). Diagnosis of involumetric issues in flat rolling of external helical gears through the use of finite-elements models. International Journal of Machine Tools & Manufacture, vol. 47, p. 1257-1262.

[10] McCormack, C., Monaghan, L. (2001).

A finite element analysis of cold forging dies using 2D and 3D models. Journal of Materials Processing Technology, vol. 118, p. 286-292.

[11] Hartley, C.S., Srinivasan, R. (1983).

Constitutive equations for large plastic deformation of metals. Journal Enginering Materials Technology, vol. 105, p. 162-167.

[12] Johnson, G.R., Cook, W.H. (1983). A constitutive model and data for metals subjected to large strains, high strain rates and high temperature. Proceedings of the 7th International Symposium on Ballistics, La Haye, p. 541-547.

[13] Tabor, D. (1951). The hardness and strength of materials. Journal of the Institute of metals, vol. 79, p. 1-18.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 674-680 DOI:10.5545/sv-jme.2010.140

Paper received: 22.06.2010 Paper accepted: 03.08.2011

Surface Topography Modelling for Reduced Friction Sedlaček, M. – Vilhena, L.M.S. – Podgornik, B. – Vižintin, J. Marko Sedlaček* – Luis Miguel Silva Vilhena – Bojan Podgornik – Jože Vižintin University of Ljubljana, Faculty of Mechanical Engineering, Slovenia

The aim of the present research was to investigate the possibility to design contact surfaces with reduced friction using surface roughness and topography analysis. For this purpose, different 100Cr6 plate samples with different surface topography were prepared. Using different grades and combinations of grinding and polishing samples with similar Ra values, but different Rku and Rsk values were prepared.

To evaluate influence of roughness parameters on friction and wear, dry and lubricated pin-on-disc tests were carried out under different contact conditions. Test results show that surfaces with high Rku and negative Rsk values results in reduced friction.

To investigate the effect of surface topography on surface roughness parameters and consequently on friction, real roughness profiles were virtually altered to achieve virtually textured surfaces. Using NIST SMATS softgauge for calculation of surface roughness parameters, virtually altered roughness profiles were investigated in terms of texture size, shape and spacing, and their influence on surface roughness parameters, especially on skewness and kurtosis. Lower diameter, higher spacing and wedge-shaped dimples were found to reflect in higher Rku and more negative Rsk parameters, which should lead to lower friction. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: surface topography, roughness parameters, friction, surface texturing 0 INTRODUCTION Controlling the friction is becoming increasingly significant due to constant demands for improved reliability and effectiveness of mechanical parts and especially the reduction in frictional loses. In recent years surface texturing was introduced as a way of reducing friction. With employment of different patterns in the form of micro dimples or groves on the surface, reduction of friction can be obtained [1] to [13]. There are some texturing parameters like the shape of the dimples, their depth and width, their area density and orientation, which exert an influence on friction and wear. However, a lot of research work has been done in the field of surface texturing, while modification of surface topography by texturing is still mainly based on trial and error approach.

A possible way of designing surface texturing parameters, which would result in contact surfaces with lower friction, is by treating surface texturing as a controlled roughness. By knowing what kind of surface topography in terms of roughness parameters results in lower friction, we would be able to select proper surface texturing parameters. However, for this knowledge about 674

the correlation between surface roughness and friction is essential.

Surface roughness and topography, which are used to characterize contact surfaces are described with surface roughness parameters.

Unfortunately, standard surface roughness parameters normally used by designers do not describe contact surfaces sufficiently, with completely different surfaces showing similar or even the same values of standard roughness parameters and the other way round – similar surfaces having much different standard roughness parameters. In addition, different standards use different parameters. In practice most commonly used parameters for surface roughness description are Ra, Rq, and Rmax.

Average surface roughness (Ra) gives a very good overall description of height variations, but does not give any information on the wavelength and it is not sensitive to small changes in profile. Root mean square deviation of the assessed profile (Rq) is more sensitive to deviations from the main line than Ra. However, there are also other roughness parameters defined

by ISO 4287 standard, which give better surface description. Rsk is defined as skewness and is sensitive on occasional deep valleys or high

*Corr. Author’s Address: University of Ljubljana, Faculty of Mechanical Engineering, Aškerčeva 6, 1000 Ljubljana, Slovenia, marko.sedlacek@ctd.uni-lj.si

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peaks. Zero skewness reflects in symmetrical height distribution, while positive and negative skewness describe surfaces with high peaks or filled valleys, and with deep scratches or a lack of peaks, respectively. On the other hand, kurtosis (Rku) describes the probability density sharpness of the profile. For surfaces with low peaks and low valleys, Rku is less than 3, and more than 3 for surfaces with high peaks and low valleys [14]. The load bearing ratio, as well as the maximum contact pressure increase with the increase of skewness and kurtosis [15]. It was also observed that parameters Rk, Rpk and Rvk tend to have influence on coefficient of friction [16]. Parameters Rk, Rpk and Rvk are based on the bearing ration curve (Abbott-Firestone Curve) and are defined by the standard ISO13565-2. Core peak-to-valley height (Rk) give a numerical summary of the information contained in the bearing ratio curve based on the division of the curve into three regions. The upper region is defined as reduced peak height (Rpk), the middle region as core region (Rk) and the lower region as reduced valley depth (Rvk). When two surfaces rub together, the peak region usually gets worn out; the core region bears the load and has influence on the life of the product and valley region which acts as a lubricant reservoir.

Therefore, the aim of the present research was to investigate surface topography of contact surfaces in terms of different surface roughness parameters and to correlate surface topography changes to friction. Furthermore, to investigate the possibility of using roughness parameters as design parameters for surface topography modification, real roughness profiles were virtually altered to achieve virtually textured surfaces and roughness parameters calculated using NIST Surface Metrology Algorithm Testing System (SMATS) softgauge [17]. Virtually altered roughness profiles were investigated in terms of influence of texture size and shape on skewness and kurtosis parameters.

1 EXPERIMENTAL For the purpose of this investigation, 100Cr6 (AISI 52100) steel samples were used.

Disc type steel samples with a diameter of 24 mm were prepared in terms of different average surface roughness. By using different grades of

grinding and polishing results with Ra values from 0.02 to 0.49 μm, and different Rku and Rsk values were obtained. With the use of different grain size and grinding parameters it is possible to obtain different surface roughness. Increasing grain size and depth of cut increased the grinding forces and surface roughness values [18].

All samples were first grounded under water for 10 min, using grinding paper Piano 120, force 250 N, and rotation speed of 300 rpm.

Samples with average roughness Ra of about 0.5 mm were further grounded in water for 7 s with SiC paper and grain number 80, force 50 N and rotation speed of 150 rpm. On the other hand, smoothest surfaces with a surface roughness of about 0.02 mm were obtained with through two step polishing. They were first polished for 10 min with 9 mm DP-plan disc, using a force of 250 N and 150 rpm, and then subsequently polished for 10 min with 6 mm DP-plan disc, at 200 N force and 150 rpm. To prepare surfaces with similar Ra values but different Rku and Rsk values, previously prepared samples with Ravalues from 0.02 to 0.45 μm were used and treated afterwards with sandpapers or polishing pads to achieve desired roughness. Samples with higher Ra values were polished in order to remove high peaks on the surface, while polished samples were rubbed against sandpapers with different grain sizes to achieve groves on the smooth surface.

Surface topography of all samples was randomly orientated.

Average surface roughness (Ra), root square (Rq), skewness (Rsk), kurtosis (Rku), core peak-to-valley height (Rk), reduced peak height (Rpk) and reduced valley depth (Rvk) of prepared samples, measured on three different positions on the disc, with each incorporating 25 individual profiles, are shown in Table 1. The first four surface roughness parameters are according to ISO 4287 standard, while last three parameters are according to ISO 13565-2 standard. Samples tested under dry conditions are denoted with D and those under lubricated conditions as L. To investigate the influence of different roughness parameters on friction, regarding similar Ra values but different Rsk and Rku values, we can mutually compare samples D1 and D2 with Ra in the range of ~0.11 μm, samples D3 to D5 (Ra ~0.16 μm), D6 and D7 (Ra ~0.47 μm), L1 and L2 (Ra ~0.02

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μm), L3 and L4 (Ra ~0.13 μm), L5 and L6 (Ra ~0.16 μm), L7 and L8 (Ra ~0.44 μm) (Table 1), Tribological testing under dry and lubricated sliding conditions was carried out on Pin on disc tribometer using a ball on disc contact. In order to concentrate all the wear and surface topography changes on the investigated steel discs polished Al2O3 ball (ϕ10 mm) was used as a counterpart. Tribological testing was then performed at 3 different sliding speeds (0.05, 0.1 and 0.2 m/s) and a normal load of 1 N, which corresponds to a nominal contact pressure of 0.56 GPa. All tests (dry and lubricated) were made at room temperature (23±2 °C) and the relative humidity of 50 ± 10%, with lubricated ones carried out under boundary lubrication conditions using pure PAO 8 oil (ν40 = 46 mm2/s). During testing, coefficient of friction was monitored as a function of time and wear of contact surface measured after the test by means of topography analysis.

Table 1. Values of surface roughness parameters for different samples

D1 D2 D3 D4 D5 D6 D7 L1 L2 L3 L4 L5 L6 L7 L8

Ra

0.11 0.12 0.15 0.16 0.16 0.49 0.45 0.02 0.04 0.12 0.13 0.16 0.16 0.44 0.45

Rq

0.20 0.16 0.23 0.23 0.21 0.74 0.59 0.02 0.05 0.17 0.19 0.21 0.26 0.6 0.59

Rsk

-1.74 -1.44 -2.51 -0.86 -0.38 -1.33 -0.90 0.25 -1.37 -0.27 -1.67 -0.38 -3.11 -1.32 -0.90

Rku

24.42 12.55 16.02 6.977 4.240 10.80 5.341 3.075 16.63 9.310 12.21 4.240 23.20 10.29 5.341

Rk

0.266 0.345 0.348 0.436 0.488 1.360 1.334 0.071 0.132 0.352 0.304 0.488 0.342 1.002 1.334

Rpk

0.18 0.14 0.13 0.17 0.18 0.81 0.40 0.02 0.03 0.07 0.18 0.17 0.17 0.65 0.40

Rvk

0.22 0.25 0.47 0.36 0.27 1.44 0.88 0.01 0.05 0.26 0.32 0.27 0.53 1.41 0.88

2 RESULTS AND DISCUSSION 2.1 Dry Sliding For all tests made under dry sliding conditions, a combination of plastic deformation and abrasion was found as the main wear mechanism (Fig. 1a), resulting in a complete change in surface topography. Consequently, wear volume and sliding distance, when steady state 676

friction conditions are reached displayed a very high degree of scattering, as shown in Table 2.

a) b) Fig. 1. Topography change after; a) dry, b) lubricated sliding test However, when comparing samples D1 and D2, with sample D1 having much higher Rku value at similar average roughness, the coefficient of friction tends to be lower for sample D1. A comparison of samples D3, D4 and D5 which have similar Ra and Rq values, but different Rsk and Rku, shows that surface with the lowest skewness and the highest kurtosis (D3) also results in the lowest friction. Furthermore, for sample D3 friction is almost independent of the sliding speed, while for samples D4 and D5 friction is reduced with increase in sliding speed. Also the sliding distance to steady-state conditions is the shortest for sample D3, which is true for all sliding speeds.

Table 2. Coefficient of friction and sliding distance to steady state conditions for dry sliding Coefficient of friction [m/s]

Sliding distance to steady state conditions [m/s]

0.05 0.89

0.1 0.92

0.2 0.91

0.05

0.1

0.2

D1

34

50

38

D2

0.96

0.93

0.93

49

28

38

D3

0.89

0.90

0.89

31

29

16

D4

0.94

0.89

0.86

37

39

30

D5

0.99

0.97

0.92

39

30

33

D6

0.82

0.85

0.87

35

39

42

D7

0.84

0.63

0.66

40

24

34

When comparing samples D6 and D7, which have the highest Ra and Rq values between all D samples, lower friction and shortest sliding distance to steady state condition were obtained for sample D7, which displays smaller values of Rsk and Rku parameters. For sample D6 sliding speed has a very minor effect on friction, while for sample D7 increase in sliding speed results in reduced friction.

Sedlaček, M. – Vilhena, L.M.S. – Podgornik, B. – Vižintin, J.

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However, experimental results show that for dry sliding in general coefficient of friction is lower when roughness is low, and it gets reduced with increase in sliding speed. Furthermore, for surfaces with high value of Rku and more negative Rsk parameter, friction also tends to be lower.

Under dry sliding high degree of wear and change in surface topography limits a proper comparison and correlation between surface roughness and tribological properties of contact surfaces.

2.2 Lubricated Sliding Under lubricated sliding original topography of the contact surfaces was preserved during the test (Fig. 1b), with higher sliding speed generally resulting in lower friction. The coefficient of friction and sliding distance to steady state conditions for lubricated testing are summarized in Table 3.

Table 3. Coefficient of friction and sliding distance to steady state conditions for lubricated sliding

L1 L2 L3 L4 L5 L6 L7 L8

Coefficient of friction [m/s] 0.05 0.1 0.2 0.11 0.09 0.08 0.09 0.08 0.07 0.14 0.14 0.14 0.14 0.13 0.12 0.15 0.15 0.14 0.10 0.10 0.10 0.14 0.14 0.13 0.16 0.14 0.14

Sliding distance to steady state conditions [m/s] 0.05 0.1 0.2 8 15 21 17 16 8 19 16 13 17 9 5 28 20 16 23 11 6 31 23 6 22 18 9

In the case of smooth surfaces (samples L1 and L2), sample L2 gave lower friction although its average roughness (Ra) was higher than for L1.

However, while sample L1 has positive skewness and kurtosis around 3, L2 has negative skewness and very high kurtosis. Besides the lower friction sample L2 also shows a reduction in the sliding distance to steady-state conditions with sliding speed as compared to sample L1 where it increases as shown in Table 3. For samples with Ra values around 0.12 μm (L3 and L4) similar friction was observed although L4 displays higher Rku value and more negative Rsk than sample L3. On the other hand, the difference in Rsk and Rku roughness parameters has an influence on sliding distance when steady state conditions are reached. Sample

L4 with higher Rku and more negative Rsk shows drastically shorter sliding distances, as shown in Table 3 When comparing samples with Ra values of about 0.16 μm (L5 and L6), effect of parameter Rku and Rsk is even more pronounced. Higher Rku and more negative Rsk values measured for sample L6 result in lower friction. Furthermore, when comparing coefficient of friction for sample L6 and samples L3 and L4, it can be seen that sample L6 shows lower friction and shorter sliding distance to steady state conditions although displaying higher Ra values. In the case of the roughest samples (L7 and L8), with Ra values ~0.45 μm, similar tendencies were observed, with higher Rku and more negative Rsk values leading to lower friction. On the other hand, for very rough surfaces sliding distance to steady state condition tends to increase with an increase in Rsk and Rku parameters (L7). Interesting results were observed when comparing samples L3 and L7.

Although sample L7 shows much higher average surface roughness, it has similar Rsk and Rku value and consequently gives similar friction under lubricated sliding. However, as expected, sliding distance to steady state condition is shorter for the smoother surface (sample L3).

It can be also noticed that all samples which resulted in lower friction (L2, L4, L6 and L7) show lower values of Rvk regarding to the mutually compared samples (Table 1).

Taking a look at the nature of surface roughness parameters Rku and Rsk, it can be seen that negative Rsk describes surfaces with deep scratches or lackness of peaks, and parameter Rku, greater than 3, surfaces with high peaks and low valleys. By combining those two descriptions, we end up with smooth surface containing deep valleys. If we treat surface texturing as ordered roughness, those deep valleys can be treated as micro-dimples. With an implementation of micro dimples a ordered surface, which should reflect in lower friction, is obtained. The main idea of surface texturing is that textures act either as micro-traps for capturing wear debris or as micro-reservoirs which enhance lubrication.

When comparing experimental results with the idea of surface texturing some correlations can be observed. Under dry sliding higher values of parameter Rku and more negative values of

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parameter Rsk led to lower friction, indicating that deep valleys act as wear particle traps. The same tendency was observed under lubricated sliding, where greater Rku and more negative Rsk values give lower friction and shorten the sliding distance to steady-state condition. In this case it can be assumed that deep valleys act as micro reservoirs, which enhance lubrication. Accordingly, a general conclusion can be drawn, that is that Rku and Rsk have an influence on friction and could therefore be used as a guideline for designing surface topography with reduced friction.

