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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 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
55 YE ARS
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no. 9 2010 56
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Cover: The photos show the development of shock waves (bright circular contours) generated in the air above the water surface (dark band in the bottom of the image) irradiated with Er:YAG laser pulses. The dark shape in the upper part of the images is the handpiece of a Er:YAG laser for dental applications. The photos are taken using a novel double exposure shadowgraph method. On each photo, the shock wave is captured at two time instances delayed by 1.56 μs. Recorded in collaboration between: UL-FS COLA and Fotona d.d.
© 2010 Strojniški vestnik - Journal of Mechanical Engineering. All rights reserved. SV-JME is indexed / abstracted in: SCI-Expanded, Compendex, Inspec, ProQuest-CSA, SCOPUS, TEMA. The list of the remaining bases, in which SV-JME is indexed, is available on the website. The journal is subsidized by Slovenian Book Agency.
President of Publishing Council Jože Duhovnik UL, Faculty of Mechanical Engineering, Slovenia 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 Print LITTERA PICTA d.o.o., Barletova 4, 1215 Medvode, Slovenia General information Strojniški vestnik – The Journal of Mechanical Engineering is published in 11 issues per year (July and August is a double issue). Institutional prices include print & online access: institutional subscription price €100,00, general public subscription €25,00, student subscription €10,00, foreign subscription €100,00 per year. The price of a single issue is €5,00. Prices are exclusive of tax. Delivery is included in the price. The recipient is responsible for paying any import duties or taxes. Legal title passes to the customer on dispatch by our distributor. Single issues from current and recent volumes are available at the current single-issue price. To order the journal, please complete the form on our website. For submissions, subscriptions and all other information please visit: http://en.sv-jme.eu/ You can advertise on the inner and outer side of the back cover of the magazine. We would like to thank the reviewers who have taken part in the peer-review process.
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9 Contents
Contents Strojniški vestnik - Journal of Mechanical Engineering volume 56, (2010), number 9 Ljubljana, September 2010 ISSN 0039-2480 Published monthly
Papers Aljaž Osterman, Matevž Dular, Marko Hočevar, Brane Širok: Infrared Thermography of Cavitation Thermal Effects in Water Zlatko Petrović, Slobodan Stupar, Ivan Kostić, Aleksandar Simonović: Determination of a Light Helicopter Flight Performance at the Preliminary Design Stage Dušan Mežnar, Momir Lazovič: The Strength of the Bus Structure with the Determination of Critical Points Antonios Kyriazopoulos, Ilias Stavrakas, Konstantinos Ninos, Cimon Anastasiadis, Dimos Triantis: Pressure Stimulated Current Emissions on Cement Paste Samples under Repetitive Stepwise Compressional Loadings Igor Solodov, Daniel Döring, Gerd Busse: Air-Coupled Lamb and Rayleigh Waves for Remote NDE of Defects and Material Elastic Properties Dejan D. Ivezić, Trajko B. Petrović: Robust IMC Controllers with Optimal Setpoints Tracking and Disturbance Rejection for Industrial Boiler Mitja Košir, Aleš Krainer, Mateja Dovjak, Rudolf Perdan, Živa Kristl: Alternative to the Conventional Heating and Cooling Systems in Public Buildings Hasan Gökkaya: The Effects of Machining Parameters on Cutting Forces, Surface Roughness, Built-Up Edge (BUE) and Built-Up Layer (BUL) During Machining AA2014 (T4) Alloy
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Paper received: 18.11.2009 Paper accepted: 01.09.2010
Infrared Thermography of Cavitation Thermal Effects in Water Aljaž Osterman* - Matevž Dular - Marko Hočevar - Brane Širok University of Ljubljana, Faculty of Mechanical Engineering, Slovenia Although the thermal effects of cavitation are believed to be negligible for cavitation in water, they were successfully experimentally measured using infrared thermography. Cavitation was generated in a small container holding about 500 ml of water. It was oscillated with ultrasonic frequencies of 42 kHz to trigger growth and collapse of bubbles. For the temperature measurements a high-speed thermovision camera was used. It captures light in infrared spectrum with wavelength of 3 to 5 μm. The frequency of temperature field acquisition was set to 600 Hz. A silicon glass, which is transparent in the infrared light spectrum, was attached to a cylinder and partially submerged into water. Bubbles, which tend to appear in the vicinity of solid surface, appeared on the submerged side of the glass. The visual path for the thermovision camera was: air – silicon glass – water. In this way, the temperatures on the submerged side of the silicon glass where bubble growth and implosions occur could be measured. With the applied thermographic method small but distinctive local decreases of temperature (with magnitudes up to 0.3 K), caused by cavitation, were detected. © 2010 Journal of Mechanical Engineering. All rights reserved. Keywords: cavitation, ultrasound, temperature, IR thermography, bubbles 0 INTRODUCTION Cavitation phenomenon, characterized by vapor generation and condensation, frequently occurs in industrial fluid flows. Often it is accompanied by effects like vibration, increase of hydrodynamic drag, changes in the flow hydrodynamics, noise, erosion, light effects such as sonoluminescence as well as thermal effects. They are often neglected since cavitation is, generally speaking (on a large scale), an isothermal phenomenon [1]. However, on a small scale (a single bubble), significant thermal effects such as a considerably high rise in temperature when the bubble collapses and local cooling of the bubble surrounding at its growth, take place. These phenomena are almost always in equilibrium (if energy for cavitation erosion, sonoluminiscence etc. is neglected). This means that the increase of the temperature must be balanced by its prior or later decrease. In this way, a quasi-isothermal nature of the cavitation is preserved. As the pressure is locally decreased, conditions for evaporation (growth of the cavitation bubble) become plausible. To initiate the growth of cavitation bubble another condition is required: the presence of a cavitation nucleus in the region, represented by gas (air), trapped gas in
wall fissures or dissolved in the fluid. From the incipient cavitation on, the nuclei are produced also by the bubble collapses. When both conditions are fulfilled, the bubble begins to grow and latent heat is sucked from the surrounding liquid, creating a thermal boundary layer. The result is a small local decrease of liquid temperature that is usually called “thermal delay” [1]. This temperature drop due to cavitation growth results in lower vapor pressure and is retarding cavitation growth because greater pressure drop is needed. After the bubble reaches its maximum size, it collapses – the temperature (and also pressure) inside it increases very rapidly and reaches extreme values (e.g. 6700 K and 848 bar [2]). Thermal delay can usually be neglected when dealing with fluids for which critical point temperature is much higher than the working temperature. In that case, heat gained during the bubble collapse is in quasi-equilibrium with the heat used for the bubble growth and the isothermal condition is fulfilled. In this case of equilibrium, the data from the study of thermal delay could be used for determining the conditions in the bubble at its collapse. On the other hand, the thermal delay becomes significant when the critical point temperature and fluid
*Corr. Author's Address: Faculty of Mechanical Engineering, Aškerčeva 6, SI – 1000 Ljubljana, Slovenia aljaz.osterman@fs.uni-lj.si
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working temperature lie close together (for example, in the case of cryogenic fluids [3]). Cavitation can be generated by subjecting a liquid to an oscillating pressure field of a particular (usually ultrasonic) frequency. The local decreases and increases of pressure are caused by the inability of a fluid (due to inertia) to follow the oscillations of a solid wall. If the amplitude and the frequency of oscillations are adequate and a cavitation nucleus is present in the region, a bubble will appear and collapse in such a region. Therefore, on one hand it is a very fast phenomenon while on the other, it happens on a very small spatial scale. Bearing in mind such a specific nature of cavitation, it is clear that an experimental approach on a local level is very difficult and not many experimental methods are available. For all types of cavitation it is also important that it reflects certain conditions of the flow so that the experimental method used does not alter the flow. Therefore, it is understandable that only few experimental data of cavitation thermal effects are available. This applies especially to the temperature measurements [4] and [5]. Some estimations of the influence of the thermal effect were done by comparing cavitation structures, based on visualization and pressure measurements [6] to [8]. In this paper a new experimental approach to cavitation is presented. A non-contact and noninvasive thermographic method was used to measure cavitation-dependent temperature fields. Temperature fields were measured because temperature plays such an important role in cavitation. Measured cavitation thermal effects were in form of local appearance of colder areas. They were connected to thermal delay but due to the difference between temporal scales of the observed phenomena and a collapse time of a single cavitation bubble [1], a reconstruction of the bubble collapse was not possible. Cavitation was generated in water using an ultrasonic cleaning device with an aim to bridge the gap between a common isothermal approach on a system scale and extreme local thermal phenomena driven by collapses. The aim was also to use the ultrasonic cleaning device as an example of how some cavitation effects have been favorably exploited. It is believed that the introduction of the thermographic method for
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cavitation measurement and the obtained results can contribute to a better understanding of the cavitation phenomenon, thus improving many practical applications where cavitation is used [9]. 1 EXPERIMENTAL SET UP The experiment was performed under atmospheric pressure (approximately 1 bar) and the water temperature was close to the ambient (23 °C). The whole experimental set-up is presented in Fig. 1.
Fig. 1. Experimental set-up A small vessel made out of stainless steel with a capacity of approximately 500 ml (dimensions are 140 × 70 × 60 mm) was used. Beneath it an oscilator that produced periodic oscillations at a frequency of about 42 kHz, was mounted. The ultrasonic frequency, causing water tank oscillations, was verified by measurements with the capacitative hydrophone Bruël & Kjær type 8103. Hydrophone was connected to the charge amplifier B&K type 2635. Data were sampled to a PC at 100 kHz which was below the upper frequency limits for the hydrophone (180 kHz) and the charge amplifier (200 kHz). Measurements have been taken during the several hours of operation and in that period short sequences of data were recorded. It was confirmed that the main frequency, causing the cavitation phenomena, was 42 kHz, to which a network frequency of 50 Hz was added. As can be seen in Fig. 3, the spectral power of the 42 kHz signal was significantly higher than that of
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the 50 Hz signal. This means that the latter practically did not influence the results. For an acquisition of the temperature field a high-speed thermovision camera CMT384SM Thermosensorik was used. Its sensitive wavelength range lies between 3 and 5 m. Temperature fields were taken at the water – observation window interface because in the measured spectrum water is not transparent to IR [10], while for the observation window IR transparent silicon with antireflective (AR) coating was used [11]. The acquisition frequency of the camera was set to 600 frames per second where the best results (in respect to the highest contrast and the lowest noise) were achieved. Frame size was 128 × 128 pixels and the integration time was 500 s. The thermocamera had constant focal length and a pixel size of 55 m was determined from a known geometry. The high-speed thermovision camera was calibrated in the environment as used for the experiment. The constraint was constant and uniform temperature distribution which was achieved by mixing during long periods of time allowed for the system to reach certain temperature ranges at which the camera was calibrated. In this way, at each calibration temperature a referential temperature field was measured. Uncertainty of mean water operating temperature was conditional on the measurements with A-class Pt100 sensor and was 0.2 K. However, our main goal was to quantify relative differences in non-uniform time-dependant temperature field. For a single element on the temperature sensor of the thermocamera (representing one pixel) the uncertainty of temperature changes relative to the operating temperature was only 0.03 K. A round silicon window 25 mm in diameter and 1 mm thick was used. The glass was coated to additionally ensure the desired absorption and reflection properties. The transmission of infrared radiation through the glass was measured and it was confirmed that the loss is minimal. It was attached to a cylinder and partially submerged into water. Bubbles, which tend to appear near the solid surface, appeared on the submerged side of the glass. The visual path for the IR camera was: air – silicon glass – water-glass interface. This way the temperatures on the submerged (wetted) side
of the silicon glass, where bubbles grow and implode, could be measured. Due to its high absorption, the water itself is not transparent in the wavelength range between 3 to 5 m – this was experimentally verified as 20 m thick film of water let trough no IR light. In a visible spectrum observations through a glass window in a side wall of the ultrasonic cleaning device were done with high-speed CCD camera MotionBLITZ Cube (Mikrotron GmbH). Acquisition rate was 1000 fps for a frame size of 1024 × 1280 pixels. The focus of the camera was on the lower (wetted) surface of the AR coated silicon window. The purpose was to monitor the appearance and behavior of bubbles in a region near the interface observed by the thermocamera. 2 RESULTS AND DISCUSSION A strong tendency of bubbles to grow and collapse in the regions with an uneven surface and sharp edges was found (for example, at the contact point between the silicon window and the seal of the cylinder). This was expected since rough and uneven surfaces act as a cavitation nuclei generators, which stimulate bubble growth. Bubbles also appeared in the central part of the IR observation window, but they were quite dispersed. It is expected that for bubbles so distant (as shown in Fig. 2), their mutual interaction is very weak. In this view, the greatest effect on a collapse of the cavitation bubble is expected because of the presence of a solid wall, so that an impinging microjet forms [12] to [14].
Fig. 2. Conventional image of several cavitation bubbles on the silicon glass From the images in the visible spectrum also a size of the bubbles was estimated. The average size was 0.084 mm. The agreement between the observed size and a theoretically
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predicted size was 93%. The difference is mostly due to visualization where the exact size of moving bubbles was difficult to determine. The theoretical size is a size of a bubble that has the same eigenfrequency as is the frequency of the ultrasound we used. A detailed relation is described with the following Eq. [15]: 1 3n ( p p v ) 2(3n 1) S 8 l2 , 4 (1) 2 l R2 l R3 R where feig is eigenfrequency of the bubble, n polytrophic constant, pv vapor pressure, p∞ referential pressure, R bubble radius, S surface tension, νl kinematic viscosity of a liquid and ρl liquid density. For example, for water at 20 °C and frequency of 42 kHz the bubble radius is 0.078 mm. The frequency of the ultrasound was obtained from hydrophone measurements of pressure oscillations in water during the ultrasonic cleaning device operation. From a time signal a frequency spectrum was obtained by using discrete Fourier transformation. The spectrum is presented in Fig. 3, in which a distinctive peak that lies at 42 kHz can be identified (Nyquist frequency is 50 kHz).
Fig. 4 shows a sequence of images captured with the IR camera. A darker color on the images represents the colder region. An area of 6 × 6 mm is presented in the sequence. A time step between the images is 2 ms.
f eig
Fig. 4. Sequence of images showing the evolution of the temperature field Spreading of a cold front, which resembles the shape of a cloud, is observed. The first frame shows the very beginning of the event, where the cloud has not been formed yet and only some small temperature deviations from the mean value, which might be related to the internal noise of the IR camera sensor, can be seen. In the next time frame, a lower temperature region appears. Its shape is random at the beginning but evolves into circular form later on. Towards the end of the sequence, the cold front propagation velocity decreases and eventually stops. After that, the cold front shape deviates from the circular form and again becomes random. The apparently warmer region that can be seen in the bottom right corner of all the images in Fig. 4, is a result of different emission coefficients of the cylinder that holds the silicon window and should not be interpreted as an actual temperature change.
Fig. 3. Frequency spectrum from hydrophone measurements 2.1 Temperature Measurements
Fields
from
IR
With the thermocamera the temperature fields of the water-silicon window interface were measured and some distinctive phenomena of temperature changes related to cavitation were detected. When there was no cavitation, no phenomena were recorded.
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Fig. 5. Evolution of temperature with time in a single point The temperature inside the cold front appears to be almost uniform. Small temperature
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gradients, which appear randomly at the end of a sequence, can be related to mixing currents inside the vessel caused by cavitation. Fig. 5 shows the temperature drop in a single point in the middle of the “cold cloud” from the sequence in Fig. 4. The evolution of the temperature is similar all over the cloud (apart from the shift in time) since its temperature is almost uniform. Initially (at t = 0), the temperature is equal to the ambient temperature (T-T0 = 0 K). At a certain point in time, a rapid drop in temperature can be seen. The temperature decreases from the initial temperature by approximately 0.35 K (T-T0 = -0.35 K) in a period of about 10 ms. The temperature then exponentially rises (the heat flows from slightly warmer silicon glass). The last part of the exponential temperature evolution is gently sloping and is left out of the diagram for a clearer presentation of the central part of the drop. Eventually the temperature reaches the initial water temperature. Small oscillations (T-T0 = 0.02 K) of temperature are the result of noise, which has not been filtered in order to preserve the complete signal.
which the fit does not cover. The rising temperature development is similar to the response of a system on a step change in a temperature. A step change could be represented by a phenomenon that caused the temperature drop, for which it should hold true that it is very short in time and has ceased after the temperature reached its minimum. Therefore, if the system answers a step change with the exponential response, then it must be described by a firstorder differential equation. This is typical where heat conduction is prevalent. Thus, the detected temperature rise from the lowest temperature to the ambient (starting) temperature might be a result of the heat, going from the silicon glass plate back to the water touching it. 2.2 Kinematic Analysis Raw images from the IR camera were processed in a way allowing the contours of the so-called clouds were extracted. They were approximated by circles and the average temperatures of the clouds were calculated (Fig. 7).
Fig. 7. Raw image and processed image with detected edge and its approximation with a circle
Fig. 6. Evolution of the average temperature of the cloud and its exponential fit In Fig. 6 a time evolution of the average temperature of a cloud (dots) and an exponential approximation for a rising part are presented. This approximation seems suitable for a heat transfer after the initial temperature drop and the results show sound agreement with the experimental data (also mentioned above in Fig. 5). The first point which lies outside the proposed fit belongs to the part where the temperature is still decreasing and
The edge of a cloud is representing a cold front, originating in microbubble generation caused by the splashing effect [16]. To find the edges, the Matlab edge function was used [17] and [18]. The edge function was used with the Canny method [19]. To study the phenomenon further, a propagation of the cold front was determined (Fig. 8). The positions of the front correspond to the temperature fields in Fig. 4. From the data of the front position in a specified time lapses, the propagation velocity could be determined. The mean velocity for the front movement in each time frame is based on a change of radius of the cloud, approximated by a circle. Velocities of the cloud spreading are presented as a function of time and the cloud radius (Fig. 9).
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Fig. 8. Propagation of the cold front, 1) initial state with almost no temperature effect, 2) first appearance of a cold cloud, 3 to 7) decelerated growth of the cloud, resulting in 8) a final size of a cold area
Fig. 9. Velocity of the cold front propagation, a) as a function of time, b) as a function of radius The collapse, causing a splashing effect responsible for the observed local cooling, strongly depends on the initial size of the bubble and its distance from the solid wall [20] and [21], therefore, for the fitting function, at least two coefficients are necessary. For this reason, an exponential function aebx was chosen without considering any other physical background. However, data presented in Fig. 9 were fitted with the agreement R2 > 0.96. Generally, sets of data (for an evolution of each single cloud) and appurtenant coefficients of fitting functions differ quite a lot, although the evolutions all share the same general shape. A possible reason why, in some cases, the shape of the cloud differs from circular is the interaction between bubbles that appear close to
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one another or some not yet damped velocity field from previous collapses so that the microjet that forms during the collapse is no longer perpendicular to the solid wall. 3 CONCLUSIONS Oscillations with ultrasound frequencies (42 kHz) were used to generate bubbles in a small tank. A silicon glass window was attached to the cylinder and partially submerged into water. Bubbles that imploded and grew in the vicinity of the window generated small temperature changes, which were recorded with a high speed thermocamera. Small but distinctive local temperature decreases in order of few 1/10 K with a duration in order of few 1/100 s were detected.
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In this way, a new approach towards an experimental classification of cavitation on a local level by using infrared thermography was presented. The motivation was set by the lack of comparable methods whereas the aim was not to capture the extreme conditions that happen in the final stage of a collapse for which the method is still not fast enough. Based on the obtained thermographic results a kinematic analysis was performed, where the expansion of the areas of temperature changes was investigated. The velocity of expansion was related to the characteristic area size as a function of two independent variables, indicating the effect of cavitation intensity and its distance from the wall. The introduced thermographic method was applied on the ultrasound induced cavitation as an example of useful cavitation effects and as it is believed that the method can contribute to a better understanding of the cavitation phenomenon, this may lead to an improved performance of many practical applications where cavitation is present. In addition, as the water sensibility to cavitation thermal effects at room temperature is very weak [4] and [15], it may also be expected that the method will provide even better results in cases when hot water or other, e.g. cryogenic fluids, are used. 4 REFERENCES [1] [2]
[3]
[4]
Franc, J.P., Michel, J.M. (2004). Fundamentals of cavitation. Kluwer Academic Publishers, Dordrecht. Fujikawa, S., Akamatsu, T. (1980). Effects of the non-equilibrium condensation of vapor on the pressure wave produced by the collapse of the bubble in a liquid. Journal of Fluid Mechanics, vol. 97, p. 481-512. Utturkar, Y., Wu, J., Wang, G., Shyy, W. (2005). Recent progress in modeling of cryogenic cavitation for liquid rocket propulsion. Progress in Aerospace Sciences, vol. 41, p. 558-608. Fruman, D.H., Reboud, J.L., Stutz, B. (1999). Estimation of thermal effects in cavitation of thermosensible liquids. Int. Journal of Heat and Mass Transfer, vol. 42, p. 3195-3204.
[5] [6]
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Ishii, T., Murakami, M. (2003). Comparison of cavitating flows in He I and He II. Cryogenics, vol. 43, p. 507-514. Franc, J.P., Rebattet, C., Coulon, A. (2004). An experimental investigation of thermal effects in a cavitating inducer. Journal of Fluids Engineering, vol. 126, p. 716-723. Yoshida, Y., Kikuta, K., Watanabe, M., Hashimoto, T., Nagaura, K., Ohira, K. (2006). Thermodynamic effect on cavitation performances and cavitation instabilities in an inducer. Proceeding of 6th International Symposium on Cavitation CAV2006, Wageningen. Yoshida, Y., Sasao, Y., Okita, K., Hasegawa, S., Shimagaki, M., Nakamura, N., Ikohagi, T. (2006). Influence of thermodynamic effect on synchronous rotating cavitation. Proceeding of 6th International Symposium on Cavitation CAV2006, Wageningen. Mason, T.J., Lorimer, J.P. (2002). Applied Sonochemistry. Wiley-VCH Verlag, Weinheim. Hale, G.M., Querry, M.R. (1973). Optical constant of water in the 200 nm to 200 μm wavelength region. Appl. Opt., vol. 12, p. 555-563. Edmund Optics http://www.edmundoptics.com (2009), accesed on 2009-10-29. Benjamin, T.B., Ellis, A.T. (1966). The collapse of cavitation bubbles and the pressures thereby produced against solid boundaries. Phil. Trans. Roy. Soc., vol. 260, p. 221-240. Plesset, M.S., Chapman, R.B. (1971). Collapse of an initially spherical vapor cavity in the neighbourhood of a solid boundary. Journal of Fluid Mechanics, vol. 47, no. 2, p. 283-290. Lauterborn, W., Bolle, H. (1975). Experimental investigations of cavitationbubble collapse in the neighbourhood of a solid boundary. J. of Fluid Mechanics, vol. 72, no. 2, p. 391-399. Brennen, C.E. (1995). Cavitation and bubble dynamics. Oxford University Press, New York. Tong, R.P., Schiffers, W.P., Shaw, J.S., Blake, J.R., Emmony, D.C. (1999). The role of “splashing” in the collapse of the lasergenerated cavity near a rigid boundary.
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[17] [18] [19]
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Journal of Fluid Mechanics, vol. 380, p. 339-361. Lim, J.S. (1990). Two-dimensional signal and image processing. Englewood Cliffs, NJ, Prentice Hall, p. 478-488. Parker, J.R. (1997). Algorithms for image processing and computer vision. John Wiley & Sons, Inc., New York, p. 23-29. Canny, J. (1986). A computational approach to edge detection. IEEE Transactions on Pattern Analysis and Machine Intelligence, vol. 8, no. 6, p. 679-698.
[20]
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Reboud, J.L., Fortes-Patella, R., Archer, A. (1999). Analysis of damaged surfaces: part i: cavitation mark measurements by 3D laser profilometry. Proceedings of the 3rd ASME / JSME Joint Fluids Engineering Conference, San Francisco. Philipp, A., Lauterborn, W. (1998). Cavitation erosion by single laser-produced bubbles. Journal of Fluid Mechanics, vol. 361, p. 75-116.
