Journal of Mechanical Engineering 2012 10

http://www.sv-jme.eu

58 (2012) 10

Since 1955

Contents Papers

Matej Rajh, Srečko Glodež, Jože Flašker, Karl Gotlih: 563 Manipulability of a Haptic Mechanism within the Cylindrical Space of an MR Scanner

David Zaremba, Christian Biskup, Thomas Heber, Nico Weckend,

Werner Hufenbach, Frank Adam, Friedrich-Wilhelm Bach, Thomas Hassel: 571 Repair Preparation of Fiber-Reinforced Plastics by the Machining of a Stepped

Peripheral Zone

Xiwen Zhang, Xiaodong Wang, Yi Luo: 578 An Improved Torque Method for Preload Control in Precision Assembly of

Miniature Bolt Joints

Arif Gok, Cevdet Gologlu, Ibrahim H. Demirci, Mustafa Kurt: 587 Determination of Surface Qualities on Inclined Surface Machining with Acoustic

Sound Pressure

Roman Kunič: 597 Vacuum Insulation Panels (VIP) - An Assessment of the Impact of Accelerated

Ageing on Service Life

Marko Perkovic, Lucjan Gucma, Marcin Przywarty, Maciej Gucma, Stojan Petelin,

Peter Vidmar: 607 Nautical Risk Assessment for LNG Operations at the Port of Koper

Janez Sušnik, Roman Šturm, Janez Grum: 614 Influence of Laser Surface Remelting on Al-Si Alloy Properties

Journal of Mechanical Engineering - Strojniški vestnik

10 year 2012 volume 58 no.

Strojniški vestnik Journal of Mechanical Engineering

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

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58 (2012) 10

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Strojniški vestnik Journal of Mechanical Engineering

Journal of Mechanical Engineering - Strojniški vestnik

F. Tasič, Marjeta Terčelj-Zorman, Matej Zajc: ronchoscopy Image Processing in the Selected Colour

year

no. 10 2012 58

volume

Cover: Cover page shows special shape of developed 3DOF MR compatible haptic mechanism and the kinematic test of the mechanism in MR environment. Figure also presents limited working space within the tunnel of MR scanner. This is the basis for manipulability analysis and development of visualization method for representation of manipulability indices in a limited space. The analysis and optimization of mechanisms shape is performed in order to achieve high quality force transmission.

Image Courtesy: Laboratory for Computer Aided Design, Faculty of Mechanical Engineering, University of Maribor

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

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Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10 Contents

Contents Strojniški vestnik - Journal of Mechanical Engineering volume 58, (2012), number 10 Ljubljana, October 2012 ISSN 0039-2480 Published monthly

Papers Matej Rajh, Srečko Glodež, Jože Flašker, Karl Gotlih: Manipulability of a Haptic Mechanism within the Cylindrical Space of an MR Scanner David Zaremba, Christian Biskup, Thomas Heber, Nico Weckend, Werner Hufenbach, Frank Adam, Friedrich-Wilhelm Bach, Thomas Hassel: Repair Preparation of Fiber-Reinforced Plastics by the Machining of a Stepped Peripheral Zone Xiwen Zhang, Xiaodong Wang, Yi Luo: An Improved Torque Method for Preload Control in Precision Assembly of Miniature Bolt Joints Arif Gok, Cevdet Gologlu, Ibrahim H. Demirci, Mustafa Kurt: Determination of Surface Qualities on Inclined Surface Machining with Acoustic Sound Pressure Roman Kunič: Vacuum Insulation Panels - An Assessment of the Impact of Accelerated Ageing on Service Life Marko Perkovic, Lucjan Gucma, Marcin Przywarty, Maciej Gucma, Stojan Petelin, Peter Vidmar: Nautical Risk Assessment for LNG Operations at the Port of Koper Janez Sušnik, Roman Šturm, Janez Grum: Influence of Laser Surface Remelting on Al-Si Alloy Properties

563 571 578 587 598 607 614

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 563-570 DOI:10.5545/sv-jme.2009.085

Paper received: 2009-07-13, paper accepted: 2012-07-12 © 2012 Journal of Mechanical Engineering. All rights reserved.

Manipulability of a Haptic Mechanism within the Cylindrical Space of an MR Scanner Rajh, M. – Glodež, S. – Flašker, J. – Gotlih, K. Matej Rajh1,* – Srečko Glodež2 – Jože Flašker3 – Karl Gotlih3

Plus d.o.o., Slovenia of Natural Science and Mathematics, University of Maribor, Slovenia 1 Faculty of Mechanical Engineering, University of Maribor, Slovenia

2 Faculty

1 Rajh

The aim of this paper is to present developed 3 DOF haptic mechanism and 3D visualization method for analysis of mechanism manipulability problems within limited space. Improvement in mechanism manipulability within cylindrical space is crucial for devices, which operate in MR tunnel. This solution enables the plotting of quantitative 3D representation for each point in the mechanism’s workspace, using selected resolution which can be determined in advance. The cross-section between the limited space and the whole arbitrary workspace shows the ability for movement execution.

Key words: manipulability, Jacobian matrix, workspace, haptic mechanism, MR compatibility, limited workspace

0 INTRODUCTION The choice of a robotic mechanism depends on the task or the type of work to be performed and is determined by the positions of the robots, their dimensions, and structure. In general, this selection is done through experience and intuition; therefore, it is important to formulate a quantitative measurement for the robotic system’s manipulation capability, which can be useful during robot control and in trajectory planning. In regard to this perspective, Yoshikawa proposed the concept of kinematic manipulability measurements as described in [1].