3 MODELLING To investigate the effect of surface texturing on surface roughness parameters and to analyse the possibility of using roughness parameters for designing contact surfaces, the real roughness profile was virtually altered to achieve virtually textured surfaces. Using NIST SMATS [17] softgauge for the calculation of surface roughness parameters, virtually altered roughness profiles were investigated in terms of the effect of texture size and shape on surface roughness parameters, especially on skewness and kurtosis.

results which emphasized Rku and Rsk as the most influencing parameters, modelling was concentrated on finding texturing parameters, which would reflect in high Rku and more negative Rsk values.

First, the effect of spacing between dimples was investigated. Rectangular dimples with a diameter of 0.12 µm and depth of 6 µm. were simulated by shifting down the original profile for the value of the dimple depth. This surface topography modelling showed that increase in spacing result in decreased Ra, Rq and Rsk values and increased Rku value. For instance, if spacing is increased from 0.12 to 0.24 mm, Rsk will decrease and Rku increase from -0.668 to -1.026, and from 1.492 to 3.028, respectively. With a further increase in spacing between dimples, Ra and Rq values are further decreased and Rku further increased. However, for Rsk parameter an increase in spacing between dimples after initial reduction leads to increase in Rsk value, as shown in Fig. 3.

Fig. 3. Surface roughness parameters in dependence of spacing between dimples (diameter of dimples 0.12mm, depth 6μm) Fig. 2. Example of virtually textured surface (diameter of dimples 0.06 mm, spacing 0.24 mm, depth 6 μm, rectangular shape) The real roughness profile was obtained using stylus profilometer. The polished sample with Ra value of 0.02 µm (Rq = 0.024 µm, Rsk = 0.24, Rku = 3.07) was used as the origin profile.

This profile was then virtually altered using Microsoft Excel to achieve virtually textured surfaces as shown in Fig. 2 and the influence of different spacing, and dimple diameter, shape and depth on the surface roughness parameters investigated.

According to experimental 678

Varying the depth of dimples was found to have very minor effect on the Rsk and Rku parameter values. For the same diameter of 0.12 mm, changing the depth from 6 to 10 μm resulted in doubled Ra and Rq values, but almost equal values of Rsk and Rku parameter. The same observations were found for smaller dimple diameters and larger dimple depths. Changing the diameter of dimples only slightly affects Ra, Rq and Rku parameters, but has a considerable effect on Rsk value. A reduction in dimple diameter results in an only minor increase in Ra, Rq and Rku values but a drastic increase in Rsk, as shown in Fig. 4.

Sedlaček, M. – Vilhena, L.M.S. – Podgornik, B. – Vižintin, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 674-680

Fig. 4. Surface roughness parameters in dependence of diameter of the dimples (spacing 0.12 mm, depth 6 μm) a)

By combining different dimple parameters, such as bigger spacing between dimples (0.36 mm), smaller diameter (60 μm) and wedge-shaped dimple (α = 82.4° and β = 24.7°), for which it has been found to increase Rku and decrease Rsk values, very high values of Rku (12.3) and negative Rsk (-2.7) can be obtained. Additionally, such texturing parameters also result in lower Ra and Rq values (0.44 and 0.92 μm, respectively).

According to the experimental results, presented in this paper, such shape and spacing of the dimple should result in lower friction. However, these findings still need to be experimentally confirmed.

Table 4. Roughness parameters vs. dimples wall angle (diameter 0.12 mm, depth 6 μm)

Ra Rq α=β=0° 1.85 1.93 α=β=59° 1.18 1.49 α=β=75° 1.56 1.67 α=β=83.7° 1.07 1.31 α=85.4° β=14.6° 1.21 1.45 α=86.2° β=14.6° 1.07 1.33 α=83.7° β=0° 1.52 1.69 α=86.3° β=0° 1.29 1.56 α=86.9° β=0° 1.14 1.44 α and β defined in Fig. 5a

b)

Rku 1.35 2.25 1.42 2.96

Rsk 0.03 -0.04 -0.26 -1.05

2.55

-0.86

3.21

-1.05

0.33

-1.69

2.61

-0.77

3.30

-0.99

Fig. 5. Profile of the dimple; a) α = β = 59° b) α = 82.4°, β = 24.7°

5 CONCLUSIONS

Finally, the shape of the dimples was investigated, using dimples with a diameter of 0.12 mm and spacing of 0.12 mm. By changing the slope of the dimple walls (Fig. 5a), an increase in the slope angle was found to result in reduced values of Ra, Rq and Rsk parameters in increased values of Rku. By shifting the bottom of the dimple to one side (α ≠ β, Fig. 5b) higher values of Rku and more negative Rsk are obtained, as shown in Table 4. Such dimple shapes should also have good hydrodynamic effects. As can be seen from Table 4, the most suitable values of Rku and Rsk parameter are obtained when angle α is close to 90º and β is very small, but not 0°.

Under dry sliding high wear rates and change in surface topography blurs the effect of surface roughness on friction. However, in general lower roughness of the contact surface and higher sliding speeds result in lower friction, with surfaces displaying higher kurtosis (Rku) values and more negative skewness (Rsk) parameter often having an advantage.

For lubricated sliding contact kurtosis and skewness were found to be the most important roughness parameters in terms of tribological behaviour. The use of surfaces with higher kurtosis and more negative skewness always reflects in lower friction and shorter distance when steady-

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state condition has been reached, even if average roughness is not the same.

Plateau-like topography of contact surfaces with high kurtosis and more negative skewness and the effect of kurtosis and skewness on friction indicate that these parameters could be used as design parameters for surface topography modification and texturing aimed to reduce friction.

By surface topography modelling, surface texturing parameters which would result in greater kurtosis and more negative skewness parameter and consequently in lower friction in boundary lubrication can be defined. These parameters are lower diameter of the dimple, higher spacing between the dimples and wedge-like shape of the dimples.

REFERENCES [1] Wang, X., Kato, K., Adachi, K., Aizawa, K. (2001). The effect of laser texturing of SiC surface on the critical load for the transition of water lubrication mode from hydrodinamic to mixed. Tribology International, vol. 34, p.

709-711.

[2] Varenberg, M., Halperin, G., Etsion, I. (2002).

Different aspects of the role of wear debris in fretting wear. Wear, vol. 252, p. 902-910.

[3] Pettersson, U., Jacobson, S. (2003). Influence of surface texture on boundary lubricated sliding contacts. Tribology International, vol.

36, p. 857-864.

[4] Wakuda, M., Yamauchi, Y., Kanzaki, S., Yasuda, Y. (2003). Effect of surface texturing on fricition reduction between ceramic and steel materials under sliding contact. Wear, vol. 254, p. 356-363.

[5] Etsion, I., Halperin, G. (2003). A laser surface textured hydrostatic mechanical seal. Sealing Technology, p. 6-10.

[6] Etsion, I. (2004). Improved tribological performance of mechanical components by laser surface texturing. Tribological Letters, vol. 17, no. 4, p. 733-737.

[7] Pettersson, U., Jacobson, S. (2004). Friction and wear properties of micro textured DLC coated surfaces in boundary lubricated sliding. Tribological Letters, vol. 17, no. 3, p.

553-559.

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[8] Schreck, S., Zum Gahr, K.-H. (2005). Laserassisted structuring of ceramic and steal surface for improving tribological properties.

Applied Surface Science, vol. 247, p. 616622.

[9] Etsion, I. (2005). State of the art in laser surface texturing. Journal of Tribology, vol.

127, p. 248-253.

[10] Kovalchenko, A., Ajayi, O., Erdemir, A., Fenske, G., Etsion, I. (2005). The effect of laser surface texturing on the transitions in lubrication regimes during unidirectional sliding contact. Tribology International, vol.

38, p. 219-225.

[11] Pettersson, U., Jacobson, S. (2007). Textured surfaces for improved lubrication at high pressure and low sliding speed of roller/piston in hydraulic motors. Tribology International, vol. 40, p. 355-359.

[12] Galda, L., Pawlus, P., Sep, J. (2009). Dimples shape and distribution effect on characteristics of Stribeck curve. Tribology International, vol. l42, p. 1505-1512.

[13] Pawlus, P., Galda, L., Dzierwa, A., Koszela, W. (2009). Abrasive wear resistance of textured steel rings. Wear, vol. 267, p. 18731882.

[14] Gadelmawla, E.S, Koura, M.M, Maksoud, T.M.A., Elewa, I.M, Soliman, H.H. (2002).

Roughness parameters. Journal of Materials Processing Technology , vol. 123, p. 133-145.

[15] Wang, W., Chen, H., Hu, Y., Wang, H. (2006).

Effect of surface roughness parameters on mixed lubrication characteristics. Tribology International, vol. 39, p. 522-527.

[16] Sedlaček, M., Podgornik, B., Vižintin, J. (2009). Influence of surface preparation on roughness parameters, friction and wear.

Wear, vol. 266, p. 482-487.

[17] NPL. Internet based surface metrology algorithm testing system, from http://syseng. nist.gov/VSC/jsp/index.jsp, accessed on 200901-03.

[18] Demir, H., Gullu, A., Ciftci I., Seker, U. (2010). An investigation into the influences of grain size and grinding parameters on surface roughness and grinding forces when grinding.

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Sedlaček, M. – Vilhena, L.M.S. – Podgornik, B. – Vižintin, J.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 681-688 DOI:10.5545/sv-jme.2010.205

Paper received: 28.09.2010 Paper accepted: 21.06.2011

Aspects of Using Tool Axis Inclination Angle

Sadílek, M. – Čep, R.– Budak, I. – Soković, M. Marek Sadílek1,* – Robert Čep1 – Igor Budak2 – Mirko Soković3 1 VŠB-Technical University of Ostrava, Faculty of Mechanical Engineering, Czech Republic 2 University of Novi Sad, Faculty of Technical Science, Serbia 3 University of Ljubljana, Faculty of Mechanical Engineering, Slovenia This contribution deals with the research and proposal to change a position of tool axis against milled surface during multi-axial milling. Our target is achieving an increase in milling efficiency (improvement of functional surface properties, increase in milling accuracy, increase in tool durability, decrease in energy load on a machine, and shortening of milling time). This research attempts to make production of shape planes more efficient. This concerns production of molds, impression dies, and other complicated parts in various engineering industries, primarily automotive and aircraft ones. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: tool axis inclination angle, ball-end milling, surface roughness, cutting forces 0 INTRODUCTION Pre-defined milling cycle possibilities are used during the programming of multiaxial milling centers in the CAM systems [1].

However, the question is whether programmers are able to utilize all the available setting options optimally. Do they put in all the significant information necessary for effective milling? Do these programmers know what values are the most effective?

These questions for programmers lead to an underestimated and neglected option to change tool axis position against the normal of milled surface. This work, among other things, attempts to find an answer to this question and confirm it scientifically. The result then should be determining an effective angle or rather an effective range of settings of the spatial angle of the tool axis position in relation to a milled surface.

All milling parameter settings have to lead to increased milling efficiency, i.e., to an increase in accuracy, improvement of functional properties of milled surface (roughness, waviness, residual stress, micro-hardness, etc.), a decrease in milling time, shortening of machine energy load, economical and ecological aspects, etc. [1] and [2].

Surface roughness and residual stress affect functional surfaces, durability and dependability of parts, their noisiness, break-in period, friction losses, electrical resistance, heat transfer, fatigue

strength, wear and corrosion resistance, etc. [1], [3] and [4].

1 MILLING BY INCLINED TOOL During standard milling with ball end milling cutters, when the material and the tool are in right angles, a spherical cutting edge has zero cutting speed at the tool axis. The tool merely pushes in the milled material at this place. Due to this, undesirable effects such as: chip contraction, increase in the cutting temperature, increase in vibrations, and increased creation of a build-up edge, can appear [4] to [6]. These phenomena result in a worsened quality of the milled surface and decreased tool durability. Fig. 1 shows the possibilities of tool inclination toward the surface normal.

The above mentioned phenomena can be eliminated by a change of the tool axis position in relation to the milled piece, i.e., by an inclination of the tool or the piece.

The effective diameter of the cutting tool during milling without tool inclination is calculated according to the following relationship [6]:

(

)

d eff = 2 ⋅ a p d - a p ,

(1)

where deff is effective tool diameter [mm], ap axial depth of cut [mm] and d tool diameter [mm].

Feed direction is highly significant. If the feed direction is (the so called) pulled one (Fig.

*Corr. Author’s Address: VŠB - Technical University of Ostrava, Faculty of Mechanical Engineering, tř. 17. listopadu 15, 708 33, Ostrava, Czech Republic, marek.sadilek@vsb.cz

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2a), the tool action is more silent and the surface of the milled material is better, as opposed to (the so called) pushed feed direction (Fig. 2b). These two ways can be used for an inclination in the feed direction and also for an inclination that is perpendicular to the feed direction.

where deff is effective tool diameter [mm], ap depth of cut [mm], βf inclination angle in feed direction [°] and d tool diameter [mm].

Milling with pushed tool corresponds with negative values of βf in the relationship (Eq. (2)). This signifies that, according to the mathematical expression, the use of a pushed tool is disadvantageous.

a) b) Fig. 1. Milling strategy with tool axis inclination angle, a) tilt in feed direction, b) tilt in pick feed direction

Fig. 3. The relationship of the effective tool diameter deff and the effective cutting speed vceff on the angle of tool inclination βn, (d = 10 mm, ap = 0.3 mm, vc = 210 m·min-1)

Fig. 2. Feed direction; a ) pulled tool, b) pushed tool

The problem of the scallop generation mechanism is quite complicated and falls into the field of applied mathematics. Many foreign publications describe this mechanism. However, these publications tend to only describe the situation when the tool and the workpiece are in a relative translation motion only (no rotation of ball-nosed cutter). The generating mechanism of the rotation ball-nosed end milling, however, is much comlicated because the orientation of the cutting edge is dynamically and periodically changed during the spindle rotation [8].

Owing to a change in position the effective tool diameter changes and so does the resulting (actual) effective cutting speed (Fig. 3).

The effective diameter of the cutting tool during milling with a pulled tool is calculated according to the following relationship [7]:

682

  d - 2 ap d eff = d ⋅ sin arccos    d

  + ² 

f

  , (2) 

Fig. 4. Occurrence of theoretical surface roughness in pick feed direction on the plane [6]

Sadílek, M. – Čep, R.– Budak, I. – Soković, M.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 681-688

Literature [9] presents a new geometrical model for the surface scallop estimation that considered the dynamic cutting edge rotation effect in the ball-nose end milling process.

Literature [10] presented the model to include the effect of the tool axis inclination.

One of many published approaches used to determine theoretical surface roughness is described in [7].

According to this approach the resulting relationship for calculation of the average arithmetic deviation of surface roughness for milling with ball end milling cutter (Fig. 4) is [7]: 2   1 ae a  R  + arc arcsin e   − arc 2 arccos  cos arcsin ⋅ ⋅ R   ae  2 2 R a 2  e 

Ra = R

 1 a a    R − sin 2 arccos  cos arcsin e + arc arcsin e    ⋅1000, ⋅ ⋅ R    2 2 R a 2  e 

(3)

where Ra is arithmetical mean deviation of the profile [µm], R cutter radius [mm] and ae depth of cut [mm].

In the case of milling on inclined plane, ae is substituted with a´e modified by the angle of the inclined plane α: a'e =

ae , cos α

(4)

where α isangle of the milled surface [°], a´e angle of the inclined plane [°] and ae -depth of cut [mm].

For practical application the calculation for the maximum height of the profile Rz, that can be found in the CSN EN ISO 4287 and CSN EN ISO 4288 standards, is sufficient:  a Rz = R ⋅ 1 − 1 − e 2  4 ⋅ R 

  , 

(5)

where: Rz maximum height of the profile [mm], R cutter radius [mm], ae depth of cut [mm].

3 EXPERIMENTAL WORK Experiment characterization: tool axis angle in pick feed direction, conventional milling and climb milling combination, • strategy of feed designated as pulled tool, • using cutting fluid, • workpiece 1.7131, • ball end milling cutter (cutting inserts, 2 flutes, coating 8040), • cutting geometry of exchangeable cutting edge: γp = 0° and γf = -7° to 14°, • cantilever length ln = 110 mm.