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 535-543 UDC:533.661:629.01
Paper received: 30.01.2008 Paper accepted: 02.07.2010
Determination of a Light Helicopter Flight Performance at the Preliminary Design Stage Zlatko Petrović - Slobodan Stupar - Ivan Kostić* - Aleksandar Simonović University of Belgrade, Faculty of Mechanical Engineering, Serbia Promising solutions for the problem of the extensively time-consuming modern urban transportation has been found in the use of light and very light helicopters. This paper presents a part of the preliminary design methodology, compiled at the Faculty of Mechanical Engineering, University of Belgrade and includes performance calculations of such helicopters. Due to limited budgets and an extremely demanding process of helicopter development, it is highly significantt that during all development stages reliable performance estimates are obtained in order to ensure assigned operational requirements. The scope of this paper is confined to the preliminary design stage, where it is customary to substitute the very complex helicopter rotor dynamics with its averaged mechanical and aerodynamic characteristics and apply certain empirically verified simplifications. Based on this approach, the independent, efficient and reliable computer programs for the calculation of different performance characteristics have been developed. In addition to their application on an actual on-going project, they have also been applied on several existing helicopters of a similar class for a more accurate determination of the empirical input parameters. The applied methodology and obtained results have been presented, verifying the overall algorithm efficiency. ©2010 Journal of Mechanical Engineering. All rights reserved. Keywords: light helicopter, flight performance, preliminary design stage 0 INTRODUCTION Proper evaluation of the most important helicopter design parameters and estimation of its basic flight performance in the initial design stages is highly significant for the overall project time and cost effectiveness. It assumes the application of mathematical models that are fairly simple, in conjunction with empirical coefficients and parameters derived from the previous successful designs. Such a requirement of simplicity is in contrast with the extremely complex helicopter rotor dynamics. During one revolution, a rotor blade in progressive flight is generally subjected to the pitching, flapping and leading - lagging motions, repeated several hundreds of times in a minute, while being exposed to the gravitational, centrifugal, inertial, and aerodynamic loads [1] and [2]. At higher progressive flight speeds, blades at the advancing azimuths are at very small pitch, with tips that might have local supersonic flow zones even in case of light helicopters, while the retreating blades are at incidences that are often beyond the static stall angle, with inner domains subjected to the reverse flow (velocity of flight is higher than local tangential velocities in this domain). The
aim of such cyclic blade motions is to keep the resultant rotor thrust acting in the plane of symmetry. Flapping motions of the blades enable tilting of the main rotor disc in certain directions when desired, to generate forward, backward or lateral thrust components. Due to all these factors, helicopter blades in progressive flight are subjected to very complex unsteady airflow patterns. In spite of that, many years of industry experience has shown that quite good and very efficient preliminary estimates of helicopter flying characteristics can be obtained by averaging some aspects of the rotor and overall helicopter dynamics and aerodynamics. Many methods have been developed so far, varying in the level of complexity and accuracy of the obtained results. Their general aim is to bring the most important design parameters of a new helicopter close enough to their optimums, so that major changes are hopefully not necessary at higher design levels, which involve very complex computational and experimental methods that are expensive and time consuming. The best results in the initial design analyses can be obtained if calculations at the preliminary level are repeated
*
Corr. Author's Address: University of Belgrade, Faculty of Mechanical Engineering, Aeronautical Department, Kraljice Marije 16, 11120 Belgrade 35, Serbia, ikostic@mas.bg.ac.rs
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in several refinement steps, with the complexity level increasing from one to another. This paper presents some of the calculation procedures applied in the preliminary flight performance estimates of the light helicopter designs currently under development at the Aeronautical Institute of the Belgrade Faculty of Mechanical Engineering. The first one is a design ordered by a foreign partner (Fig. 1.), while the second one is aimed to be its simplified technology demonstrator version, with the take off mass limited to 650 kg, for which the results are presented in this paper. The design process, in accordance with Certification Specifications for Small Rotorcraft CS-27 and Certification Specifications for Very Light Rotorcraft CSVLR, was initiated with certain performance requirements. Some of these requirements considered low gross weight, payload larger than 180 kg, range over 450 km, high value of hover ceiling and cruising speed at 1000 m ISA+15, higher than 160 km/h. Applied calculation algorithms have been compiled with an aim to establish a proper balance between the required simplicity and time effectiveness on one hand, and the expected accuracy on the other. For the verification purposes, the same calculations have been applied on several existing light helicopters. Those results were then used to improve some of the empirical parameters initially applied in new helicopter calculations. The obtained results have proven to be very valuable inputs for the following higher level calculations. Considering the fact that this is an on-going project, only the results from the initial calculation stage will be presented in this paper. 1 CALCULATION PROCEDURES For the here presented analyses, the take off mass of m 650 kg has been considered as a constant input value. For operational purposes, mass should be varied within the predefined range. Initially, an optimization procedure had been applied to determine the most relevant calculation inputs. Considering the main rotor, the most relevant parameters that were obtained are: number of blades n 2 , rotor radius R 3.8 m, blade chord length c 0.205 m, solidity factor of the rotor n c / R 0.0343 , rotor disc
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area A 45.96 m2, number of revolutions per minute N 440 rpm = const. for all flight regimes, giving blade tip tangential velocity VT 175 m/s = const. The reciprocating power plant gives the maximum output of P0 max = 147 kW (200 HP) at the sea level. The engine power at other altitudes is estimated as: PH P0 max (1.11 / 0 0.11) , (1) where 0 1.2255 kg/m3 represents the density of air at H 0 m, and is density at a given altitude, defined by equation: 20000 H 0 , (2) 20000 H in which the altitude H is expressed in meters. Also, for the purpose of this paper, the optimum fuel/air mixture at all altitudes has been assumed. In operational design work, the actual engine characteristics for different altitudes should be used. At this level of helicopter performance calculations, a standard approach is that the aerodynamic characteristics of the blades are averaged over the main rotor disc. All presented analyses, based on [1] to [3] have been done using a custom developed software for solving the sets of equation that will briefly be presented within the oncoming sections. 1.1 The Average Main Rotor Blade Lift Coefficient The average blade lift coefficient is determined as: C L 6CT / , (3) while the thrust coefficient in hovering and level flight is given by: T CT , (4) R 2 VT2 and T W m g. For initial estimates, Eq. (4) can be used both for hovering and progressive flight. At higher calculation levels, this equation should be refined by including the disc slope angle and collective pitch for the given mass for example to determine a more accurate CT for the given progressive flight regimes.
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Fig. 1. Helicopter CAD model and some examples of the components manufactured for the laboratory tests and the production technology verifications (higher development stages of the on-going project) 1.2 Power Required Horizontal Flight
for
Hovering
and
(7)
and:
Total power which is required for hovering and horizontal flight of a helicopter can be obtained as the main rotor power increased by approximately 10% to take into account the required tail rotor power and transmission losses P 1.1 C P AVT3 . The main rotor power coefficient C P in hovering is calculated from:
CP h CT
i h vi / vh
CD , (5) 8 0 while for horizontal flight it becomes: R CP i CT CD 0 (1 k 2 ) D CT , (6) W 8 where represents a coefficient which takes into account the induced velocity distribution irregularities over the main rotor disc. For hovering 1.15 , while for forward flight 1.2 ; V / VT , where V is progressive flight velocity. Constant k 4.65 is an empirical value, used by the Westland Helicopters [3]. Drag of the helicopter, except the rotor, is 2 RD 0.5 CD SV 2 0.5 f AV 2 , where f A 1 m is
an estimate of the flat plate equivalent area, a usual rounded value applied in the initial step for small helicopter designs. Using the notation vh and vi for velocities induced by the main rotor in hovering and horizontal flight, the induced velocity coefficient in progressive flight i vi / VT is obtained from equations:
vi / vh V / vh vi / vh 4
2
2
1 0 .
(8)
Parameter h vh / VT CT / 2 represents the induced velocity coefficient in hovering. For the calculated average rotor disk lift coefficients determined using Eq. (3), the averaged blade profile drag coefficients CD 0 for several characteristic altitudes have been obtained using the steady state polar curve shown in Fig. 2. Knowing that vibrations during the flight generally affect the boundary layer generation over the rotor blades [1], a dominant turbulent layer has been assumed, and standard roughness polar for Reynolds number 1.8×106 has been used [4]. It should be noted that, according to Fig. 2, the average CL for H = 5 km is practically at the maximum the steady-state lift curve. On the other hand, the retreating blades under operational conditions will require higher local lift coefficient values. Since the maximum lift coefficient of an airfoil under dynamic flow conditions encountered on helicopter rotors is always higher than in steady flow [2], this value can also be used in formal averaged calculations. It should be noted that because of the actual engine operational restrictions, kinematic limitations of the main rotor (still not known at this design level) and similarly, the preliminary results considering the domain of the absolute ceiling must be taken with reservations as they might be overoptimistic to a certain extent.
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Fig. 2. Steady-state lift coefficient and standard roughness polar curve for the NACA 8-H-12 airfoil [4] 1.3 Rate of Climb in Progressive Flight
1.4 Rate of Descent in Autorotation
Main rotor power required for climbing is P CP AVT3 . In this case, the power coefficient is given by: f 1 CP i CT CD 0 (1 k 2 ) 3 A 8 2 A (9) c CT . In Eq. (9), c w / VT is relative climbing velocity, where w is the actual rate of climb in meters per second. Parameter 1.3 takes into account additional losses caused by the changes of relative flow direction in climb, while other parameters have the same meaning as already mentioned. The highest rates of climb at a given flight velocity V and altitude H can be reached when maximum available engine power PAV for this altitude is applied. Since the main rotor receives approximately 90% of the total engine power, the equation for c becomes:
c
09 CPAV
CT
3 fA . 2 CT A
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i CD (1 k 2 ) 8 CT 0
(10)
PAV
Even in case of total engine power loss 0 kW , helicopters are able to land the
autorotation regime. For this case, Eq. (10) transforms into: CD 0 3 fA d i (1 k 2 ) , (11) 8 CT 2 CT A where d represents relative descending velocity. For this purpose, the value 1.0 gives more reliable results. The rate of descent in autorotation waut for given V at H is then obtained as: CD 0 waut VT i (1 k 2 ) 8 CT 1 3 fA VT . 2 CT A
(12)
1.5 Height-Speed Envelope, Optimum Speeds and Maximum Rates of Climb in Progressive Flight After calculating the powers required for progressive flight and the available powers for different altitudes using the presented algorithms, at their crossing points, the minimum Vmin and maximum Vmax flight speeds of flight have been determined. To achieve sufficiently small altitude
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steps, required values of C L and C D have been 0
interpolated or extrapolated using the data from Table 1. At the points of maximum corresponding power differences, the velocities Vopt for given altitudes have been defined. Maximum rates of climb that correspond to the calculated optimum speeds are obtained as described in section 1.3. 1.6 Ground-Effect Influence on the Main Rotor Power Required for Hovering
In the proximity of ground, the power required for hovering becomes smaller. The induced velocity decreases progressively with the ground proximity according to the equation: vhG 05 1 , 2 vh (13) H 1 4 R where H represents the main rotor disc height from the ground, and is the ground effect coefficient. The equation for the power coefficient for the calculation of the main rotor power (only) in the ground proximity PG CPG AVT3 is: CPG h CT
CD . (14) 8 0 In case of light helicopters, at heights of the order of 15 m, Eq. (12) practically takes the form of Eq. (5). The total power required should include an additional 10% for the tail rotor and transmission losses. 1.7 Acceleration in Horizontal Flight
Expressions which define acceleration in horizontal flight are derived from the Second Law of Newton: dV m X , (15) dt
dV XV PAV P . dt from wich the acceleration is: dV PAV P ahor . dt mV mV
(16)
(17)
2 RESULTS AND DISCUSSION Figs. 3 to 7 show the results that were obtained in the initial stage of the preliminary analyses of the here presented helicopter project, using algorithms explained in previous sections. All diagrams, except Fig. 6, have been obtained using the induced velocity distribution coefficient 1.2 for the progressive flight in the whole domain. On the other hand, as mentioned in section 1.2, for V = 0 km/h the hovering value 1.15 should be applied. In order to avoid a singularity jump at speeds just above zero, in curves involving this coefficient, an interpolation should be made in latter refinement steps in the domain of small progressive flight speeds. For example, considering the power required at H 0 m, the difference between the application of the two values of the coefficient results in the difference of the order of about 2.5 kW at V = 0 km/h. This could be verified using the main rotor power value P uncorrected for the ground influence in Fig. 6 (dashed line). To get the total power required, this value should be multiplied by factor 1.1, and then compared with the H = 0 km curve in Fig. 3, which leads to the above mentioned value. Considering the diagram shown in Fig. 9, it is obvious that values of acceleration in horizontal flight at small velocities, when V 0 , tend to infinitely large values. This is a natural consequence of the application of a simple approach described in section 1.7, which is good enough in preliminary analyses, but results for very small speeds of flight must be ignored.
Table 1. Definition of the averaged profile drag coefficient for some characteristic altitudes for NACA 8H-12 airfoil, derived from Fig. 2 H [m] 0 1000 2000 3000 4000 4500 5000 CL 0.654 0.723 0.799 0.885 0.981 1.034 1.090 CD 0
0.0120
0.0126
0.0134
0.0147
0.0181
0.0250
Determination of a Light Helicopter Flight Performance at the Preliminary Design Stage
0.0393
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Fig. 3. Variation of the total power P for horizontal flight and the available power PAV, with altitude
Fig. 4. Variation of the rate of climb with speed
Fig. 6. Main rotor hovering power including ground effect
Fig. 5. Rate of descent in autorotation (H = 0 m)
Fig. 7. Height-speed envelope and Vopt
Algorithms which are applied in this paper are influenced by empirical factors. Thus, it is good engineering practice to verify them on several existing helicopters whose overall design characteristics are as close as possible to the category of the new model under development. Some results obtained for the Robinson R22 Beta
II are shown in Fig. 10 and in Table 2. Parameters [5] to [7] applied in the calculations simulating the preliminary design level of this helicopter were: m = 621 kg, n = 2, R = 3.85 m, c = 0.18 m, main rotor blade airfoil NACA 63-015 [8] and [9], etc. Its Lycoming O320B2C engine whose nominal maximum power is 160 HP, is derated to
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the maximum of 131 HP at lower altitudes in order to extend the engine and transmission life time. For the here presented calculations it has been assumed that P = 131 HP can be maintained constant up to the altitude of 1780 m, after which it begins to decrease (value obtained from Eq. (1), using P0 max 160 HP; the actual engine can develop 131 HP up to H 2250 m [7]).
Fig. 8. Variation of the maximum rate of climb with altitude
Fig. 9. Acceleration in horizontal flight at H=0m Keeping in mind that we are talking about the accuracy at the preliminary analyses level, agreements between the calculated values and the existing data are good. Again, the absolute ceiling is most probably overestimated to a certain extent, but this result can not be compared with the operational data. It is actually a theoretical value which is generally not flight-tested for helicopters because it could lead to a disaster. On the other hand, the operational ceiling can be assigned by the manufacturer only after a vast number of rigorous test flights. Still, Fig. 10a leads to a conclusion that this helicopter would have a very reasonable speed range and power
reserve for safe operations if flown the at the altitude of 4270 m (oxygen system is not a part of the R22 standard equipment). Performing such "reverse-engineering" analyses of the existing models, together with the calculations for a new helicopter, is very useful for fine adjustments of the applied empirical coefficients in calculation algorithms. For example, a proper match for the Robinson's maximum speed has been achieved using f A 0.8 m2 instead of the 1 m2 value from section 1.2, which has been applied in the here presented initial calculations of the new helicopter. Therefore, the calculations had to be repeated with f A 0.8 to 0.9 m2 for more realistic estimates of its fuselage drag of the here analyzed project. Such relatively simple calculation techniques are obviously extremely valuable for quick and efficient relative comparisons with the existing designs. In addition, the obtained results provide a good initial insight in the capability of the new design to satisfy certain requirements prescribed by air regulations for the given helicopter category. Preliminary performance calculations, such as the ones presented in this paper, are most often done using the data for the original airfoil (or airfoils) applied in rotor design. On the other hand, it is known that for composite rotor blades, the airfoil must be modified to comprise a fixed flat tab along the whole trailing edge, primarily in order to enable proper merging of the upper and lower blade surface plies during the manufacturing. This modification can affect the airfoil profile drag [1] to a certain amount. For the here presented new helicopter, it has been shown [8] that the most unfavorable expected tab design from the aspect of drag increase should not absorb more than 1.5% of the maximum available engine power at H = 0 m, so at the preliminary design level this influence can be ignored. On the other hand, if an asymmetrical airfoil is used in the blade design, even a small error in determining a proper angular position of the tab can seriously affect the moment about the aerodynamic center [8]. In case of helicopter blades this value must remain small enough, so this particular issue should also be very carefully considered at higher design levels.
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Fig. 10. Calculations for the Robinson R22 Beta II: a) height-speed envelope (1 - power derated to 131 HP = const. up to 1780 m; 2 - operational ceiling value taken from Table 2); b) variation of the rate of climb with speed Table 2. Comparison of the existing and calculated data for Robinson R22 Beta II – hovering ceiling; 2 – absolute ceiling; 3 – maximum operating altitude (service ceiling); 4 – maximum rate of climb at sea level; 5 – maximum rate of climb at 3 km altitude; 6 – maximum speed in level flight 1
2
3
4
5
6
Robinson R22 m = 621 kg
Hmax hover [M]
Hmax [M]
Hmax oper [M]
wmax (H = 0 km) [m/s]
wmax (H=3 km) [m/s]
Vmax [km/h]
Existing data
2867* [5]
/
4270 [5]
> 5.1 [5]; 6.1 [6]
> 3.05 [5]
180 [7]
Calculations
2889
5433
/
6.07
4.03
178÷188 **
*
in ground effect; out of ground effect, a bit smaller value would be obtained for H = 0 to 1780 m, assuming that derated power of P = 131 HP is kept constant
**
3 CONCLUSION In this paper some of the most important issues considering a light helicopter flight performance, confined to the preliminary design level, have been analyzed. An approach common at this project level, based on the averaged rotor and helicopter dynamic and aerodynamic characteristics, has been applied. For the presented calculations of the power required for hovering and horizontal flight, the rate of climb in progressive flight and the rate of descent in autorotation, height-speed envelope, optimum speeds and maximum rates of climb, ground effect influence and acceleration in horizontal flight, sets of equations have been carefully selected to achieve proper balance between the simplicity, time effectiveness and the required accuracy at this design level. Custom developed software has shown the ability to perform
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efficient analyses and enable variations of the design parameters in the required ranges, giving stable and smooth solutions. Knowing the fact that such calculations are based on certain empirical parameters, parallel calculations have been performed for several similar existing helicopter designs, and some of them for the Robinson R22 Beta have also been presented in this paper. This particular example has shown the need for certain adjustments of the fuselage drag calculations of the presented helicopter project in the oncoming design steps. After verifying all other applied parameters for the given helicopter category in a similar manner, the presented calculations can provide extremely useful inputs for further, much more complex and time consuming analyses at higher design levels where detailed helicopter dynamics and aerodynamics must be taken into account. The presented approach in preliminary calculation metodology
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can noticeably contribute to the overall project effectiveness. 4 REFERENCES [7] [1] [2] [3] [4]
[5]
[6]
Leishman, J.G. (2000). Principles of helicopter aerodynamics. Cambridge University Press, Cambridge. Bramwell, A.R.S., Done, G., Balmford, D. (2001). Bramwell's helicopter dynamics,2nd ed.. Butterworth-Heinemann, Oxford. Seddon, J. (1990). Basic helicopter aerodynamics. BSP Professional Books, Oxford. Stivers, L.S., Rice, F.J. (1946). Aerodynamic characteristics of four NACA airfoil sections designed for helicopter rotor blades. NACA RB, No. L5K02. Robinson helicopter company, USA, R22 Beta II specifications and dimensions, from http://www.robinsonheli.com/r22 specs.htm, accessed on 2007-22-11. Robinson helicopter company, USA, R22 Pilot’s operating handbook and FAA
[8]
[9]
[10]
approved rotorcraft flight manual, p. 2-2, & p. 7-17., 1996; from Wikipedia, Internet encyclopedia, http://en.wikipedia.org/wiki/ Robinson_R22 2, accessed on 2007-22-11. Airliners.net, Aircraft technical data & specifications - Robinson R22, from http://www.airliners.net/info/stats.main?id =339, accessed on 2007-23-11. Lednicer, D., Incomplete guide to airfoil usage (aircraft-airfoil database under progressive development), Analytical Methods Inc., Redmond WA, from http://amber.aae.uiuc.edu/~m-selig/ads/ aircraft.html, accessed on 2004-06-05. Abbot, I.H., Von Doenhoff. A.E. (1959). Theory of Wing Sections, including a summary of airfoil data. Dover Publications Inc., New York. Kostić, I. (2007). Numerical evaluation of the aerodynamic influence of the helicopter composite blade trailing edge tabs, Archive of Applied Mechanics, Springer-Verlag, vol. 77, p. 893-909.
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Paper received: 22.09.2009 Paper accepted: 24.05.2010
The Strength of the Bus Structure with the Determination of Critical Points Dušan Mežnar* - Momir Lazovič Tovarna Vozil Maribor d.o.o., Skupina Viator&Vektor, Slovenia A monocock structure of an airport bus is a very demanding product as regards its strength. With the application of the FEA (Finite Element Analysis) method the allegedly critical points of the framework were determined; these especially occurred on the door frameworks. The experimental methods of measuring mechanical deformations confirmed the presumptions that maximum deformations measured at the points which were previously analysed with the FEA method. The driving regime with a maximum speed of 40 km/h in a circle with a minimum turning radius and a changeable regime of acceleration and braking proved critical. The measurements led to appropriate construction amendments, additional strengthening of the framework and other measures which fulfilled the required strength criteria. ©2010 Strojniški vestnik - Journal of Mechanical Engineering. All rights reserved. Keywords: buses, motor vehicles, bodies, reinforcements, strength 0 INTRODUCTION The framework of a bus is a self-bearing construction made from steel profiles welded together in a monocock. The monocock consists of a chassis frame and the frameworks of the front and rear panels, the left and the right panels, the roof, the dash board and other components. The chassis and the body are one and the same structure. The self-bearing monocock bus construction means that it is completely prepared for the installation and fitting of the bus chassis and the body systems and sets. The monocock construction predetermines the places where the engine, the transmission gear, the suspension, the axles, the steering system and the other equipment will be fitted. A sufficient rigidity of the construction ensures elastic deformations of the bus framework, but keeps them within the maximum permissible limits so that they do not affect the functions of individual aggregates and systems. Above all this refers to the door function; their opening and closing, a limited deformation of the air-conditioning device mounted on the roof, and the unaffected function of the propulsion installations, the engine etc. There should be a favourable ratio between the rigidity and the weight of the bus which also serves as a criterion for proving and estimating the *
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successfulness of the construction in reference to its carrying capacity and own weight [1]. A special feature of the airport bus regarding the wheelbase of dimension 8400 mm was taken into account. Therefore specific and especially extreme loads were included in the project and construction analyses. The strength analysis provides a comparison with the data on similar buses which in the past suffered from the occurrence of cracks, roof waving, deflections of the tracks, cracks on the door corners etc. 1 THE BASIC STRUCTURE A monocock consists of four modules: the chassis frame, the left and right panel frameworks, the front and real panel frameworks, the roof. The basic part of the monocock is the chassis frame onto which the panels and the roof are mounted. The positioning of the main points for the installation of the systems and aggregates is controlled vertically with regard to the floor, laterally with regard to the centre of the axles and longitudinally with regard to the determined distances. The roof framework is the final part of the monocock and is positioned according to the positions of the side and other panels in relation to the chassis frame. This stage of the technological process is followed by adhering the
a) b) c) d)
Corr. Author's Address: Tovarna Vozil Maribor d.o.o., Skupina Viator&Vektor, Cesta k Tamu 33, 2000 Maribor, Slovenia, menarduan456@gmail.com
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metal sheets to the monocock structure which provides the latter with reinforcement and at the same time solves the issues of noise and temperature isolation, vibrations, uneven surfaces etc. 2 A MONOCOCK, THE SELF-BEARING BUS CONSTRUCTION Monocock is the self-bearing construction for fitting of the propulsion system of the vehicle: the engine, the clutch, the transmission gears, the suspension, the axles and individual sets of the braking, steering and other systems. Due to this function, the monocock is an essential part of the bus as it has to be designed in a way which enables an uninterrupted and simple fitting of all the systems and aggregates to their pre-determined installation positions. The selfbearing monocock construction is a step forward and represents a pre-defined form, design and function. In addition, all analyses of the crash impact tests are taken into account.