Magnetic resonance (MR) compatible haptic devices have certain special structural and operational properties because of the magnetic field and MR scanner shape. A few papers relating to this problem have been published recently ([1] to [4]). Dovat et al. described a mechanical interface to use in conjunction with functional magnetic resonance imaging (fMRI). Two designs were retained and implemented from MR compatible materials. They suggested that simpler interfaces using potential mechanical energy can produce position dependent force fields during arm movements, and can be used to study the brain mechanisms. In another paper from Dovat et al. was presented a 2 degrees of freedom (DOF) haptic interface, which is driven by hydrostatic transmission separated into a master and an MR compatible slave system. Khanicheh et. al. presented the design, fabrication and preliminary testing of a novel, one degree of freedom, MR compatible, computer controlled, variable resistance hand device that may be used in brain MR imaging during hand grip rehabilitation. A novel feature of the device is the use of Electro-Rheological Fluids (ERFs) to achieve

tunable and controllable resistive force generation.

Moreover, the development of the first magnetic resonance imaging (MRI)-compatible robotic system capable of automated brachytherapy seed placement was introduced by Muntener et. al.

However, the above quoted papers do not consider the manipulability of MR mechanisms as one of the most significant design attributes for quality force transmission in haptic interfaces when considering a limited working space. This can play an important role in the mechanism’s design because operation close to the singularity point could lead to serious operating problems. Only rare articles ([5] and [6]) are involved in the area of manipulability analyses in haptic devices, but they do not include the particularity of limited space and MR compatible mechanisms.

In addition, when we talk about operation within a limited workspace any manipulability problem becomes more complex and any improvements in mechanism design or the position of the mechanism’s base could be even more significant.

In this work, the limited workspace is presented by a block placed inside the MR scanner bore with cross-section of 200×350 mm or with the diagonal length equal to human forearm as shown in Fig. 1. The main constraint of this workspace is its very high density of magnetic field. For this reason, safety is crucial during the robot operation within an MR environment exposed to a strong magnetic field of 1.5 to 3 T. The robot, must therefore, be insensitive to the imaging sequence and should not disturb the imaging itself. Functional magnetic resonance imaging scanning sequences are even more sensitive to inhomogenities of the magnetic field than MRI sequences [7] to [9].

*Corr. Author’s Address: Rajh Plus d.o.o., Črešnjevec 143, 2310 Slovenska Bistrica, Slovenia, matej.rajh@rajh.eu

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handle to interact with the explored world, the user feels the apparent mass, compliance and friction of the interface. This distortion introduced between the operator and the virtual environment must be identified in order to enhance the design of the device and develop appropriate control laws. The device’s workspace should be large enough to cover or exceed the required workspace. Furthermore, a compromise is needed between the haptic device workspace size and the available output forces [16]. In order to reach good “virtual feeling” transmissions, the device must move very easily. Ideally, weight, friction, backlash, slip, material deformations etc would be absent.

Therefore, good manipulability is also very important.

1.2 MR Compatibility and Choice of Materials Fig. 1. Red colour indicates the task space inside the MR tunnel

The main aim of this work was to design a new 3DOF fMRI compatible haptic mechanism that can be driven by electric motors, despite usually being banned from the MR environment. However, the motors have excellent control possibilities. Some unique changes have been made in mechanism design, which become reasonable when considering the mentioned structural and geometrical limitations.

A kinematic model to optimize mechanism characteristics has been proposed to provide good manipulability of the mechanism and clearly present dexterity properties. The model visualizes manipulability characteristics in the tunnel of the MR device and enables the interpretation of possible improvements in sections of the volume.

1 MR COMPATIBLE HAPTIC MECHANISM 1.1 Haptics An haptic interface is a force feedback device, which enables its user to interact with a virtual world or a remote environment explored by a slave device. It aims at matching the force and displacements given by the user and those applied to the virtual world. Such systems are in growing demand for applications such as force feedback remote-control systems for extreme environments, man-machine interaction, training in professional operating procedures and rehabilitation ([10] to [16]).

Usually, haptic interfaces make use of a mechanically actuated structure whose distal link is equipped with a handle. When manipulating this 564

Over the past few years, fMRI has established itself as a major research tool for investigating the brain mechanisms of motor control and cognition.

Performing arm movements in controllable dynamic environments during fMRI could provide important insights into human motor control and related dysfunctions, and enable therapists to quantify, monitor, and improve physical rehabilitation [1] to [3], [17].

The major problem when creating fMRI compatible robots is the strong magnetic field needed for MRI (1.5 to 3 T) precluding the use of conventional materials or actuators close to the scanner bore. This requirement prevents the use of conventional robotic interfaces. MR environment refers to the general environment near an MR scanner. In particular, it includes the area encompassed by the 10 mT line. This may or may not include the entire magnet room and surrounding support areas [18].

By referring to the device design criteria for fMRI environment introduced by Hartwig et al. [18], we decided that this new mechanism would be composed of composite materials, ceramic passive mechanical devices and ropes from high strength polymer fibres.

Carbon and advanced ceramics fit well within the MR environment even inside an MR tunnel [19].

Plastics are easily machined but do not have the strength needed to build a light structure with low inertia. Non-ferrous metals (Aluminium, Beryllium, Copper, etc.) are desirable for their non-magnetic properties and strength. Non-ferrous metals may contain some impurities, which would cause certain magnetic properties. Heating and stress to the structure may also introduce some undesirable magnetic properties and could cause some artifacts within an image [20].

Rajh, M. – Glodež, S. – Flašker, J. – Gotlih, K.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 563-570

The main problem with composites, high strength fibres and ceramic mechanical elements is their cost, availability of elements, and lack of appropriate data regarding exact mechanical properties.