Representative examples of the surface roughness parameter Rz dependency on the angle of tool inclination (βn) have been selected from many performed experiments (Figs. 5 and 6).

The biggest maximum height of the profile Rz was measured in the feed direction and in the direction perpendicular to it. The lines are shown with the expanded combined uncertainty Uc. Measurements of the surface roughness parameters were performed on Hommel – Tester T2000.

The graphs in Figs. 5 and 6 show that the most suitable cutter inclination is βn around 15°. When the inclination angle is larger surface roughness increases, especially due to the change of cutting geometry, where there is cutting outside of the so called transitional edge. The best longitudinal (feed direction) roughness was achieved with the inclination βn = 10°.

Deviations of theoretically calculated roughness from the actually measured value are presented on Fig. 7. The graph shows smaller measured roughness as opposed to the theoretical one for larger cut widths and feed per tooth. This result is related to the problem of the minimum chip thickness. For double cut widths and feed per • •

Table 1. Cutting conditions collection depth of of cut surfaces AMF a b

diameter spindle of rev. endmill

ap [mm]

d [mm]

0.3

10

cutting speed

n vc [min-1] [m·min-1] 6685

210

theoretic surface roughness pick feed direction Rz Ra [µm] 0.78 0.2 1.56 0.4

Aspects of Using Tool Axis Inclination Angle

feed direction Rz

Ra [µm] 0.78 0.2 1.56 0.4

width of cut

feed per tooth

fz ae [mm] [mm] 0.1765 0.1765 0.2497 0.2497 683

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Fig. 5. Surface roughness (Rz) dependence on tool axis inclination angle, collection of surfaces “a”

Fig. 6. Surface roughness (Rz) dependence on tool axis inclination angle, collection of surfaces “b”

Fig. 7. Deviations of theoretical surface roughness (Rz) at different tool axis angles, pick feed direction (d = 10 mm, ae, fz = 0.18 mm and ae, fz = 0.25 mm, collection of surfaces „a and b“) 684

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tooth (ae and fz) relatively equal surface roughness can be achieved.

Bearing length ratio changes more in longitudinal direction then in transversal one (Fig. 8). The most advantageous values of bearing length ratio appear in the longitudinal direction with the tool inclinations of 5 and 15°. Here the bearing length ratio is the largest. These results indicate a beneficial impact of the tool inclination on functional aspects of milled surfaces.

The question is to what extent the variable tool inclination influences the change in surface properties. Shape and curvature of the tool edge

have the biggest influence on surface integrity.

Out of cutting parameters the fundamental influence is shown primarily by the cutting speed [1], [2] and [10].

Residual stress at the workpiece surface layer areas is a manifestation of the used machining technology. One of the possibilities how to indicate structural and tension states of ferromagnetic materials in the surface layer is using the magnetoelastic method.

From the experiments it follows that the cutter inclination results in a decrease in the residual stress, see Fig. 9. During perpendicular

Fig. 8. Curves of the profile bearing length ratio (level 30%) at different tool axis angles

Fig. 9. Barkhausen noise relation to tool axis angle - feed direction and pick feed direction (material X3CrNiMo13-4, d = 25 mm, vc = 153 mm, ap = 0.3 mm, ae = 0.6 mm, fz = 0.6 mm) Aspects of Using Tool Axis Inclination Angle

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position of the ball end milling cutter (inclination = 0°) in relation to a milled surface there is a relatively high value of Barkhausen noise (i.e., undesirable tensile stress that needs to be eliminated) due to pushing-in at the tool axis with zero cutting speed.

Already a tool inclination of five degrees results in a significant decrease of the residual stress influence on the surface (milled) layer. The Barkhausen noise values do not change with a further change of the tool position significantly.

Curves of the values measured in feed and pick feed directions are similar.

Residual stress was measured in 0.01 to 0.04 mm depth.

In order to obtain stress values in [N] a calibration needs to be performed and the BS values should be recalculated using the calibration curves. This calibration, however, is economically demanding and time consuming, therefore it is not absolutely necessary for the confirmation of influence of inclination on residual stress.

Fig. 10. Direction and size of cutting forces result (Fv) dependence on cutting time (toll axis angle βn = 0° and 15°, d = 10 mm, ap= 0.3 mm, fz= 0.2 mm, vc=250 m·min-1)

Fig. 11. Resulting cutting force depending on tool axis inclination angle during climb milling (workpiece 1.2343, 47 – 51 HRC , d = 10 and 20 mm, ap = 0.3 mm, fz = 0.2 mm, vc=250 m·min-1) 686

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Dynamometer Kistler 9255B and DASYLab, Excel, and Matlab programs were used for measuring of cutting forces (Fig. 10).

As an example out of many measurements the curves of resultant forces during climb milling are shown in Fig. 11. Lower ranges of individual cutting force components were achieved by the cutting tool inclination angle changes. This leads to cutting process stability. For conventional milling several fold decreases of value range can be shown.

The course of the values component Fy (feed direction) decreases with tool inclination and then grows. Such a course (similar bathtub curve) is also noted for the results of measurements of roughness parameters, depending on the tool axis inclination angle. This confirms the fact that the decline in component Fy causes the improvement of surface roughness.

4 DISSCUSION AND CONCLUSIONS After the final experiment evaluation it can be concluded that the tool inclination has a significant influence on the longitudinal and transversal surface roughness. With increasing cutter diameter smaller inclination can be selected.

Benefits of using tool axis inclination angle include: • an increase in cutting speed, • a decrease in surface roughness in both directions (pick feed direction and feed direction), • a decrease in cutting time (using bigger ae, fz by the same surface roughness), • an increase in the durability of cutting tool, • an increase in the accuracy of cutting, • constant cross sectional area of chip, • constant cutting conditions, • a decrease in size of cutting forces components, • favorable orientation of cutting forces direction, • an increase in functional surface properties of the machined surface, • inhibition of self-excited oscillations, and • a decrease in cutting temperature.

During milling it is necessary not to exceed the maximum tool inclination in the relation to a workpiece; the inserts geometry needs to be

maintained. When this inclination is exceeded the result is increased surface roughness. This eventuality can occur while milling steeper surfaces. With used inserts there was cutting outside the so called transition edge during higher inclinations (e.g., for d = 10 mm the critical βn = 17°).

For most surface groups there was also considerable dependency of the tool inclination on residual stress.

The tool inclination against a workpiece has a significant influence on the size and direction of individual cutting force components. This can be used for the optimization of tool inclination spatial angle in order to achieve a better quality of milled surface, an increase of tool durability, and lower energy demands during milling. Based on the measured cutting forces the tool inclination spatial angle can be optimized online, which leads to adaptive optimization [10], [11] and [12] .

The application research area is the milling of shape surfaces with suppression of grinding operation. The integrity of ground surface is often unsatisfactory, for example from the point of view of heat and tension (residual stress) influence on surface layers [6]. Therefore, research of the proposed technology contributes to the effort to eliminate grinding (i.e., the final operation of manual finishing of shape surfaces), or to minimize it.

5 REFERENCES [1] Župerl, U., Čuš, F., Reibenschuh, M. (2011).

Neural control strategy of constant cutting force system in end milling, Robotics and Computer-Integrated Manufacturing, vol. 27, no. 3, p. 485-493.

[2] Zetek, M., Řehoř, J., Strnad, T. (2006).

Increasing cutting tool efficiency. PhD thesis, University of West Bohemia, Plzeň, p. 1-9.

[3] Adamczak, S., Miko, E., Čuš, F. (2009). A model of surface roughness constitution in the metal cutting process applaying tools with defined stereometry. Strojniški vestnik Journal of Mechanical Engineering, vol. 55, no. 1, p. 45-54.

[4] Xu, Ch., Chen, H., Liu, Zh., Cheng, Zh. (2009). Condition monitoring of milling tool wear based on fractal dimension of

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[5]

[6] [7]

[8]

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vibration signals. Strojniški vestnik - Journal of Mechanical Engineering, vol. 55, no. 1, p.

15-25.

Wei, Z.C., Wang, M.J., Zhu, J.N., Gu, L.Y. (2011). Cutting force prediction in ball end milling of sculptured surface with Z-level contouring tool path. International Journal of Machine Tools & Manufacture, vol. 51, no. 5, p. 428-432. Čep, R., Neslušan, M., Barišić, B. (2008).

Chip formation analysis during hard turning.

Strojarstvo, vol. 50, no. 6, p. 337-345.

Peterka, J. (2004). The new approach of calculating the arithmetic mean deviation of surface roughness of machined surfaces during ball end milling. Engineering Technology, vol. IX, no. 2, p. 28-32. (in Czech) 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 & Manufacture, vol. 45, p. 1077-1084.

[9] Mizugaki, Y., Hao, M., Kikkawa, K. (2001).

Geometric generating mechanism of machined surface by ball-nosed end milling.

Annals of the CIRP, vol. 50, no. 1, p. 69-72.

[10] Mizugaki, Y., Kikkawa, K., Teral, H., Hao, M. (2003). Theoretical estimation of machined surface profile based on cutting edge movement and tool orientation in ballnosed end milling. Annals of the CIRP, vol.

52, no. 1, p. 49-52.

[11] Kopač, J., Kržič, P. (2008). CAM algorithm as important element by achieving of good machined surface quality. Strojniški vestnik Journal of Mechanical Engineering, vol. 54, no. 4, p. 280-287.

[12] Čuš, F., Župerl, U. (2011). Real-time cutting tool condition monitoring in milling.

Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 2, p. 142-150.

Sadílek, M. – Čep, R.– Budak, I. – Soković, M.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 689-696 DOI:10.5545/sv-jme.2011.035

Paper received: 14.02.2011 Paper accepted: 08.06.2011

Influence of Input Parameters on the Characteristics of the EDM Process Shabgard, M. – Seyedzavvar, M. – Oliaei, S.N.B. Mohammadreza Shabgard1,* – Mirsadegh Seyedzavvar1 – Samad Nadimi Bavil Oliaei2 1Department of Mechanical Engineering, University of Tabriz, Iran 2Department of Mechanical Engineering, Middle East Technical University, Turkey

This paper presents the results of experimental studies carried out to conduct a comprehensive investigation on the influence of Electrical Discharge Machining (EDM) input parameters on the characteristics of the EDM process. The studied process characteristics included machining features, embracing material removal rate, tool wear ratio, and arithmetical mean roughness, as well surface integrity characteristics comprised of the thickness of white layer and the depth of heat affected zone of AISI H13 tool steel as workpiece. The experiments performed under the designed full factorial procedure, and the considered EDM input parameters included pulse on-time and pulse current. The results of this study could be utilized in the selection of optimum process parameters to achieve the desired EDM efficiency, surface roughness, and surface integrity when machining AISI H13 tool steel. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: electrical discharge machining, material removal rate, tool wear ratio, surface roughness, white layer thickness, depth of heat affected zone 0 INTODUCTION Considering the challenges brought on by advanced technology, the Electrical Discharge Machining (EDM) process is one of the best alternatives for machining an ever increasing number of high-strength, non-corrosion, and wear resistant materials [1] and [2]. AISI H13 tool steel is considered as a significant material that has a widespread application in mold industries [3].

Electrical discharge machining utilizes rapid, repetitive spark discharges from a pulsating direct-current power supply between the workpiece and the tool submerged into a dielectric liquid [4]. The thermal energy of the sparks leads to intense heat conditions on the workpiece causing melting and vaporizing of the workpiece material. Due to a high temperature of the sparks, not only work material is melted and vaporized, but the electrode material is also melted and vaporized, which is known as tool wear. The tool wear process is quite similar to the material removal mechanism of the workpiece as the tool and the workpiece are considered as a set of electrodes in the EDM process. Due to this wear, tool loses its dimensions resulting in inaccurate cavities formed on the workpiece. Consequently, during the EDM process, the main machining

output parameters are the material removal rate (MRR), tool wear ratio (TWR) and surface roughness (Ra) of the workpiece. It is desirable to obtain the maximum MRR with minimum TWR and surface roughness [5].

Furthermore, at the end of each discharge, depending on the plasma flushing efficiency (% PFE) or the ability of plasma channel in removing molten material from the molten material crater, collapsing of the plasma channel causes very violent suction and severe bulk boiling of some molten material and removing them from the molten crater [6]. The material remained in the crater re-solidifies, which is called the “white layer” or “recast layer”, and develops a residual stress that often causes micro cracks. An annealed Heat Affected Zone (HAZ) lay directly below the recast layer. The micro cracks created in the white layer could penetrate into the HAZ. Additionally, this layer is softer than the underlying base material. This annealed zone could weaken prematurely and cause the material to develop stress fractures that could lead to anything from a minor malfunction to a catastrophic failure.

Since the quality of an ED machined surface is becoming more and more important to satisfy the increasing demands of sophisticated component performance, longevity and reliability [7] and

*Corr. Author’s Address: Tabriz University, Department of Mechanical Engineering, Tabriz, Iran, mrshabgard@tabrizu.ac.ir

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Table 1. Mechanical and physical properties of AISI H13 [10] Electrical Modulus of Temperature Density Specific heat resistivity elasticity [°C] [kg/dm3] [J/(kg·K)] [Ω·mm2/m] [N/mm2] 20°C 7.80 460 0.52 215×103 500°C 7.64 550 0.86 176×103 600°C 7.60 590 0.96 165×103 Liquidus temperature 1454 °C Solidus temperature [8], the optimum utilization of the EDM process requires the selection of an appropriate set of machining parameters that would result in the minimum thickness of the recast layer and depth of heat affected zone [9].

This paper aims to fill the gap in the existing literature with respect to the processing of AISI H13 tool steel with EDM. In particular, EDM machining experiments were conducted on AISI H13 samples having a hardness of 52.7 HRC using copper electrode to investigate the correlations between the EDM parameters (pulse on-time and current) and the EDM characteristics of such a workpiece. The output factors investigated were the material removal rate, tool wear ratio, surface roughness, as well as the thickness of white layer and depth of heat affected zone of EDMed workpiece. This experimental study results in the selection of optimum process parameters to achieve the desired EDM efficiency, surface roughness, and surface integrity when machining such a workpiece material.

1 EXPERIMENTAL SETUP AND PROCEDURE The workpiece material used in this study was AISI H13 tool steel. Prior to EDM processing, the workpiece was cut in a cylindrical shape with a length of 20 mm and a diameter of 20 mm. The main mechanical and physical properties of such workpiece material at different temperatures are given in Table 1.

The tool material was forged commercial pure copper with the main properties given in Table 2. The experiments were performed on a die sinking EDM machine (CHARMILLES ROBOFORM200) which operates with an iso-pulse generator. Machining tests were carried out at five pulse current settings, as well as four pulse ontime settings. As a result, 20 experiments could be 690

Thermal conductivity [W/m·K] 24.30 27.70 27.50 1315 °C

designed. Each machining test was performed for 15 minutes. Table 3 presents the experimental test conditions.

Table 2. Physical properties of copper electrode [11] Physical properties Thermal conductivity [W/m·K] Melting point [°C] Boiling temperature [°C] Specific heat [cal/g·°C] Specific gravity at 20°C [g/cm3] Coefficient of thermal expansion [×10-6 (1/°C)]

Copper 380.7 1083 2595 0.092 8.9 17

A digital balance (CP2245-Surtorius) with a resolution of 0.1 mg was used for weighing the workpieces before and after the machining process. The tool wear ratio is defined as the volume of material removed from the tool (VE) divided by the volume of material removed from the workpiece (VW). Eqs. (1) and (2) show the calculations used for assessing the values of MRR and TWR.

MRR = (M1 – M2) / ( ρw·T) , (1) TWR = (VE / VW)·100% ,

(2)

where M1 and M2 are the weight of workpiece before and after machining [g], respectively. ρw is the density of workpiece [g/mm3], and T is the machining time [min].

According to Lee and Tai [12], the amount of white layer thickness (WT) has been measured by measuring this layer’s thickness at 30 different points by utilizing VEGA\\ TESCAN scanning electron microscopy (SEM) and accounting for their average (Figs. 1 to 3). Therefore, the machined specimens were sectioned transversely by a wire electrical

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discharge machine and prepared under a standard procedure for metallographic observation. Etching was performed by immersing the specimens in 5% Nital reagent.

harness tester. The values of WT and HD are represented in Table 4.