More specific data are only obtained experimentally using the strain gauge, i.e. by applying the method of measuring the strength with electrical resistance. The FEA model only served to determine the locations where the measuring tapes will be applied (Fig. 1). On the most loaded locations 32 points were determined at which measuring tapes were adhered to the framework at an early stage of the monocock construction. In this phase the starting values of deformations and strength were determined which had a »zero value« in the finished bus – the system was re-set and the bus's own deformations were neutralised after the installation of the bus equipment; so in measurements only absolute values were taken into account. Combinations of measuring tapes (Fig. 2) which registered deformations in all directions, i.e. in x, y, z orientations, were used. These were linear measuring tapes and rosettes which are normally used in measuring composed loads and deformations.
3 THE APPLICATION OF THE FEA METHOD A linear modular analysis FEA is used to determine the basis for the analysis and testing of the strength with an experimental method. The objective is determination or a rough estimation of the load and deformations of the monocock.
Fig. 2. Adherence of the measuring tapes 4 PREPARATION OF THE CONSTRUCTION FOR MEASUREMENT
Fig. 1. FEA with a local concentration of tension
The self-bearing bus construction was tested using the measuring tapes on 32 locations (Fig. 3) which were determined after a preliminary analysis and a FEA calculation. The most critical points were the corners of the six bus doors which was also indicated by the tension measurement [2]. The bus in the monocock form was tested statically and dynamically. The testing was carried out for different cases of the framework
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loading, i.e. for different regimes of driving of an unloaded and loaded vehicle. The following analyses were made [3]: analysis of the vehicle's statics, analysis of acceleration and driving ahead, analysis of braking, analysis of driving in a circle with acceleration and braking. The maximum permissible tension with regard to the used material and the manner of use of the bus was 150 MPa. An example of the tension calculation is given according to the equation Eq (1) . E 2.1105 850 178.5 N/mm2 (MPa) . (1) Table 1 presents partial results of the measurements of the bus structure strength for the different driving conditions.
5 THE METHODOLOGY OF MEASUREMENTS AND MARGIN CONDITIONS Maximum tension values were achieved at the measuring points during the driving of a loaded bus in a circle with braking. The test surface on which the experiment was carried out was uneven with holes and bumps. The tension oscillations were especially emphasised due to the impact loads at passing through obstacles [4] and [5]. We arrived to the following conclusions: a) The testing, i.e. measuring of the propulsion strength of the bus framework was carried out under special conditions which are not suitable for the regular use of the bus. The driving regime of a fully loaded bus with sharp braking in a bend presented a critical test of the strength.
a)
b)
a) 1 – A rear door pillar, 6 – A middle door pillar, 2 – B front door pillar
b)26 -27 – A front door pillar, 21 – A middle door pillar, 16 – H holder
Fig. 3. Measuring points: a) on the left, b) right side of the framework Table 1. Maximum deformation values (m/m) Point Driving regime 01 unloaded – on the spot without zero adjustment 03 unloaded - driving 07 unloaded – driving – bend + breaking 08 loaded 10 loaded – driving + braking 12 loaded – bend + braking 13 loaded – on the spot – elevated at the back 14 loaded – driving – without H-holder 15 loaded – driving + braking – without H holder
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1 6 155 -600 15,5 -37 120 180 -70 340 -300 620 320 660 1075 350 -695 625 -710 700
Mežnar, D. - Lazovič, M.
12 16 25 140 -20 5 295 310 30 5 70 380 -400 380 -360 -140 -200 250 -340 275
21 -670 320 530 850 400 570 610
26 90 11 110 -90 -260 -280 -185 -200 -200
27 -60 7 160 130 330 440 250 360 490
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b) In the case of an unloaded bus the deformations and tensions at point 16 (65.1 MPa) and at point 12 (61.95 MPa), both under 150 MPa which is supposed to be the maximum permissible value of the strength of the framework material. c) The design relating to the propulsion strength derives from the loading conditions which were determined according to the following criteria: 70% driving on the airport: o 40% acceleration and driving, o 30% normal braking (extreme braking 2 to 5%), o 30% driving in a circle; 30% waiting – staying at standstill. 6 THE MEASURES FOR INCREASING THE RESISTANCE MODULUS IN THE LONGITUDINAL DIRECTION Regardless of the test results the following measures for increasing the resistance modulus in the longitudinal direction were taken [6]:
Deformation [µm/m]
strengthening of individual parts of the existing construction: the track, the door pillars, the roof etc.,
new solutions on the framework of the second prototype which included additional joints in the framework, other measures – modifications of the framework. All critical points which were determined on the basis of FEA were constructionally analysed and appropriately strengthened. Figs. 4 to 7 show deformations of the critical points on the framework and in the area of the opening for the middle bus door [7]. The measuring points 6 to 10 determined deformations on the left side of the middle door; the measuring points 21 to 25 determined deformations on the right side of the middle door in the case of a fully loaded bus. Position 6 was a linear strain gauge on the lower part of the horizontal profile near the front pillar of the right middle door. Position 7 was a linear measuring tape on the rear side of the front pillar of the right middle door near the upper horizontal profile. The measuring points 8, 9 and 10 were joined in one strain gauge – a rosette, where two of the three measuring tapes were adhered at 900 angle (horizontal and vertical network) and the third measuring tape was adhered at 450 angle [8] to [9]. The rosette was located in the middle of the front strengthening plate to the right of the middle door [10].
Time Device_1 [s] Fig. 4. Strains at measuring point 6 to 10 of a loaded vehicle
Mežnar, D.- Lazovič, M.
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Time Device_1 [s]
Deformation [µm/m]
Fig. 5. Strains at measuring point 21 to 25 of a loaded vehicle
Time Device_1 [s] Fig. 6. Dilatation at measuring point 6 to 10 during the driving and braking of a loaded vehicle
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Deformation [µm/m]
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 544-550
Time Device_1 [s] Fig. 7. Strains at measuring point 21 to 25 during the driving and braking of a loaded vehicle The location of the measuring tapes on the right side of the framework followed the same principle as on the left side which is shown in Fig. 3. The deformations of the framework at critical points did not exceed 350 m/m at static loading [11] and achieved maximum 800 m/m during the driving of a loaded bus [12] and [13,14]. 7 CONCLUSION The experimental methods of measuring mechanical deformations confirmed the presumptions that maximum deformations were measured at the points which were previously analysed with the FEA method. The driving regime with a maximum speed of 40 km/h in a circle with a minimum turning radius and a changeable regime of acceleration and braking proved critical. The measurements led to appropriate construction amendments, additional strengthening of the framework and other measures which fulfilled the required strength criteria.
8 REFERENCES [1] Kušar, J., Duhovnik, J., Tomaževič, R., Starbek, M. (2007). Finding and evaluating customers needs in the product-development process. Strojniški vestnik – Journal of Mechanical Engineering, vol. 53, no. 2, p. 78-104. [2] Puklavec, B. (2004). Preparation of measuring points for strain measurement on the prototype of the Neoplan airportbus. University of Maribor, Faculty of Civil Engineering, Institute of Civil and Traffic Engineering, Maribor. Documentation of Tovarna vozil Maribor. [3] Puklavec, B. (2004). Results of strain measurements on the prototype of the Neoplan airportbus. University of Maribor, Faculty of Civil Engineering, Institute of Civil and Traffic Engineering, Maribor. Documentation of Tovarna vozil Maribot. [4] Kušar, J., Bradeško, L., Duhovnik, J., Starbek, M. (2008). Project management of product development. Strojniški vestnik – Journal of Mechanical Engineering, vol. 54, no. 9, p. 588-606.
The Strength of the Bus Structure with the Determination of Critical Points
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[5] Gnosa, S., Bartha, E. (2004). Testing protocol III. Neoplan Bus GmbH. [6] Lazović, M. (2004). Measuring protocols for airport bus Neoplan 11.11.2004, Car Factory Maribor. [7] Lazović, M. (2005). Integral testing of airport buses, 7th Conference and exhibition of Inovative automobile technology, IAT'05. Bled, 21.-22. April, p. 271-281. [8] Chen, L.F. (2004). Comparative analysis for bus side structures and lightweight optimization. ImechE, Proc. Instn. Mech. Engrs., vol. 218, part D: J. Avtomobile Engineering. [9] Mc.Manus, K.J., Mann, A., Evans, R.P. (1998). The analysis of road roughness and the perception of road roughness. Proceedings of 5th international symposium on heavy vehicle weights and dimensions in conjunction with 5th engineering foundation conference on vehicle-infrastructure interaction, Maroochydore, Queensland.
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[10] Matolcsy, M. (2001). Body section rollover test as an approval method for required strength of bus superstructures. Automotive and Transportation Technology Congress and Exhibition (P-371), Barcelona. [11] de Aquiar, F.H.V., Gimenez, M.C., Pazian, A., Spinelli, D.M. (2002). Frame structure optimization for bus chassis. SEA International. [12] Sharp, J., Pesheck, E., Nelson, D., McLellan, G.A. (2000). Orion bus industries: Virtual prototyping of a transit bus to predict service life. 15 International ADAMS User Conference, Rome ,Italy, 15.nov. 2000 [13] Gu, P., Slevinsky, M. (2003). Mechanical bus for modular product design. CIRP Annals-Manufacturing Tehnology, vol. 52, no. 1, p. 113-116. [14] Bartha, E. (2005). Virtual prototyping of a transit bus to predict service life. MAN Richtlinien, München.
Mežnar, D. - Lazovič, M.
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 551-556 UDC 620.173.25:621.742.48
Paper received: 28.02.2008 Paper accepted: 29.10.2009
Pressure Stimulated Current Emissions on Cement Paste Samples under Repetitive Stepwise Compressional Loadings Antonios Kyriazopoulos*, Ilias Stavrakas, Konstantinos Ninos, Cimon Anastasiadis, Dimos Triantis Laboratory of Electric Properties of Materials, Department of Electronics, Technological Educational Institution of Athens, Greece The electric signals detection technique that is described here was applied on several geomaterials in the past and on cement based materials lately. In this work cement paste samples were studied regarding electric signal emissions during axial stress application processes and specifically when the samples were subjected to repetitive loadings and unloadings in the range where crack opening and propagation processes are established. It was observed that the electric signal was emitted in two stages. Initially, current was emitted simultaneously with the stress step in the form of a spike which gradually returned to its background level. A secondary current emission was recorded while the stress was maintained constant at the high level of each stress step. ©2010 Journal of Mechanical Engineering. All rights reserved. Keywords: Pressure Stimulated Current; electric current emissions; cement paste; uniaxial compressional stress 0 INTRODUCTION The study of the properties of cement products has spurred scientific interest as they constitute the main structural materials. Several techniques have been applied in order to monitor the health of cement constructions. Some of them involve the electrical properties of cement. Since health monitoring does not provide the flexibility to extract cement samples from the constructions, non-destructive testing methods are the most suitable for application. For many years it has been known that electric and electromagnetic (EE) signals can be observed when solids and especially non-metallic materials, are mechanically stressed. Such signals have been reported by [1] to [5]. Micro- and macro-cracking processes are often accompanied by these signals. Several mechanisms for the EE signal generation have been discussed in literature. Rapid movement of electric charges, separation of electric charges at crack formation and their recombination to form a miniature spark discharge, rapid movement of electric double layers under the action of the mechanical loading or piezoelectric phenomena are some which have been reported and studied by [2], [6] to [9]. In previous works, processes of electrical emission in rock samples like marble and amphibolite were studied by [10] to [14]. The emitted current during the temporal stress
variation that leads to catastrophic processes in the bulk of the samples and finally to their fracture, has been rendered under the term Pressure Stimulated Current (PSC). The technique applied to detect and record such electric emissions is mentioned as the PSC technique. The relevant literature refers to electric signal emissions observed with similar techniques regarding electrical emission in mortars under low compressive loading [15] and the appearance of an electric current that increases nonlinearly with compressional stress [9]. During this work cement paste samples were subjected to stress adequate to lead them to the Crack Propagation Zone (CPZ). For the used samples this zone is estimated to be reached with compression at around 11 MPa. Consequently, repetitive mechanical loadings and unloadings were applied on the samples. Between each loading and unloading the stress level was maintained at its high value for a relatively long time. The emitted PSC during this process was measured and is presented and discussed here. 1 EXPERIMENTAL CONFIGURATION A set of cement paste samples was prepared for the measurements. The dimensions of the samples were 40 x 40 x 40 mm. The proportion of the contents of the OPC (Ordinary Portland Cement) was 2:1. The drying time of the
*
Corr. Author's Address: Laboratory of Electric Properties of Materials, Department of Electronics, Technological Educational Institution of Athens, Athens, 12210, Greece, akyriazo@teiath.gr
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samples was 90 days. Conducting preliminary systematic tests with various stress modes such as monotonically increasing stress at a constant rate, or maintaining high stress levels for a long time, or applying sequential stepwise stress increments up to failure, the average ultimate compressional strength of the sample was found to vary around 25 ± 5 MPa. Consequently, samples prepared from the same mixture were used to conduct the experiments.
Strain [%]
Fig. 1. A representative curve that describes the relative compressive stress with respect to strain for the used samples The relative compressional stress value (Fig. 1) is given as ˆ / max where max corresponds to the ultimate compressional strength of the sample. It is evident that it can be characterized by a linear behavior at least up to a
stress of approximately 80%, of the ultimate compressional strength (i.e. ˆ 0.8 ). When ˆ 0.8 approximately, the material is driven to a range of non-linear deformation and eventually into the localized failure zone. Fig. 2 shows the experimental installation. For the implementation of this experimental technique a pair of gold plated copper electrodes was attached at the perpendicular axis of the stress. The measurements were recorded using a Keithley electrometer (model 6514). Electric measurements were stored in a computer hard disk through a GPIB interface while the load cell and the strain gages bridge were guided to an A/D Keithley DAQ. The stressing system comprised a uniaxial hydraulic load machine (Enerpac– RC106) that applied the load to the samples. The experiments were conducted in a Faraday shield to prevent electric noise. The sample under test was slowly loaded up to a value of approximately 50% of the ultimate compressional stress strength. Consequently, a stress increase was applied at a relatively high rate and the stress maintained its high value for 10 min. Afterwards the stress was removed until the level of 50% of the ultimate stress strength was reached. This procedure was repeated three times. Consequently, the stress was further increased in the vicinity of the fracture and after some time the sample failed without further increasing the stress. During this entire process the emitted PSC was recorded.
axial stress electrodes Data Acquisition through GPIB
sample
load cell keithley 6514
Fig. 2. Schematic diagram of the experimental setup
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 551-556
2 RESULTS AND DISCUSSION
Stress [MPa]
Fig. 3 shows the temporal variation of the three repetitive mechanical loadings and the corresponding emitted PSC. Specifically, the upper plot (a) corresponds to the applied compressional stress and is graded in MPa, while the three lower (b, c and d) are graded in pA and correspond to the temporal variation of the emitted PSC during the three repetitive loadings. Two kinds of PSC emissions can be seen in Fig. 3. Specifically, a primary current that is emitted simultaneously with the stress increase from the lower to the higher level is observed. This current is restored relatively fast. A secondary current emission that takes place while the stress is maintained practically constant is also observed. It is obvious that both PSC emissions during each following loading become weaker. The primary PSC emission is attributed to the crack formation and propagation processes
that are measured by means of the corresponding deformation. The reduction of the peak value of the PSC can be attributed to the electric emission memory effect that has already been discussed and interpreted in previous works that refer to PSC emissions from rock samples like marble [14] and [16] and amphibolite [13]. Fig. 3 also shows the secondary PSC emissions that take place after stabilizing the stress at the corresponding high level of its stress increase. The stress level that was maintained after each stress step was approximately 16.5 MPa. It becomes obvious that despite the fact that there is no stress variation, a significant PSC is emitted. This can be attributed to the fatigue of the sample due to the opening of new cracks formation or propagation of the existing ones since the material is already in the Crack Propagation Zone (CPZ).
PSC [pA]
t [s]
PSC [pA]
t [s]
PSC [pA]
t [s]
t [s]
Fig. 3. a) Plot of a representative stress step which after a 3-fold repetition produced primary and secondary PSC emissions during the first; b) the second; c) and the third; d) stress step, respectively
Pressure Stimulated Current Emissions on Cement Paste Samples under Repetitive Stepwise Compressional Loadings
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contrast to the PSC emission of the first stress step (Fig. 4c) so that the changes in the form of the emitted PSC slightly before fracture become clear. The deformation, after the application of a compressional stress step, continues to increase (hysteresis). This phenomenon becomes more intense as the sample reaches the ultimate stress strength. The above findings become obvious in Fig. 5 where the temporal variation of the emitted PSC (Fig. 5a) and the corresponding temporal development of the deformation are depicted (Fig. 5b). Here the deformation continues to increase despite the fact that the stress is maintained practically constant. During this process the PSC becomes more intense due to the fact that the strain increases. Slightly before failure and while the strain increases at a gradually higher rate a brushlike PSC indicates the upcoming failure.
Stress [MPa]
Fig. 4 depicts the last stress step (Fig. 4a) that was performed from 16.5 up to 22 MPa approximately. The sample suffered this loading for 5 min and consequently it failed due to fatigue. During this compressional stress step a significant PSC emission was observed and lasted until the sample failed. During this final stress step the primary PSC emission cannot be distinguished from the secondary and they seem to overlap. The fact that the duration of the emissions is long (from the stress change to fracture) and has a high magnitude in combination with the fact that the primary emission was never restored, are the factors that predict the fracture of the sample which finally took place at the time tf = 270 s (see Fig. 4a). Another observation is that before the sample fracture the secondary emission has a brush-like form introducing the upcoming fracture. This prior-to-fracture PSC emission (Fig. 4b) is put in
PSC [pA]
t [s]
PSC [pA]
t [s]
t [s]
Fig. 4. a) The final stress step in the vicinity of fracture and b) the corresponding emitted PSC; (c) the PSC emission during the initial stress step at lower stress level in order to observe the time window between the primary and the secondary PSC emissions
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Strain [%]
StrojniĹĄki vestnik - Journal of Mechanical Engineering 56(2010)9, 551-556
t [s]
Fig. 5. a) The temporal variation of the PSC during the last stress step and b) the corresponding temporal recording of the deformation 3 CONCLUDING REMARKS
4 REFERENCES
Cement paste samples were studied using the PSC technique. The recorded PSC showed a primary emission detected during the stress increase and a secondary one recorded while the stress was maintained constant at the higher level of each stress step. The PSC emissions are attributed to the crack formation and propagation processes and the consequent deformation. The experimental results were discussed according to this theoretical background. The primary PSC emission was simultaneously with stress attributed to the deformation increase and the secondary emission was attributed to the deformation hysteresis mechanisms. Another experimental observation was a lower value that the PSC reached after each stress application and this result was put in contrast to similar previous results recorded and discussed for geomaterials like marble and amphibolite. In conclusion, the PSC technique and the qualitative characteristics of the emissions observed can become a significant factor in monitoring the health state of cement paste using a non destructive method.
[1] Enomoto, J., Hashimoto, H. (1990). Emission of charged particles from indentation fracture of rocks. Nature, vol. 346, p. 641-643. [2] Nitsan, U. (1997). Electromagnetic emission accompanying fracture of quartz-bearing rocks. Geophys. Res. Lett., vol. 4, p. 333-337. [3] Ogawa, T.K., Miura, T. (1985). Electromagnetic radiation from rocks. J. Geophys. Res., vol. 90, p. 6245-6249. [4] O’Keefe, S.G., Thiel, D.V. (1995). A mechanism for the production of electromagnetic radiation during fracture of brittle materials. Phys. Earth Planet. Int., vol. 89, p. 127-135. [5] Vallianatos, F., Tzanis, A. (1998). Electric current generation associated with the deformation rate of a solid: Preseismic and coseismic signals, Physics and Chemistry of the Earth, vol. 23, p. 933-938. [6] Brady, B.T., Rowell, G.A. (1986). Laboratory investigation of the electrodynamics of rock fracture. Nature, vol. 321, p. 448-492. [7] Vallianatos, F., Triantis, D., Tzanis, A., Anastasiadis, C., Stavrakas, I. (2004). Electric earthquake precursors: from laboratory results to field observations. Physics and Chemistry of the Earth, vol. 29, p. 339-351.
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[8] Varotsos, P., Alexopoulos, K. (1986). Thermodynamics of point defects and their relation with bulk properties. North-Holland, Amsterdam. [9] Sun, M., Li, Z., Song, X. (2004). Piezoelectric effect of hardened cement paste. Cement & Concrete Composites, vol. 26, p. 717-720. [10] Stavrakas, I., Anastasiadis, C., Triantis, D., Vallianatos, F. (2003). Piezo Stimulated currents in marble samples: Precursory and concurrent – with – failure signals. Natural Hazards and Earth System Sciences, vol. 3, p. 243-247. [11] Anastasiadis, C., Triantis, D., Stavrakas, I., Vallianatos, F. (2004). Pressure stimulated currents (PSC) in marble samples after the application of various stress modes before fracture. Annals of Geophysics, vol. 47, p. 2128. [12] Stavrakas, I., Triantis, D., Agioutantis, Z., Maurigiannakis, S., Saltas, V., Vallianatos, F., Clarke, M. (2004). Pressure stimulated currents in rocks and their correlation with
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[13]
[14]
[15]
[16]
mechanical properties. Natural Hazards and Earth System Sciences, vol. 4, p. 563-567. Triantis, D., Anastasiadis, C., Vallianatos, F., Kyriazis, P., Nover, G. (2007). Electric signal emissions during repeated abrupt uniaxial compressional stress steps in amphibolite from KTB drilling. Natural Hazards and Earth System Sciences, vol. 7, p. 149-154. Anastasiadis, C., Triantis, D., Hogarth, C.A. (2007). Comments on the phenomena underlying pressure stimulated currents (PSC) in dielectric rock materials. Journal of Materials Science, vol. 42, p. 2538-2542. Sun, M., Liu, Q., Li, Z., Wang, E. (2002). Electrical emission in mortar under low compressive loading. Cement and Concrete Research, vol. 32, p. 47-50. Kyriazis, P., Anastasiadis, C., Triantis, D., Vallianatos, F. (2006). Wavelet analysis on pressure stimulated currents emitted by marble samples. Natural Hazards and Earth System Sciences, vol. 6, p. 889-894.
Kyriazopoulos, A. – Stavrakas, I. – Ninos, K. – Anastasiadis, C. – Triantis, D.