Another problem besides image artefacts is radiofrequency (RF) noise. It often appears as static on the image and can be caused by an electrical device located anywhere in the MR procedure room. Some members of our work group conducted an experiment on the same fMRI scanner using a reshaped Phantom haptic device [12], which was far enough from the scanner bore and had a lengthened handle for MR influence exclusion. The exact results have not yet been published, but this experiment did not show any influence on image quality.

2 KINEMATIC DESIGN OF THE PROPOSED MECHANISM 2.1 Kinematic Requirements When operating in an MRI tunnel, it is quite difficult to decide which mechanical structure is appropriate because the workspace is limited and it is impossible to incorporate a parallel mechanism with a high level of stiffness. On the other hand, according to Lee and Lee [21], serial manipulators are unsuitable for haptic devices, because of low level stiffness. Therefore, we searched for a combined mechanism, which had to be driven on its base and with actuators mounted away from the tunnel, as far as possible. It is for this reason that MR compatible haptic devices are mostly developed as a 2DOF devices [1] to [3], [9], [22] and [23].

has following dimensions: d1 =140 mm, a2 = 470 mm, a3 = 115 mm, a4 = 565 mm, b4 = –80 mm.    px  C1 ( a2 C2 + ( a3 + a4 ) C23 − b4 S 23 )  P =  p y  =  S1 ( a2 C2 + ( a3 + a4 ) C23 − b4 S 23 )  , (1)    pz   d − b C − a S − ( a + a ) S  1 4 23 2 2 3 4 23  

with Si = sin(qi), Ci = cos(qi), Cij=cos(qij), Sij=sin(qij) and qij= qi+ qj .

2.2 Mechanism Design The kinematical background as described previously can serve as a basis during the design process of a treated haptic mechanism and also for its final production and exploitation in a MR compatible haptic interface. An important problem to be solved is how to ensure that the mechanism is as efficient as it is in an open space. In order to do this, the shape of the mechanism and handle must be carefully designed. The mechanism consists of four bars with innovative curved shapes, as shown on Fig. 2, that assure accessibility to each experimental point in the MR tunnel, high stiffness and good manipulability characteristics. The shape is adapted to the tunnel of the MRI scanner. So far, an aluminium mock-up has been built for verification of kinematic suitability (Fig. 3). Moreover, this same mechanism can be upgraded with drives and controls to test a full haptic performance. For the prototype are used Maxon DC motors and special haptic controls developed by Laboratory for Robotics at the Faculty of Electrical Engineering in Ljubljana.

Fig. 2. Schematic view of curved haptic mechanism

A schematic view of the treated haptic mechanism is shown on Fig. 2, where “0” indicates global coordinate system and P (Eq. (1)) is the position at the top end of the mechanism. It was derived using Denavit-Hartenberg parameters [24]. The mechanism

Fig. 3. Manufactured aluminium mechanism used for kinematic tests

Manipulability of a Haptic Mechanism within the Cylindrical Space of an MR Scanner

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Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 563-570

The mechanism is composed of two main segments, which are driven by a capstan or cable drive, which is the most widely used driving solution in haptic devices due to its low backlash (almost zero), stiffness, backdrivability and simplicity [12], [16], [25] and [26]. Backdrivabililty is a measure of how accurately a force or motion applied at the top end is reproduced at the input end.

In a mechanical robot-like linkage, good backdrivability means that a person can grasp the tip of the linkage and move it around effortlessly [16].

Fig. 4 shows the placement of a complete mechanism within a MR environment and its position in regard to the patient.

the generation of force and moment along a certain direction should be done easily in order to deliver a precise force to the operator. Therefore, it is necessary to ensure consistent manipulability characteristics within the workspace [21] and [28].

The velocity and force transmission characteristics of a manipulator at any posture can be represented geometrically as ellipsoids (Fig. 5). The velocity ellipsoid is a useful tool for visualizing the velocity transmission characteristics of a manipulator at a specific posture.

3 MANIPULABILITY In the literature manipulability is originally defined as a measure of a robotic structure’s performances, normally given in the force domain by means of manipulability ellipsoids or polytopes [5]. The manipulability index proposed by Yoshikawa [27] is defined by the following equation: w = det ( J ) , (2)

which represents the volume of the velocity ellipsoid.

Fig. 5. Velocity ellipsoid with principal axes Consider n-degree of freedom manipulator with configuration space coordinates q (internal n coordinates), with qi ∈ ℜ , and a task described m by the vector x, x j ∈ ℜ , and m ≤ n. The geometric transformation from configuration space to task space (external coordinates) is:

Fig. 4. The model of MRI device with patient and mounted mechanism

Close to the singularity point the volume of the ellipsoid is still large and imprecisely reflects the closeness of singularity because the parameter is not sensitive enough. This is the main reason for the shortest ellipsoid axes being used as the quantitative measurement of the closeness of a manipulator to singularity.

A haptic device should have excellent mobility to facilitate the operator giving orders. In addition, 566

x = x (q) .

(3)

Differentiating x with respect to the time, we obtain:

x = Jq , (4)

where J is the m × n Jacobian matrix which can be obtained as follows:

Rajh, M. – Glodež, S. – Flašker, J. – Gotlih, K.

J=

∂x , (5) ∂q

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 563-570

and consequently for our mechanism according to Eq. (8).