Table 3. Experimental test conditions

2.1 Effect of Pulse on-Time and Pulse Current on Machining Characteristics

Generator type

Iso-pulse (ROBOFORM 200) Oil Flux ELF2 Normal submerged

Dielectric fluid Flushing type Power supply voltage 200 [V] Reference voltage [V] 70 Pulse current [A] 8; 12; 16; 20; 24 Polarity Positive Pulse on-time [µs] 12.8; 25; 50; 100 Pulse interval [µs] 6.4 Tool material Commercial pure copper Cylindrical Tool shape (Ø18.3 mm and L=20 mm) On the other hand, according to Hascalyk and Caydas [13], since there are not many significant differences between HAZ and parent material in the microscopic images that could be identified, measuring micro-hardness is a reasonable way to obtain the depth of heat affected zone (HD). With this in mind, micro-hardness from the cross-section of machined specimens was measured to determine the depth of the heat affected zone. The micro-hardness of specimens was measured by the OLyMPUS LM700 micro-

2 RESULTS AND DISCUSSION

The correlation between machining characteristics and pulse on-time in machining of AISI H13 tool steel using copper electrode are shown in Figs. 4 to 6. According to these figures, an increase in the pulse on-time causes an increase in the MRR and Ra, but a decrease in the TWR. By the increase in pulse on-time, the discharge energy of the plasma channel and the period of transferring of this energy into the electrodes increase. This phenomenon leads to a formation of a bigger molten material crater on the workpiece which results in a higher surface roughness.

However, the dimension of plasma channel and the effect of thermal conductivity of electrodes in dispersing the thermal from the spark collision position increase by the increase in pulse on time.

Consequently, by dispersing more heat from the spark stricken position and increasing the amount of heat transferred from the plasma channel to the electrodes, the plasma channel efficiency in removing molten material from the molten crater at the end of each pulse decreases, while the dimensions of the molten crater on the electrodes increases. This effect is more pronounced for copper electrode, since its thermal conductivity

Table 4. The average values for the white layer thickness (WT) and depth of heat affected zone (HD) at different machining settings Settings [A] [µs] 8 12.8 8 25 8 50 8 100 12 12.8 12 25 12 50 12 100 16 12.8 16 25

Average WT [µm] 7.3 8.6 19.3 23.4 7.5 11 18.8 22.3 7.7 10.7

Average HD [µm] 12.0 15.7 24 34.4 12.5 16.5 23 34.8 13 17.8

Settings [A] [µs] 16 50 16 100 20 12.8 20 25 20 50 20 100 24 12.8 24 25 24 50 24 100

Average WT [µm] 17.75 22.5 7 10 16 20 6.5 8.3 14.2 20.5

Influence of Input Parameters on the Characteristics of the EDM Process

Average HD [µm] 23.5 32.7 12 16.2 21.5 30.2 11 15 21 29.6 691

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is much higher than that of the workpiece. As a result, tool wear ratio decreases by increase in pulse on-time.

Figs. 7 to 9 show that MRR, TWR, and Ra increase with augments of the pulse current. Such results were expected as it is obvious that a higher current causes a stronger spark, which results in more eroded material for both electrodes.

Fig. 3. SEM micrograph showing the white layer of EDMed workpiece (I = 24 A and Ti = 100 µs)

Fig. 1. SEM micrograph showing the white layer of EDMed workpiece (I = 8 A and Ti = 25 µs)

At a low current, a small quantity of heat is generated and a substantial portion of it is absorbed by the surroundings. As a result, the amount of utilized energy in melting and vaporizing the electrodes is not so intense. However, with an increase in pulse current and with a constant amount of pulse on-time, a stronger spark with higher thermal energy is produced [14], and a substantial quantity of heat will be transferred into the electrodes. Furthermore, as the pulse current increases, discharge strikes the surface of the sample more intensely, and creates an impact force on the molten material in the crater and causes more molten material to be ejected out of the crater, so the surface roughness of the machined surface increases.

2.2 Effect of Pulse on-Time and Pulse Current on Surface Integrity

Fig. 2. SEM micrograph showing the white layer of EDMed workpiece (I = 24 A and Ti = 50 µs) 692

The increase in the thickness of white layer and depth of heat affected zone by the increase in pulse on-time can be clearly seen from the experimental results (Figs. 10 and 11).

The justification for this phenomenon is that the plasma flushing efficiency has a strict effect on the white layer thickness. With an increase in pulse on-time, plasma flushing efficiency decreases.

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Fig. 4. MRR vs. pulse on-time

Fig. 7. MRR vs. pulse current

Fig. 5. TWR vs. pulse on-time

Fig. 8. TWR vs. pulse current

Fig. 6. Ra vs. pulse on-time

Fig. 9. Ra vs. pulse current

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Fig. 10. WT vs. pulse on-time

Fig. 11. HD vs. pulse on-time

Fig. 12. WT vs. pulse current 694

As a result, the ability of plasma channel for ejecting the molten material from the molten puddle decreases. Subsequently, this remained molten material in the molten puddle re-solidifies forming a white layer upon the machined surface.

Furthermore, the increase of discharge duration increases the amount of the conducted heat into the workpiece during each discharge, and consequently, more underlying material is affected by the high temperature. Overly, this phenomenon causes the increase in the white layer thickness and heat affected zone.

In other words, the amount of molten material which can be flushed away at the end of each discharge is dependent on the plasma flushing efficiency (%PFE). Clearly the %PFE depends on the discharge energy (W), energy gradient (dW/dt), geometrical dimensions of the gap and molten material crater, pressure of the gap (P), and gap pressure gradient (dP/ dt). Depending on the amount of the mentioned parameters, plasma flushing efficiency decreases as pulse on-time increases. The cause of this phenomenon could be justified by the fact that the increase in pulse on-time causes to decrease in the energy changing rate, as this causes a major increase in diameter while not much increase in the average temperature of the plasma channel, which leads to decrease in the pressure of the gap and its changing rate. Therefore, regarding the mechanism of bulk boiling phenomena, the amount of molten material, which is ejected from the molten material crater at the end of discharged, decreases and as a result, the %PFE decreases.

From Figs. 12 and 13 it is clear that increasing the pulse current has a very small effect on the white layer thickness and the depth of the heat affected zone. Although an increase in pulse current leads to an increase in the dimensions of the molten crater and the heat penetrating depth, the plasma flushing efficiency increases as pulse current increases. The increase in plasma flushing efficiency causes more molten material to be swept away from the molten crater, therefore, a thinner layer of re-deposited material appears on the surface of the workpiece. Since an increase in the penetrating depth of heat into the workpiece and plasma flushing efficiency counterbalance the effect they have on each other, an increase in the

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pulse current has no significant effect on the depth of the heat affected zone.

4. A slight decrease could be observed in the white layer thickness by an increase in the pulse current.

5. By constant level of discharge energy, high pulse current and low pulse on-time leads to a reduction in the white layer thickness and depth of heat affected zone on the surface of EDMed workpiece.

4 ACKNOWLEDGEMENTS

Fig. 13. HD vs. pulse current Furthermore, with an increase in the pulse current and with a constant amount of pulse on-time, causing a sharp rise in the average temperature of the plasma channel [15], the energy gradient increases, which leads to an increase in the pressure of gap. Therefore, regarding the mechanism of bulk boiling phenomenon, the amount of molten material, which is ejected from the molten puddle at the end of each discharge, increases and as a result, the %PFE increases [16] as the reports of Marafona et al. prove [10].

3 CONCLUSION Results from an experimental investigation on the effect of machining parameters on EDM process characteristics have been presented. The leading conclusions are as follows: 1. The increase in pulse on-time leads to an increase in the material removal rate, surface roughness, as well the white layer thickness and depth of heat affected zone.

2. The increase in pulse current leads to a sharp increase in the material removal rate and surface roughness.

3. The tool wear ratio decreases by the increase of pulse on-time, and increases by the increase in the pulse current.

The authors of this study are indebted to the Razi Metallurgical Laboratory, Metallurgical Laboratory of Sahand University of Technology, universal workshop of Training Center of Iran Tractor Manufacturing Company, and advance machining workshop of Manufacturing Engineering Department of University of Tabriz.

Also, we would like to appreciate the help of authors Professors J. Khalil Allafy, T.B. Navid Chakharlu, as well Mr. A. Nejat Ebrahimi for their invaluable technical support.

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[2] Merdan, M.A.E.R., Arnell, R.D. (1991). The surface integrity of a die steel after electrodischarge machining, 2. residual stress distribution. Surface Engineering, vol. 7, p.

154-158.

[3] Castro, G., Fernandez-Vicente, A., Cid, J. (2007). Influence of the nitriding time on the wear behavior of an AISI H13 steel during a crankshaft forging process. Wear, vol. 263, p.

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[4] Abdullah, A., Shabgard, M.R. (2008). Effect of ultrasonic vibration of tool on electrical discharge machining of cemented tungsten carbide (WC-Co). International Journal of Advanced Manufacturing Technology, vol.

38, p. 1137-1147.

[5] Khan, A.A. (2008). Electrode wear and material removal rate during EDM of aluminum and mild steel using copper and brass electrodes. International Journal of

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Advanced Manufacturing Technology, vol.

39, p. 482-487.

[6] Dibitoto, D.D., Eubank, Ph.T., Patel, M.R., Barrufet, M.A. (1989). Theoretical models of the electrical discharge machining process, I. A simple cathode erosion model. Journal of Applied Physics, vol. 66, no. 9, p. 4095-4103.

[7] Mamalis, A.G., Vosniakos, G.C., Vaxevanidis, N.M. (1987). Macroscopic phenomena of electro-discharge machined steel surface: an experimental investigation. Journal of Mechanical Working Technology, vol. 15, p.

335-356.

[8] Boujelbene, M., Bayraktar, E., Tebni, W., Ben Salem, S. (2009). Influence of machining parameters on the surface integrity in electrical discharge machining. Archive of Materials Science and Engineering, vol. 37, p. 110-116.

[9] Rebelo, J.C., Dias Morao, A., Kremer, D., Lebrun, J.L. (1998). Influence of EDM pulse energy on the surface integrity of martensitic steel. Journal of Materials Processing Technology, vol. 84, p. 90-96.

[10] Böhler Edelstahl, http://www.bohleredelstahl.at, accesed on 2008-08-23.

[11] Fischer, U., Heinzle, M., Näher, F., Paetzold, H., Gomeringer, R., Kilgus, R., Oesterle, S., Stephan, A. (2008). Tabellenbuch Metall.

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Verlag Europa-Lehrmittel, Nourney, Vollmer GmbH & Co. KG, Haan-Gruiten.

[12] Lee, H.T. Tai, T.Y. (2003). Relationship between EDM parameters and surface crack formation. Journal of Materials Processing Technology, vol. 142, p. 676-683.

[13] Hascalyk, A., Caydas, U. (2004).

Experimental study of wire electrical discharge machining of AISI D5 tool steel.

Journal of Materials Processing Technology, vol. 148, p. 362-367.

[14] Petropoulos, G.P., Vaxevanidis, N.M., Radovanoviü, M., Zoler, C. (2009).

Morphological - Functional Aspects of Electro-Discharge Machined Surface Textures. Strojniški vestnik - Journal of Mechanical Engineering, vol. 55, no. 2, p.

95-103.

[15] Kansal, H.K., Singh, S., Kumar, P. (2008).

Numerical simulation of powder mixed electric discharge machining (PMEDM) using finite element method. Mathematical and Computer Modeling. vol. 47, p. 12171237.

[16] Das, S., Klotz, M., Klocke, F. (2003). EDM simulation: finite element-based calculation of deformation, microstructure and residual stresses. Journal of Materials Processing Technology, vol. 142, p. 434-451.

Shabgard, M. – Seyedzavvar, M. – Oliaei, S.N.B.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 697-703 DOI:10.5545/sv-jme.2010.258

Paper received: 23.12.2010 Paper accepted: 17.06.2011

Application of Numerical Simulations in the Deep-Drawing Process and the Holding System with Segments’ Inserts Volk, M. – Nardin, B. – Dolšak, B. Mihael Volk1,* – Blaž Nardin1 – Bojan Dolšak2 1 Gorenje Orodjarna d.o.o., Slovenia 2 University of Maribor, Faculty of Mechanical Engineering, Slovenia

The demands for complicated products have increased dramatically over the last few years taking into consideration the utilisation of sheet metal, product quality and process conditions. For reliable product development and stable production process, the use of FEM is necessary.

One of the most significant parameters in the sheet metal forming process is the blank holding force. In the research work, the optimisation of the blank holding force was performed with the help of FEM analysis. For the optimisation the geometry and the structure of the blank holder was optimised.

The best results were obtained with flexible, segmented blank holders, which enables wider technological window for good parts. ©2011 Journal of Mechanical Engineering. All rights reserved.

Keywords: Sheet metal forming, deep drawing, segmented holding system, finite elements methods (FEM), optimization

0 INTRODUCTION Sheet-metal forming and deep-drawing are well-known manufacturing processes. However all developers are of the same opinion that deepdrawing is a very complicated process. There are several factors, such as nonlinearity, large deformation, friction and material characteristics that have a direct influence on the process and are sensitive to each other. In addition, the tolerances of the input materials are very rough, which is an extra challenge for the developers. Therefore, tryouts of the dies are required within an actual industry, in order to find a technological window of good parts (Fig. 1). FEM analyses reduce this set-up time and the subsequent improvement in product quality without cracks, wrinkles and scratches is significant [1].

As can be seen from Fig. 1, holding-force is one of the more important parameters that influence the deep-drawing process and can be calculated using FEM analyses [2]. Therefore, with an appropriate approach, a better quality of the workpiece could be achieved, and the reaction forces on the main parts of the tool are smaller, thus achieving a longer life-time of the tool. FEM analyses can calculate different types of forming, such as hot-forming [3] and hydro-forming but this study only conducted conventional cold *Corr. Author’s Address: Gorenje Orodjarna d.o.o., Partizanska 12, SI-3503 Velenje, Slovenia, mihael.volk@siol.net

forming. There are also different holding systems such as holding with constant force, holding with time-dependent forces, and with segmented or distributed-holding forces [4] to [6]. A segmented holding system with segments’ inserts was chosen for this study. These holding systems are rarely used in current household appliance production and the automotive industry [7].

Fig. 1. Technological window [1] The purpose of this work and analyses was to identify sensitive matrix between the holding forces and the qualities of the workpieces.

Analyses were carried out with a FEM, and the Pam-Stamp software package was used for 697

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calculation. The first goal of this research was to determine the most appropriate holding system and the second to optimize blank holding forces (BHF) for the chosen holding system, by applying fuzzy logic [8].

1 NUMERICAL ANALYSES OF THE DEEP DRAWING PROCESSES 1.1 Product being Investigated An asymmetrical workpiece was chosen from household appliances industry (Fig. 2). This workpiece is one of the component parts from a cooking device and was chosen because it looks simple. However, there were a lot of problems with it, especially with critical corner areas. This part has a valid special criterion because the part is visible on the end-product. No wrinkles, scratches or cracks are allowed.

The workpiece has a different depth of draw on both sides; on the higher side 36 mm and on the lower side 10 mm. The dimensions of the initial blank are 660×220×0.7 mm. The computer models are simplified and are shown in Fig. 3.

The main parts of the computer model contain die, blank, blank holder, segments’ inserts, and punch.

1.2 Basic Parameters The surfaces of the die, punch, blank holder and segments’ inserts were discretized, mainly by quadrangular surface elements and were assumed to be perfectly rigid. The blank was discretized by quadrangular surface elements and the plastic behaviour was discretized by Hollomon’s hardening law.

The dynamic explicit approach was chosen for the calculation of the forming-process.

Appropriate parameters such as friction, punchvelocity and drawing-radii were selected from the metal-forming handbooks and from previous research. These invariably geometrical parameters and process parameters are shown in Table 1. The maximal punch stroke or drawing height was 36 mm; any workpieces made with a drawing height of less than 36 mm were unacceptable.

Table 1. Geometrical and process parameters Parameter Punch velocity Drawing radii Punch stroke Friction coefficient

Value 5 m/s 1.8 mm 36 mm die/blank 0.1; punch/blank 0.12; holder/blank 0.1

1.3 Material Properties of Sheet Metal

Fig. 2. Workpiece

A commercially-available DC04 sheet metal with a thickness of 0.7 mm was used for the blank material and tensile tests were conducted to determine the material properties. For the calculations Hill48 with orthotropic anisotropy was used. The material model coefficients were identified based on stress-strain curves (Table 2).