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 557-564 UDC 544.57:539.3
Paper received: 22.08.2009 Paper accepted: 04.03.2010
Air-Coupled Lamb and Rayleigh Waves for Remote NDE of Defects and Material Elastic Properties Igor Solodov* - Daniel Döring - Gerd Busse Institute for Polymer Technology, NDE Group (IKT-ZfP), Germany Conventional air-coupled ultrasound (ACU) is a well-established tool for acoustic NDT and material characterization. Its major shortcoming is concerned with a week penetration into solid materials due to a severe impedance mismatch at the air-solid interface. A dramatic rise in acoustic coupling is obtained by using acoustic mode conversion into plate and surface waves in slanted configurations. In our experiments, an increase of the ultrasound amplitude by up to one order of magnitude was observed in various materials (metals, wood, concrete, composite) under phase matching conditions. On this basis, fully air-coupled configurations are developed and applied for non-contact NDT. The methods based on this principle enable precise measurements of fibre directions and quantification of in-plane anisotropy in composites and natural materials, elastic depth profiling, drying of coatings, advanced imaging of cracked defects and delaminations. ©2010 Journal of Mechanical Engineering. All rights reserved. Keywords: air-coupled ultrasonic testing, Lamb and Rayleigh waves, material characterization, elastic anisotropy, elastic depth profiling, process monitoring, NDT imaging 0 INTRODUCTION Air-coupled ultrasound has become a routine tool for non-destructive testing and material characterization [1]. The conventional through-transmission or normal transmission mode (NTM) is based on conversion of the incident ACU into longitudinal acoustic waves which propagate through the bulk of material and interact with defects. A modified configuration of the focused slanted transmission mode (FSTM) [2] employs the ACU conversion into plate acoustic waves (PAWs, also known as “Lamb waves”) which are found to be highly sensitive to surface-breaking cracked defects and delaminations as well as applicable to evaluation of specimen properties like thickness, stiffness and in-plane anisotropy. The interaction of the PAW with a thin liquid or solid layer on the plate substrate can be used for real-time monitoring of elastic properties of films and coatings, for example of drying of paint. The PAW amplitude excited depends strongly on the angle of incidence: the resonance PAW generation at the optimal angle enhances substantially the signal-to-noise ratio of the ACUNDT systems. The inverse PAW-ACU conversion results in the radiation of a pair of ultrasonic waves leaking into air at the same (optimal) angles from both surfaces of the
specimen. The wave radiated on the excitation side is used in the focused slanted reflection mode (FSRM) in which both sending and receiving transducer are positioned on the same side of the object under inspection. Despite some scale distortion due to an elongated probing area, this mode enables flexible single-sided scanning of large specimens. The single-sided configuration provides an opportunity to expand the family of the ACUNDT methods based on the mode conversion to include surface acoustic waves (SAWs, also “Rayleigh waves”). These waves propagate within a thin surface layer (~ one wavelength) of solids, thus being useful for selective testing of this area for its elastic properties and defects. Since SAWs exhibit no dispersion in homogenous semi-infinite solids, any change in phase velocity as a function of frequency is either an indication of inhomogeneity (surface hardening, inhomogeneous porosity, etc.) or a violation of the “thick-against-the-wavelength” condition for the substrate (e.g. sub-surface voids and delaminations). The use of PAWs and SAWs in NDT&E is not a new development, but existing solutions for the excitation/detection normally require contact configurations (wedge transducers, phased arrays), which are also limited to rather fast materials (SAW velocity > 1500 m/s). Other
*
Corr. Author's Address: Institute for Polymer Technology, NDE Group (IKT-ZfP), University of Stuttgart, Pfaffenwaldring 32, D-70569 Stuttgart, Germany, igor.solodov@ikt.uni-stuttgart.de
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techniques are based on materials with special properties (conducting (EMAT) or piezo-electric (interdigital transducers)) that confines substantially the area of guided wave applications in NDT. Much greater flexibility is obtained with the air-coupled FSRM-methodology which does not require particular material qualities, contact or couplant for the transducers, and extends the range of materials for evaluation down to those with SAW velocity as low as ~ 400 m/s. The new materials inspected with SAW in this paper include technical and natural fibre composites, plastics and metals. For the NDT application, a scanning FSRM system was developed and applied to a wide range of materials. 1 METHODOLOGY 1.1 Principle: Phase-Matched Coupling
guided wave velocity which carries information on the elastic material parameters. The SAW velocity is dominated by the ratio of sheer modulus G and mass density with minor contribution of Poisson’s ratio . An approximate expression for the Rayleigh wave velocity has the form [4]:
vSAW
G.
(2)
The PAW family includes a number of different modes which are all dispersive in the frequency-thickness product. For low frequencies / thin plates, the lowest-order antisymmetric mode, which is of the highest importance for ACU testing, the velocity dispersion can be approximated using a flexural-wave approach [4]:
vao 4
An efficient mode conversion (air-coupled ultrasound into transversal bulk waves, PAWs, SAWs, etc.) can only occur if the projection of kvector for each of the waves on the plane of incidence takes on the same value.
0.87 1.12 1
E D . 3 (1 2 ) 2
(3)
In many cases, the contribution of Poisson’s ratio in the denominator is neglected so that Eq. (1) can be used for deriving the value of the material Young’s modulus. 1.2 Application: Focused Slanted Transmission Mode (FSTM)
Fig. 1. Schematic representation of the coincidence rule for phase-matched coupling In the case of a guided wave to be exited, this can be regarded as a coincidence of the wave fronts at the surface (see Fig. 1) [3] which takes place for particular (resonance) angle of incidence Θo :
sin o
vair .
vguided
(1)
This relation shows that the value of the resonance angle can be used to measure the
558
The primary method to apply the mode conversion is based on the slanted transmission using two air-coupled transducers in a co-axial setup (Fig. 2a). This is a suitable configuration for a rotation around the measurement point in the specimen to adjust the incidence angle to obtain maximum transmitted wave amplitude (Fig. 2b). If the specimen is also rotated in the azimuth plane, such an adjustment made for each value of the azimuth angle (Fig. 2c) enables to quantify and map the in-plane stiffness anisotropy on the basis of Eq. (1). 1.3 Application: Focused Slanted Reflection Mode (FSRM) A propagating plate wave radiates aircoupled ultrasound from both sides of the specimen, so that this signal can also be detected on the excitation side. As the specular reflection is normally much stronger than the re-radiated signal, a beam shield of absorbent material positioned close to the specimen surface is used
Solodov, I. - Döring, D. - Busse, G.
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 557-564
(Fig. 3). For longer propagation paths, the signals can also be separated in time domain: the PAW velocity is usually in the range of 600 to 2000 m/s as compared to the 340 m/s for ultrasound in air. The single-sided configuration also provides a remote access to the Rayleigh wave which does not penetrate into the bulk of a thick specimen and thus can only be excited and detected in the single-sided setup.
Fig. 3. Setup for FSRM scanning 1.4 Application: Air-Coupled Time of Flight (DTOF)
Differential
The accuracy of guided wave velocity measurements can be significantly increased by tracing the wave along the specimen surface. When the receiver in the FSRM configuration is shifted parallel to the surface in the direction of wave propagation, only the guided wave path changes (Fig. 4a).
a)
a) b)
c) Fig. 2. a) FSTM co-axial configuration, b) transmitted amplitude as a function of incidence angle in paper, c) bi-axial rotation for mapping in-plane stiffness anisotropy
b) Fig. 4. a) A scheme and b) laboratory setup for air-coupled time-of-flight measurements Thus, the measurement of the relative phase difference in the received signal and
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the change in propagation path x is sufficient for a precise calculation of the guided wave phase velocity at the frequency f:
v ph 2f
x .
(4)
In the experiment, the phase shift is determined from the recorded A-scans by using a discrete Fourier transformation. This enables to recover the value confidently even from data with a low signal-to-noise-ratio. 2 MEASUREMENTS AND RESULTS All experiments used commercial aircoupled ultrasonic testing equipment with piezocomposite transducers and a 3-axis stepped scanning table. High-voltage (200 V) square wave bursts were used for excitation of aircoupled ultrasound in the frequency range of 200 to 450 kHz. The higher harmonic components of the input signal were filtered out due to a low bandwidth of the transducers (~ 10 %) so that the ultrasonic signal was monochromatic. 2.1 FSTM for Non-Destructive Air-Coupled Ultrasonic Imaging Surface breaking tight cracks are known to be a rather difficult NDT subject for bulk-wave ultrasonics. This is confirmed by the air-coupled NTM image presented in Fig. 5 (left): the contrast of the image produced by a longitudinal wave propagating parallel to the crack faces is
measured to be only ~2%. On the contrary, the flexural wave in the FSTM mode (Fig. 5, right) is scattered by the crack much stronger resulting in much higher contrast (~ 80%) and signal-to-noise ratio of the image. 2.2 Experimental Study of Slanted-Mode Wave Conversion Efficiencies The mode conversion in slanted configuration is expected to enhance the ultrasound penetration into solid materials due to the resonance excitation of PAW/SAW. To measure the efficiency of mode conversion, the output signal in the FSTM/FSRM configuration was compared with the amplitude of the ACU transmitted directly between the transmitter and receiver. For ACU-PAW conversion, the additional losses strongly depend on the value of Θo (Table 1). The data of the table confirm that mode conversion provides a substantial enhancement in elastic coupling compared with conventional NTM. The gain obtained increases along with the values of Θo, which indicates a contribution of space resonance. In fact, the length of the excitation area (in wave-lengths) changes as (W / air )tg 0 and rises sharply for large values of Θo. The measurements of ACU-SAW mode conversion were carried out in the FSRM configuration at 390 kHz and a SAW propagation distance of ~6 cm.
550
1400
540
1200
530
800
520
600
510
200
NTM
FSTM
Fig. 5. C-Scan images (top) and amplitude distributions across the images (bottom) for a surface breaking crack in polycarbonate measured in normal and slanted transmissions
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Table 1. Experimental results of ACU-PAW mode conversion (frequency 450 kHz) Thickness [mm]
Material
Resonance Angle 0
PAW velocity a
[degrees]
[m/s]
0
Gain Conversion Conversion (NTM- FSTM) losses losses [dB] (NTM) [dB] (FSTM) [dB]
Paper I (A4)
0.1
52 ± 1
430 ± 10
44
20.5
23.5
Paper II
0.17
48 ± 1
460 ± 10
47
28.5
18.5
Al-foil
0.1
32 ± 1
640 ± 20
53.5
38.5
15
Polystyrene
1.1
21 ± 1
950 ± 60
64.5
53
11.5
Wood (spruce, L/LT)
0.65
19 ± 1
1040 ± 60
53
43
10
Table 2. Experimental results of ACU-SAW mode conversion Conversion losses SAW acoustic (FSRM) [dB] impedance (Mrayl)
0 [degrees]
SAW [m/s]
Fir (L/LT)
17 ± 0.5
1160 ± 30
37
0.9
PMMA
16 ± 0.5
1230 ± 40
40
1.5
Graphite
14 ± 0.5
1400 ± 50
45
2.1
Concrete
9 ± 0.5
2200 ± 100
61
5.3
Aluminum
7 ± 0.5
2800 ± 200
54
7.8
Copper
10 ± 0.5
2000 ± 100
59
18.8
Steel
6.5 ± 0.5
3000 ± 200
63
23.4
Material
Fig. 6. SAW velocity anisotropy a) in the RL-plane of beech, b) in unidirectional CFRP
Air-Coupled Lamb and Rayleigh Waves for Remote NDE of Defects and Material Elastic Properties
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Thus, a few dB of dissipation should be subtracted from the values of conversion losses given in Table 2. All values of Θo are smaller in this case, so that impact of the length of the excitation area is insignificant. Instead, a good correlation between conversion losses and the value of acoustic impedance of materials for SAW is observed. The only deviation in the case of concrete is apparently due to higher propagation losses (due to scattering in multi-scale inhomogeneities) in this material. Table 3. SAW velocities measured by DTOF at 200 kHz Material Orientation α vSAW [m/s]
CFRP
Wood (Beech, RL- cut)
0°(in fibre)
2160 ± 20
22.5°
1950 ± 20
45°
1550 ± 10
67.5°
1395 ± 10
90° (across)
1369 ± 7
0° (L)
1216 ± 15
22.5°
1048 ± 15
45°
815 ± 12
67.5°
727 ± 12
90° (R)
697 ± 12
2.3 Experimental Study of SAW Anisotropy Using the DTOF methodology, the SAW velocity was measured in a natural (wood) and an engineering material (unidirectional CFRP laminate) with a high elastic anisotropy as a function of propagation direction (azimuth angle ) relative to the fibres. The degree of the velocity anisotropy becomes more apparent when the measured data is plotted in a polar diagram (see Fig. 6). In both cases it reflects the reinforcement characteristics induced by the fibres with an almost two-fold increase in velocity which implies approximately four-fold increment in material stiffness.
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2.4 Air-Coupled Single-Sided Imaging for Non-Destructive Testing A number of light-weight aerospace components comprise sandwich-type structures consisting of light (and relatively weak) core materials (foam, honeycomb, balsa wood) which join adhesively the high-strength CFRP liners. The adhesion of the honeycomb to the liners is critical, as it is the only factor that prevents them from buckling under compressive load. The FSRM with a plate wave propagating in the CFRP liner provides a technique for remote testing of adhesion to the core. Two specimens containing artificial delaminations between the CFRP liners and two types of honeycomb structures were fabricated and tested with air-coupled ultrasound. The FSRM scans (Fig. 7) clearly show the delamination-simulating inserts as well as some core cells filled with epoxy resin. All defects in the image are stretched in the scanning (horizontal) direction due to initial distance between the transducers (~5 cm) characteristic for the FSRM setup. 2.5 Non-Destructive Testing with Air-Coupled SAW Natural sandstone is a material in which a macro-scale elastic anisotropy is hardly expected. Nevertheless, it might be produced by the sedimentation process involved in the formation of the stone, with either deposition of layers of different properties or a dominant orientation of non-symmetrical individual particles. The direction of increased stiffness is found to be parallel to the sediment layers (0° in Fig. 8, sometimes visible as colour pattern), resulting in a SAW velocity increase by 10% compared to the weak (90°) direction (Table 4). A change in SAW frequency increases the penetration depth of the wave, but does not significantly change the SAW velocity on a freshly-cut surface (Table 4). If the surface is infiltrated with a silicabased resin to improve the properties ̶ mainly to harden it against pollution corrosion ̶ the SAW velocity is increased by 20 to 30%, depending on frequency and propagation direction.
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Table 4. DTOF measurements of SAW velocities of variegated sandstone as a function of propagation direction and surface treatment vSAW vSAW Side Orientation at 200 kHz at 200 kHz [m/s] [m/s] Freshly cut Hardened
0°
1850 ± 20
1820 ± 30
90°
1670 ± 20
1690 ± 30
0°
2240 ± 10
2400 ± 40
90°
2220 ± 10
2300 ± 40
From Table 4 it is also evident that the infiltration process mostly cancels the cause of the elastic anisotropy of the sedimentation structure, which suggests that the small amount of additional substance equalizes the contact between the individual grains. In addition, a closer comparison of the velocities at 200 and 400 kHz shows a somewhat higher value for the shorter wavelength (400 ~ 6 mm). This can be considered as an estimate of the hardened layer thickness; the 200 kHz wave penetrates deeper (200 ~ 12 mm) and in the bulk of the material that results in a slower SAW velocity. By using a wider range of probing frequencies, a non-contact elastic depth profiling should be possible.
2.6. Remote Process Monitoring: Drying of Paint The SAW/PAW velocity and dissipation also depend on stiffness and viscosity of coatings. As a result, the FSTM- and FSRM- output signals are sensitive to changes in the physical state of films and coatings (hardening, polymerization, drying, etc.). Both amplitude and phase of air-coupled SAW/PAW can be used for real time non-contact monitoring of such processes on-site in an industrial environment. An example of SAW application for the monitoring of paint drying is illustrated in Figure 9. It shows that drying of identical paints develops differently for concrete and PMMA substrates. In concrete, the reaction proceeds more intensively (higher values of phase derivative in Fig. 9, b) and faster (drying time ~4500 s against ~8000 s in PMMA) because the paint solvent not only evaporates, but also diffuses into the porous cementitious material. 3 CONCLUSIONS The slanted mode methodologies with mode conversion to plate and surface waves were shown to enhance significantly the efficiency of penetration of air-coupled ultrasonic waves in solid materials as compared to the established normal transmission mode. hardened
freshlycut
Teflon inserts
epoxy filled honeycomb cells
Fig. 7. PAW FSRM C-scan (200 kHz) of a carbon-composite specimen with aluminium (left) and Nomex (right) honeycomb core; teflon foils between core and CFRP liner and resin-filled core cells simulate production defects; specimen provided by FACC, Austria
sedimentation layers (=0°)
Fig. 8. Sedimentation in natural stone
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SAW output phase [degrees]
200
100
PMMA 0
Concrete -100
Phase derivative [degree/s]
0.15 300
Concrete
0.10
PMMA 0.05
0.00 -200 0
2000
4000
6000
0
8000
1000 2000 3000 4000 5000 6000 7000 8000
Time [s]
Time [s]
a) b) Fig. 9. Air-coupled SAW monitoring of drying paint: a) output phase; b) and its derivative variations in time The use of air-coupled flexural waves improves the detection of surface-breaking cracks while a single-sided non-contact configuration enables to image critical delamination-type flaws in sandwich compounds. Differential time-of-flight measurement of guided wave velocities is a sensitive instrument for an evaluation of elastic anisotropy and depthresolved profiling. Non-contact excitation/detection of aircoupled guided waves is a basis for remote and non-invasive monitoring of fluid-solid phase transitions (drying, hardening, etc.) applicable in an industrial environment.
4 REFERENCES [1]
[2]
[3]
[4]
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Rogovsky, A.J. (1991). Development and application of ultra-sonic dry-contact and air-contact C-scan systems for nondestructive evaluation of aerospace components. Material Evaluation, vol. 50, p. 1491-1497. Solodov, I., Stößel, R., Busse, G. (2004). Material characterization and NDE using focused slanted transmission mode of aircoupled ultrasound. Research in NonDestructive Evaluation, vol. 15, p. 1-21. Cremer, L. (1947). About the analogy between angle of incidence and problem of frequency. Archive of Electric Transmission, vol. 1, no. 28. (in German) Victorov, I.A. (1967). Rayleigh and lamb waves. Physical theory and applications. Plenum Press, New York.
Solodov, I. - Döring, D. - Busse, G.
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 565-574 UDC 621.182
Paper received: 29.01.2008 Paper accepted: 31.08.2010
Robust IMC Controllers with Optimal Setpoints Tracking and Disturbance Rejection for Industrial Boiler Dejan D. Ivezić1,* Trajko B. Petrović2 University of Belgrade, Faculty of Mining and Geology Engineering, Department for Mechanical Engineering, Serbia 2 University of Belgrade, Faculty of Electrical Engineering, Department for Control Engineering, Serbia 1
Robust controllers based on Internal Model Control (IMC) theory are developed in this paper to improve the robust performance of industrial boiler system against uncertainties and disturbances. A simplified model of a boiler’s drum unit is developed and transfer matrix realization of its dynamics is obtained for a nominal operational condition. Controllers parameters are selected in accordance with the frequency domain optimization method based on -optimality frameworks. The proposed controllers are robust for reference signals and/or for disturbances. Finally, a comparison between the performances of the closed-loop system with designed IMC controllers is obtained. ©2010 Journal of Mechanical Engineering. All rights reserved. Keywords: industrial boiler, robust control, internal model control 0 INTRODUCTION The main aim of this work is to present the problem and to devise a method for designing robust IMC controllers of industrial boiler subsystem using the frequency domain optimization method based on -optimality framework. Various control techniques have been applied to boiler or boiler–turbine controller design, e.g., inverse Nyquist array [1], linear quadratic Gaussian (LQG) [2], LQG/loop transfer recovery (LTR) [3], mixed-sensitivity approach [4], loop-shaping approach [5], and predicative control [6]. The -optimality framework takes into account the system's nominal plant model, incorporating real and complex uncertainties, which describe interested parameter variation range to ensure that the closed-loop with the controller is stable with a certain degree of performance over all possible plants. A brief introduction to the robust control theory and its application on the distillation column, dc/dc converters and solid-fuel boiler are given in [7] to [15]. The boiler studied is an industrial boiler system with a normal steam production of 8.7 kg/s and with an outlet steam pressure and a steam temperature of 18×105 Pa and 400C. Set of nonlinear equations for describing the boiler’s subsystem (steam-water part) dynamics is presented. The Boiler model, in the form of transfer functions matrixes is developed by *
linearization around the operating point. For this multivariable model, IMC controllers (IMCr,0, IMC0,d, IMCr,d) are designed. The controllers are proposed for three opposite goals. The first controller (IMCr,0) is designed for optimal setpoint tracking, the second (IMC0,d) for optimal disturbance rejection and the third (IMCr,d) for optimal overcome of the trade-off between these opposite demands. The final goal is to compare the robustness of closed-loop systems with IMCr,0, IMC0,d, IMCr,d controllers using frequency analysis and to verify the results using transient analysis. 1 THE PROCESS AND ITS MODEL In the literature, modeling of boilers has been treated in many different ways, [16] to [20]. The boiler process consists of water heater, steam drum, downcomers tubes, mud drum, riser tubes, and superheater (Fig. 1). However, in this paper only the steam-water part (i.e. steam drum, downcomers and risers) is taken into account (Fig. 2), because the water heater and the super heater system are weakly coupled to the steamwater system and it is natural to treat the three systems separately. The input variables are the powder coal flow rate, the feedwater flow rate and the steam flow rate. The output variables are the drum level and the drum pressure.
Corr. Author's Address: University of Belgrade, Faculty of Mining and Geology, Djušina 7, 11000 Belgrade, Serbia, ivezic@rgf.bg.ac.rs
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Feedwater
Superheater steam
Water heater
Steam
d (Vds ds ) X 0 M 0 M s . dt
Steam drum
As the volume of the drum is constant, an increase of steam volume results in a decrease of water volume and vice versa:
2
water
d d Vds Vdw 0 . dt dt
Risers
Downcomers
(5)
A combination of Eqs. (1) and (4) gives: d 1 ds [( M 0 M w M dow M s ) dt Vdt Vdw
Mud drum
Fig. 1. Industrial boiler system Ms
d ( dwVdw H dw ) M 0 (1 X 0 ) H rw dt M dow H dow M w H ewo .
ds, Vds Mw Hdw, dw, Vdw M0, X0, Hrm
Mcoal, Qrw
Hrw
dwVdw
Fig. 2. A simplified description of the steam-water part of the boiler system The mass balance of the water in the drum determines the dynamics of the water mass: d 1 Vdw ((1 X 0 ) M 0 M w M dow ) . (1) dt dw The drum water level is given by:
,
(6)
(7)
It is presumed that the feedwater temperature is constant, i.e. Hevo = const. This is a rational proposition as the water heater has its own control system. The combination with Eq. (1) gives the dynamics of the drum water enthalpy:
Mdow, Hdow
Vd Vdw Ad hd
( ds dw ) d Vdw ]. dw dt
From the steam table, for known ρds, it is possible to find a corresponding drum pressure. The water in the drum is not in the saturation state, and the energy balance is:
pd
(2)
where Vd is the reference volume of water in the drum at nominal point. The mass flow rate of steam condensing in the drum is neglected, that is: d d (Vdw dw ) dw Vdw . (3) dt dt The dynamics of the steam density is taken from the mass balance in the drum:
566
(4)
d H dw M 0 (1 X 0 )( H rw H dw ) dt M w ( H ewo H dw ).
(8)
Water density, ρdw is determined with an assumption of a saturation condition and ρdw is then the function of drum pressure. The energy balance of the steam-water mixture in the raisers is: d ( rm H rm )Vr M dow H dow dt (9) Qrw M 0 ( H rw X 0 r ) . Hrw and r (evaporation heat) are functions of drum pressure. Heat flow to the risers is assumed to be a function of the powder coal flow rate: Qrw k rw M coal .
(10)
The enthalpy of steam-water mixture is a function of the steam quality:
Ivezić, D.D. - Petrović, T.B.
H H rw ( H rs H rw ) X
.