The Jacobian is simply a linear transformation that maps the joint velocity in ℜn into task velocity in ℜm . The unit sphere in ℜn is defined by:

q&

2

= q&12 + q&22 + L + q&n2 ≤ 1, (6)

and is mapped into an ellipsoid in ℜm defined by:

−1 x T JJ T x ≤ 1. (7)

(

)

The principal axes of the velocity ellipsoid coincide with the eigenvectors of ( J J T ) –1 and the length of a principal axes is equal to the reciprocal of the corresponding eigenvalue’s square root [29].

 − S1 ( a2 C2 + ( a3 + a4 ) C23 − b4 S 23 ) −C1 ( b4 C23 + a2 S 2 + ( a3 + a4 ) S 23 ) C1 ( b4 C23 + a2 S 2 + ( a3 + a4 ) S 23 )    J =  C1 ( a2 C2 + ( a3 + a4 ) C23 − b4 S 23 ) − S1 ( b4 C23 + a2 S 2 + ( a3 + a4 ) S 23 ) S1 ( −b4 C23 − ( a3 + a4 ) S 23 )  .   0 −a2 C2 − ( a3 + a4 ) C23 − b4 S 23 − ( a3 + a4 ) C23 + b4 S 23     A vector v, with its origin in the centre of the velocity ellipsoid (TCP – tool centre point, which is equal to the top of the mechanism in our case) and magnitude equal to the length from the ellipsoid centre to a point on the surface of the ellipsoid, represents the velocity vector in this particular direction. If the length of the vector is equal in all directions this mechanism posture is isotropic. In this case, the velocity ellipsoid is a sphere. In some particular mechanism postures, the ellipsoid loses one of its dimensions. The ellipsoid degenerates into an ellipse. These particular postures are the singular postures of the mechanism. In the cases of these singularities, the TCP can only move in directions within the plane of the ellipse.

The velocity ellipsoid indicates the manipulability of the manipulator in any posture. The highest manipulability is obtained in the direction of the longest principal axes of the ellipsoid. The lowest manipulability is in the direction of the shortest principal axes.

The robot’s working space is velocity anisotropic [30]. This fact restricts the working space to a subset where the required TCP velocities can be performed and others where the required velocities can not be obtained.

The parameter for measuring the manipulability of the robot in any posture is the shortest length of the principal axis of the velocity ellipsoid. If the length of the shortest principal axis at a point within the working space is long enough, the robot will be able to perform the required manipulability, and the robot’s posture and the corresponding point are acceptable. If not, this point within the working space is unacceptable.

The optimal direction for effecting velocity is along the major axis of the ellipsoid, where the

(8)

transmission ratio is at a maximum. Conversely, the velocity is most accurately controlled along the minor axis of the ellipsoid, where the transmission ratio is at a minimum. Velocity is the most accurately controlled in the direction where the manipulator can resist large disturbance forces, and force is most accurately controlled in the direction where the manipulator can quickly adapt its motion [16] to [18], [21], [23] to [25], [28] and [31].

4 VISUALIZATION OF MANIPULABILITY INDICES 4.1 3D Representation of Manipulability A computational model was developed, within the framework of the presented research which is able to analyze the manipulability properties of the mechanism with a predefined Jacobian matrix.

Volumetric representation of the manipulability index is used for the representation of calculated data.

In the developed model, the mechanism workspace is discretised. An equidistant mesh of points is a substitute for the whole set. The manipulability of the mechanism at each point of the mesh has a scalar value between 0 and 1 and is defined as the shortest axis of the velocity ellipsoid.

A value for manipulability index of 1 means that the mechanism has the highest manipulability at the treated discretised point. If the value is 0, the mechanism has a singular position.

To find the best position for task space within the mechanism’s workspace, the algorithm moves the task space through a set of discrete points within the workspace. At each position it counts the number of points regarding the mechanism’s workspace that

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are located within the task space. From all possible positions, it chooses the position within task space that has the highest number of counted points with adequate manipulability.

21 slice planes (Fig. 7) and the best plane from the manipulability point of view is plane number 16 (Fig. 7).

4.2 Graphical Manipulability Analysis within a Limited Space When a human moves the mechanism handle inside the tunnel it is very important that there is the possibility to adapt the best dexterity characteristics within the area where the movements are to be performed. Therefore, this method is very useful for the adaptation of a mechanism’s base position when the task changes. The operator can quickly define a new optimal position, obtain better force transmission, and more accurate results.

4.3 Results The developed program is able to show the mechanism’s workspace (Fig. 6) and appurtenant indices of manipulability within an open space or indices in a predefined limited workspace (Fig. 7). In the case of an MRI scanner, the entire workspace is half of a cylinder with 60 cm diameter, but it is usually reduced to the task space. Tests for different shapes were also performed, mainly for sphere and block and showed great results for arbitrary shapes.

The lowest manipulability index in the useful area of MR tunnel is 0.5, such a high value indicates that the mechanism’s design is appropriate for haptic applications. In this particular case, there are

Fig. 6. 3D-plot of manipulability indices in an open workspace at slice plane 15 of 21 (the white point indicates the origin coordinates of the mechanism and the green frame indicates limited space)

Every task of a patient’s hand movements has different requirements and positions inside the tunnel.

With the help of 3D graphical representation, it is possible to evaluate manipulability characteristics in different positions, thus enabling the user to adapt the mechanism’s base position according to a direction for better manipulability index.

5 CONCLUSIONS A new mechanism as a force-feedback device to be used for evaluating brain activation by movements

Fig. 7. Slice planes of workspace (left) and 3D-plot of manipulability indices in limited workspace of MR tunnel (slice plane xy - number 16)

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of a human’s upper extremity has shown good manipulability properties inside a limited space.

This mechanism should be installed within an MR environment and for this reason it has to be MR compatible and all its active mechanical elements placed outside the critical zone using an outstandingly high density of magnetic field.