The material was defined by Hollomon’s hardening law, and is given by i.e. (1):

σ f = C ⋅ ϕen ,

(1)

where σf is yield stress, C is strength coefficient, φe is true strain, and n is hardening exponent.

Tensile tests were carried out on a Zwick/ Roell 1474 machine based on SIST standard. The values are average values gained from five tests.

Fig. 3. Computer model 698

Volk, M. – Nardin, B. – Dolšak, B.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 697-703

Table 2. Material properties of sheet metal Blank material Yield strength Tensile strength Elastic module Strength coefficient Hardening exponent Coefficient of anisotropy

0° 45° 90°

DC04 188.9 N/mm2 298.4 N/mm2 210000 N/mm2 384.98 N/mm2 0.21 1.971 1.538 2.079

2 RESULTS The most significantly controlled parameter was the holding force. The first analyses were made with constant holding force on the blank holder. The holding force was optimized but good parts were still not achieved. Good workpieces are considered to be workpieces without wrinkles, cracks, scratches, and dimensional errors. The critical regions of the workpieces were in the

corner areas where radial tensile stresses and tangential compressive stresses are present due to the material’s retention property. These compressive stresses often lead to flange wrinkles.

Apart from wrinkles, another typical problem in sheet metal stamping is the generation of cracks.

Workpieces made by holding forces of less than 400 kN had wrinkles present and those with holding forces of more than 400 kN had cracks present. From Fig. 4 critical areas where wrinkles and fissures occur can be seen.

In order to achieve good parts additional changes had to be made on the computer model.

Therefore, segment holding systems were used with distributed holding forces. Optimizing techniques were used for testing the shapes of the segment inserts and the sizes of the local holding forces. Different shapes of segment inserts were numerically calculated and qualities of the workpieces estimated. The shapes of the segments and half of the holding system are shown in Fig 5.

All the holding systems consisted of 10 segments

a) b) c) Fig. 4. Critical areas of the workpiece; a) colour scheme, b) wrinkling trend and c) cracks

a) b) c) d) Fig. 5. a) Conventional blankholder and b), c), d) three versions of segmented blankholder Application of Numerical Simulations in the Deep-Drawing Process and the Holding System with Segments’Inserts

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b)

c)

a)

Fig. 6. Flange wrinkling; a) whole model, b) colour scheme, c) detailed view and one guiding blank holder plate, as used for holding in the intermediate areas and for guiding the segment inserts. By using these systems, it is possible to distribute holding forces into 10 areas and each segment’s force could be separately controlled.

The results showed that the best segment shape was option d (Fig. 5). The shape of the segments in this case was designed in such a way that the front parts of the straight segments are holding and the rear parts of the segments are guiding. The corner segments did not need an extra guiding. The holding areas in all options were identical.

The results from these computer analyses are shown above. The main problems occurred on those areas shown in Fig. 6. Strong wrinkling tendencies occurred on the areas between the corners and the straight segments.

This flange-wrinkling may have an effect on the end-product, therefore the proposed holding systems did not satisfy primary demands.

In order to solve this problem, the guiding blank holder was also supported by the holding force. The range of this holding force was from 10 to 100 kN. The results proved to be much better and the workpiece already satisfied the presumed demands of good parts. As the results with guiding blank holder forces of both 10 and 100 kN were similar, the guiding blank holder force had a big 700

influence but the value of the force did not. The main reason is that intermediate areas (areas between segments) are relatively small, therefore even a small holding force could decrease wrinkling tendencies in these areas.

Finally, the segmented blankholder forces were optimized. The BH area was divided into ten segments; however this workpiece was symmetrical in one plane and, therefore only 6 segments were actually optimized. The defects were recognized and localized by a human and the desired trajectory of the holding force was adjusted only for those segments situated in that part where tearing or wrinkling occurred. The extent of increasing or decreasing the values of holding forces depended on the sizes of the defects. As mentioned above, optimization was performed using Fuzzy logic rules, which consisted of IFTHEN sentences which were chosen based on the previous results of FEM analysis. Cracks and wrinkles were recognized based on the forming limit diagram.

Table 3. IF-THEN sentences IF IF

Cracking Wrinkling

THEN THEN

Decrease BHF Increase BHF

Using these rules all 6 segment holding forces were optimized. For determining the rough technological window first the increment of 10 kN

Volk, M. – Nardin, B. – Dolšak, B.

StrojniĹĄki vestnik - Journal of Mechanical Engineering 57(2011)9, 697-703

was used but later for more prĂŠcising computing this increment was 1 kN. Numerous numerical simulations should be performed to obtain the correct trajectories for holding forces. Good workpieces with no wrinkling and no cracks present on the model were made with a holding force from 306 to 734 kN (Fig. 7).

(option a in Fig. 5) and in Fig. 9b was made with a new holding system with segment inserts and with optimized holding forces (option d in Fig. 5). In the technological window with segmented holding force, the holding force is the sum of all 10 segmented forces as well as the blankholder plate force. a)

a)

b)

b)

Fig. 7. Good workpieces made by segmented holding system a) with segmented holding forces 306 kN b) with segmented holding forces 734 kN End-workpieces with a drawing-height of 36 mm made by a holding force of less than 306 kN had wrinkles present and workpieces made by a holding force of more than 734 kN had cracks present. The ranges for the holding forces on each segment are shown in Fig. 8.

Fig. 9. Technological window for a) conventional holding system and b) segmented holding system

Fig. 8. Ranges for the holding forces A new technological window was constructed and is shown in Fig. 9b. The technological window in Fig. 9a was made with a constant holding force and a conventional holding system with only one holding plate

The differences in the technological window regarding the corner areas and the straight areas are shown in Fig. 10. From these diagrams it can be seen that the corners areas are more critical than straight areas.

From all diagrams (Figs. 9 and 10) it is clear that the wrinkles in the corner areas had already occurred in the earlier steps of the deepdrawing process. However, later some of the wrinkles disappear which was also confirmed by the experiment (Fig. 11).

Application of Numerical Simulations in the Deep-Drawing Process and the Holding System with Segmentsâ&#x20AC;&#x2122;Inserts

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 697-703

b)

a)

Fig. 10. Technological window for; a) corner areas and b) straight areas b)

a)

b)

a)

Fig. 11. Wrinkling trend; a) step 4 (FEM), b) step 10 (FEM), c) early step (experiment) and d) final draw (experiment) In order to avoid wrinkling tendencies over all steps time-dependent profiles for corner segmented holding forces will have to be applied, which needs to be carried out in further research.

3 CONCLUSION Different holding systems were evaluated and the advantages and disadvantages are 702

presented. The presented results for deep drawing show that the quality of a workpiece can be improved with a better holding system. It is evident that even small changes in BHF can lead to failure during the process. These failures can be avoided if a variable BHF is applied, but the correct trajectories need to be chosen. This work shows that the results using guiding blank holder forces of 10 and 100 kN were similar. Therefore,

Volk, M. – Nardin, B. – Dolšak, B.

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, 697-703

it can be concluded that because of a relatively small area where guiding blank holder force is in contact the blank had a big influence but value of the force did not.

Taking into account the geometric restrictions of the die, the shape of the segment inserts was also highlighted. Finally, the FEM analyses segmented holding forces were optimized. Numerous numerical simulations were performd and technological windows for each area were given.

However, the analyses did not consider spring-back. Spring-back in sheet metal forming can be described as the change in sheet metal shape compared with the shapes of tools after the forming process. Currently, FEM calculations for spring-back are not relevant, therefore further research should be carried out to find out the effects of different holding systems on the springback effect during sheet metal stamping processes.

4 ACKNOWLEDGMENT Practical part of the research was made in Gorenje Orodjarna d.o.o.

The operation part was financed by the European Union, European Social Fund.

This work is part of the 7th framework program, research and development project, Self-Learning sheet metal forming system (LearnFORM), Grant agreement N° NMP2SL-2009-228346.

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[2] Pepelnjak, T., Kampuš, Z. (2001). Analysing the Quality of Sheet-Metal holding during deep Drawing. Strojniski vestnik – Journal of Mechanical Engineering, vol. 47, no. 2, p.

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Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9 Vsebina

Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 57, (2011), številka 9 Ljubljana, september 2011 ISSN 0039-2480 Izhaja mesečno

Povzetki člankov Franci Pušavec, Janez Kopač: Ocena trajnostnega kriogenega odrezavanja Ni zlitine Inconel 718 Blaza Stojanović, Nenad Miloradović, Nenad Marjanović, Mirko Blagojević, Lozica Ivanović: Spreminjanje dolžine zobatih jermenov med uporabo Van Tuan Do, Ui-Pil Chong: Zaznavanje in diagnostika napak pri asinhronskih motorjih na osnovi modela signalov z uporabo značilnosti signala vibracij v dvorazsežnostni domeni Eduard Niţu, Monica Iordache, Luminiţa Marincei, Isabelle Charpentier, Gaël Le Coz, Gérard Ferron, Ion Ungureanu: Modeliranje hladnega prečnega valjanja krožnih žlebov po metodi končnih elementov Marko Sedlaček, Luis Miguel Silva Vilhena, Bojan Podgornik, Jože Vižintin: Modeliranje topografije površine z namenom zmanjšanja trenja Marek Sadílek, Robert Čep, Igor Budak, Mirko Soković: Vidiki uporabe kota nagiba osi orodja Mohammadreza Shabgard, Mirsadegh Seyedzavvar, Samad Nadimi Bavil Oliaei: Vpliv vhodnih parametrov na značilnosti procesa EDM Mihael Volk, Blaž Nardin, Bojan Dolšak: Uporaba numeričnih simulacij pri procesu globokega vleka in pridrževalni sistem s segmentnimi vstavki

SI 123 SI 124 SI 125 SI 126 SI 127 SI 128 SI 129 SI 130

Navodila avtorjem

SI 131

Osebne vesti Doktorati, magisteriji, specializacije in diplome

SI 133

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 123

Prejeto: 13.12.2010 Sprejeto: 03.08.2011

Ocena trajnostnega kriogenega odrezavanja Ni zlitine Inconel 718 Pušavec, F. ‒ Kopač, J. Franci Pušavec* ‒ Janez Kopač Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija Delo obravnava prednosti trajnostnega kriogenega odrezavanja, ki kaže velik potencial za zmanjšanje proizvodnih stroškov procesov odrezavanja in izboljšanje konkurenčnosti izdelkov, ob zmanjšani porabi virov in manjši količini ustvarjenih odpadkov.

Vprašanja trajnosti, ki se postavljajo na ravni industrijskih tehnologij in procesov odrezavanja, se obravnavajo z ocenjevanjem trajnostnih lastnosti oziroma s t.i. oceno življenjskega cikla. Predpostavka dela je, da predstavlja kriogeno odrezavanje prihodnost za doseganje trajnostne proizvodnje, saj omogoča manjše obremenjevanje okolja in manjša zdravstvena tveganja, hkrati pa povečanje produktivnosti in prihodkov. V delu je tudi eksperimentalno analizirana obdelovalnost visokotemperaturne Ni zlitine (Inconel 718) in predstavljena primerjava proizvodnih stroškov za konkreten primer ob uporabi kriogenega in konvencionalnega odrezavanja.

V prvem delu raziskave so obravnavani ekonomski, socialni in okoljski izzivi v proizvodnji s stališča tehnologij odrezavanja. Natančneje, opravljena je analiza življenjskega cikla (LCA). Rezultati kažejo, da je kriogeno odrezavanje v primerjavi s konvencionalno obdelavo trajnostno alternativna tehnologija. LCA je pokazala, da prehod od hladilno-mazalnih sredstev (HMS) na bazi olj na uporabo tekočega dušika vodi do znatnega zmanjšanja obravnavanih cenilk: trdnih odpadkov, porabe vode, potenciala globalnega segrevanja, zakisljevanja in energetske potratnosti. Ugotovljeno je bilo, da je bolj kot ekstrakcija tekočega dušika iz zraka energetsko potraten proces proizvodnje sodobnih karbidnih rezalnih orodij. Na podlagi primerjalne analize in izračuna celotne porabe energije je bilo dokazano, da je lahko celostno gledano kriogeno odrezavanje energetsko učinkovitejše od konvencionalnih odrezovalnih procesov. To je moč pripisati podaljšanju obstojnosti orodij in tako manjši porabi rezalnih orodij pri kriogenem odrezavanju.

V drugem delu prispevka je izvedena analiza obdelovalnosti zlitine Inconel 718. Opravljene so bile meritve obrabe in določena je bila obstojnost karbidnih rezalnih orodij. Ob znani obstojnosti so bili določeni stroški proizvodnje in uporabe hladilno-mazalnih sredstev (tekočega dušika in HMS na bazi olj). Izkazalo se je, da zamenjava običajnih HMS s kriogenim fluidom zaradi nižjih stroškov, povezanih z odpadki in daljšo obstojnostjo orodij, drastično zmanjša celotne proizvodne stroške na izdelek (do 30%).

To potrjuje, da lahko kljub večjim začetnim stroškom opreme za kriogeno odrezavanje pričakujemo očitne trajnostne koristi s krajšimi proizvodnimi cikli, nižjimi stroški in večjo produktivnostjo.

Za industrijsko aplikacijo predstavljene ideje je razen stroškov ter okoljskih in zdravstvenih vplivov zelo pomemben tudi vpliv procesa na integriteto oz. kakovost obdelane površine. Avtorji članka so ta del obravnavali v enem od nedavno objavljenih člankov, na katere se sklicujejo v tem delu.

Delo predstavlja prvo celostno evalvacijo kriogenega odrezavanja z vidika stroškov in trajnosti uporabe takšnega procesa. Prikazano je, da so stroški rezalnih orodij prevladujoč delež skupnih stroškov proizvodnje in da kriogeno odrezavanje ponuja čist in stroškovno učinkovit način za izboljšanje trajnosti odrezovalnih tehnologij v industrijskih aplikacijah. Rezultati več kot dokazujejo potencial in potrebnost uporabe take tehnologije v orodjarski, avtomobilski in letalski proizvodni industriji, kot nadgradnje obstoječih odrezovalnih procesov. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: odrezavanje, visokotemperaturne zlitine, kriogenika, evalvacija, trajnost, stroški

*Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za strojništvo, Aškerčeva 6, SI-1000 Ljubljana, Slovenija, franci.pusavec@fs.uni-lj.si

SI 123

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 124

Prejeto: 19.03.2010 Sprejeto: 13.06.2011

Spreminjanje dolžine zobatih jermenov med uporabo

Stojanović, B. ‒ Miloradović, N. ‒ Marjanović, N. ‒ Blagojević, M. ‒ Ivanović, L. Blaza Stojanović* ‒ Nenad Miloradović ‒ Nenad Marjanović ‒ Mirko Blagojević ‒ Lozica Ivanović Univerza v Kragujevcu, Fakulteta za strojništvo, Srbija Zobati jermenski pogoni so nova vrsta pogonov, ki se je pojavila v 50. letih prejšnjega stoletja.

Prenos moči in gibanja pri zobatih jermenih poteka z obliko in s trenjem. Podrobne kinematične analize so pokazale, da na prenos moči in gibanja vpliva večje število dejavnikov, pri tem pa je obveljala predpostavka, da je vpliv trenja zanemarljiv. Spreminjanje dolžine zobatega jermena je osnovni pokazatelj stanja zobatega jermena. V članku je predstavljena analiza spreminjanja dolžine zobatega jermena (delitve) med uporabo zaradi triboloških procesov, ki potekajo na njegovih kontaktnih površinah.

Za doseganje zastavljenih ciljev je bila opravljena podrobna analiza prenosa moči in gibanja pri zobatih jermenskih pogonih. Zobje jermena med prenosom moči prek utorov vprijemajo z jermenico in jermen se pri tem premika v tangencialni, radialni in aksialni smeri. Boki zob jermena pridejo po vprijemu v stik z boki zob na jermenici. V periodičnem stiku so tudi notranja površina utorov jermena in zunanja površina jermenice ter jermen in obroč prirobnice jermenice.