(11)
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 565-574
Hrs is a function of the drum pressure and if steam quality is considered as linearly distributed along the raisers, the average enthalpy in raisers is: rX H rm H rw 0 . (12) 2 Eq. (9) can now be written as: d 1 H rm [Qrw M dow ( H rm H dow ) dt Vr rm X M 0r 0 ] . 2
(13)
The density of the steam-water mixture is given by: 1 1 1 1 X( ). (14) rm rw rs rw Mass flow rate in downcomers is shown by the Bernouli’s equation: M dow k c dw rm ,
(15)
where kc represents the inverse of the circulation losses. Steam-water flow at the top of the risers (M0) can be obtained from mass and energy balance of the risers. To simplify, it is assumed that it can be approximated by the empiric expression: M 0 M dow M 0 k1M oil k3 M s ,
(16)
where ΔM0 is transient contribution to M0: d 1 M 0 k1M oil k 2 M s M 0 , (17) dt T where time (T) and gain (k1, k2) factors are load dependent, and can be estimated from plant recordings. The set of nonlinear Eqs. (1), (6), (8), (13) and (17) presents state space description of the industrial boiler subsystem. The state vector’s elements are: hd ̶ drum water level; Hdw ̶ drum water enthalpy; ds ̶ drum steam density; Hrm ̶ average enthalpy of steam-water mixture in risers; M0 ̶ transient contribution to the steam-water flow rate in risers. Transfer matrix, as a description of boiler subsystem dynamics is obtained by linearization for nominal working conditions, given in Table 1. ~ (18a) y P ( s )u Pd d ,
Table 1. Boiler’s working conditions Steam drum Input water flow rate Mw [kg/s] Input water enthalpy Hewo [kJ/kg] Water level hd [m] Volume Vd [m3] Water surface size Ad [m2] Water volume Vdw [m3] Pressure pd [Pa] Water density
dw [kg/m3]
Water temperature
Tdw [C] Hdw [kJ/kg]
Water enthalpy Steam density
8.7 463 0.75 9.54 8.10 4.76 20105 849.90 212.37
ds [kg/m ]
908.5 10.04
3
Steam flow rate Downcomers Water enthalpy Water flow rate
Ms [kg/s]
8.70
Hdow [kJ/kg] Mdow [kg/s]
2,100 0.18
Risers Water enthalpy Steam quality Volume
Hrw [kJ/kg] X0 [kg/kg] Vr [m3]
2,100 0.75
Mixture enthalpy Mixture flow rate Transient contribution
Hrm [kJ/kg] M0 [kg/s]
M0 [kg/s]
0.03300 2,326 0.18 0.11
Heat flow Coal flow rate
Qrw [kW] Mcoal [kg/s]
16,208 1.866
0.048s ~ ( s ) P ( s) 0.386s ( s )
1.119s 2 0.8247s 1000 ( s) , 1.92s 2 1.71s 0.2007 1000 ( s)
( s ) 100 s 3 10.8s 2 0.08s ,
(18b)
(18c)
0.0002437 s 0.008 Pd ( s ) 0.0409 (18d) ( s 0.006) , s ( s 0.1) The following notations are used: Output vector y = [y1 y2]T y1 – pd, drum pressure; y2 – hd, drum-water level; Input vector u = [u1 u2]T
Robust IMC Controllers with Optimal Setpoints Tracking and Disturbance Rejection for Industrial Boiler
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u1 – Mcoal, fuel flow rate; u2 – Mw, water flow rate; Distrubance d – Ms, steam flow rate. The goal of the control law is to generate u1 and u2 so that it maintains y1 and y2 close to setpoint r = [y1ref y2ref]T and insensitive to disturbance d. 2 THE MODEL UNCERTAINTY The description of the model uncertainty arises from the fact that process plant operates with certain flow rates (coal and water) on its inputs. Changes on inputs are effected by servocontrolled valves, which rely on the measurement of the flow. Existing of 1% error in flow measure produces 10% error in required variation [8]. Thus, our plant model, which describes changes about some operating point is subject to errors of up to 10% on each input channel. Since the error on each input channel is independent of the others, the suitable representation of the disagreement between the real plant P and model ~ P is described by a multiplicative input perturbation L and structured model uncertainty ~ description of multivariable model P . P P ( I L) P ( I lu u ), (u ) 1. (19) lu = diag(l,l), l = 0.1, represents the uncertainty weighted operator (the frequency dependent magnitude bound of u) and u = diag (1, 2) is unknown unity norm bounded block diagonal perturbation matrix that reflects the structure of the uncertainty. Also, such description of uncertainty covers the neglected heat capacity of the riser metal, i.e. it was included in the uncertainty of fuel flow.
the following term of weighted sensitivity operator: 50 s 1 W p 0.25 I. (21) 50 s The weight (21) implies that we require an integral action (Wp(0) = ) and allow an amplification of disturbances at high frequencies by a factor four at most (Wp() = 0.25). A particular sensitivity function, which matches the performance bound (21) exactly for low frequencies, is 200 s E I. (22) 200 s 1 This corresponds to a first order response with time constant 200 s. 4 IMC IN THE -OPTIMALITY FRAMEWORK The IMC structure developed [7] as an alternative to the classic feedback structure. Its main advantage is that closed-loop stability is guaranteed simply by choosing a stable IMC controller. This concept is based on an equivalent transformation of the standard feedback structure into IMC structure shown in Fig. 3a. The synthesis and analysis of robust IMC controllers, based on the structured singular value approach, impose a forming interconnection matrix by rearrangement of the block diagram of the IMC control structure shown in Fig. 3a in general G-Δ form (Fig. 3b) necessary for analysis. The interconnection matrix G in Fig. 3b is partitioned into four blocks consistent with the dimensions of the two input and the two output vectors: G12 G G 11 . G21 G22
3 THE PERFORMANCE OBJECTIVE The sensitivity weighting operator Wp is selected by a designer to give a preferred shape to the sensitivity operator E. The feedback system satisfies robust performance if the -norm of the weighted sensitivity operator is unity bounded for any perturbation u of the plant: Wp E
sup (W p E ) 1 , for any u .
(20)
The input vector of G consists of the outputs from the uncertainty block Δu and the desired external input v (setpoints, disturbance or both). The output vector of G is formed by the inputs to the uncertainty block Δu and the weighted error e’. Form of matrix Δ is:
The required limiting values of the closedloop time constant of the closed-loop system give
568
(23)
Ivezić, D.D. - Petrović, T.B.
u 0
0 , p
(24)
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 565-574
d' Controller C r
e -
u
d
~ P
Plant P
WP e'
Pd
lu
u
Q ~ P
p u
y
G e’
v
b)
a)
Fig. 3. The block diagram of IMC control structure a) and G-Δ form; b) in accordance with -optimality framework where Δp is a full unity norm bounded matrix ( p ) 1 . The -optimality framework adopts measures of robust stability (RS) and robust performance (RP) as suitable objectives [7] and [10], which define the performance of the multivariable feedback system in the presence of structured uncertainty: RS rs ( ) G11 sup (G11 ) 1 , (25)
RP rp ( )
G
sup (G ) 1 .
(26)
The operator is a structured singular value (-norm) computed according to the blockdiagonal structure of [7], [8] and [10] -norm is the natural extension of -norm when both the bound and the structure of model uncertainty are known. The upper bound of (G) is defined as
(G ) inf ( DGD 1 ) , D
(27)
where D is any real positive diagonal matrix with the structure diag(diIi) where each block (i.e., the size of Ii) is equal to the size of the blocks i. Ideally, the goal is to find a controller C that satisfies Eqs. (25) and (26). These objectives ensure stability and performance of the closedloop system in the presence of all expected uncertainties. The demand is not always reachable, especially with controllers of simple structure. The convenient objective for the synthesis of the robust controller C using -norm is min G C
,
(28)
G11
1.
(29)
Thus, the optimal controller minimizes the performance index RP, i.e. the -norm of the weighted sensitivity operator for all possible plants. 5 THE OPTIMAL PERFORMANCE The design of the robust IMC controller is a two-step design procedure. In the first step it is ~ assumed that the model is perfect (P = P ) and an ~ optimal controller Q that minimizes a performance objective is designed. It is implied that an H2-optimal (minimum variance) controller should be selected, which minimizes the nominal performance, i.e. the 2-norm of the weighted ~ nominal sensitivity operator E . . min W E min W ( I PQ (30) Q
p
2
p
Q
2
~ As the nominal plant P is stable but the nonminimum phase transfer function, it has to be factored into a stable all pass portion PA and a minimum phase portion PM such that: (31) P = PAPM . ~ The controller Q that solves Eq. (30) satisfies [7]: ~ Q PM1W p1 W p PA1 . (32)
Here the operator {.} denotes that after a partial fraction expansion of the operand all terms involving the poles of PA-1 are omitted.
with constrain:
Robust IMC Controllers with Optimal Setpoints Tracking and Disturbance Rejection for Industrial Boiler
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~ In the second step of the algorithm, Q has to be augmented by a low-pass filter F for robustness: ~ (33) Q QF .
The structure of F is predetermined and the parameters should be selected in order to minimize the RP measure. In this approach the following definitions ~ of the nominal sensitivity operator E are adopted, according to choices of appropriate inputs of the system. The transfer function of ~ nominal sensitivity operator E r ,0 referring to ~ setpoints (v=r), E0,d referring to disturbance ~ (v=d), and E r ,d referring to all inputs v=(rT, dT)T are defined in the following terms: ~ ~ ~ Er , 0 ( I P Q ) E , ~ ~ ~ E0, d ( I P Q) Pd EPd ,
(34) (35)
) I P E I P . E r ,d ( I PQ d d
The optimal design problem is: min G Q
,
(40)
with constraint on robust stability: ~ QP lu 1.
(41)
6 THE IMC FILTER The structure of filter F is fixed and with a few tuning parameters adjusted to obtain desired robustness properties Eqs. (40) and (41). If the zero steady-state error for step inputs is required, then it is necessary that F(s=0)=I. The simple filter with a unity steady-state gain and with four tuning parameters is: (k1s 1) n (k s 1) n 2 F 2 0
, (k 3 s 1) n (k 4 s 1) n 2 0
(42)
(36) ~ E r ,0 is adopted for optimal setpoint ~ tracking, E0,d for optimal disturbance rejection ~ E r ,d is adopted to optimal overcome of the trade-
where k1, k2, k3 and k4 are the filter’s tuning parameters and the n selected is large enough to make Q proper. In this case n = 1 is selected and the corresponding controller Q is:
off between setpoint tracking and disturbance rejection. The corresponding interconnection matrix Gr,0, G0,d and Gr,d referring to sensitivity operators (34), (35) and (36) are obtained as follows: ~ QP lu Q Gr , 0 (37) ~~ ~ , W p EP lu W p E
(43)
~ QP lu G0,d ~~ W p EP lu Gr , d
~ QP lu ~~ W p EP lu
QPd ~ , W p EPd
(38)
QPd ~ . ~ W p E W p EPd
(39)
Q
k1s 1 ~ (k s 1) 3 Q Q 2 0
. k3 s 1 (k 4 s 1) 3 0
7 IMC CONTROLLER DESIGN
The controllers (43) are designed using [21] and [22]. The measures of performance of the closed-loop system rp with all here designed controllers, using Gr,0, G0,d and Gr,d respectively are computed and shown in Fig. 4. The values of robust stability (RS) and robust performance (RP) are given in Table 2.
Table 2. Performance and stability measures of closed-loop system with IMC controllers Controller k1 k2 k3 k4 RS RP(Gr,0) RP(G0,d) IMCr,0 12.10 4.33 16.64 3.64 0.19 0.97 9.60 IMC0,d 300.86 27.87 213.23 44.48 0.50 2.74 0.61 IMCr,d 82.28 72.76 195.10 40.37 0.25 1.28 2.24
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RP(Gr,d) 6.81 2.74 1.34
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 565-574
3
3
2.5
gain
a)
gain
2.5
gain
b)
c)
2.5
2
IMC0,d 2
2 1.5
IMC0,d
1.5 1
1.5
IMCr,d IMC0,d
1
IMCr,0
-3
10
10
-2
10
-1
frequency [rad/s]
10
0
10
1
0 -4 10
IMCr,0
1
0.5
0.5 0 -4 10
IMCr,d IMCr,0
IMCr,d
0.5
-3
10
10
-2
10
-1
frequency [rad/s]
10
0
10
1
0 -4 10
-3
10
-2
10
10
-1
frequency [rad/s]
10
0
1
10
Fig. 4. Frequency responses of rp( ) evaluating for a) Gr,0; b) G0,d; c) Gr,d It is evident that robust stability condition is satisfied for all the proposed controllers. According to the interconnection matrix used for robust performance determination, optimal results are obtained with different controllers, i.e. IMCr,0 is optimal for RP(Gr,0), IMC0,d for RP(G0,d) and IMCr,d for RP(Gr,d) measure. Such results were expected, but it is interesting to analyze the change of RP measure over the frequency range of interest. High values of RP for IMCr,0 controller with an inclusion of disturbance in the interconnection matrix are characteristic for a very low frequency range, so the consequence could be visible in a steady state. Over the higher frequency range the RP characteristic of this controller is quite similar to the IMCr,d controller. The IMC0,d controller has the maximum of the RP measure in the range 0.01 to 0.1 rad/s, for interconnection matrixes with included reference signals, so the consequence could be expected in a transient response. The IMCr,d controller has overall optimal values for all the frequency range and for all interconnection matrix. 8 THE TRANSIENT ANALYSIS Computer simulations are performed to evaluate the performance of the proposed robust controllers. Time responses of closed-loop system, using IMCr,0, IMC0,d, IMCr,d are compared in order to observe how the system tracks setpoint changes and rejects external disturbance. The most typical time responses of outputs of nominal and perturbed plant are shown in Figs. 5 to 7. For a satisfactory confirmation of evaluated robust performances, time responses for the worst-case assumption are computed and shown for the following cases:
~ 1.1 0 P(s) P (s) , 0 0.9
(44)
~ 0.9 0 P(s) P (s) (45) . 0 1.1 As maximal perturbations are used, simulations only have a theoretical importance, because a reasonably large variation which shifts the operation of the system from its nominal point exists in those cases. However, such time domain simulations are the best tools for a remarkably good confirmation of computed robust performances. In real cases, uncertainties will be smaller, so that the corresponding time responses will have similar though less distinctive features.
9 DISCUSSION Time responses of nominal plant are shown in Fig. 5. The achievement of diagonal dominance in the reference tracking due to a specific structure, Eqs. (32) and (33), of IMC controllers is evident. Nominal system behavior is optimal for reference tracking with IMCr,0 controller, i.e. with IMC0,d controller for disturbance rejection. The behavior of the system with IMCr,d is somewhere between them, but for the nominal plant all proposed controllers are acceptable. Time responses of perturbed plants are shown in Figs. 6 and 7. Again, the IMCr,0 controller is optimal for reference tracking and IMC0,d controller for disturbance rejection, but the problem with those two controllers is their unsatisfactory characteristics for the opposite mission. The IMCr,0 controller used to reject disturbance is not able to provide zero steady state errors. Transient response in reference
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tracking with IMC0,d is with significant overshoot. Physically this means that the system with the IMCr,0 controller will not be able to achieve insensitivity of drum water level to change in the steam flow rate and that the system with IMC0,d controller will have a significant (and possibly unacceptable) variation in drum pressure and drum water level, especially with a change of fuel flow rate.
The correlation of perturbed plant time response with robust performance measures is evident. The highest rp measures of the IMC0,d controller with Gr,0 and Gr,d have the obvious influence on time response in reference tracking. Also, the worst rp measure with G0,d at low frequencies is confirmed in problems with disturbance rejection.
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Fig. 7. Time responses of perturbed plant (45) to unity step signal: a and b) in input u1; c and d) in input u2; e and f) in input d The IMCr,d controller is the best solution for overcoming the trade-off between setpoint tracking and disturbance rejection. Time responses of perturbed plant with that controller confirm such a conclusion. It is obvious that IMCr,d is somewhat slower than IMCr,0 for setpoint tracking, i.e. has longer settling time than IMC0,d for disturbance rejection, but without their disadvantage in initial variations and steady-state error. 10 CONCLUSIONS Steam-water part (i.e. a steam drum, downcomers and risers) of an industrial boiler system has been modeled and the input model uncertainty is defined. The controllers are designed using -analysis control theory and they achieve robustness against uncertainties and disturbance. The designed controllers are robust for setpoints and/or disturbance. The tuning parameters are selected so that they minimize the robust performance objective. The performances of the closed-loop system with the designed controllers are evaluated by simulations and compared. The achieved results have shown remarkable robustness of the proposed controllers i.e. the closed-loop stability and a satisfactory degree of performance over all the possible plants. Using IMC framework, optimal controller switches effectively overcome the trade-off
between setpoint tracking and disturbance rejection over a wide range of possible plants that are investigated. REFERENCES [1]
[2] [3]
[4]
[5]
[6]
Johansson, L., Koivo, H.N. (1984). Inverse Nyquist array technique in the design of a multivariable controller for a solid-fuel boiler. International Journal of Control, vol. 40, no. 6, p. 1077-1088. Cori, R., Maffezzoni, C. (1984). Practical optimal control of a drum boiler power plant. Automatica, vol. 20, no. 2, p. 163-173. Kwon, W.H., Kim, S.W., Park, P.G. (1989). On the multivariable robust control of a boiler-turbine system. Proceedings of Symposium on Power Systems and Power Plant Control, Seoul, p. 219-223. Pellegrinetti, G., Bentsman, J. (1994). H controller design for boilers. International Journal of Robust and Nonlinear Control, vol. 4, no. 5, p. 645-671. Tan, W., Marquez, H.J., Chen, T. (2002). Multivariable Robust Controller Design for a Boiler System. IEEE Transactions on Control Systems Technology, vol. 10, no. 5, p. 735-742. Hogg, B.M., Rabaie, N.M.E. (1991). Multivariable generalized predictive control
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[7] [8] [9] [10] [11] [12]
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of a boiler system. IEEE Transactions on Energy Conversion, vol. 6, no. 2, p. 282-288. Morari, M., Zafiriou, E. (1989). Robust Process Control. Englewood Cliffs/Prentice Hall, New York. Maciejowski, J.M. (1989). Multivariable Feedback Design. Addison-Wesley, Wokingham. Doyle, J.C., Francis, B., Tannenbaum, A. (1991). Feedback control theory. Macmillan, New York. Zhou, K., Doyle, J.C., Glover K. (1996). Robust and Optimal Control. Prentice Hall, New Jersey. Green, M., Limbeer, D.J.N. (1996). Linear robust control. Prentice Hall, New York. Petrović, T., Vasilić, S. (1997). Robust decentralized controllers with disturbance rejection for parallel operating DC/DC converters. Control and Computers, vol. 25, no. 3, p. 80-87. Petrović, T., Ivezić, D., Debeljković, D. (1999). Multivariable frequency response methods for designing robust decentralized control of a solid-fuel boiler. Engineering simulation, vol. 16, p. 671-687. Petrović, T., Ivezić, D., Debeljković, D. (2000). Robust IMC controllers for a solidfuel boiler. Engineering simulation, vol. 17, p. 211-224.
[15] Moradi, H., Bakhtiari-Nejad, F., SaffarAvval, M. (2009) Robust control of an industrial boiler system; a comparison between two approaches: Sliding mode control & H∞ technique. Energy Conversion and Management, vol. 50, p. 1401-1410. [16] Rhine, J.M., Tucker, R.J. (1991). Modeling of Gas-Fired Furnaces and Boilers. McGraw-Hill, New York. [17] Flynn, M.E., O’Malley. M.J. (1999). Drum boiler model for long term power system dynamic simulation. IEEE Transactions on Power Systems, vol. 14, no. 1, p. 209-217. [18] Astrom, K., Bell, R. (2000). Drum-boiler dynamics. Automatica, vol. 36, p. 363-378. [19] Kim, H, Choi, S. (2005). A model on water level dynamics in natural circulation drumtype boilers. International Communications in Heat and Mass Transfer, vol. 32, p. 786796. [20] Rusinowski, H., Stanek, W. (2007). Neural modelling of steam boilers. Energy Conversion and Management, vol. 48, p. 2802-2809. [21] Chiang, R.Y., Safonov, M.G. (1992). Robust Control Toolbox. MA-USA: The Math Works Inc. [22] Balas, G.J., Doyle, J.C., Gloven, K., Packard, A., Smith, R. (1993). -Analysis and Synthesis Toolbox. MA-USA: The Math Works Inc.
Ivezić, D.D. - Petrović, T.B.
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Paper received: 01.12.2009 Paper accepted: 02.03.2010
Alternative to the Conventional Heating and Cooling Systems in Public Buildings Mitja Košir* - Aleš Krainer - Mateja Dovjak - Rudolf Perdan - Živa Kristl University of Ljubljana, Faculty of Civil and Geodetic Engineering, Slovenia The paper presents an alternative system for heating and cooling in public buildings. The system was designed for the retrofitted building of the Slovene Ethnographic Museum (SEM) where it was also extensively tested. The installed system includes radiant wall mounted panels for heating and cooling, localized automated tangential fans for cooling and ventilation and a centralized building management system for the regulation and supervision of the performance. The efficiency of the system was thoroughly investigated through a series of experiments conducted prior to the renovation of the building as well as after the museum was put into service. The application of the described system resulted in substantial reduction of energy consumption, better internal thermal conditions and lower investment costs for the Heating, Ventilation and Air Conditioning (HVAC) system of the entire building. © 2010 Journal of Mechanical Engineering. All rights reserved. Keywords: heating, cooling, ventilation, low temperature system, radiant panels 0 INTRODUCTION The paper presents a system for indoor temperature regulation with the use of lowtemperature radiant heating/cooling panels and automated natural ventilation. Extensive experimentation with the low-temperature wall mounted heating/cooling panels was conducted before and during the retrofitting of the Slovene Ethnographic Museum (SEM) in Ljubljana, Slovenia [1]. The in-situ experimental results, simulations and later measurements of the building performance in real time conditions proved high efficiency of the system. The wall mounted low-temperature radiant heating and cooling panels present an alternative to the air heating and air cooling systems originally proposed for the museum building. Because the majority of heat transport in the heated or cooled spaces equipped with low-temperature systems is conducted by radiation and not by convection, a smoother temperature profile preferred by the majority of users is achieved [2]. Lowtemperature heating systems that operate close to the environmental temperatures are in addition to low energy also low exergy systems [3], although the use of high exergy fuels (e.g. electricity or fossil fuels) where low exergy work is needed somewhat reduces this effect. The previously proposed mechanical centralized ventilation was replaced by a localized automated ventilation
system utilizing small tangential fans integrated into the window sills. The system enabled the necessary physiological ventilation during museum opening hours as well as cooling via the night ventilation.
Fig. 1. North-east external view of the retrofitted museum building The geometry of the museum building is presented in Fig. 2. Exhibition spaces equipped with panels are located in the east wing of the ground floor and on the 1st, 2nd and 3rd floors of the building; the total floor area is 2575 m2, while the floor area of the entire building is 5214 m2. The existing exterior walls were composed of external rendering applied to a 50 cm thick brick wall with internal rendering removed. Floors
*
Corr. Author's Address: University of Ljubljana, Faculty of Civil and Geodetic Engineering, Chair for Buildings and Constructional Complexes, Jamova cesta 2, Ljubljana, Slovenia, MKosir@fgg.uni-lj.si
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were mostly brick vaulted, gravel filled and finished with wood decking. 1 CONCEPT OF HEATING/COOLING AND VENTILATION SYSTEM The main objectives of the SEM – Museums project were to assure optimal conditions for exhibitions and for the storage of museum’s exhibits in accordance with international standards, to assure appropriate environment for visitors from the visual and from the thermal point of view and rational use of energy without reducing the quality of functional use. When studying various options for achieving these goals, the decision was made that the problems had to be dealt with holistically (and not every aspect of internal conditions solved with a separate system) and that the rational use of energy for heating and cooling was to be achieved by reduction of operating costs controlled by a building management system (BMS). There were seven main spheres of activities resulting in the framework of the following interventions: thermal energy with heating and cooling part, ventilation, daylight (not presented in this paper), control and management, constructional complexes, simulations, testing and measurements.
1.1 Heating and Cooling System The design and performance of the wall mounted low-temperature heating/cooling system progressed through four distinct experimental phases. The conducted measurements spanned over four years, from those executed prior to the building renovation to those carried out during the first year of building operation [4] and [5]. The first phase of measurements (conducted during the summer and autumn months of 2000) encompassed the recording of values for the visual and thermal environment, while the building was un-refurbished and in “free run”. The measurements showed that during summer season internal temperatures were maintained mostly within the 18 to 25 °C zone. Thermal mass of the building compensated for night lows (T 10 K) and day highs (T 4 K). During this time the possibility of installing a low-temperature heating and cooling system was discussed. The decision was made to investigate the effectiveness of such a system by conducting experiments in real environment of the building. The experience based on preliminary measurements and simulations also put in question the necessity of an air conditioning system proposed in the original design. During the second phase of experiments
Fig. 2. 1st floor of the renovated SEM building
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(conducted from the end of autumn 2000 till March 2001) the response of test rooms to various heating modes was tested. Two experimental test rooms were completely renovated in the SW corner in the 1st and 2nd floors (the position of the rooms relative to the retrofitted state of the building is shown in Fig. 2). Both rooms were additionally thermally insulated and outfitted with new double glazed windows with low-e coatings as well as equipped with electrically powered heating units (conventional electrical radiators), as shown in Fig. 3. Measurements and computer simulations showed that annually the system would use 10 kWh/m2 of heating energy less than the proposed air system [1], [4] and [5], if an alternative radiant system was to be used. In time for the beginning of the third phase of experiments (conducted from the end of March 2001 till July 2002) the prototypes of wall mounted heating/cooling panels were constructed and installed in the renovated test room on the 1st floor (dubbed the model room), while the second room on the 2nd floor (dubbed the reference room) remained the same as in the second phase (Fig. 4). Thorough measurements of wall panels were executed under real time environmental conditions. Example of results acquired during winter testing is shown in Fig. 5. For the model room the average energy consumption for heating was 12% lower than in the reference room. Difference in energy consumption between the two test rooms is the result of different temperatures of heating media and modes of heat transport. The cooling energy consumption was measured during two summer seasons, from August 2001 till June 2002. Different set-point temperature series were tested: 24 h/day continuous cooling and two modes of intermittent cooling, from 08 to 20 h with 22.5 and 25 oC setpoint temperatures, respectively. Collected results were derived into seasonal specific cooling energy consumption between 10 kWh/m2 (setpoint 25 °C) and 15 kWh/m2 (set-point 22.5 °C) for intermittent cooling and 25 kWh/m2 (set-point 25 °C) to 30 kWh/m2, (set-point 22.5 °C) for continuous cooling. Such low set-point temperature was defined in order to test the cooling performance of the system and any possible occurrence of surface condensation. The acquired data showed important differences in energy consumption between the reference room
and the model room for heating and small specific energy consumption for cooling.