Low inertia and the ability to move very easily are the main design requirements owing to their haptic functions. Low inertia is achieved by using a light structure and high strength materials. On the other hand, it is easy to move the mechanism because of manipulability optimization. The method of 3D representation is very useful for imaging the properties of different placements of the mechanism within a limited space. It helps us to reach better mechanism positions and assures higher manipulability indices inside the workspace of the MR tunnel.

The developed computer program is also useful as a design tool for robot production cells as it enables better insight into workspace or more efficient use of workspace. The aim of our work in the future is the development of user friendly interface for production cells and haptic interface design, and their off-line programming by considering optimal manipulability characteristics. Additional work can also be done on the export of 3D plots in different CAD formats, thus providing the possibility of importing workspaces with common manipulability indices into all major PLM systems.

6 ACKNOWLEDGEMENTS The authors thank professor Marko Munih and Aleš Hribar for their advice in the area of haptic mechanisms and MR compatibility, and all the staff at the University Medical Centre Ljubljana for enabling the testing of the mechanism’s mock-up.

7 REFERENCES [1] Dovat, L., Gassert, R., Chapuuis, D., Ganesh, G., Burdet, E., Bleuler, H. (2005). A haptic interface based on potential mechanical energy to investigate human motor control using fMRI. Proceedings of the IEEE Engineering in Medicine and Biology, 27th Annual Conference Shanghai, p. 5021-5024.

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25-34, DOI:10.1007/11552246_3.

[10] Payandeh, S., Dill, J., Zhang, J. (2007). Using haptic feedback as an aid in the design of passive mechanisms.

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[11] Hirabayashi, T., Akizono, J., Yamamoto, T., Sakai, H., Yano, H. (2006). Teleooperation of construction machines with haptic information for underwater applications. Automation in Construction, vol. 15, no.5, p. 563-570, DOI:10.1016/j.autcon.2005.07.008.

[12] Selected OEMs & System Integrators of SensAble Products (2008). Retrieved at 2008-12-18, from http:// www.sensable.com/oems-integrators.htm.

[13] Wang, P., Becker, A.A., Jones, I.A., Glover, A.T., Benford, S.D., Greenhalgh, C.M., Vloeberghs, M. (2006). A virtual reality surgey simulation of cutting and retraction in neurosurgery with force-feedback.

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[16] Mali, U., Munih, M. (2006). HIFE-Haptic Interface for Finger Exercise. IEEE/ASME Transactions on Mechatronics, vol. 11, no. 1, p. 93-102, DOI:10.1109/ TMECH.2005.863363.

[17] Baudendistel, K., Schad, R.L., Wenz, F., Essig, M., Schröder, J., Jahn, T., Knopp, V.M., Lorenz, J.W. (1996). Monitoring of task performance during functional magnetic resonance imaging of sensorimotor cortex at 1,5T. Magnetic Resonance Imaging, vol. 14, no.1, p. 51-58, DOI:10.1016/0730-725X(95)02052-U. [18] Hartwig, V., Vanello, N., Gaeta, G., Sgambelluri, N., Scilingo, E.P., Bicchi, A. (2008). Design of fMRI Compatible Actuators. Retrieved at 2008-12-09, from http://www.touch-hapsys.org/.

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[22] Meneses, J., Castejón, C., Corral, E., Rubio, H., GarciaPrada, J.C. (2011). Kinematics and dynamics of the quasi-passive biped “PASIBOT”. Strojniški vestnik – Journal of Mechanical Engineering, vol. 57, no. 12, p.

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[24] Tsai, L.W. (1999). Robot analysis, The Mechanics of Serial and Parallel Manipulators. John Wiley & Sons, New York.

[25] Gosselin, F., Bidard, C., Brisset, J. (2005). Design of a high fidelity haptic device for telesurgery.

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[26] Hayward, V., Oliver, R.A., Hernandez, M.C., Grant, D., Robles-De-La-Torre, G. (2004). Haptic interfaces and devices. Sensor Review, vol. 24, no. 1, p. 16-29, DOI:10.1108/02602280410515770.

[27] Yoshikawa, T. (1985). Manipulability of robotic mechanisms. Hanafusa, H., Inoue, H. (eds.): Proceedings of Robotic Research: the 2nd International Symposium, p. 439-446().

[28] Rajh, M., Glodež, S., Flašker, J., Gotlih, K., Kostanjevec, T. (2011). Design and analysis of an fMRI compatible haptic robot. Robotics and ComputerIntegrated Manufacturing, vol. 27, no. 2, p. 267-275, DOI:10.1016/j.rcim.2010.06.007.

[29] Chiu, S.L. (1988). Task compatibility of manipulator postures. The International Journal of Robotic Research, vol. 7, no. 5, p. 13-21, DOI:10.1177/027836498800700502.

[30] Gotlih, K., Kovač, D., Vuherer, T., Brezovnik, S., Brezočnik, M., Zver, A. (2011). Velocity anisotropy of an industrial robot. Robotics and ComputerIntegrated Manufacturing, vol. 27, no. 1, p. 205-211, DOI:10.1016/j.rcim.2010.07.010.

[31] Fonseca Ferreira, N.M., Tenreiro Machado, J.A. (2008).

Manipulability analysis of two-arm robotic system, retrieved at 2008-12-19, from http://ave.dee.isep.ipp. pt/~gris/_private/Nuno_INES00.PDF.

Rajh, M. – Glodež, S. – Flašker, J. – Gotlih, K.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 571-577 DOI:10.5545/sv-jme.2012.305

Paper received: 2012-01-05, paper accepted: 2012-08-23 © 2012 Journal of Mechanical Engineering. All rights reserved.