Na osnovi opravljene kinematične analize so bili določeni osnovni tribomehanski sistemi zobatih jermenskih pogonov. Laboratorijski preizkusi jermenskega pogona so bili opravljeni na posebnem preizkuševališču, ki deluje po principu odprte zanke moči. Postavljena je bila tudi ustrezna merilna veriga z vso potrebno opremo in senzorji. Za realno sliko o triboloških značilnostih zobatega jermena so bile opravljene eksperimentalne meritve sprememb geometrijskih vrednosti in parametrov hrapavosti. Med vsemi geometrijskimi parametri ima največji vpliv na življenjsko dobo zobatih jermenskih pogonov delitev jermena, ki določa njegovo dolžino.

Rezultati meritev kažejo, da se merjene geometrijske vrednosti med uporabo občutno spreminjajo, največje spremembe pa so opazne pri delitvi jermena. Spremembe delitve in s tem dolžine so posledica triboloških procesov na kontaktnih površinah jermena in jermenice ter plastičnih deformacij nosilnega elementa. Jermen se je v danih pogojih preizkušanja podaljšal za 23 mm, pri čemer je bilo 70 % podaljška posledica plastičnih deformacij nosilnega elementa, preostanek pa elastomerne obrabe bokov zob jermena.

Eksperimentalni preizkusi v laboratorijskih pogojih potrjujejo teoretične predpostavke. Vpliv trenja na prenos moči in gibanja ni zanemarljiv in se neposredno odraža v spremembah geometrije jermena, zlasti njegove delitve in dolžine. Laboratorijske preizkuse bi bilo mogoče razširiti tudi na druge vrste jermenov ter na preizkušanje jermenov do uničujoče obrabe.

V članku je predstavljena identifikacija osnovnih tribomehanskih sistemov zobatih jermenskih pogonov in njihova analiza v povezavi s prejšnjimi raziskavami sklopljenih sistemov in porazdelitev obremenitev. Z eksperimentalno in teoretično analizo je bilo ugotovljeno, da ima sprememba delitve (dolžine) jermena kot neposredna posledica procesov na kontaktnih površinah neposreden vpliv na življenjsko dobo in zmogljivost jermenov. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: zobati jermenski pogoni, delitev jermena, trenje, obraba, tribologija, tribomehanski sistemi

SI 124

*Naslov avtorja za dopisovanje: Univerza v Kragujevcu, Fakulteta za strojništvo, Sestre Janjić 6, 34000 Kragujevac, Srbija, blaza@kg.ac.rs

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 125

Prejeto: 20.07.2010 Sprejeto: 27.07.2011

Zaznavanje in diagnostika napak pri asinhronskih motorjih na osnovi modela signalov z uporabo značilnosti signala vibracij v dvorazsežnostni domeni Do, V.T. – Chong, U.P. Van Tuan Do1,* – Ui-Pil Chong2 1 Finski tehnični raziskovalni center VTT, Finska 2 Univerza v Ulsanu, Oddelek za računalništvo in informacijske tehnologije, Južna Koreja

Sistemi za zaznavanje in diagnosticiranje okvar strojev v splošnem in posebej za asinhronske motorje so kritičnega pomena za industrijo, saj lahko okvare strojev povzročijo visoke stroške vzdrževanja, nesprejemljive izdelke itn. V zadnjem času so bile razvite številne tehnike za zaznavanje in diagnostiko napak asinhronskih motorjev na osnovi enorazsežnostnega (1D) signala vibracij. Značilnosti v enorazsežnostni domeni je mogoče pridobiti s tehnikami kot so Fourierjeva transformacija, valčna transformacija, analiza glavnih komponent (PCA), empirična modalna dekompozicija (EMD) in druge.

Te značilnosti je nato mogoče uporabiti v nevronskih mrežah, metodah razvrščanja in metodah podpornih vektorjev (SVM) za sprejemanje odločitev ob pojavu napak, klasifikacijo stopenj napak itd. Avtorji so mnenja, da so za zaznavanje napak in za diagnostiko zelo zanimive tudi značilnosti, pridobljene iz signala v dvorazsežnostni (2D) domeni. Za delo v dveh razsežnostih je treba izvirne podatke o vibracijah pretvoriti v dvorazsežnostne podatke.

V članku je predstavljen pristop k vzpostavitvi sistema zaznavanja in diagnostike napak asinhronskih motorjev na osnovi signala vibracij. Pristop sestoji iz dveh zaporednih procesov: procesa zaznavanja napak in procesa diagnosticiranja napak. V procesu zaznavanja napak se pridobivajo signifikantne značilnosti iz dvorazsežnostnih signalov vibracij s pomočjo algoritma SIFT, izhod procesa pa so simptomi napak.

V procesu diagnosticiranja napak je nato na simptomih napak uporabljena tehnika razvrščanja vzorcev.

Signal vibracij je tako namesto analiziranja za ugotavljanje napak asinhronskega motorja možno razvrstiti v ustrezno kategorijo napak asinhronskega motorja. Predstavljen je tudi okvir tehnike razvrščanja vzorcev, ki je uporabna za vzorce SIFT, in narejena je primerjava z dvema drugima pristopoma iz naših prejšnjih del. Rezultati preizkušanja kažejo, da ima naš predlagani prostop bistveno boljšo natančnost razvrščanja napak in boljše rezultate kot prejšnji pristopi.

Rezultati visoke natančnosti razvrščanja s predlaganim pristopom jasno izkazujejo potencial za izkoriščanje dvodimenzionalnih podatkov signala vibracij v sistemu zaznavanja in diagnosticiranja napak z algoritmom SIFT. V prihodnjih delih bomo uporabili druge metode 1D- in 2D-pretvorbe, kot je Hanklova matrica, ter druge napredne funkcije v 2D, rezultate pa bomo primerjali s predlaganim pristopom. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: zaznavanje in diagnostika napak, SIFT, vektor značilnosti, slovar tekston, dvorazsežnostna domena, natančnost razvrščanja

*Naslov avtorja za dopisovanje: Finski tehnični raziskovalni center VTT, FI-02044, Espoo, Finska, dtuan@ualberta.ca

SI 125

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 126

Prejeto: 03.12.2010 Sprejeto: 27.07.2011

Modeliranje hladnega prečnega valjanja krožnih žlebov po metodi končnih elementov Niţu, E. – Iordache, M. – Marincei, L. – Charpentier, I. – Le Coz, G. – Ferron, G. – Ungureanu, I. Eduard Niţu1,* – Monica Iordache1 – Luminiţa Marincei1 – Isabelle Charpentier2 – Gaël Le Coz3 – Gérard Ferron2– Ion Ungureanu1 1 Univerza Piteşti, Romunija 2 LEM3, Univerza Paul Verlaine-Metz, Francija 3 LEM3, ParisTech-Metz, Francija

V članku so predstavljeni rezultati raziskave, ki je bila opravljena v Romuniji v okviru projekta št.

ID_711/2008 pod nazivom “Analitično in numerično modeliranje procesov izdelave kompleksnih profilov z volumetričnim hladnim tokom”.

Raziskava, predstavljena v tem članku, je namenjena razvoju numeričnega modela za analizo procesa hladnega radialnega valjanja izdelkov z žlebovi. Članek opisuje razvoj tridimenzionalnih modelov po metodi končnih elementov s pomočjo zakonov odvisnosti napetosti in deformacij, ugotovljenih v tlačnih preizkusih, in optimalne mreže modela obdelovanca za doseganje natančnih rezultatov s sprejemljivim številom elementov in računskim časom.

Proces je bil simuliran s FE-kodo Abaqus/Explicit, validacija rezultatov pa je bila opravljena na osnovi meritev radialnih sil in mikrotrdote.

Vedenje materiala je popisano z dvema zakonoma: nizkohitrostni tlačni preizkus (LST) z zakonom deformacijskega utrjanja, ki ima pet parametrov ter združuje Hollomonov in Vocejev zakon, visokohitrostni tlačni preizkus (HST) pa s stopnjo deformacij iz prej omenjenega zakona in temperaturno občutljivostjo po Johnson-Cookovem zakonu. Obe simulirani radialni sili imata zelo podoben razvoj, vrednost maksimalne sile pa je podobna eksperimentalno določeni sili. Podobnost med radialnimi silami, izračunanimi s pomočjo krivulj odvisnosti napetosti in deformacij za primera LST in HST, je mogoče razložiti z dejstvom, da je zelo velikim deformacijam izpostavljena le tanka površinska plast, večina notranje energije, porabljene med valjanjem, pa je v področjih z zmernimi deformacijami. Razvoj sile se zato ravna predvsem po zakonu odvisnosti napetosti in deformacij pri majhnih in srednje velikih deformacijah.

Optimalna mreža obdelovanca je bila določena z metalografsko analizo: gosta mreža v območju deformacij blizu površine surovca in precej bolj groba mreža v notranjosti surovca. Mreža ima tri področja v aksialni smeri in dve področji v radialni smeri profila. Dimenzije elementov v teh petih področjih so bile določene glede na glavno značilnost profila – korak p med sosednjimi žlebovi.

Raziskave se lahko nadaljujejo z analizo procesa hladnega valjenja za različne materiale in za podobne profile izdelkov.

Opravljene so bile simulacije procesa hladnega valjenja po metodi končnih elementov z optimizirano mrežo in ekstrapolacijo odvisnosti med napetostmi in deformacijami, določenih s tlačnimi preizkusi. Rezultati kažejo, da dajejo simulacije po metodi končnih elementov dragocene informacije za izboljšanje zmogljivosti izdelkov. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: profil žleba, hladno valjanje, mikrotrdota, modeliranje z metodo končnih elementov

SI 126

*Naslov avtorja za dopisovanje: Univerza Piteşti, Str. Târgul din Vale, nr. 1, Piteşti, Romunija, eduard.nitu@upit.ro

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 127

Prejeto: 22.06.2010 Sprejeto: 03.08.2011

Modeliranje topografije površine z namenom zmanjšanja trenja

Sedlaček, M. – Vilhena, L.M.S. – Podgornik, B. – Vižintin, J. Marko Sedlaček* – Luis Miguel Silva Vilhena – Bojan Podgornik – Jože Vižintin Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija Nadzorovanje tornih lastnosti kontaktnih površin postaja vse bolj in bolj pomembno zaradi nenehnih zahtev za izboljšanje zanesljivost in učinkovitost mehanskih delov in predvsem zmanjšanje trenja. V zadnjih letih se je obličenje površin uveljavilo kot tehnika inženiringa površin za zmanjšanje trenja. Zmanjšanje trenja je bila doseženo z izdelavo različnih vzorcev v obliki mikro vdolbinic ali utorov na površini. Sprememba parametrov obličenja, kot so oblika vdolbinic oz. utorov, njihove globine in širine, gostote porazdelitve in usmeritve, ki vplivajo na trenje in obrabo, za področje mejnega mazanja še vedno večinoma temelji na pristopu poskusov in napak. Možna ideja za oblikovanje kontaktnih površin, ki bi se odrazila v manjšem trenju je, da se obličene površine obravnava kot urejeno hrapavost. Poznavanje povezave med parametri hrapavosti in tornim obnašanjem, bi nam omogočila pravilno izbiro parametrov obličenja in s tem ustrezno pripravo površin za določene kontaktne pogoje. Poznavanje korelacije med hrapavostjo površine in trenjem je tako bistvenega pomena za doseganje tega cilja.

Cilj te raziskave je bil prepoznati povezavo med standardnimi parametri hrapavosti in trenjem ter možnosti oblikovanja obličenih kontaktnih površin z manjšim koeficientom trenja z uporabo standardnih parametrov hrapavosti.

V ta namen so bili pripravljeni jekleni ploščati preizkušanci (100Cr6) z različnimi topografijami površin. Z uporabo različnih vrst in kombinacijami brušenja ter poliranja so bili pripravljeni vzorci s podobnimi vrednosti parametra hrapavosti Ra, vendar različnimi vrednostmi Rku in Rsk. Za preučitev vpliva parametrov hrapavosti na trenje in obrabo, so bili izdelani suhi in namazani pin-on-disk testi pri različnih kontaktnih pogojih. Rezultati preizkusov kažejo, da se površine z visoko vrednostjo parametra hrapavosti Rku in negativno vrednostjo Rsk odrazijo v nižjem trenju.

Na podlagi teh izsledkov, smo raziskali vpliv obličenja površin na parametre hrapavosti. To smo naredili virtualno s spreminjanjem realnega profila hrapavosti v obliki različnih velikosti in oblik vdolbinic, in izračunavanjem parametrov hrapavosti površine s pomočjo programske opreme NIST SMATS softgauge. Rezultati so pokazali, da parametri obličenja, ki se odražajo v višjih vrednosti parametrov hrapavosti Rku in bolj negativnih vrednosti Rsk, ter posledično manjšim trenjem, manjša širina vdolbinic, večja razdalja med vdolbinicami in klinast profil vdolbinic.

Prednost pristopa, da obličene površine obravnavamo kot urejeno hrapavost je ta, da s poznavanjem povezave med standardnimi parametri hrapavosti in trenjem, lahko načrtujemo obličenje površin, ki se je za področje mejnega mazanja, do sedaj večino izvajalo po principu preizkusa in napake. ©2011 Strojniški vestnik. vse pravice pridržane.

Keywords: topografija, parametri hrapavosti, trenje, obličenje površin

*Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za strojništvo, Aškerčeva 6, 1000 Ljubljana, Slovenija, marko.sedlacek@ctd.uni-lj.si

SI 127

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 128

Prejeto: 28.09.2010 Sprejeto: 21.06.2011

Vidiki uporabe kota nagiba osi orodja

Sadílek, M. – Čep, R.– Budak, I. – Soković, M. Marek Sadílek1,* – Robert Čep1 – Igor Budak2 – Mirko Soković3 1 VŠB-Tehnična univerza Ostrava, Fakulteta za strojništvo, Republika Češka 2 Univerza v Novem Sadu, Fakulteta za tehnične vede, Srbija 3 Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija Prispevek se ukvarja z raziskavami in predlogi za spremembo položaja osi orodja proti frezani površini med več osnim frezanjem. Cilj raziskav je doseči povečanje učinkovitosti frezanja (izboljšanje funkcionalnih lastnosti površine, povečanje natančnosti frezanja, povečanje obstojnosti orodja, zmanjšanje obremenitev stroja in skrajšanje časa frezanja).

Vnaprej opredeljene možnosti za proces frezanja se uporabljajo pri programiranju več osnih frezalnih strojev v CAM sistemih. Vendar, ali so programerji sposobni popolnoma uporabiti vse razpoložljive možnosti nastavitev? Ali zajamejo vse pomembne informacije potrebne za učinkovito frezanje? Ali ti programerji vedo katere vrednosti so najbolj učinkovite? Ta vprašanja za programerje vodijo do podcenjene in zanemarjene možnosti, da spremenijo položaj osi orodja glede na normalo na frezano površino. Med drugim, ta prispevek skuša najti odgovor na vprašanje in ga znanstveno potrditi.

Rezultat je nato določitev učinkovitega kota oziroma območje nastavitve prostorskega kota, položaj osi orodja glede na frezano površino.

Vse nastavitve parametrov frezanja privedejo do večje učinkovitosti frezanja, a to so: povečanje natančnosti, izboljšanje funkcionalnih lastnosti frezane površine (hrapavost, valovitost, zaostale napetosti, mikrotrdota …), skrajšanje časa frezanja, zmanjšanje obremenitev stroja, ekonomski in ekološki vidiki, itd.

Med standardnim frezanjem z krogelnimi frezali, ko sta obdelovanec in orodje v delovnih položajih, sferični rezalni rob ima rezalno hitrost enako nič v osi orodja. Orodje samo potiska frezani material na tem mestu. Zaradi tega se pojavijo neželeni učinki, kot so: krčenje odrezka, povišanje temperature rezanja, povečanje vibracij in povečana inteziteta oblikovanja »nalepka« na rezalnem robu. Ti pojavi povzročijo poslabšanje kakovosti frezane površine in zmanjšanje obstojnosti orodja. Omenjene pojave je mogoče odpraviti s spremembo položaja osi orodja v razmerju do frezane površine, to je z nagibom orodja ali obdelovanca.

Na osnovi analize poskusov se lahko sklepa, da naklon orodja ima pomemben vpliv na vzdolžno in prečno hrapavost obdelane površine, zaostale napetosti ter velikost in smer posameznih komponent sile rezanja.

Izsledke se lahko uporablja za optimiranje naklona osi orodja, za doseganje boljše kakovosti frezane površine, povečanje obstojnosti orodja in znižanje potrebne energije med frezanjem. Na osnovi izmerjene sile rezanja se lahko naklon osi orodja optimira on-line, kar vodi k adaptivni optimizaciji.