Fig. 3. Reference room on the 2nd floor of the SEM building heated with conventional radiator system
Fig. 4. Model room on the 1st floor equipped with the prototype heating/cooling wall mounted panels In Fig. 5 daily temperature profiles for the model room during winter period (December) in intermittent heating mode are presented. The outside air temperatures were very low, between -15 and -3 oC which is low even for Ljubljana. In the diagram also the temperature profile for the reference room exhibiting a “wave” pattern is presented. In the model room the temperature profile is smooth and follows very well the prescribed set-point temperature profile. The experiments showed that the panels reacted well in winter, summer and mid-season conditions and consistently maintained the indoor temperatures close to the set-point temperatures without any difficulties.
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Fig. 5. Winter time experiment conducted in the model room during the 25th of December; shown are: external air temperature (Toutside), internal air temperature in the model room (TIAverage), vertical surface temperature distribution at the bottom of the panels (measuring nodes 7/1, 7/3, 7/5, 7/7, 7/9), core temperature of the panel (5/3), inlet (3/1) and outlet (3/2) water temperatures; as a reference the internal air temperature in the reference room (TIIAverage) and the set-point temperature (Tset-point) are shown 1.2 Ventilation System For the ventilation of exhibition rooms new automated and localized natural ventilation system was designed. Small tangential fans were integrated into the window sills (Fig. 6) of new windows and were controlled with the BMS. The integration of the ventilation system with the window was necessary due to the minimal impact on the building exterior prescribed by the strict conservation standards. The window integrated ventilation system is used for the necessary physiological ventilation during opening hours and for cooling purposes (night cross ventilation), when the conditions are
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favourable. The cooling by ventilation system is harmonised with the functioning of the wall cooling system. Generally the wall cooling system is activated if the outside conditions do not enable cooling of the building with ventilation. The ventilation system also uses the microclimatic conditions surrounding the building in a way that optimizes the use of fresh air on internal comfort conditions. This means that during heating season the air is supplied from the south façade and expelled on the north side of the building. The situation is reversed during cooling season, as the supplied air is taken from the north side and expelled on the south side of the museum (Fig. 7).
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2 FUNCTIONING OF SYSTEM IN THE BUILDING 2.1 Heating and Cooling In the starting phase of the project no requirements for thermal insulation of buildings under historical monument preservation protection were foreseen. Firstly, the decision was made to place a 10 cm thermal insulation with the corresponding vapour barrier and plaster boards on the inner side of the existing outer brick wall. On one hand, this reduced the vast thermal mass of the building and on the other hand it reduced the U value of the outer wall from 1.16 W/m2K to 0.30 W/m2K. With this intervention several benefits were gained. First, quick thermal response of the building was achieved, which enables effective intermittent heating and cooling. At the same time the outer side of the protected facade was not touched and it could retain its original structure an appearance. Second, lowtemperature wall mounted heating/cooling vertical system was used. Third, non-manageable thermal mass of the original wall was excluded from the wall’s thermal conduction transport system, but at the same time it was replaced by a designed thermal mass in the reinforced concrete wall panels separated from the other parts of the outer wall structure. Forth, thermal comfort was improved due to the surface temperature to air temperature relation and lateral radiation effect. Fifth, consideration of the new design of combined heating/cooling wall panel system resulted in the decision to omit the designed air conditioning system. This sets free 158 m2 of space for depository area, and the investment was reduced by about 100,000 to 150,000 euros. 2.2 Ventilation Each floor is divided into two zones (east and west zones). A set of tangential fans is integrated into the sills of windows on the north and south side of the building. In zone 1 (ground floor, SE part of the building) there is a CO2 sensor, triggering physiological ventilation during working hours when critical levels (700 ppm) are reached. The functioning of the ventilation in other parts of the museum is operated according to time dependent ventilation protocols that were
Fig. 6. Scheme of the tangential fan integrated into the window sill
Fig. 7. Functioning of the localized ventilation system according to the seasonal microclimatic conditions (W – winter operation, S – summer operation, continuous line – supplied air, dotted line – expelled air) derived from the measurements of CO2 concentration in the SE part of the ground floor. The fans of the ventilation system are also used for cross ventilation and cooling of exhibition rooms in the case of convenient outside temperature and relative humidity conditions. Cooling with ventilation is enabled when the external air temperature is 1 K lower than the internal set-point air temperature. If the cooling by ventilation is not sufficient (the internal air temperature is 1 K higher than the internal set-point air temperature), the system switches to the wall mounted cooling panels. In this case the fans are activated according to the physiological ventilation protocols.
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2.3 Control and Management Computer simulations of the building’s energy consumption in the pre-retrofitted state showed the annual value of 156 kWh/m2 of energy consumption for heating and cooling. This number was used as reference for the evaluation of energy performance of the building after the proposed interventions had been carried out. It was evident that the energy efficiency of the building could be improved with the application of the proposed interventions to the building envelope (installation of thermal insulation) and by using low-temperature heating and cooling system. If these interventions were considered in the TRNSYS [6] simulations, the reduction of heating and cooling energy consumption of 46.5% could be achieved in comparison to the reference state (Table 1). Simulation results predicted an annual energy consumption of 73 kWh/m2 for heating and 10.5 kWh/m2 for cooling, totalling at combined consumption 83.5 kWh/m2 annually. The simulated energy consumption for the ventilation was predicted at 2.37 kWh/m2 annually. For the purposes of BMS the exhibition area with the total floor surface of 2884 m2 is divided into seven zones: the East wing of ground floor and in the 1st, 2nd and 3rd floors (East and West zone in each floor), which are separately controlled by BMS. The scheme of the central control system and an example of opening BMS screen for heating-cooling panel system are presented in Figs. 8 and 9, respectively. The heating system of the building is connected to the city district heating system. It is divided into 7 control zones (East part and West part of the building in each floor). Temperature/time/season sensitive BMS system is used and enables establishing of different setpoint temperature profiles during opening and non-opening hours. During heating season the inlet water temperature was 35 °C with occasional peaks reaching 40 °C at times of extreme loads. For the cooling the wall mounted panels are connected to a common cooling plant (McQuay AGF-XN 070.2, cooling power: 218 kW, electric 88 kW, 2 compressors, 4 steps of 25, 50, 75, 100%). The temperature of cooled inlet water was typically kept around 15 °C with occasional lows of 11 °C. The same division of spaces as for heating is used for the control of
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cooling. Temperature/time/season sensitive control wall mounted system is supplemented and harmonized with the ventilation cooling system. The activation of cooling panels is enabled when the external conditions do not allow cooling of the building via ventilation. Both systems are linked and harmonized. The possibility of manual override is foreseen for all zones. In addition to cooling purposes the localized automated ventilation system is also used for physiological ventilation of the museum. For physiological requirements daily/weekly regime of performance is executed. During visiting hours the air exchange level is set to 0.5 h-1. This means that all fans in zones 2 to 7 (1st, 2nd and 3rd floor, East and West wings) are switched on every 15 minutes for the duration of 15 minutes. When the museum is closed, the ventilation is switched off. The East wing on the ground floor (zone 1) is ventilated according to the levels of CO2 concentration.
Fig.8. Scheme of the BMS
Fig.9. BMS screen with controls of the heatingcooling panels for the East wing of the ground floor
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Table 1. Results of simulations and measurements conducted during 2004 for the annual energy consumption of the exhibition part of the SEM; the pre-retrofitted reference state (simulated) energy consumption for heating and cooling was 156 kWh/m2 annually SIMULATED RESULTS [kWh/m2] Qh Qc Qv Qh + Qc Jan. 17.0 0.0 0.18 17.0 Feb. 12.0 0.0 0.18 12.0 Mar. 9.0 0.0 0.18 9.0 Apr. 3.3 0.0 0.18 3.3 May 1.0 0.0 0.18 1.0 Jun. 0.0 2.1 0.25 2.1 Jul. 0.0 4.2 0.25 4.2 Aug. 0.0 4.2 0.25 4.2 Sep. 0.0 0.0 0.18 0.0 Oct. 6.0 0.0 0.18 6.0 Nov. 9.0 0.0 0.18 9.0 Dec. 16.0 0.0 0.18 16.0 Annual 73.0 10.5 2.37 83.5 Annual reduction compared to the reference state [%] -46.5
The BMS permanently controlled and collected the quantities: Microclimate and indoor comfort: ambient air temperature, humidity and lighting levels. Energy systems: heating consumption (district heating, each zone separately and total consumption), cooling consumption (electricity), lighting consumption (electricity), total electricity consumption. Measurements of the whole building performance that were performed during the year 2004 were collected using installed BMS (Johnson Controls Metasys with FX15 Controllers) and the following sensors/meters: • Air temperature: JC Series A99 sensors. • Air humidity: JC Series HT-9000 sensors. • Heating/cooling energy: Allmess, type CF Echo. • Electricity: Iskra Instruments, d.d., type WS1202. • CO2: Siemens, type QPA63.1. Protocols for on-line monitoring of the control system during the operation of the building were prepared for different day-night, summer-winter regimes, for different sources: heating-cooling panels for heating and cooling function, physiological ventilation, for the combination of cooling and relative humidity with corresponding descriptions of interventions in tabular form. The following information is available and stored by the BMS during the operation of the building: • Review of conditions on PC: outside air temperature and relative humidity, temperature and relative humidity in zones, temperature of
Qh 21.0 12.7 5.6 1.3 1.0 0.0 0.0 0.0 0.0 1.6 3.8 4.4 50.4
MEASURED RESULTS [kWh/m2] Qc Qv Qh + Qc 0.0 0.06 21.0 0.0 0.06 12.7 0.0 0.06 5.6 0.0 0.06 1.3 0.0 0.06 1.0 0.0 0.06 0.0 5.6 0.19 5.6 5.6 0.19 5.6 0.0 0.06 0.0 0.0 0.06 1.6 0.0 0.06 3.8 0.0 0.06 4.4 11.2 0.98 61.6 -60.5
heating medium by zones, energy use (heating, cooling), daily/seasonally by zones, electrical energy use daily/seasonally by zones, condition of panels and ventilators, both for the part of the building treated in the framework of the museums project and for the other part of the SEM exhibition building. • Possibilities of data storage on PC: outside temperature and relative humidity (hourly average), temperature and relative humidity by zones (hourly average), temperature of heating medium by zones (hourly average), energy use (heating, cooling), full hourly data, electrical energy use, full hourly data, condition of panels and ventilators. There were two different energy use patterns during the measurements conducted in 2004, the first in the beginning and the second at the end of the year. In the first part energy consumption was higher by 8.6% in the part of the museum with installed wall panels than in the rest of the building. During this period the controller maintained constant heating media temperature for 24 hours per day for the whole building. In the second part of the heating season the wall panels functionality was optimized by the BMS which resulted in 5 times smaller consumption in December 2004 compared to January 2004 (Fig. 10). 3 CONCLUSION Low-temperature radiative heating and cooling systems represent an efficient solution for creating good thermal environment. Low-
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temperature systems enable the transport of heat through radiation and this eliminates the problems of user discomfort due to annoying air movement [2]. In addition to better thermal comfort of users such systems also exhibit improved energy efficiency due to utilization of lower temperatures of heating and/or cooling medium, which results in direct energy savings due to better boiler efficiency and lower thermal losses of the entire system [7]. Nonetheless, systems that utilize low temperatures of heating and cooling medium have special configuration compared to conventional systems. The most obvious difference is that large surfaces have to be heated or cooled for efficient functioning. In the case of the SEM optimal relation between air temperature and surface temperature in the museum building is achieved with the use of heating/cooling wall mounted panels. They represent the main intervention in the framework of the construction. The system is connected to the district heating system for heating purposes in winter and to cooling plant for cooling purposes in summer. Window integrated BMS controlled ventilation system (small tangential fans) is used for the necessary ventilation during opening hours and for cooling purposes (night ventilation), harmonised with the wall cooling system. Due to “high risk” nature of the proposed innovative heating/cooling system a series of
experiments were conducted in various phases of its development. On the basis of these findings the decision for using the proposed system in the exhibition area of the building was adopted. After the execution of the planned interventions and the installation of the heating/cooling and ventilation system, the building was put into operation. The performance of the building was closely monitored during the whole first year (2004) of the museums operation. The actual measured energy consumption of the heating system was even lower than had been indicated by the computer simulations, as the exhibition area that encompasses approximately one half of the museum floor spaces consumed only 14% of the total heating energy consumption of the whole building. The most important, difficult and problematic part of the project was design and tuning of harmonised control of temperatures, relative humidity, CO2 and cooling oriented ventilation with the application of central control system designed specially for this project. In the end the gross energy demand for heating and cooling is reduced by 60.5% (Table 1), from 156 kWh/m2 annually (simulated pre-renovation state with presumed continuous heating) to 61.6 kWh/m2 annually (measured - combined energy consumption for heating and cooling).
Fig. 10. Heating energy consumption of the low-temperature wall panels installed in the exhibition space compared to the overall consumption of the whole SEM building
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The average measured consumption in similar buildings is usually more than 140 kWh/m2 annually (based on simulated cases). The selection of a localized automated ventilation system integrated under the window resulted in a negligible quantity of only 0.98 kWh/m2 of energy used for ventilation purposes per year. The quantity of the foreseen blown-in air was reduced from 36500 m3/h (predicted in the original project) to 10000 m3/h (implemented system). The power of heating station was reduced by 110 kW and for cooling by 62 kW. The result of the project is also the reduction of the investment budget in the field of HVAC from 530,319 to 434,075 eur [1], [4] and [5]. As a result of the installation of the wall mounted heating/cooling system and window integrated ventilation system 158 m2 of floor space were liberated for critical deposit area. The introduction of heating-cooling wall panels resulted besides in considerable energy reduction also in better indoor comfort because of large vertical heating and cooling areas with optimal surface temperatures. The obtained results pointed to the importance of proper system regulation and automatic control [8], as realised energy savings would not be possible without sufficient control provided by the BMS. In the end the use of wallmounted heating/cooling system in a renovated building also showed that successful use of lowtemperature systems can be achieved in retrofitting projects if they are well coordinated throughout the project activities.
4 ACKNOWLEDGEMENT This project has been supported by the European Commission 5th Framework Programme MUSEUMS Energy Efficiency and Sustainability in Retrofitted and New Museum Buildings, NNE5/1999/20, Ministry of Education,
Science and Sport, Republic of Slovenia, Research Programme Renewable Energy Sources, P0-504-0792-02 and Chair for Buildings and Constructional Complexes, Faculty of Civil and Geodetic Engineering, University of Ljubljana, Slovenia. 5 REFERENCES [1] Krainer, A., Rudi, P. (2005). Slovene Etnographic Museum. Slovene Etnographic Museum, Ljubljana. [2] Imanari, T., Omori, T., Bogaki K. (1999). Thermal comfort and energy consumption of the radiant ceiling panel system: Comparison with the conventional all-air system. Energy and Buildings, vol. 30, no. 2, p. 167-175. [3] Asada, H. (2004). Exergy analysis of low temperature radiant heating system. Building Service Engineering Research and Technology, vol. 25, no. 3, p. 197-209. [4] Krainer, A., Perdan, R., Krainer, G. (2007). Retrofiting of the Slovene Ethnographic Museum. Bauphysik, vol. 29, no. 5, p. 350365. [5] Krainer, A., Perdan, R., Krainer, G. (2006). Slovene Ethnographic Museum: SEM, a case study. International Journal of Sustainable Energy, vol. 25, no. 3, p. 131-151. [6] TRNSYS. A transient System Simulation Program, version 15.0. Solar Energy Laboratory, The Centre Scientifique et Technique du Batiment, Transsolar Energietechnik GmBH and Thermal Energy Systems Specialists, 2002. [7] Paić, Z. (2002). Surface heating and cooling systems (Sustavi površinskog grijana i hlađenja). Energetika marketing, Zagreb. (In Croatian) [8] Košir, M. (2003). Integrated regulating system of internal environment on the basis of fuzzy logic use: doctoral thesis, Ljubljana: University of Ljubljana, Faculty of Civil and Geodetic Engineering.
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Paper received: 14.07.2009 Paper accepted: 31.08.2010
The Effects of Machining Parameters on Cutting Forces, Surface Roughness, Built-Up Edge (BUE) and Built-Up Layer (BUL) During Machining AA2014 (T4) Alloy Hasan Gökkaya* Department of Mechanical Engineering, Karabuk University, Turkey
*
Tool wear, formed in cutting tool during machining processes, affects the surface roughness of the work piece, cutting forces and other output parameters. The effects of the machining parameters cutting speed (Vc) and the feed rate (f) on built-up edge (BUE), built-up layer (BUL), main cutting force (Fc), and surface roughness (Ra) is investigated in this study. The effects of the cutting parameters on cutting force and surface roughness has been examined by the use of Variance Analysis (ANOVA); and their optimum and critical cutting parameters were determined accordingly. AA2014 aluminum alloy was machined with uncoated carbide tools using Computer Numerical Control (CNC) turning machine under dry cutting conditions. Four different cutting speeds (200 m/min, 300 m/min, 400 m/min, and 500 m/min), five different feed rates (0.10 mm/rev, 0.15 mm/rev, 0.20 mm/rev, 0.25 mm/rev, and 0.30 mm/rev) and a constant depth of cut were selected as the machining parameters. BUE and BUL in the cutting tool were formed most at cutting speed 200 m/min and feed rate 0.30 mm/rev. The lowest cutting force was determined as 137 N at cutting speed 500 m/min and feed rate 0.10 mm/rev. The lowest average surface roughness, however, was determined as 0.93 µm at 500 m/min cutting speed and feed rate 0.10 mm/rev. ©2010 Journal of Mechanical Engineering. All rights reserved. Keywords: AA2014 alloy, built-up edge (BUE), built-up layer (BUL), cutting force, surface roughness, machining 0 INTRODUCTION Al-Cu alloys are one of the indispensable materials of the current industry because of their superior properties over other metal alloys. They can also undergo aging heat treatment which engenders them to be widely used in the industry [1] to [6]. Two different types of cutting methods (orthogonal-oblique) are implemented in the machining processes. Although most of the cutting processes are oblique, orthogonal cutting techniques are used in experiments to determine the effects of the parameters since mechanical behavior of the work piece is two-dimensional [7 to 9]. Tool rigidity, cutting speed, feed rate, depth of cut and tool geometry are also important factors for the determining of ideal machinability behaviors in addition to mechanical properties of a work piece [10] to [13]. AA2014 alloy, as an Al-Cu alloy, is generally shaped by using machining methods. BUE formation, occurring when aluminum alloys are machined at low cutting speeds, causes surface roughness (Ra) to increase [14] and [15]. *
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Increasing the cutting speed during machining causes the cutting forces to decrease due to low frictional forces on the tool rake face at high cutting speeds [7], [8] and [11]; and consequently resulting an elimination of BUE formation improving surface roughness of the work piece [16]. BUE formation sometimes positively affects the surface roughness of the work piece since BUE formation increases tool nose radius [16]. In this study, AA2014 (T4) with uncoated carbide tools was implemented by using five feed rates and four different cutting speeds. AA2014 aluminum alloy was then machined with orthogonal cutting method with a constant depth of cut. The effects of cutting and feed rates on the main cutting force, average surface roughness, and BUE and BUL formations were researched in the study. 1 MATERIALS AND METHOD 1.1 Materials The effects of machining parameters on the main cutting force, average surface roughness,
Corr. Author's Address: Department of Mechanical Engineering, Karabuk University, 78100, Karabuk, Turkey, hgokkaya@hotmail.com
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and BUE-BUL formation were investigated and a correlation between these parameters was determined in this experimental study. Cutting speed and feed rate were used as machining parameters. Test specimens, prepared by using the AA2014 aluminum alloy with 80 mm diameter and 500 mm length, were used for the experiments in this study. Chemical compositions of the test specimens can be seen in Table 1. AA2014 (T4) alloy was measured to have 415 N/mm2 tensile strength and 98 BHN (“Reicherter Brinell” Hardness) values. Hardness value measurements were implemented by using 10 mm depth test specimens in “Reicherter Brinell” hardness measuring device. 5 mm ball points and 125 kg weight were used for hardness measurement experiments. Average value of 10 trial measurements, from outside through the center, was used when material hardness values were determined. 1.2 Machining Parameters, Cutting Tool and Tool Holder The tests were carried out at 20±1 ºC ambient temperature using changeable carbide inserts which have CCGT 120404FN-ALU geometry and K10 quality degree. CSRNR 2525 M12, the tool holder used for these tests, was appropriate to ISO 5608 and had a 90° approaching angle. Rake angle and clearance angles of the cutting tools were 7 and 5°, respectively. Cutting parameters used for the experiments and their levels can be seen in Table 2. Four different cutting speeds (200, 300, 400, and 500 m/min), five different feed rates advised by ISO 3685 (0.1, 0.15, 0.2, 0.25 and 0.30
mm/rev), and 1.5 mm constant depth of cut were selected for all cutting speeds. A total of 20 experiments were conducted according to cutting parameters and machining levels shown in Table 2. 1.3 Machine Tool and Measuring Equipment All the machining processes were done with a “JOHNFORD T35” industrial type Computer Numerical Control (CNC) turning machine having 10 KW power and revolving capability of 50-3500 rev/min. Kistler 9257B dynamometer was used to measure all cutting forces (Fc, Ff, Fp). Fc represented the main cutting force during tests whereas Ff was the feed force and Fp was the ploughing force. MAHRPerthometer M1 measuring device was used to measure surface roughness of the work piece material. These measurements were repeated three times for its precision. Cut-off length and sampling length assumed in order to measure surface roughness were 0.8 and 5.6 mm, respectively. Finally, JEOL-JSM 6060 scanning electron microscope (SEM) was used for the analysis of BUE and BUL formed on the cutting tool. 1.4 Statistical Analysis A multiple analysis of variance (ANOVA) was used for identifying the factors significantly affecting the performance measures of main cutting force and surface roughness during machining AA2014 (T4) alloy. Duncan test was further applied to the findings with changes to find the significance level of the changes.
Table 1. Chemical compositions of the test specimens (% weight) Si Fe Cu Mn Mg 0.672 0.512 4.33 0.564 0.401
Zn 0.168
Al Balance
Table 2. Cutting parameters used for the tests Level
Cutting speed VC [m/min]
Feed rate f [mm/rev]
1 2 3 4
200 300 400 500
0.10, 0.15, 0.20, 0.25, 0.30
Depth of cut [mm]
Cutting Tool, Grade, Form
Tool Holder
1.5
Uncoated Carbide K10 CCGT 120404FNALU
CSRNR 2525 M12
The Effects of Machining Parameters on Cutting Forces, Surface Roughness, Built-Up Edge (BUE) and Built-Up Layer (BUL) 585 During Machining AA2014 (T4) Alloy
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materials. BUE and BUL regions can be seen clearly in Fig. 1.