Repair Preparation of Fiber-Reinforced Plastics by the Machining of a Stepped Peripheral Zone Zaremba, D. – Biskup, C. – Heber, T. – Weckend, N. – Hufenbach, W. – Adam, F. – Bach, Fr.-W. – Hassel, T. David Zaremba1,* – Christian Biskup1 – Thomas Heber2 – Nico Weckend2 – Werner Hufenbach2 – Frank Adam2 – Friedrich-Wilhelm Bach1 – Thomas Hassel1 1 Leibniz

Universität Hannover, Germany Universität Dresden, Germany

2 Technische

Today’s standard procedures for the repair of fiber reinforced plastics are not optimized for the structural rearrangement of the original material properties. Alongside lap repair and scarf repair, a newly introduced method for the machining of a stepped peripheral zone is discussed. For this purpose, the methods of dry ice blasting and snow blasting as well as the water jet are being investigated. Reference material is carbon fiber reinforced plastic (CFRP) compliant to laminates used in the aviation industry.

It was found that snow blasting and dry ice blasting were not suitable for this purpose in the experimental set-up. In contrast, the water jet allowed a precise control of the material removal. Subsequently to the feasibility study, a parameter study was carried out to determine applicable parameters for the surface preparation of CFRP. After successful machining of a stepped peripheral zone, a repair experiment was carried out with promising results.

Keywords: surface preparation, carbon fiber reinforced plastic, CFRP, water jet, dry ice blasting, snow blasting, CFRP repair, stepped peripheral zone

0 INTRODUCTION Initiated by the lightweight construction trends during the last few years, a certain group of composite materials has gained an increased interest in the industry: fiber-reinforced plastics (FRP) with a thermo- or duroplastic matrix. Materials out of this group find extensive use in aeronautical and aerospace engineering, as well as in vehicle construction and many other application areas, where high strength and low weight are required.

Up to now, a big disadvantage of fiber-reinforced plastics has been their reparability in case of damage [1]. This is primarily based on the fact that the individually designed mechanical properties of such a material are dependent of shape and orientation of the fiber material, which have to be rebuilt.

For the repair of carbon fiber reinforced plastics (CFRP) different mechanisms have been discussed.

These mechanisms include different forms of lap repair, where additional repair plies are attached to the surrounding surfaces [2]. Ahn et al. presented different designs methods for this repair technique (stepped lap, uniform lap, single/double-sided repair). Certainly, the basic structural properties of fiber reinforced plastics are not reconstructed using these methods. Instead, loads are led to bypass the damaged or empty section through the additional material.

Unlike lap repair, scarf repair methods are intended to restore the original load distribution and are therefore the repair method of choice. Repairs according to this approach have been made with soft patches [3], as well as with hardened patch material

[4]. Ahn et al. found that the use of prepreg patches results in higher failure loads than the use of wet layups at room temperature. For the additional support of scarf repairs, external repair plies can be applied to seal the repair area.

As introduced by Ahn and Baker, the applied methods are nowadays mostly based on manual machining of the damaged parts. In addition, Baker investigated the adoption of moulded and CNCmachined hard-patches using scarf repair in order to avoid material inhomogenities caused by manual handling. Unfortunately, this method requires a high machining effort and is not practicable. Furthermore, no mechanical values of repaired material have been published.

Using scarf repair, even for small defects a large bonding area is required in order to minimize the stiffness difference between the damaged material and the patch [5]. Beyond that, the risk of material property variations is given because of a high manual machining effort. Consequently, there is a need for automated processing procedures.

However, a reproducible preparation of the composite material is necessary to investigate the aspect of quality in the repair process sufficiently.

The jet machining of 3-dimensional free-form surfaces has successfully been researched by Öjmertz [6] and Borkowski [7]. Due to the use of ceramics and aluminium as the workpiece, the material removal was managed by the abrasive water jet (WAIS) in both publications. When machining FRP, this method would result in an immediate cut-through of the workpiece. Therefore, less abrasive methods are

*Corr. Author’s Address: Institute of Materials Science, Leibniz Universität Hannover, An der Universität 2, D-30823 Garbsen, Germany, zaremba@iw.uni-hannover.de

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researched. Pilot testing has shown that a selective removal of individual CFRP and GFRP layers is possible in general. This gives the new opportunity to adjust the geometric shape of the peripheral repair zone to the patch material. By removing single material layers in a stepped form, a defined surface for the integration of the pre-impregnated (prepreg) patch material is created.

1 MATERIALS AND METHODS Several jetting methods with different materialremoval mechanisms offer a high potential for the selective removal of individual FRP layers.

Promising methods include snow blasting, dry ice blasting and pure water jetting. The qualification of these techniques for the manufacturing of a stepped peripheral zone (Fig. 1) will be analyzed and discussed in this paper.

1.2 Snow Blasting Similar to the dry ice blasting, CO2 snow blasting can be classified as a compressed-air blasting method.

Liquid carbon dioxide is led to a chamber, where it is subsequently expanded through a restrictor. At this point, the Joule-Thomson effect causes an abrupt cooling down to temperatures lower than -78.5 °C. By this effect, a part of the carbon dioxide freezes to solid snow particles. The particles are gripped by the air blast and accelerated towards the work piece through a convergent-divergent nozzle.

1.3 Water Jet In the last decades, the high pressure water jet technology has expanded to an important industrial branch. The main applications are the cleaning and cutting (abrasive water jet as well as pure water jet) of a wide-spread material variety [9], other common methods are e.g. peening and surface preparation [10].