Uporabnost raziskovanega področja je pri frezanju kompleksnih oblik površin brez naknadnega brušenja. Integriteta brušene je pogosto neustrezna, na primer s stališča vpliva toplote in nateznih (zaostalih napetosti) v površinski plasti. Zato opravljene raziskave predlagane tehnologije prispevajo k prizadevanju za odpravo brušenja (npr. končne operacije ročne obdelave kompleksno oblikovanih površin), ali pa vsaj njegovemu zmanjšanju.

Te raziskave gredo v smeri, da bi izdelavo kompleksnih ploskev naredili bolj učinkovito. To velja za proizvodnjo kalupov, orodij za tlačno litje in druge kompleksne izdelke v različnih industrijskih panogah, predvsem v avtomobilski in letalski industriji. ©2011 Strojniški vestnik. Vse pravice pridržane Ključne besede: krogelna frezala, kot nagiba osi orodja, hrapavost površine, zaostale napetosti, sila rezanja, več osno frezanje

SI 128

*Naslov avtorja za dopisovanje: VŠB-Tehnična univerza Ostrava, Fakulteta za strojništvo, tř. 17. listopadu 15, 708 33, Ostrava, Republika Češka, marek.sadilek@vsb.cz

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 129

Prejeto: 14.02.2011 Sprejeto: 08.06.2011

Vpliv vhodnih parametrov na značilnosti procesa EDM

Shabgard, M. – Seyedzavvar, M. – Oliaei, S.N.B. Mohammadreza Shabgard1,* – Mirsadegh Seyedzavvar1 – Samad Nadimi Bavil Oliaei2 1Oddelek za strojništvo, Univerza v Tabrizu, Iran 2Oddelek za strojništvo, Tehnična univerza Bližnjega vzhoda, Turčija V članku so predstavljeni rezultati eksperimentalne študije in podrobne raziskave vpliva vhodnih parametrov elektroerozijske obdelave (EDM) na značilnosti procesa EDM. Med preučevanimi značilnostmi procesa so parametri obdelave, kot so stopnja odvzema materiala, stopnja obrabe orodja in aritmetični srednji odstopek profila, ter debelina bele plasti in globina toplotno vplivanega področja kot značilnosti integritete površine pri obdelovancu iz orodnega jekla AISI H13.

Glavni izhodni parametri procesa elektroerozijske obdelave so stopnja odvzema materiala (MRR), stopnja obrabe orodja (TWR) in površinska hrapavost (Ra) obdelovanca. Zaradi vse strožjih zahtev po zmogljivosti, dolgi življenjski dobi in zanesljivosti komponent pridobiva na pomenu tudi kakovost površin, ustvarjenih z elektroerozijsko obdelavo. Za optimalno izrabo procesa EDM je nujna izbira ustreznega nabora parametrov obdelave, ki zagotavlja največji odvzem materiala pri najmanjši možni obrabi orodja in površinski hrapavosti, kakor tudi minimalno debelino pretaljene plasti in globino toplotno vplivanega področja.

Eksperimenti so bili opravljeni po tovarniškem postopku pri petih nastavitvah jakosti impulznega toka in pri štirih nastavitvah trajanja impulza. Orodni material je bil kovani baker komercialne čistoče, material obdelovanca pa je bilo orodno jeklo AISI H13. Obdelovanec je bil pred elektroerozijsko obdelavo odrezan v valjasto obliko dolžine 20 mm in premera 20 mm. Preizkusi obdelave so bili opravljeni pri petih nastavitvah jakosti impulznega toka in pri štirih nastavitvah trajanja impulza. Vsak preizkus obdelave je trajal 15 minut. Debelina bele plasti (WT) je bila izmerjena v 30 različnih točkah s pomočjo vrstičnega (SEM) elektronskega mikroskopa VEGA\\TESCAN, nato pa je bilo izračunano povprečje. Obdelani preizkušanci so nato bili prerezani v prečni smeri s strojem za žično elektroerozijo in pripravljeni za metalografske preiskave po standardnem postopku. Preizkušanci so bili jedkani s potopitvijo v 5-odstotni reagent Nital. Mikrotrdota po prerezu obdelanih preizkušancev je bila izmerjena z napravo za merjenje mikrotrdote Olympus LM700, s čimer je bila ugotovljena globina toplotno vplivanega področja. Rezultati raziskav kažejo, da povečanje trajanja impulza povzroči povečanje stopnje odvzema materiala in površinske hrapavosti, kakor tudi debeline bele plasti in globine toplotno vplivanega področja. Povečanje trajanja impulza pa povzroči tudi zmanjšanje stopnje obrabe orodja. Povečanje jakosti toka povzroči povečanje stopnje odvzema materiala, površinske hrapavosti in stopnje obrabe orodja. Ob povečanju jakosti toka pa je bilo ugotovljeno tudi rahlo zmanjšanje debeline bele plasti. Z drugimi besedami: pri konstantni energiji razelektritve se z večjo jakostjo toka in krajšim trajanjem impulzov zmanjšata debelina belega sloja in globina toplotno vplivanega področja na površini obdelovanca. Pojav je mogoče pojasniti s tem, da ima učinkovitost plazemskega čiščenja (sposobnost plazemskega kanala za odstranjevanje raztaljenega materiala iz kraterja ob koncu impulza) pomemben vpliv na debelino bele plasti. S podaljšanjem trajanja impulza se zmanjša čistilna zmogljivost plazemskega kanala, s tem pa tudi sposobnost plazemskega kanala za odstranjevanje raztaljenega materiala. Preostali raztaljeni material se v kopeli spet strdi in tvori belo plasti na obdelani površini. Podaljšanje trajanja razelektritev poveča tudi vnos toplote v obdelovanec ob vsaki razelektritvi, vpliv visokih temperatur pa zato seže globlje v obdelovanec. Ta pojav tako povzroči povečanje debeline bele plasti in globine toplotno vplivanega področja. Članek izpolnjuje vrzel v obstoječi literaturi o korelacijah med vhodnimi parametri ter debelino bele plasti (WT) in globino toplotno vplivanega področja (HD) pri elektroerozijski obdelavi orodnega jekla AISI H13. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: elektroerozijska obdelava, stopnja odvzema materiala, stopnja obrabe orodja, površinska hrapavost, debelina bele plasti, globina toplotno vplivanega področja *Naslov avtorja za dopisovanje: Oddelek za strojništvo, Univerza v Tabrizu, Tabriz, Iran, mrshabgard@tabrizu.ac.ir

SI 129

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 130

Prejeto: 23.12.2010 Sprejeto: 17.06.2011

Uporaba numeričnih simulacij pri procesu globokega vleka in pridrževalni sistem s segmentnimi vstavki Volk, M. – Nardin, B. – Dolšak, B. Mihael Volk1,* – Blaž Nardin1 – Bojan Dolšak2 1 Gorenje Orodjarna d.o.o., Slovenija 2 Univerza v Mariboru, Fakulteta za strojništvo, Slovenija

V zadnjem času se stalno povečuje potreba po kompleksnejših izdelkih iz pločevine, kar je še posebej značilno za avtomobilsko industrijo. Poleg tega je trend tudi v uporabi novih, trših materialov, ki pa se težje preoblikujejo. V takih primerih je postala nujna uporaba numeričnih simulacij. S pomočjo numeričnih simulacij se skrajša čas razvoja procesa, proces preoblikovanja pa je tudi bolj obvladljiv. V nekaterih primerih pa tudi to ni dovolj, zato se vedno znova pojavljajo tudi potrebe po novih, naprednejših ter bolj prilagodljivih sistemih. Eden takšnih sistemov je tudi pridrževalni sistem, ki je opisan v tem članku.

Eden najpomembnejših parametrov pri preoblikovanju pločevine je pridržalna sila. Pridržalna sila je še posebej pomembna pri procesu globokega vleka in je včasih celo edini parameter, na katerega lahko občutno vplivamo. S pridržalno silo nadzorujemo tok materiala, ki ima neposreden vpliv na nastanek razpok ter gub, torej na kakovost vlečenih pločevinastih izdelkov.

V članku je zato opisana optimizacija pridržalne sile ter pridržalnega sistema s segmentnimi vstavki.

Optimizacija je bila narejena s pomočjo metode končnih elementov in mehke logike. Takšna kombinacija optimizacije v preteklosti še ni bila narejena in prav zato takšnega pridrževalnega sistema s segmentnimi vstavki ne najdemo nikjer drugje.

Omenjeni segmentni način pridrževanja je v članku primerjan z običajnim postopkom pridrževanja.

Za eksperiment je bil uporabljen industrijski primer iz bele tehnike. Rezultati so pokazali, da je mogoče z uporabo krajevno odvisne pridržalne sile občutno izboljšati kakovost izdelkov. Izdelki, narejeni s segmentnim pridrževanjem, so bili bistveno kakovostnejši in brez napak kot so gube ali razpoke. Rezultati numeričnih simulacij so bili potrjeni tudi z eksperimentom.

Sistem se je izkazal kot izredno prilagodljiv. Posebnost tega sistema je, da so segmentni vstavki vedno v neposrednem stiku s pločevinasto platino, zaradi česar imamo boljši nadzor nad procesom. S tem sistemom je mogoče nadzorovati tok materiala tako krajevno kot tudi časovno, kar z običajnimi sistemi ni mogoče, hkrati pa so rezultati občutno boljši. Tehnološko okno kakovostnih izdelkov je v tem primeru širše.

Za optimalno delovanje pridržalnega sistema potrebujemo poseben sistem hidravličnih blazin stiskalnic, ki pa ga zaenkrat v proizvodnji najdemo le redko. Sistem pa ima veliko prednosti in zato je pričakovati, da se bo v industriji vse bolj uveljavljal.

Analiza v članku je narejena samo s krajevno odvisno pridržalno silo, in sicer za 10 segmentnih vstavkov. Ni pa bila uporabljena časovno odvisna pridržalna sila, ki lahko privede še do bistveno boljših rezultatov. V prihodnje je zato treba več pozornosti posvetiti tudi uporabi časovno odvisne pridržalne sile.

Rezultate, predstavljene v tem delu, je mogoče uporabiti pri razvoju novih orodij za preoblikovanje pločevine. Opisani sistem pridrževanja je uporaben predvsem za zahtevnejše izdelke, ki jih ni mogoče kakovostno izdelati z do sedaj znanimi sistemi pridrževanja. Optimizacija pridržalne sile s kombinirano uporabo numeričnih simulacij in mehke logike lahko občutno skrajša čas optimizacije. ©2011 Strojniški vestnik. Vse pravice pridržane.

Ključne besede: preoblikovanje pločevine, globoki vlek, segmentni sistem pridrževanja, metode končnih elementov (MKE), optimiranje

SI 130

*Naslov avtorja za dopisovanje: Gorenje Orodjarna d.o.o., Partizanska 12, SI-3503 Velenje, Slovenija, mihael.volk@siol.net

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 131-132 Navodila avtojem

Navodila avtorjem Članke pošljite na naslov: Strojniški vestnik Journal of Mechanical Engineering Aškerčeva 6, 1000 Ljubljana, Slovenija Tel.: 00386 1 4771 137 Faks: 00386 1 2518 567 E-mail: info@sv-jme.eu strojniski.vestnik@fs.uni-lj.si 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 oglatih oklepajih navedena tudi enota. V tabelah naj se ne podvajajo podatki, ki se nahajajo v besedilu.

SI 131

Strojniški vestnik - Journal of Mechanical Engineering 57(2011)9, SI 131-132

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.

[1] Zadnik, Ž., Karakašič, M., Kljajin, M., Duhovnik, J. (2009). Function and Functionality in the Conceptual Design Process. Strojniški vestnik – Journal of Mechanical Engineering, vol. 55, no. 7-8, p. 455-471.

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 Industry. MOTSP 2009 Conference Proceedings, p. 422-427.

SI 132

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 3.000 - 3.500 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)9, SI 133-137 Osebne objave

Magisteriji, specialistična dela in diplome MAGISTERIJ ZNANOSTI Na Fakulteti za strojništvo Univerze v Ljubljani je z uspehom zagovarjal svoje magistrsko delo: dne 31. avgusta 2011 Vlado SCHWEIGER z naslovom: »Karakteristike vetrne turbine z nasproti vrtečo izvedbo rotorjev« (mentor: prof. dr. Branko Širok).

SPECIALISTIČNO DELO Na Fakulteti za strojništvo Univerze v Ljubljani so z uspehom zagovarjali svoje specialistično delo: dne 17. avgusta 2011 Ina ČARIĆ z naslovom: »Izboljšanje razumljivosti govora v učilnicah« (mentor: prof. dr. Mirko Čudina, somentor: doc.dr. Jurij Prezelj).

DIPLOMIRALI SO Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 24. avgusta 2011: Eva OBLAK z naslovom: »Določanje nastanka mejnih filmov z meritvami na nano in makro skali« (mentor: prof. dr. Mitjan Kalin); Salih TUBEISHAT z naslovom: »Inženirska podpora pri investicijskih objektih« (mentor: prof. dr. Jožef Duhovnik); Mitja VARL z naslovom: »Računalniško podprto konstruiranje velikih električnih transformatorjev« (mentor: izr. prof. dr. Jože Tavčar somentor: prof. dr. Jožef Duhovnik); Maša ZALAZNIK z naslovom: »Tribološke lastnosti materialov za medicinske implantante« (mentor: prof. dr. Mitjan Kalin); dne 26. avgusta 2011: Mitja FRANKO z naslovom: »Razvoj aktuatorjev za sistem kontrole zdrsa za simulator vožnje v razmerah mejnega zdrsa« (mentor: doc. dr. Jernej Klemenc somentor: prof. dr. Matija Fajdiga); Jure GRUDNIK z naslovom: »Naslov: Izdelava poenostavljenega modela končnih elementov za sestav pnevmatike s platiščem za

uporabo pri simulacijah trkov vozil« (mentor: doc. dr. Jernej Klemenc, somentor: prof. dr. Matija Fajdiga); Kristjan KREBELJ z naslovom: »Dinamika sistema togih teles z uporabo Poissonovega zakona trka s trenjem v okviru razširjene Lagrangeeve metode« (mentor: prof. dr.

Miha Boltežar, somentor: doc. dr. Janko Slavič); Tijan MEDE z naslovom: »Numerična analiza dinamskih efektov pri impulzni obremenitvi jeklene pločevine« (mentor: prof. dr.

Miha Boltežar, somentor: doc. dr. Janko Slavič); Matjaž MRŠNIK z naslovom: »Širokospektralne vibracijske obremenitve prožnih struktur« (mentor: prof. dr. Miha Boltežar, somentor: doc. dr. Janko Slavič); dne 31. avgusta 2011: Tomaž GROŠIČAR z naslovom: »Rezanje mehkih materialov z žico« (mentor: doc. dr. Henri Orbanić, somentor: prof. dr. Mihael Junkar); Martin PETKOVŠEK z naslovom: »Eksperimentalno modeliranje kavitacijske erozije« (mentor: prof. dr. Branko Širok, somentor: doc. dr. Matevž Dular); Peter ŠTEBLAJ z naslovom: »Fleksibilni sistem za kontrolo geometrije izdelkov v realnem času« (mentor: prof. dr. Alojzij Sluga); Blaž ŠTEFE z naslovom: »Analiza in optimizacija oblikovnih parametrov Francisove turbine« (mentor: prof. dr. Branko Širok); Miha TRDIČ z naslovom: »Prepoznavanje strukture gibanja trdnih delcev v transportnem sistemu sipkega materiala« (mentor: prof. dr.

Branko Širok). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 12. avgusta 2011: Klavdija KARLOVČEC z naslovom: »Trirazsežno modeliranje ciljno zasnovanih kolekcij nakita« (mentor: izr. prof. Vojmir Pogačar); Viki PAVLIČ z naslovom: »Koncept orodja za obrezovanje robov travnatih površin« (mentor: izr. prof. Vojmir Pogačar); dne 25. avgusta 2011: SI 133

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Tomaž IRGOLIČ z naslovom: »Model vodenja operativnega procesa v podjetju Var d.o.o.« (mentor: doc. dr. Marjan Leber, somentor: izr. prof. dr. Borut Buchmeister); Aljaž KOVAČIČ z naslovom: »Karakterizacija mehanskih lastnosti poroznega gradiva z vzdolžnimi porami s parametričnimi računalniškimi simulacijami« (mentor: prof. dr.