2 RESULTS AND DISCUSSION 2.1 Formation of BUE and BUL Tool life is generally determined by tool wear in machining processes. Adhesion wear occurs by the detachment of tool particles from the cutting tools with the help of metallic chips. The formation of BUE and BUL on the cutting tool is caused by tool-tool chip interface temperature and extreme pressure. The work piece material adheres/emanates to the rake face of the cutting tool in two different forms. The first method involves the formation known as BuiltUp Edge (BUE); which is the emanation of the work piece material to the cutting edge of the tool. In the second method, known as the formation of Built-Up Layer (BUL), the material is adhered by pouring to wider areas on the rake face of the tool. This second type formation is frequently observed in machining ductile
BUE BUL
Chip Tool
WorkPiece Fig. 1. Schematic image of cutting tool with BUE and BUL [18]
a)
b)
Fig. 2. SEM image of AA2014 (T4) alloy’s BUE and BUL formation on uncoated sementite carbide surface at 200 m/min and 0.30 mm/rev: a) SEM image of the tool rake face, b) 3D SEM image of the cutting tool 586
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A piece from the cutting tool also detaches from the tool with the detachment of the gradually hardened BUE formed in the cutting tool and therefore breaks occur via adhesion wear mechanisms. Cutting edge gradually wears since the formation of BUE occurs periodically during machining. A strong conjunction between the work piece material and cutting tool is known to occur during the machining aluminum alloys [17]. Accordingly, BUE and BUL formation should certainly be considered during the machining of such materials. Four different cutting speeds and five different feed rates were used to determine the effects of cutting speed on BUE and BUL formation in this part of the study. SEM images of BUE and BUL formations at 0.30 mm/rev feed rate was evaluated, since the highest BUE and BUL formation was observed to be occurring at this rate.
a)
BUE
SEM photograph depicting BUE and BUL formation of AA2014 aluminum alloy on uncoated sementite carbide cutting tool at 200 m/min and 0.30 mm/rev can be seen in Fig. 2. BUL formation was observed to be accumulated on the tool surface whereas BUE was formed on main and collateral cutting edges of the tool by accumulating on the tool rake face. The major part of BUE formation was observed at the tool main cutting edge and where the chip makes contact with the air from the tool nose through the tool holder in Fig. 2. The cutting tool nose radius also increased due to BUE formation. A formation of BUE in lesser amounts near tool nose radius can be attributed to a lower temperature around the tool nose radius. SEM images of BUE and BUL formation regions on tool surface at various cutting speeds (300, 400 and 500 m/min) are depicted in Figs. 3 to 5.
BUE BUL
BUL
b)
Fig. 3. BUE and BUL image of machining AA2014 (T4) alloys at 300 m/min cutting speed and 0.30 mm/rev feed rate: a) SEM image of tool rake face, b) 3D SEM image of the cutting tool The Effects of Machining Parameters on Cutting Forces, Surface Roughness, Built-Up Edge (BUE) and Built-Up Layer (BUL) During Machining AA2014 (T4) Alloy
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a)
BUE
BUL
b)
Fig. 4. BUE and BUL image of machining AA2014 (T4) alloys at 400 m/min cutting speed and 0.30 mm/rev feed rate a) SEM image of tool rake face, b) 3D SEM image of the cutting tool BUE formation can be clearly observed at the tool nose radius and the main cutting edge when Figs. 2 and 3 are analyzed. Furthermore, BUE and BUL formation at 300 m/min cutting speed was observed to be smaller than the BUE and BUL formation at 200 m/min when the two figures are compared. Likewise BUE and BUL formation was observed to be decreasing when the cutting speed was increased to 400 m/min (Fig. 4). This situation can be connected to the temperature increase in the second deformation region [12] and [19]. A decrease in the formation of BUE and BUL, compared to the formation at 400 m/min cutting speed, was also observed at 500 m/min cutting speed (Fig. 5). According to the conducted tests, lower cutting speeds (200 and 300 m/min) were observed to cause formation of BUE and BUL in higher amounts during the machining of AA2014 alloy. In conclusion, 500 m/min or higher cutting speeds were determined
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to prevent BUE and BUL formation for the machining of AA2014 alloy. 2.2 Changes in Cutting Forces and Surface Roughness 2.2.1 Change in Cutting Forces The main cutting force (Fc) and surface roughness (Ra) values determined from the experiments depending on cutting speed and feed rate cutting parameters, can be seen in Table 3. The effects of cutting speed and feed rate on the main cutting force were evaluated by applying analyses of multiple variances on the determined data. Significant changes with a confidence level of 95% were determined between all the factors according to the results of the analysis of variance. The analysis of variance implemented to determine the effects of cutting speed and feed rate can be seen in Table 4.
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b)
a) BUE
BUL
BUL
BUE
Fig. 5. BUE and BUL image of machining AA2014 (T4) alloys at 500 m/min cutting and 0.30 mm/rev feed rate a) SEM image of tool rake face, b) 3D SEM image of the cutting tool
a)
b)
Fig. 6. Main cutting force values for AA2014 (T4) alloy determined during machining with uncoated carbide cutting tool: a) depending on cutting speed, b) depending on feed rate P-values shown in Table 4 are the realized significance levels, associated with the F-tests for each source of variation. The sources with a P value less than 0.10 are considered to have a statistically significant contribution to the performance measures [20]. The minimum main
cutting force value according to the feed rate was determined at 0.10 mm/rev; and the minimum value for the cutting force regarding the cutting speed was determined at 500 m/min. A detailed graphic of the main cutting force values, obtained
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from the experiments, depending on cutting speed and feed rate is depicted in Fig. 6. It was observed that lower cutting speed values decreased as the values of cutting speeds were increased from a lower cutting speed (200 m/min) to a higher cutting speed (500 m/min) as can be analyzed in the graphic in Fig. 6. Increasing the cutting speed to obtain smaller values of cutting forces is the most frequent method used in the literature [7] and [8]. High temperature at the flow zone and decreasing surface area are considered to be the reasons of this inverse proportion. The reduction amount in cutting forces can depend on work piece material, working conditions, and cutting speed ranges. A decrease in average main force values was observed depending on the increase on the tested values of cutting speeds in this study. An increase of 150% in cutting speed caused 11.67% decrease (266.75 N) while working on the lower speeds (200 m/min). Maximum cutting force value (390 N) was determined at 200 m/min cutting speed and 0.30 mm/rev feed rate when machining AA2014 alloy in this study. The minimum cutting force (137 N), on the other hand, was determined at 500 m/min cutting speed and 0.10 mm/rev feed rate. Higher feed rates were observed to be causing higher cutting forces when the results of the experiments conducted with five different feed rates are analyzed. A direct relation between the tested feed rates and the main cutting force was determined. Therefore, feed rate values should be decreased in order to decrease the main cutting force; as the increase in the main cutting force depending on the feed rate is an expected result [21] and [22]. An increase of 200% at 0.1 mm/rev feed rate caused 146.25% increase (362 N) in the main cutting force. The maximum cutting force value (362 N) depending on the feed rate was determined at 0.30 mm/rev. The minimum cutting force (147 N), on the other hand, was determined at 0.10 mm/rev feed rate. 2.2.2 Change in Surface Roughness The surface roughness (Ra) values determined from experiments when machining AA2014 alloy with uncoated carbide cutting tool can be seen in Table 3. The effects of cutting speed and feed rate on surface roughness were
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evaluated by applying analyses of multiple variances on the determined data. The analysis of variance implemented to determine the effects of cutting speed and feed rate can be seen in Table 5. The minimum average value of surface roughness depending on the cutting speed was determined at 500 m/min whereas the minimum average value depending on the feed rate was determined at 0.10 mm/rev. Maximum average surface roughness value (5.34 µm) was determined at 300 m/min cutting speed and 0.30 mm/rev feed rate in this experimental study. Minimum average surface roughness value (0.93 µm), on the other hand, was determined at 500 m/min cutting speed and 0.10 mm/rev feed rate (Table 3). A detailed graphic of average surface roughness values, obtained from the experiments, depending on the cutting speed and feed rate is depicted in Fig. 7. Table 3. Main cutting force (Fc) and average surface roughness (Ra) values depending on cutting speed (Vc) and feed rate (f) Vc f Fc Ra Test No [m/min] [mm/rev] [N] [µm] 1 200 0.10 152 1.15 2 200 0.15 212 1.68 3 200 0.20 270 1.81 4 200 0.25 336 3.72 5 200 0.30 390 5.19 6 300 0.10 154 1.29 7 300 0.15 207 2.31 8 300 0.20 260 2.93 9 300 0.25 313 4.10 10 300 0.30 363 5.34 11 400 0.10 145 1.10 12 400 0.15 201 1.58 13 400 0.20 255 2.37 14 400 0.25 307 3.65 15 400 0.30 355 5.14 16 500 0.10 137 .93 17 500 0.15 190 1.44 18 500 0.20 244 2.04 19 500 0.25 293 3.17 20 500 0.30 340 4.44
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Table 4. Variance analysis (ANOVA) regarding main cutting force (Fc) Degree of Source of Sum of Squares Freedom Variance Factor A 1153620.000 4 Factor B 25240.809 3 A*B 5239.625 12 Error 120.000 179 Total 14208780.000 199 Adjusted Total 1189634.271 198 Factor A; Feed rate (0.1, 0.15, 0.20, 0.25, 0.30 mm/rev) Factor B; Cutting speed (200, 300, 400, 500 m/min)
Variance
F Value
P Value
288405.000 8413.603 436.635 .670
430204.125 12550.291 651.314
.000 .00 .00
Variance
F Value
P Value
98.873 5.289 .459 6.667E-05
1483090.500 79333.000 6890.500
.000 .00 .00
Table 5. Variance analysis (ANOVA) regarding the surface roughness (Ra) Source of Degree of Sum of Squares Variance Freedom Factor A 395.491 4 Factor B 15.867 3 A*B 5.512 12 Error 1.200E-02 180 Total 1950.354 200 Adjusted Total 416.882 199 Factor A; Feed rate (0.1, 0.15, 0.20, 0.25, 0.30 mm/rev) Factor B; Cutting speed (200, 300, 400, 500 m/min)
a)
b)
Fig. 7. Average surface roughness (Ra) values for AA2014 (T4) alloy determined during machining with uncoated carbide cutting tool: a) depending on cutting speed, b) depending on feed rate Minimum average surface roughness values were determined at 500 m/min cutting speed as can be seen in Fig. 7a. Average surface roughness was increased when the cutting speed was increased from 200 to 300 m/min. This was because BUE formations at 200 m/min cutting speed increased the tool nose radius causing an
increase in surface roughness (Fig. 2). These results show similarity with the literature. BUE formation sometimes improves the surface roughness since it increases the tool nose radius [12] and [23]. An improvement was also observed at 400 and 500 m/min cutting speeds depending on the increase in cutting speeds. This
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improvement can be attributed to a simpler deformation process, easier deformation of the working piece around the cutting edge and tool nose radius, formation of flow zone (Fz) due to a higher temperature, and the prevention of BUE formation [10] and [22]. Minimum average surface roughness (1.12 µm) depending on feed rate was determined at 0.10 mm/rev feed rate. Maximum average surface roughness (5.03 µm), on the other hand, was determined at 0.30 mm/rev. A detailed graphic of average surface roughness values, obtained from the experiments, depending on feed rate is depicted in Fig. 7b. It was observed that average surface roughness was increasing as the feed rate was increased when average surface roughness values at 0.10, 0.15, 0.20, 0.25 and 0.30 mm/rev feed rates were analyzed. An increase of 200% at 0.1 mm/rev feed rate caused 349.3% increase (5.03 µm) in the average surface roughness. A direct relation was determined between the tested feed rates and average surface roughness. Therefore, feed rate values should be decreased in order to decrease the average surface roughness as the increase in the main cutting force depending on the feed rate is an expected result [22]. 3 CONCLUSION Conclusions drawn from the results, within the testing limits defined in the aim of the study, are given below. BUE formation on the cutting tool caused an increase in the tool nose radius on the cutting tool surface. This BUE formation at lower cutting speeds (200 m/min) positively affected the surface roughness. BUE and BUL formed at greater amounts on the surface of the cutting tool at 200 and 300 m/min cutting speeds, whereas BUE and BUL formation was observed at lesser amounts at 400 m/min and 500 m/min cutting speeds. Maximum BUE and BUL formation occurred during the machining process at 200 m/min cutting speed. 500 m/min and higher cutting speeds were determined for the selection of cutting speed to prevent BUE and BUL formation when machining AA2014 (T4) alloy with an uncoated sementite carbide cutting tool. The minimum cutting force depending on the cutting speed and feed rate (137 N) was
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determined at 500 m/min cutting speed and 0.10 mm/rev in this study. The maximum cutting force value (390 N), on the other hand, was determined at 200 m/min cutting speed and 0.30 mm/rev feed rate. The minimum average surface roughness value (0.93 µm) was determined at 500 m/min cutting speed and 0.10 mm/rev feed rate in this experimental study. The maximum average surface roughness value (5.34 µm), on the other hand, was determined at 300 m/min cutting speed and 0.30 mm/rev feed rate. 4 REFERENCES [1] Dogan, M. (1989). Heat treatment of aluminum alloys. Master’s Thesis, Marmara University, Istanbul. (in Turkish) [2] Yuksel, E.E. (1991). Opportunities to control the properties of 7075 aluminum alloys by the use of stepwise aging, Master’s Thesis, Istanbul Technical University, Istanbul. (in Turkish) [3] Su, S. (1988). The effect of solid to molten temperature and duration on the post aging properties of 2XXX alloys. Master’s Thesis, Selcuk University, Konya. [4] Boyer, H.E. (1989). Practical Heat Treating. American Society for Metals, p. 213-217. [5] ASM Handbook (1991). Heat Treating, vol. 4, ASTM. [6] Ozcatalbas, Y., Aydın, B. (2006). The effect of mechanical properties and cutting geometry to the machinability properties of AA2014 alloy. Gazi University Journal of Engineering and Architecture Department, vol. 21, no. 1, p. 21-27. (in Turkish) [7] Aydin, B., Ozcatalbas, Y. (2003). The effect of cutting tool geometry to the machinability properties of AA2014 (T6) alloy. Journal of machinery design and fabrication, vol. 2, no. 5, p. 89-95. (in Turkish) [8] Ozcatalbas, Y. (1998). The effect of built-up edge on the machinability properties of low alloy steel. 8th International Machinery design and fabrication Convention, ODTU, p. 25-33. (In Turkish) [9] Trent, E.M. (1988). Metal Cutting. Taner Ltd., London. [10] Modern Metal Cutting, Practical Handbook (1994). Sandvik Coromant.
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[11] Directorate of Etibank Aluminum Enterprise (1995). Product Catalogue. Ankara. (in Turkish) [12] Hong, S.Y., Ding, Y., Jeong, W. (2001). Friction and cutting forces in cryogenic machining of Ti–6Al–4V. International Journal Machine Tools & Manufacture, vol. 41, p. 2271-2285. [13] Pušavec, F., Krajnik, P., Kopač, J. (2006). High-speed cutting of soft materials. Strojniški vestnik – Journal of Mechanical Engineering, vol. 52, no. 11, p. 706-722. [14] Dae, E.K., Dong, H.H. (1998). Experimental investigation of the contact sliding behavior of metals. Journal of Manufacturing Science and Engineering, vol. 120, p. 395-400. [15] Jeelani, S., Musial, M. (1986). Dependence of Fatigue Life on the Surface Integrity in the Machining of 224-T 351 Aluminum Alloy Unlubricated Conditions. Journal of Materials Science, vol. 21, p. 155-160. [16] Oishi, K. (1996). Mirror, Cutting of Aluminum with Sapphire Tool. J Mater Proc Tech, vol. 62, no. 4, p. 331-334. [17] Rubio, E.M., Camacho, A.M., SánchezSola, J.M., Marcos, M. (2005). Surface roughness of AA7050 alloy turned bars, analysis of the influence of the length of machining. J Mater Proc Tech, vol. 162163, p. 682-689. [18] Gokkaya, H., Taskesen, A. (2008). The effects of cutting speed and feed rate on Bue-Bul formation, cutting forces and surface roughness when machining AA6351 (T6) alloy, Strojniški vestnik ̶
[19]
[20]
[21]
[22]
[23]
Journal of Mechanical Engineering, vol. 54, no. 7-8, p. 521-530. Courbon, C., Kramar, D., Krajnik, P., Pušavec, F., Rech, J., Kopač, J. (2009). Investigation of machining performance in high-pressure jet assisted turning of Inconel 718: an experimental study. International Journal of Machine Tools and Manufacture, vol. 49, no. 14, p. 1114-1125. Aslan, E., Camuscu, N., Birgoren, B. (2007). Design optimization of cutting parameters when turning hardened AISI 4140 steel (63 HRC) with Al2O3 + TiCN mixed ceramic tool. Materials & Design, vol. 28, no. 5, p.1618-1622. Altin, A., Gokkaya, H., Nalbant, M. (2006). The effect of cutting speed on the machinability properties of Inconel 718 Super alloy. Gazi University Journal of Engineering and Architecture Department, vol. 21, no. 3, p. 581-586. (in Turkish) Gokkaya, H., Sur, G., Dilipak, H. (2004). Experimental study of the effect of machining parameters of PVD and CVD coated cementite carbide cutting tools on surface roughness. Zonguldak Karaelmas University Journal of Technical Education Faculty, vol. 7, no. 3, p. 473-478. (in Turkish) Adamczak, S., Miko, E., Čuš, F. (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.
The Effects of Machining Parameters on Cutting Forces, Surface Roughness, Built-Up Edge (BUE) and Built-Up Layer (BUL) During Machining AA2014 (T4) Alloy
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 594-595 Instructions for Authors
Instructions for Authors For full instructions see the Authors Guideline section on the journal's website: http://en.sv-jme.eu/ Send to: University of Ljubljana Faculty of Mechanical Engineering SV-JME Aškerčeva 6, 1000 Ljubljana, Slovenia Phone: 00386 1 4771 137 Fax: 00386 1 2518 567 E-mail: info@sv-jme.eu strojniski.vestnik@fs.uni-lj.si All manuscripts must be in English. Pages should be numbered sequentially. The maximum length of contributions is 10 pages. Longer contributions will only be accepted if authors provide justification in a cover letter. Short manuscripts should be less than 4 pages. Prior or concurrent submissions to other publications should be included in the cover letter. We also ask you to specify the typology of the manuscript – you can define your paper as an original, review or short paper. The required contact information (e-mail and mailing address) and a suggestion of at least two potential reviewers should be provided in the cover letter. Reasons for a certain reviewer to be excluded from the review process can also be provided in the cover letter. Every manuscript submitted to the SV-JME undergoes the course of the peer-review process. THE FORMAT OF THE MANUSCRIPT The manuscript should be written in the following format: - A Title, which adequately describes the content of the manuscript. - An Abstract should not exceed 250 words. The Abstract should state the principal objectives and the scope of the investigation, as well as the methodology employed. It should summarize the results and state the principal conclusions. - 6 significant key words should follow the abstract to aid indexing. - An Introduction, which should provide a review of recent literature and sufficient background information to allow the results of the article to be understood and evaluated. - A Theory or experimental methods used. - An Experimental section, which should provide details of the experimental set-up and the methods used for obtaining the results.
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, 594-595
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Industry. MOTSP 2009 Conference Proceedings, p. 422-427. Standards: Standard-Code (year). Title. Organisation. Place. [5] ISO/DIS 16000-6.2:2002. Indoor Air – Part 6: Determination of Volatile Organic Compounds in Indoor and Chamber Air by Active Sampling on TENAX TA Sorbent, Thermal Desorption and Gas Chromatography using MSD/FID. International Organization for Standardization. Geneva. www pages: Surname, Initials or Company name. Title, from http://address, date of access. [6] Rockwell Automation. Arena, from http://www.arenasimulation.com, accessed on 2009-09-07. COPYRIGHT Authors submitting a manuscript do so on the understanding that the work has not been published before, is not being considered for publication elsewhere and has been read and approved by all authors. The submission of the manuscript by the authors means that the authors automatically agree to transfer copyright to SV-JME and when the manuscript is accepted for publication. All accepted manuscripts must be accompanied by a Copyright Transfer Agreement, which should be sent to the editor. The work should be original by the authors and not be published elsewhere in any language without the written consent of the publisher. The proof will be sent to the author showing the final layout of the article. Proof correction must be minimal and fast. Thus it is essential that manuscripts are accurate when submitted. Authors can track the status of their accepted articles on http://en.sv-jme.eu/. PUBLICATION FEE For all articles authors will be asked to pay a publication fee prior to the article appearing in the journal. However, this fee only needs to be paid after the article has been accepted for publishing. The fee is 180.00 EUR (for articles with maximum of 6 pages), 220.00 EUR (for articles with maximum of 10 pages), 20.00 EUR for each addition page. Additional costs for a color page is 90.00 EUR.