Basically, the water jet is generated by decompression from the operating pressure in a nozzle. It can either be used for an erosive impact directly on the work piece, or, in case of the abrasive water injection jet (AWIJ), to accelerate an abrading medium [11]. For the designated purpose, the pure water jet is used.

Fig. 1. Stepped peripheral zone for optimal patch integration (step width b and gap size a)

1.1 Dry Ice Blasting Dry ice blasting can be separated into two groups depending on how the dry ice supply is provided. The usage of dry ice blocs requires a breakup into particles immediately before the blasting process. The variety of particle sizes in this group is comparatively high. In order to provide a preferably homogenous process, the application of ready-to-use dry-ice pellets is the more suitable option [8].

The pellets are fed into an air blast by a vibrating plate with an attached metering unit. The resulting mixture of pressurized air and dry ice pellets is led to the nozzle through a flexible tube. Subsequently, the blast is led to the work piece, where the dry ice acts as the blasting agent.

572

Fig. 2. Schematic states of a water jet, according to [8] and [9]

Caused by environmental effects as well as inner turbulences, the water jet alters its geometrical form and its fluid dynamical characteristics in dependence of its distance to the water nozzle. The basic assumption is that the forming water jet has a constant

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velocity profile directly in the nozzle section. Induced by friction with the environment, behind the nozzle a mixing zone develops subsequently and grows larger with increased distance to the nozzle [9]. As a result, the jet expands (compact jet, see Fig. 2).

For larger nozzle distances, the water jet changes its characteristics again and forms droplets [9]. The impact of single droplets can significantly generate higher peak loads than a continuous jet, because every impact causes stress to the work piece for a short time [10].

Experiments were made with three different blasting nozzles. These included two round nozzles (diameter 6 and 11 mm) and one flat jet nozzle (12×3 mm).

The constant CO2 bottle pressure of 5.7 MPa at a temperature of 20 °C allowed the determination of the fill level by weight measurement. The flow was controlled by an integrated flow regulator. A pressurereducing valve made the regulation of the air blast pressure possible. Used parameters are listed in Table 2.

Table 2. Parameters for snow blasting

1.4 Experimental Setup The test material for all experiments was a symmetrical (0°/90°/90°/0°) CFRP layup (Fig. 3). Each prepreg layer had a thickness of 0.3 mm. Corresponding to common methods in the aviation industry, the laminates were hardened using an autoclave.

Parameter Air blast pressure [MPa] CO2 mass flow rate [g/s] Feed rate orientation [°] Working distance [mm] Feed rate [mm/s]

min.

0.2 1.67 0 30 0

max.

1.5 5 90 100 15

For the feasibility study, the blasting pistol was oriented vertically to the work piece. The axis manipulation was started, when erosion on the work piece was evident.

For the water jet experiments, a “Baldor” CNC guiding machine was used with a “Uhde HP19/37” high pressure intensifier pump. Used nozzles were standard sapphire nozzles with diameters from 0.1 mm up to 0.25 mm. The varied parameters are listed below (Table 3).

Fig. 3. Microsection of an intact (90°/0°/0°/90°) CFRP layup

The dry ice blasting unit used was a “Linde Cryoclean Cryomax Plus”, attached to a 3-axis industrial robot. Used parameters are listed in Table 1. The feed rate of 0 mm/s is equivalent to a stationary jet. All experiments were carried out with a 27×5 mm flat nozzle, both with and without masking. The tool was guided orthogonally to the work piece.

Table 1. Parameters for dry ice blasting Parameter Air blast pressure [MPa] Shaker pressure [MPa] Feed rate orientation [°] Working distance [mm] Feed rate [mm/s] Cycles

min.

0.2 0.1 0 30 0 1

max.

1.5 0.15 90 100 160 100

For the CO2 snow blasting a “CryoSnow SJ-10” unit with an attached “CryoSnow JP-10” blasting pistol was used. It was attached to the same guiding machine as in the dry ice blasting application.

Table 3. Water jet parameters Parameter Jet pressure [MPa] Feed rate orientation [°] Working distance [mm] Feed rate [mm/s] Nozzle diameter [mm]

min.

30 0 45 0 0.1

max.

400 90 100 15 0.25

The material removal of all specimens was measured using a “Rodenstock RM-600” laser measuring device. All experiments were carried out in two feed rate directions, depending on the orientation of the first fiber layer. In the first instance, the design of experiments scheduled a feasibility study for all three blasting/jetting methods. In the next step, an ensuing parameter test was done for the method with the best test results.

The repaired laminates were prepared for tension testing using the abrasive water injection jet. The material testing was made using a universal tension/ pressure testing machine with a 250 kN load cell. The testing speed was 2 mm/min.

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2 FEASIBILITY STUDY RESULTS 2.1 Dry Ice Blasting Results The experimental findings of the dry ice blasting tests showed that the method offers enough power for the layer removal of CFRP. During the stationary operation, a complete removal of all layers was reached even at low pressures (0.2 to 0.5 MPa).

Blasting pressures of more than 1.1 MPa resulted in a large-area material damage, caused by the high kinetic energy of the dry ice particles. The bottom layer was detached from the laminate and started to vibrate. This caused an uncontrolled delamination. The best results were reached either with a high feed rate and medium pressure (0.6 to 1 MPa), or with a low feed rate and low pressure (0.2 to 0.5 MPa). Both variants required multiple cycles.

The edge sharpness of the blasting results could be improved significantly by using a metal mask in the process. Although the blasting results with a fiber-corresponding feed direction were slightly more continuous than in the orthogonal direction, none of the tested parameter sets could provide a homogenous and defined material removal. Position and depth of the abrasion track were subject to fluctuations, which partially included multiple layers (Fig. 4). For this reason, no further experiments were performed using this method.