Zoran Ren, somentor: doc. dr. Matej Vesenjak); Andrej LIKEB z naslovom: »Vpliv termične obdelave na lomno žilavost jekla« (mentor: prof. dr. Nenad Gubeljak, somentor: doc. dr. Jožef Predan); Matej PAULIČ z naslovom: »Optimizacija postopka obdelave konzole menjalnika« (mentor: prof. dr. Franci Čuš); Marko ŠORI z naslovom: »Dimenzioniranje polžnega gonila za pogon sledilnika sonca« (mentor: prof. dr. Srečko Glodež); * Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva: dne 24. avgusta 2011: Matej CVETKO z naslovom: »Izdelava programa teoretičnega in praktičnega usposabljanja za letenje s sistemom za nočno opazovanje – NVG« (mentor: pred. mag. Primož Škufca, somentor: izr. prof. dr. Tadej Kosel); Matija FOJKAR z naslovom: »Načrtovanje letališč, teorija in praksa na primeru Letališča Jožeta Pučnika Ljubljana« (mentor: pred. mag.

Andrej Grebenšek, somentor: izr. prof. dr. Tadej Kosel); Klemen POVŠE z naslovom: »Zasnova in preračun ultralahkega jadralnega letala« (mentor: izr. prof. dr. Tadej Kosel); Simona TOBIAS z naslovom: »Ugrabitve letal in varovanje letalskega prevoza« (mentor: viš. pred. mag. Aleksander Čičerov, somentor: izr. prof. dr. Tadej Kosel); Miha TRILER z naslovom: »Zamenjava transportnih helikopterjev srednjega razreda na podlagi temeljnega razvojnega programa Letalske policijske enote« (mentor: pred. mag. Primož Škufca, somentor: izr. prof. dr. Tadej Kosel); Jan VIHER z naslovom: »Priročnik za izpraševalce praktične in strokovne usposobljenosti pilotov, ki izhaja iz zahtev JAR SI 134

FCL 1 in 2« (mentor: pred. mag. Primož Škufca, somentor: izr. prof. dr. Tadej Kosel); dne 26. avgusta 2011: Jure BROVČ z naslovom: »Uvajanje predgrevanja v tehnologijo stiskanja s steklenimi vlakni ojačane poliesterske smole« (mentor: izr. prof. dr. Zlatko Kampuš); Rok DOLINAR z naslovom: »Izboljšanje prenosa toplote pri vrenju z uporabo nanofluidov« (mentor: prof. dr. Iztok Golobič).

Rok JUDEŽ z naslovom: »Obratovalni parametri procesnih sistemov v farmacevtski industriji« (mentor: prof. dr. Iztok Golobič); Andrej KASTELIC z naslovom: »Merjenje toplotne karakteristike ploščnega aluminijastega prenosnika toplote« (mentor: prof. dr. Iztok Golobič); Gašper POTOKAR z naslovom: »Univerzalna miza za varilnik korit« (mentor: prof. dr. Marko Nagode); dne 29. avgusta 2011: Anže ABUNAR z naslovom: »Zaščita vodnega curka pri Peltonovi turbini« (mentor: prof. dr. Branko Širok); Andrej PANGERŠIČ z naslovom: »Tipizacija mehanskega dela modularnih tiskanih vezij« (mentor: prof. dr. Jožef Duhovnik); Jan RUDOLF z naslovom: »Vpliv vgradnje zaznaval pri diagnostiki kavitacije v turbinskih strojih« (mentor: prof. dr. Branko Širok, somentor: izr. prof. dr. Marko Hočevar); Rastko DEBEVEC z naslovom: »Varnost civilnega letalstva - analiza uredbe (ES) št.

216/2008 in Konvencije o mednarodnem civilnem letalstvu« (mentor: viš. pred. mag. Aleksander Čičerov, somentor: izr. prof. dr. Tadej Kosel); Rok KRIVEC z naslovom: »Razkislinjenje starega papirja s pomočjo vakuumskega sušenja« (mentor: prof. dr. Iztok Golobič); Samo MAŽGON z naslovom: »Izboljšan konvektivni prenos toplote pri LED svetilkah« (mentor: prof. dr. Iztok Golobič); Peter VINTAR z naslovom: »Merjenje temperature površine pri generiranju toplote v triboloških kontaktih« (mentor: prof. dr. Mitjan Kalin, somentor: prof. dr. Iztok Golobič); Tine ZAJŠEK z naslovom: »Prevoz nevarnega tovora v luči mednarodnega letalskega prava« (mentor: viš. pred. mag. Aleksander Čičerov, somentor: izr. prof. dr. Tadej Kosel); Matej ZORKO z naslovom: »Toplotne izgube ploščnega prenosnika toplote« (mentor: prof. dr. Iztok Golobič);

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dne 30. avgusta 2011: Matej AVBAR z naslovom: »Eksperimentalno ovrednotenje cevnega kondenzatorja toplotne črpalke zrak - voda« (mentor: prof. dr. Iztok Golobič); Sandi GROBELNIK z naslovom: »Pretakanje hidravličnih kapljevin skozi zaslonke in dušilke« (mentor: doc. dr. Jožef Pezdirnik); Janez KIDRIČ z naslovom: »Izdelava prototipa avtomatskega in energetsko neodvisnega regulacijskega ventila za krmiljenje pretoka deponijskega plina « (mentor: doc. dr. Primož Podržaj); Andrej LUKŠIČ z naslovom: »Kritična gostota toplotnega toka pri nasičenem in podhlajenem mehurčkastem vrenju vode« (mentor: prof. dr. Iztok Golobič); Marko MEDEN z naslovom: »Razvoj novega krmiljenega sedežnega hidravličnega ventila« (mentor: doc. dr. Jožef Pezdirnik); Andrej POTOČNIK z naslovom: »Frekvenčni valj in hidravlični blok za preskuševališče zvezno delujočih ventilov« (mentor: doc. dr. Jožef Pezdirnik); Izidor RESMAN z naslovom: »Eksperimentalno določanje toplotnih mostov v fasadnem panelu« (mentor: prof. dr. Iztok Golobič); dne 31. avgusta 2011: Rok JUSTIN z naslovom: »Industrializacija plavajoče zaščitne zavese iz novih materialov« (mentor: prof. dr. Mirko Soković);

Andrej LEVAK z naslovom: »FMEA analiza v proizvodnji vijakov« (mentor: prof. dr.

Mirko Soković); Jernej LUKANČIČ z naslovom: »Zamenjava stružnice z novo CNC stružnico« (mentor: prof. dr. Janez Kopač, somentor: doc. dr.

Davorin Kramar); Kristijan MARC z naslovom: »Analiza uvedbe tehnologije konturnega rezanja v podjetje za izdelavo kovinskega pohištva« (mentor: doc. dr. Henri Orbanić, somentor: prof. dr. Mihael Junkar); Bogdan VRLINIČ z naslovom: »Vzpostavitev vhodne kontrole v srednje velikem podjetju« (mentor: prof. dr. Mirko Soković); * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva: dne 25. avgusta 2011: Jernej KNEZ z naslovom: »Modeliranje in osnovni preračun rotacijske varilne mize« (mentor: izr. prof. dr. Bojan Dolšak, somentor: doc. dr. Aleš Belšak); Gašper MIKEK z naslovom: »Testna naprava za preizkušanje obdelovalnih glav EMAG« (mentor: doc. dr. Darko Lovrec, somentor: doc. dr. Samo Ulaga); Toni STANOJEVIĆ z naslovom: »Razvoj orodja za element armaturne plošče« (mentor: izr. prof. dr. Stanislav Pehan).

SI 135

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Prof. dr. Miran Oprešnik 1925 - 2011

Dne 27. 7. 2011 nas je zapustil prof. dr.

Miran Oprešnik, upokojeni redni profesor na Fakulteti za strojništvo Univerze v Ljubljani.

Rojen je bil 9. 7. 1925 v Ljubljani, kjer je tudi dokončal osnovno šolo. Njegova mladost ni bila lahka. V času šolanja na gimanziji je med šolskimi počitnicami pomagal svoji družini pri delu v majhni obrtni delavnici. S prisluženim denarjem si je vsako jesen kupoval knjige, ki jih je nujno potreboval pri šolanju. Okupacija ga je zatekla v šestem razredu gimnazije. Ker so stanovali na periferiji Ljubljane, zaradi zastražene bodeče žice, s katero je bilo mesto obdano, ni mogel več hoditi v šolo. Zaradi simpatiziranja z OF so ga zaprli v Ljubljani, pozneje pa internirali v Italijo. Po vrnitvi iz internacije 1943. leta je opravil izpite za 7. in 8. razred gimnazije in veliko maturo. Po osvoboditvi je študiral v Ljubljani in l. 1953 diplomiral na Fakulteti za strojništvo.

Kot odličen student je že v času dodiplomskega študija postal pomožni asistent za predmet Termodinamika. To delo je opravljal vse do diplome, z vmesno prekinitvijo, ko je bil na strokovni praksi v Franciji in Nemčiji. Leta 1954 je postal redni asistent na Katedri za teoretično strojeslovje in toploto pri prof. Rantu. V tem času je krepko pomagal pri oranju ledine na področju eksergijske analize termodinamskih procesov, kot so n.pr. tehnološko uparjanje, kristalizacija, kalcinacija in parni delovni procesi. Uvajanje novih računskih metod v tehniško prakso je bilo zahtevno delo. Kot je to v tehniški praksi običajno, potreben vložek ur izjemno naraste pri prehodu od zasnove do realizacije. Pri vseh aplikacijah so bili potrebni podatki o entropijah snovi, ki so bile podvržene preobrazbam. Ti podatki so bili takrat v svetu zelo pičli. Tako skoraj ni bilo tehniško uporabnih podatkov o entropijah vodnih raztopin SI 136

raznih soli. Prof. Oprešnik je izdelal diagrame za dve pomembni raztopini: za raztopino kuhinjske soli in za raztopino natrijevega sulfata. Izračun je zahteval cel niz integracij kompliciranih funkcij.

Oba diagrama sta bila publicirana v okviru posebne zbirke termodinamičnih diagramov pri založbi Steinkopf in v svetovno znanem tabelarnem delu Landolf-Börnstein. Prof. Oprešnik je kot prvi konstruiral tudi diagram za živo srebro s pomočjo koordinatne transormacije, ki omogoča direkne informacije o vrsti termodinamskih veličin v povezavi z eksergijo. Prav tako je originalno obdelal vprašanja optimalnih parametrov sveže vodne pare v parnih delovnih procesih. Leta 1960 je bil izvoljen za docenta in, po odhodu prof. Ranta v Nemčijo, leta 1962 prevzel vse obveznosti v zvezi s termodinamiko na Fakulteti za strojništvo v Ljubljani. Leta 1969 je bil izvoljen v izrednega profesorja za Termodinamiko. V tem času se je osredotočil na kvalitativno ocenjevanje procesov. Na tem področju je leta 1968 tudi doktoriral s tezo: Splošne in posebne zakonitosti pri kvantitativni in kvalitativni oceni toplotnih črpalk. Sama tematika je z vidika varčevanja z energijo še kako aktualna tudi v današnjem času.

V letu 1971 je prof. Oprešnik koncipiral nov predmet Termodinamika II, ki je zajemal zmesi in raztopine. V l. 1972 je pričel s predavanjem Procesne tehnike. Leta 1980 je bil izvoljen za rednega profesorja za področje Termodinamika in Procesna tehnika. V tem zadnjem obdobju svojega raziskovalnega dela se je prof. Oprešnik osredotočil na analize s področja procesne tehnike. Vsebina znanstvenih in strokovnih sestavkov objavljenih v domačih in tujih revijah je večinoma vezana na probleme toplotnih črpalk in hladilnih naprav, ocenjenih s termodinamskega vidika. Obravnavani so eksergijski izkoristki pri

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termokompresiji in hladilnih in grelnih procesih, eksergijske izgube hladilnih naprav, eksergijske izgube toplotnih črpalk ter eksergijske izgube s posebnim poudarkom na delovnem sredstvu, temperaturi in na razmerah prestopa toplote. Z eksergijsko oceno procesov in naprav v določenih pogojih njihovega delovanja je prof. Oprešnik nakazoval ukrepe za zmanjšanje izgub, s čimer je posredno še pred “energetsko krizo” opozarjal na potrebnost smotrnega gospodarjenja z energijo. V tem pogledu njegova dela še danes predstavljajo zelo aktualne prispevke k znanosti.

Kljub izčrpujočemu urniku na visokošolskem in višješolskem rednem študiju, študiju ob delu in dislociranih centrih v Kopru in Novi Gorici, je prof. Oprešnik vedno našel čas za delo s svojimi študenti. Pri njem je diplomiralo več kot 70 inženirjev, bil pa je tudi mentor 18 magistrom oz. doktorjem tehniških znanosti. Za potrebe poučevanja in tudi za širše potrebe je izdelal celo vrsto učnih pripomočkov in učbenikov: Sto nalog iz termodinamike (1957) Termodinamične tabele (1958), Osnove termodinamike (1961), Termodinamične tabele in diagrami (1962, 1966, 1970, v soavtorstvu z M. Oparo: 1977, 1983, 1987), Naloge in rešitve iz termodinamike (1967, 1968,), Termodinamika (1974, 1978, 1983, 1987), Termodinamika zmesi (1974, 1978, 1988). Ogromno število knjig in dopolnjenih izdaj kaže na obsežno delo, ki ga je prof. Oprešnik prispeval v slovenskem prostoru. Učbeniki so prilagojeni potrebam študija strojništva z obilico grafičnega gradiva in vsebujejo avtorjeve lastne izsledke. S tem pa se njegov prispevek k pedagoškemu delu ne zaključuje. Pri organizacijskem delu je vse od 1957. leta do upokojitve sodeloval v številnih odborih in komisijah fakultete. Kdor je zasledoval razvoj študijskih zamisli, ki se porodijo pri vsakem spreminjanju študijskega programa bo razumel, koliko ur je potrebnih za izdelavo številnih predlog študijskih načrtov. Poleg fakultetnih

zadolžitev je vedno požrtvovalno sprejemal tudi zadolžitve na univerzi in v širšem prostoru. Med drugim je bil član Univerzitetne študijske komisije in pet let tudi njen predsednik, član Univerzitetne podiplomske komisije, član Komisije za študijska vprašanja pri Skupnosti jugoslovanskih univerz in dve leti tudi predsednik te skupnosti. Še bi lahko naštevali. Med drugim je bil tudi član uredništva Strojniškega vestnika, revije, kjer mu lahko danes zapišemo samo še te besede v zahvalo.

Prof. Oprešnik je imel izjemen posluh za sodelavce in študente, še posebej se je zavzemal za rešitev problemov, če je nekdo zašel v težave ali v težko obdobje pri zahtevnem študiju. Imel je visoko razvit čut do humanega ravnanja in se z vsemi močmi boril za njegovo uveljavitev. Imel je zelo močno voljo, a tega na zunaj ni želel kazati ali uveljavljati v osebno korist. Spominjam se vožnje, ko sva se peljala iz Kopra v Ljubljano.

Takrat mi je predal predavanja v dislociranem centru. Na vrhniškem klancu mi je zaupal, da si želi, da bi se enkrat lahko udeležil avtomobilske dirke v Semmeringu. Želja sesti za krmilo dirkalnika se mu ni izpolnila. Krmilo svojega življenja pa je držal z njemu lastno preciznostjo in nepopustljivostjo v svojih rokah. Vrsto let pred upokojitvijo mi je večkrat dejal, da se bo po upokojitvi posvetil izključno svoji družini.

Tako se je tudi zgodilo. 28. februarja 1991 mi je predal ključe svojega kabineta in ni več prestopil praga fakultete. Ostali so le še privatni razgovori ob naključnih srečanjih. Na fakulteti pa je ostalo še nekaj. Ostala je njegova beseda v učbeniku Termodinamika. Ko so zaloge pričele pohajati, smo v Laboratoriju za Dinamiko fluidov in termodinamiko odkupili 200 izvodov, ki jih vsako leto brezplačno posojamo študentom v pomoč pri študiju. Tako jim profesor Oprešnik s svojim minulim delom lajša življenjske izdatke, ki njemu že v rani mladosti niso bili prizanešeni.

Prof. dr. Iztok Žun

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