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9 Vsebina
Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 56, (2010), številka 9 Ljubljana, september 2010 ISSN 0039-2480 Izhaja mesečno
Povzetki člankov Aljaž Osterman, Matevž Dular, Marko Hočevar, Brane Širok: Infrardeča termografija kavitacijskih termičnih učinkov v vodi Zlatko Petrović, Slobodan Stupar, Ivan Kostić, Aleksandar Simonović: Ugotavljanje zmogljivosti lahkega helikopterja v preliminarni fazi razvoja Dušan Mežnar, Momir Lazovič: Trdnost konstrukcije avtobusa z določitvijo kritičnih mest Antonios Kyriazopoulos, Ilias Stavrakas, Konstantinos Ninos, Cimon Anastasiadis, Dimos Triantis: Tlačno stimulirana emisija toka v vzorcih cementne paste pri ponavljajočih se stopničastih tlačnih obremenitvah Igor Solodov, Daniel Döring, Gerd Busse: Oddaljene neporušne preiskave napak in elastičnih lastnosti materialov z Lambovi in Rayleighovimi valovi, ki se prenašajo po zraku Dejan D. Ivezić, Trajko B. Petrović: Robustni regulatorji IMC za industrijske kotle z optimalnim sledenjem nastavljene vrednosti in izravnavo vpliva motenj Mitja Košir, Aleš Krainer, Mateja Dovjak, Rudolf Perdan, Živa Kristl: Alternativni sistem ogrevanja in hlajenja javnih stavb Hasan Gökkaya: Vpliv parametrov obdelave na rezalne sile, površinsko hrapavost in nastanek nalepljenega roba in nalepljene plasti med obdelavo zlitine AA2014 (T4)
SI 115 SI 116 SI 117 SI 118 SI 119 SI 120 SI 121 SI 122
Navodila avtorjem
SI 123
Osebne vesti Doktorati, magisteriji, specializacija in diplome
SI 125
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 115 UDK 532.528:544.57:536.5
Prejeto: 18.11.2009 Sprejeto: 01.09.2010
Infrardeča termografija kavitacijskih termičnih učinkov v vodi Aljaž Osterman* - Matevž Dular - Marko Hočevar - Brane Širok Fakulteta za strojništvo, Univerza v Ljubljani, Slovenija Kljub temu, da se za temperaturne učinke kavitacije v vodi meni, da so zanemarljivi, so bili eksperimentalno uspešno izmerjeni z infrardečo termografijo. Kavitacija je bila vzbujana s pomočjo ultrazvoka v majhni posodi, ki je vsebovala pribl. 500 ml vode. Frekvenca ultrazvoka, ki je povzročila rast in kolaps mehurčkov, je bila 42 kHz. Temperatura je bila merjena s hitro termokamero, ki deluje v infrardečem območju od 3 do 5 μm. Termokamera je temperaturna polja zajemala s frekvenco 600 Hz. Silicijevo steklo, ki je prozorno v merjenem delu IR spektra, je bilo pritrjeno na cev, delno potopljeno v vodo. Mehurčki, ki se pojavljajo v bližini trdne površine, so se pojavljali na omočeni strani stekla. Optična pot merjenega sevanja je bila tako zrak-silicijevo steklo-voda. Na ta način je bilo možno izmeriti temperature na omočeni strani silicijevega stekla, kjer je prihajalo do rasti mehurčkov in njihovih implozij. Na podlagi uporabljene termografske metode so bili izmerjena majhna, a izrazita lokalna znižanja temperature (z velikostjo do 0,3 K), ki so bila posledica kavitacije. © 2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: kavitacija, ultrazvok, temperatura, IR termografija, mehurčki
Slika 1. Postavitev eksperimenta
*Naslov odgovornega avtorja: Univerza v Ljubljani, Fakulteta za strojništvo, Aškerčeva 6, SI – 1000 Ljubljana, Slovenija, aljaz.osterman@fs.uni-lj.si
SI 115
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 116 UDC 533.661:629.01
Prejeto: 30.01.2008 Sprejeto: 02.07.2010
Ugotavljanje zmogljivosti lahkega helikopterja v preliminarni fazi razvoja Zlatko Petrović - Slobodan Stupar - Ivan Kostić* - Aleksandar Simonović Univerza v Beogradu, Fakulteta za strojništvo, Srbija Lahki in ultralahki helikopterji predstavljajo obetavno rešitev za problem vse bolj časovno potratnega sodobnega urbanega transporta. V članku je predstavljen del metodologije preliminarne konstrukcije, ki je bila razvita na Fakulteti za strojništvo Univerze v Beogradu in vključuje tudi izračune zmogljivosti takšnih helikopterjev. Zaradi omejenega proračuna in izjemne zahtevnosti procesa razvoja helikopterja je zelo pomembno, da v vseh fazah razvoja razpolagamo z zanesljivimi ocenami zmogljivosti in tako izpolnimo zahteve. Članek se omejuje na fazo preliminarne konstrukcije, kjer se zelo zahtevna dinamika helikopterskega rotorja običajno nadomesti s povprečenimi mehanskimi in aerodinamičnimi značilnostmi in se uporabijo nekatere empirično preverjene poenostavitve. Na osnovi tega pristopa so bili razviti samostojni, učinkoviti in zanesljivi računalniški programi za izračun različnih delovnih značilnosti. Razen pri projektu, ki se trenutno izvaja, so bili ti programi uporabljeni tudi pri več obstoječih helikopterjih podobnega razreda za natančnejše ugotavljanje empiričnih vhodnih parametrov. Predstavljena je uporabljena metodologija in rezultati, ki potrjujejo celotno učinkovitost algoritma. ©2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: lahek helikopter, zmogljivost zrakoplova, preliminarna faza konstrukcije
Sl. 1. CAD-model helikopterja in nekaj primerov komponent, izdelanih za laboratorijske preizkuse in verifikacijo proizvodnih tehnologij (poznejše stopnje razvoja pri projektu, ki je v teku)
SI 116
*Naslov odgovornega avtorja: Univerza v Beogradu, Fakulteta za strojništvo, Katedra za aeronautiko, Kraljice Marije 16, 11120 Beograd 35, Srbija, ikostic@mas.bg.ac.rs
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 117 UDK 629.34:629.021
Prejeto: 22.09.2009 Sprejeto: 24.05.2010
Trdnost konstrukcije avtobusa z določitvijo kritičnih mest Dušan Mežnar* - Momir Lazovič Tovarna Vozil Maribor d.o.o., Skupina Viator&Vektor Samonosna karoserija letališkega avtobusa je z vidika trdnosti zelo zahteven izdelek. Z uporabo metode MKE so bila ugotovljena domnevno kritična mesta na okvirju, zlasti na okvirjih vrat. Eksperimentalne metode za merjenje mehanskih deformacij so potrdile domnevo, da se bodo maksimalne deformacije pojavile na mestih, predhodno ugotovljenih z metodo MKE. Kot kritičen se je izkazal režim vožnje v krog z največjo hitrostjo 40 km/h, najmanjšim polmerom obračanja in izmenljivim pospeševanjem in zaviranjem. Na osnovi meritev so bile opravljene ustrezne spremembe konstrukcije, dodatne ojačitve okvirja in drugi ukrepi, ki zagotavljajo izpolnjevanje zahtevanih trdnostnih kriterijev. ©2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: samonosna karoserija, analiza po metodi MKE, deformacije
Slika 1: Analiza po metodi MKE kaže lokalno koncentracijo napetosti
* Naslov odgovornega avtorja: Tovarna Vozil Maribor d.o.o., Skupina Viator&Vektor, 2000 Maribor, Slovenija, dusan.meznar@gmail.com
SI 117
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 118 UDK 620.173.25:621.742.48
Prejeto: 28.02.2008 Sprejeto: 29.10.2009
Tlačno stimulirana emisija toka v vzorcih cementne paste pri ponavljajočih se stopničastih tlačnih obremenitvah Antonios Kyriazopoulos*, Ilias Stavrakas, Konstantinos Ninos, Cimon Anastasiadis, Dimos Triantis Laboratorij za električne lastnosti materialov, Oddelek za elektroniko, Tehniški izobraževalni institut v Atenah, Grčija
PSC [pA]
napetost [%]
Tehnika zaznavanja električnih signalov, ki je opisana v članku, je bila v preteklosti uporabljena pri različnih geomaterialih, v zadnjem času pa se uporablja tudi pri cementnih materialih. V delu je raziskana emisija električnih signalov v cementni pasti pri aksialnih obremenitvah, zlasti pri ponavljajočih se obremenitvah in razbremenitvah vzorca v področju, kjer se pojavi nastanek in napredovanje razpok. Ugotovljeno je bilo, da prihaja do emisije električnega signala v dveh stopnjah. Prva emisija toka v obliki konice se pojavi sočasno s stopničasto obremenitvijo in se nato postopoma vrne na raven ozadja. Sekundarna emisija toka pa je bila zabeležena pri konstantni visoki vrednosti napetostne stopnice. ©2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: tlačno stimuliran električni tok, emisije električnega toka, cementna pasta, enoosna tlačna napetost
t [s]
Slika 5. a) Časovni potek PSC med zadnjo napetostno stopnico in b) ustrezen časovni potek deformacije
SI 118
* Naslov odgovornega avtorja: Laboratorij za električne lastnosti materialov, Zavod za elektroniko, Tehniški izobraževalni institut v Atenah, Atene, 12210, Grčija, akyriazo@teiath.gr
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 119 UDK 544.57:539.3
Prejeto: 22.08.2009 Sprejeto: 04.03.2010
Oddaljene neporušne preiskave napak in elastičnih lastnosti materialov z Lambovi in Rayleighovimi valovi, ki se prenašajo po zraku Igor Solodov* - Daniel Döring - Gerd Busse Inštitut za tehnologijo polimerov, skupina NDE (IKT-ZfP), Nemčija Običajne ultrazvočne preiskave z zrakom kot spojnim medijem (ACU) so dobro uveljavljeno orodje za akustične neporušne preiskave in karakterizacijo materialov. Njihova glavna slabost je šibka penetracija v trdne snovi zaradi neujemanja impedance na stiku med zrakom in trdno snovjo. S pretvorbo akustičnega načina v ploščne in površinske valove v nagnjeni konfiguraciji je možno doseči dramatično povečanje akustičnega prenosa. Eksperimenti v pogojih faznega ujemanja so pokazali do eno velikostno stopnjo večjo amplitudo ultrazvoka pri različnih materialih (kovine, les, beton, kompoziti). Na tej osnovi so bile razvite in uporabljene konfiguracije za brezkontaktno neporušno preizkušanje z zrakom kot prenosnim medijem. Metode, zasnovane po tem principu, omogočajo natančno merjenje smeri vlaken in kvantifikacijo ravninske anizotropije pri kompozitih in naravnih materialih, določanje globinskega profila elastičnosti, merjenje sušenja prevlek, napredno snemanje razpok in preučevanje delaminacije. ©2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: ultrazvočno preizkušanje z zrakom kot spojnim medijem, Lambovi in Rayleighovi valovi, karakterizacija materialov, elastična anizotropija, določanje globinskega profila elastičnosti, spremljanje procesov, posnetki za neporušne preiskave
Slika 1. Shematski prikaz pravila sovpadanja za prehod s faznim ujemanjem
*Naslov odgovornega avtorja: Inštitut za tehnologijo polimerov, NDE Group (IKT-ZfP), Univerza v Stuttgartu, Pfaffenwaldring 32, D-70569 Stuttgart, Nemčija, igor.solodov@ikt.uni-stuttgart.de
SI 119
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 120 UDK 621.182
Prejeto: 29.01.2008 Sprejeto: 31.08.2010
Robustni regulatorji IMC za industrijske kotle z optimalnim sledenjem nastavljene vrednosti in izravnavo vpliva motenj Dejan D. Ivezić1*, Trajko B. Petrović2 Univerza v Beogradu, Rudarsko-geološka fakulteta, Oddelek za strojništvo, Srbija 2 Univerza v Beogradu, Fakulteta za elektrotehniko, Oddelek za regulacijsko tehniko, Srbija 1
V članku je predstavljen razvoj robustnih regulatorjev na osnovi metode vodenja z notranjim modelom (IMC), ki izboljšujejo robustnost regulacije industrijskega kotla z ozirom na negotovosti in motnje. Razvita sta poenostavljen model kotelnega bobna in prenosna matrika njegove dinamike za imenske pogoje obratovanja. Parametri regulatorja so bili izbrani s pomočjo metode -optimizacije v frekvenčni domeni. Predlagani regulatorji so robustni tako za referenčne signale kakor tudi za motnje. Končno je podana tudi primerjava zmogljivosti zaprtozančnega sistema in razvitih regulatorjev IMC. ©2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: industrijski kotel, robustna regulacija, vodenje z notranjim modelom
Ms pd ds, Vds Mw Hdw, dw, Vdw M0, X0, Hrm
Mcoal, Qrw
Mdow, Hdow
Hrw
Slika 2. Poenostavljen opis parno-vodnega dela sistema kotla
SI 120
Naslov odgovornega avtorja: Univerza v Beogradu, Rudarsko – geološka fakulteta, Oddelek za strojništvo, Djušina 7, Beograd 11000, Srbija, ivezic@rgf.bg.ac.rs
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 121 UDK 628.8:697.353:069
Prejeto: 01.12.2009 Sprejeto: 02.03.2010
Alternativni sistem ogrevanja in hlajenja javnih stavb Mitja Košir* - Aleš Krainer - Mateja Dovjak - Rudolf Perdan - Živa Kristl Univerza v Ljubljani, Fakulteta za gradbeništvo in geodezijo, Katedra za stavbe in konstrukcijske elemente, Slovenija V pričujočem članku je predstavljen alternativni sistem ogrevanja in hlajenja javnih stavb. Predstavljen sistem je bil načrtovan in ekstenzivno preizkušen v prenovljeni stavbi Slovenskega etnografskega muzeja (SEM). Vgrajen sistem sestavljajo radiacijski ogrevalno-hladilni stenski paneli, tangencialni ventilatorji za hlajenje in lokalizirano prezračevanje ter centralni nadzorni sistem za upravljanje in nadzorovanje delovanja. Učinkovitost delovanja predlaganega ter kasneje vgrajenega sistema je bila temeljito preverjena skozi serijo eksperimentov izvedenih na stavbi pred in po izvedeni prenovi. Uporaba opisanega sistema v prenovljenih prostorih muzeja je posledično doprinesla k izrazitem zmanjšanju porabljene energije, boljšemu notranjemu toplotnemu okolju ter nižjim investicijskim stroškom za HVAC sistem celotne stavbe. © 2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: ogrevanje, hlajenje, prezračevanje, nizko temperaturni sistem, radiacijski paneli
Slika 10. Raba energije za ogrevanje nizko temperaturnih stenskih panelov vgrajenih v razstavnem delu muzeja v primerjavi s porabo celotne stavbe SEM. V prvih štirih mesecih leta so paneli delovali brez avtomatske regulacije, v zadnjih treh mesecih pa so bili vodeni s strani centralnega nadzornega sistema
*
Naslov odgovornega avtorja: Univerza v Ljubljani, Fakulteta za gradbeništvo in geodezijo, Katedra za stavbe in konstrukcijske elemente, Jamova cesta 2, Ljubljana, Slovenija, MKosir@fgg.uni-lj.si
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 122 UDK 621.91:669.71
Prejeto: 14.07.2009 Sprejeto: 31.08.2010
Vpliv parametrov obdelave na rezalne sile, površinsko hrapavost in nastanek nalepljenega roba in nalepljene plasti med obdelavo zlitine AA2014 (T4) Hasan GÖKKAYA Fakulteta za strojništvo, Univerza Karabuk, Turčija Obraba rezalnega orodja med procesom obdelave vpliva na površinsko hrapavost obdelovanca, rezalne sile in druge izhodne parametre. V članku je preučen vpliv parametrov obdelave rezalna hitrost (Vc) in podajanje (f) na nastanek nalepljenega roba (BUE), nalepljene plasti (BUL), glavno rezalno silo (Fc) in površinsko hrapavost (Ra). Vpliv rezalnih parametrov na rezalno silo in površinsko hrapavost je bil raziskan po metodi analize variance (ANOVA), določeni pa so bili tudi optimalni in kritični rezalni parametri. Aluminijeva zlitina AA2014 je bila obdelana na CNC-stružnici z neoplaščenimi trdokovinskimi orodji in v suhem. Kot parametri obdelave so bile izbrane štiri različne rezalne hitrosti (200 m/min, 300 m/min, 400 m/min in 500 m/min), pet različnih vrednosti podajanj (0,10, 0,15, 0,20, 0,25 in 0,30 mm/vrt.) in konstantna globina reza. Do največjega tvorjenja nalepka BUE in BUL na rezalnem orodju je prišlo pri rezalni hitrosti 200 m/min in hitrosti podajanja 0,30 mm/vrt. Najmanjša rezalna sila je bila 137 N pri rezalni hitrosti 500 m/min in vrednosti podajanja 0,10 mm/vrt. Najmanjša povprečna površinska hrapavost pa je bila 0,93 µm pri rezalni hitrosti 500 m/min in podajanju 0,10 mm/vrt. © 2010 Strojniški vestnik. Vse pravice pridržane. Ključne besede: zlitina AA2014, nalepljeni rob (BUE), nalepljena plast (BUL), rezalna sila, površinska hrapavost, obdelava
a)
b)
Slika 2. SEM-posnetek nastanka nalepka BUE in BUL na površini neoplaščene karbidne trdine pri obdelavi zlitine AA2014 (T4) z 200 m/min in 0,30 mm/vrt.; a) SEM-posnetek cepilne površine orodja, b) 3D SEM-posnetek rezalnega orodja *
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Naslov odgovornega avtorja: Fakulteta za strojništvo, Univerza Karabuk, 78100, Turčija, hgokkaya@hotmail.com
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 123-124 Navodila avtorjem
Navodila avtorjem Navodila so v celoti na voljo v rubriki "Informacija za avtorje" na spletni strani revije: http://en.sv-jme.eu/ Članke pošljite na naslov: Univerza v Ljubljani Fakulteta za strojništvo SV-JME 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. V spremnem dopisu navedite podatke o predhodnem ali hkratnem predlaganju članka v objavo drugje. Prosimo, da članku določite tudi tipologijo – opredelite ga lahko kot izvirni, pregledni ali kratki članek. Navedite vse potrebne kontaktne podatke (poštni naslov in email) in predlagajte vsaj dva potencialna recenzenta. 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. ms-1, 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. 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]
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 123-124
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. 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 revije 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. (Editors), 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. Standardi: Standard (leto). Naslov. Ustanova. Kraj.
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[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. Rockwell Automation. Arena, from http://www.arenasimulation.com, accessed on 200909-27. 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 Avtorji vseh sprejetih prispevkov morajo za objavo plačati prispevek v višini 180,00 EUR (za članek dolžine do 6 strani) ali 220,00 EUR (za članek dolžine do 10 strani) ter 20,00 EUR za vsako dodatno stran. Dodatni strošek za barvni tisk znaša 90,00 EUR na stran.
Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 125-126 Osebne vesti
Doktorati, magisterij in diplome DOKTORATI Na Fakulteti za strojništvo Univerze v Ljubljani je z uspehom obranil svojo doktorsko disertacijo: dne 24. avgusta 2010 Zoran BERGANT z naslovom: "Lasersko navarjanje in pretaljevanje jekla maraging" (mentor: prof. dr. Janez Grum). V doktorskem delu smo na vzorcih iz orodnega maraging jekla 1.2799 testirali in analizirali nov visokoproduktirni postopek pri popravilu orodnih jekel z laserskim pretaljevanjem predhodno nabrizgane Fe-Ni-CoMo prevleke. Kakovost nanosa, izdelanega z novim postopkom, smo primerjali z nanosom na vzorce s postopkom laserskega navarjanja s koaksialnim dovajanjem praškastega materiala. V analizo laserskih sledi so zajete geometrijske lastnosti prečnega preseka, mikrostrukturna analiza s podporo meritev mikrotrdote, mikrokemijske analize in meritev zaostalih napetosti pred in po topilnem in izločevalnem žarjenju pod različnimi pogoji. Z raziskavo smo pridobili številne tehnološke podatke o postopkih plamenskega nabrizgavanja, laserskega pretaljevanja in navarjanja, ki dajejo smernice za uporabnike maraging jekel v industriji in pri popravilu poškodovanih površin s predlaganim tehnološkim postopkom. DIPLOMIRALI SO Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 25. avgusta 2010: Žiga GOSAR z naslovom: "Eksperimentalna in računska določitev premičnega in napetostnega stanja v nosilcih po teoriji III. reda" (mentor: prof. dr. Franc Kosel); Luka KNEZ z naslovom: "Razvoj biodinamičnega testirnega sistema za antivibracijske rokavice" (mentor: prof. dr. Miha Boltežar, somentor: doc. dr. Janko Slavič); Marko POLAJNAR z naslovom: "Vpliv površinskih energij in omočljivosti površin na koeficient trenja v mazanih kontaktih" (mentor: izr. prof. dr. Mitjan Kalin); Matej SITAR z naslovom: »Stabilnost sistema, sestavljenega iz togih in nelinearnih elastičnih palic« (mentor: prof. dr. Franc Kosel);
Špela BOLKA z naslovom: »Simuliranje eksplicitne dinamike trka deformabilnih teles« (mentor: prof. dr. Miha Boltežar, somentor: doc. dr. Janko Slavič); Simon GRIŽONIČ z naslovom: »Razvoj in krmiljenje pnevmatičnega robotskega manipulatorja« (mentor: doc. dr. Niko Herakovič); Klemen RUPNIK z naslovom: »Vrednotenje dinamičnih lastnosti uporovnih temperaturnih zaznaval po metodi z notranjim vzbujanjem« (mentor: izr. prof. dr. Ivan Bajsić, somentor: doc. dr. Jože Kutin); dne 27. avgusta 2010: Marko BEK z naslovom: »Kontrola kvalitete sintranih izdelkov narejenih s tehnologijo brizganja prahu« (mentor: prof. dr. Igor Emri); Janez PEČNIK z naslovom: »Vzdrževanje strojev in naprav v podjetju Akrapovič d.d.« (mentor: prof. dr. Jožef Vižintin); Miha PIPAN z naslovom: »Merjenje neokroglih merjencev z laserskim mikrometrom brez predhodnega pozicioniranja« (mentor: prof. dr. Mihael Junkar, somentor: doc. dr. Henri Orbanić); Jernej REBERŠAK z naslovom: »Stabilnost oblike prereza tankostenih votlih pravokotnih nosilcev« (mentor: izr. prof. dr. Janez Kramar, somentor: doc. dr. Boris Jerman); dne 30. avgusta 2010: Jure JERINA z naslovom: »Vpliv topografije na adhezijske lastnosti kontaktnih površin« (mentor: izr. prof. dr. Bojan Podgornik, somentor: prof. dr. Jožef Vižintin); Marko MATIČIČ z naslovom: »Razvoj in izdelava prototipa kompresorja za lesni plin« (mentor: doc. dr. Jernej Klemenc, somentor: doc. dr. Jurij Prezelj); Aleš PETROVČIČ z naslovom: »Analiza vplivnih faktorjev na odjem toplote iz sistema daljinskega ogrevanja« (mentor: prof. dr. Alojz Poredoš, somentor: doc. dr. Andrej Kitanovski); Primož POREDOŠ z naslovom: »Prenos in shranjevanje toplote v napravi za naravno ogrevanje in hlajenje stavb« (mentor: izr. prof. dr. Sašo Medved, somentor: dr. Ciril Arkar); dne 31. avgusta 2010: Marko BOJINOVIĆ z naslovom: »Numerično modeliranje toplotnih razmer pri
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Strojniški vestnik - Journal of Mechanical Engineering 56(2010)9, SI 125-126
laserskem kaljenju« (mentor: prof. dr. Boris Štok, somentor: doc. dr. Nikolaj Mole); Matej KLEMENČIČ z naslovom: »Vpliv vbrizgavanja LPG na zmogljivosti tlačno polnjenega dizelskega motorja« (mentor: doc. dr. Tomaž Katrašnik); Gašper ŠKULJ z naslovom: »Fleksibilno modeliranje proizvodnih procesov« (mentor: prof. dr. Alojzij Sluga); Roger Martinez MARESCO z naslovom: »Vpliv vbrizgavanja utekočinjenega naftnega plina na zmogljivosti in emisije tlačno polnjenega dizelskega motorja / Influence of LPG injection on the performance and emissions of turbocharged diesel engine« (mentor: doc. dr. Tomaž Katrašnik). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 1. julija 2010: Nikola KIDESS z naslovom »Linija zamrzovalnih naprav in hladilnikov "Kombo"« (mentor: doc. dr. Vojmir Pogačar); Anton PERNAT z naslovom »Pregled in analiza sodobnih postopkov izsekovanja pločevine« (mentor: izr. prof. dr. Ivan Pahole, somentor: doc. dr. Mirko Ficko); dne 3. avgusta 2010: Janko FERČEC z naslovom »Zasnova in preračun dvoosnega sledilnika sonca« (mentor: prof. dr. Srečko Glodež, somentor: doc. dr. Janez Kramberger); Darko JAGARINEC z naslovom »Projektiranje tlačnega reaktorja 1000 L« (mentor: prof. dr. Nenad Gubeljak, somentor: doc. dr. Jožef Predan); Rok PLEVNIK z naslovom »Trdnostni preračun osi in gredi po DIN 743« (mentor: prof. dr. Srečko Glodež, somentor: prof. dr. Zoran Ren); dne 27. avgusta 2010: Peter GSELMAN z naslovom »Razvoj in načrtovanje proizvodnje medicinskega vozička« (mentor: prof. dr. Andrej Polajnar, somentor: doc. dr. Marjan Leber); Lovro KRAJNC z naslovom »Izdelava prilagojenih lobanjskih vsadkov z dodajalnimi tehnologijami« (mentor: izr. prof. dr. Igor Drstvenšek, somentor: mag. Tomaž Brajlih);
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Simon MARČIČ z naslovom »CNC izdelava visokokvalitetnih komponent za motorna kolesa« (mentor: prof. dr. Jože Balič); Aleksander PRAPER z naslovom »Oblikovalska vizija avta ALPRA GT« (mentor: doc. dr. Vojmir Pogačar); Damijan RUPNIK z naslovom »Konstruiranje priprave za vpenjanje vpenjalne glave večosnega stružnega stroja« (mentor: prof. dr. Zoran Ren); Janko SINREIH z naslovom »Skrajšanje pretočnih časov izdelkov v podjetju NIEROS METAL d.o.o.« (mentor: prof. dr. Andrej Polajnar, somentor: izr. prof. dr. Borut Buchmeister); * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva: dne 1. julija 2010: Florjan SLOVENC z naslovom »Izboljšava sklopa za kompenzacijo aksialne sile na večstopenjski radialni črpalki pomožne napajalne vode« (mentor: doc. dr. Samo Ulaga, somentor: doc. dr. Darko Lovrec); dne 27. avgusta 2010: Branko JERLAH z naslovom »Obratovanje vročevodnih kotlov brez stalnega nadzora v skladu s smernicami TRD 604« (mentor: izr. prof. dr. Aleš Hribernik); Andrej VEHAR z naslovom »Fazni model prevzema novih izdelkov v serijsko proizvodnjo« (mentor: doc. dr. Marjan Leber, somentor: prof. dr. Andrej Polajnar); Andrej ZADRAVEC z naslovom »Analiza energijskega učinka toplotne črpalke voda/voda in zrak/voda za ogrevanje in hlajenje večnamenskega objekta« (mentor: doc. dr. Matjaž Ramšak); ***** POPRAVEK Popravek napake v številki Strojniškega vestnika - Journal of Mechanical Engineering 56(2010)7-8, stran SI 111. Za napako se opravičujemo. Pravilna objava se glasi: Na Fakulteti za strojništvo Univerze v Ljubljani je pridobil naziv diplomirani inženir strojništva: dne 11. junija 2010: Peter SAMBOL z naslovom »Statične in dinamične značilnosti hidravličnega Wheatstonovega merilnega mostiča« (mentor: izr. prof. dr. Ivan Bajsić, somentor: doc. dr. Jože Kutin);
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