Fig. 4. Exemplary abrasion track after dry ice blasting of CFRP

2.2 Snow Blasting Results By the first experiments it could be shown that the snow blasting method met its performance limits at the abrasion of CFRP. At low blasting pressures (0.2 to 0.5 MPa) no appreciable erosion was recognized with any parameter combination. Medium blasting pressures (0.6 to MPa) in combination with low working distances (30 to 40 mm) led to a slow removal of the first CFRP layer in stationary operation. The second layer could be laid open after a blasting time of 4 minutes. Unfortunately, even the slowest feed rate of 0.25 mm/s was not sufficient for a satisfactory material removal during moving operations. The removal of the first layer in movement could only be 574

reached at high blasting pressures (1.1 to 1.5 MPa) combined with a low feed rate of 0.25 mm/s. The best results using this method could be achieved with the 6 mm round nozzle, which offers the highest power density. An example is pictured in Fig. 5.

Fig. 5. Exemplary abrasion track after snow blasting of CFRP

Basically, it was determined that the tested parameter sets do not permit a constant and geometrically controlled material removal. Especially, the reproducibility was not given in repeated experiments. This can be attributed to the discontinuity of the blast, which is influenced by the breaking down of icings from the nozzle. The geometric position and depth of the erosion could not be controlled.

2.3 Water Jet Results With regard to the snow- and dry ice- blasting processes, the water jet offers a significantly higher erosive potential. Pressures of more than 100 MPa in combination with the tested standard sapphire nozzles effected an immediate cut-through. After the pressure range was reduced to 30 to 45 MPa, it was possible to remove single material layers with high feed rates.

The combination of the water jet’s small processing area together with the exact positioning of the guiding machine allows a precise control of the material removal. Also according to the constancy this method shows a comparatively continuous jet form.

In the experiments it was found that the quality of the produced surfaces depends on multiple factors.

A notable finding was that the tool manipulation in fiber direction (0°) resulted in more constant abrasion tracks than with orthogonal manipulation (90°). It was possible to lay open both, the second and the third material layer in one working step. Additionally, the surface quality was investigated in dependence of a track offset Δx and pressure p. For the 0.25 mm round nozzle, which offers a significantly higher performance than the 0.1 mm nozzle, the optimum was found to be Δx = 0.3 mm (Fig. 6).

The nozzle standoff distance was 60 mm. An offset that is too big results in a complete removal of

Zaremba, D. – Biskup, C. – Heber, T. – Weckend, N. – Hufenbach, W. – Adam, F. – Bach, Fr.-W. – Hassel, T.

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the first layer and a severe structuring of the second one. A pressure reduction in combination with 0° manipulation and the use of a 0.1 mm nozzle resulted in a micro-structuring of the first layer. This allowed detaching entire fiber bundles from the second layer.

Fig. 6. Tool path for the surface preparation of an area

3 PARAMETER STUDY ANALYSES AND DISCUSSION Generally speaking, very constant surfaces could be produced using the water jet method. Based on these promising results, the surface preparation of CFRP with the water jet was subject of a parameter study in the following.

Hereby the influence of important process parameters could be investigated in view of the main criteria “abrasion depth” and “damaging”. The target was to determine a parameter set, which permits an optimal layer removal as well as a minimal variation of the results. The considered parameters were: • Standoff distance z; • Water pressure p; • Feed rate f.

Additionally, the influence of the manipulating direction was investigated (0° respectively 90° to the first fiber layer). The nozzle diameter of 0.25 mm remained constant. For the parameter studies, fivelayered CFRP laminates with a layer thickness of 0.3 mm and a (0°/90°/0°/90°/0°) layer orientation were used. These laminates were manufactured using the hot press method. Every material removal track was measured in five positions.

Regarding Fig. 7 it becomes apparent that the influence of the standoff distance is affected by the manipulating direction in a severe way. This can be reasoned with a selective layer removal, which is favored at feed directions in fiber orientation. In this case, the first layer (thickness: 300 µm) is steadily removed, while the second fiber layer is not damaged by the decreased energy of the water jet. In case of an orthogonal feed rate direction in reference to the upper fiber orientation, the second layer is in favored position. Consequently, the remaining energy is still sufficient to cause an appreciable material removal depth of almost 150 µm to this layer.

Fig. 7. Kerf depth in dependence of the standoff distance (pressure: 45 MPa, feed rate: 11.67 mm/s)

Fig. 8. Kerf depth in dependence of the water pressure (standoff: 60 mm, feed rate: 11.67 mm/s)

As pictured in Fig. 8, a rise of the kerf depth can be observed with increasing water pressure. This fact confirmed the expectations based on the dependence between jet velocity and pressure. Furthermore, the test results confirm the assumptions of the selective layer removal made in the last paragraph.

Fig. 9. Kerf depth in dependence of the feed rate (standoff distance: 60 mm, pressure: 45 MPa)

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The third main factor in the parameter study was the feed rate, which was varied in five steps. The feed rate of 11.67 mm/s presented the maximum value specific for the guiding machine used. Founded on the assumption that the contact time defines the material removal rate at a constant material removal potency, a decreasing curve progression was expected for increasing feed rates (Fig. 9).

In reference to the other parameter studies in respect of standoff distance and water pressure, the measured kerf depths in Fig. 9 are slightly higher. This can be explained by the way the parameter tests were arranged. Since the low feed rates were tested at first and the following experiment’s starting point matched the former end point, the erosion was initiated in a more damaged location at the beginning. Through this approach a higher material removal rate by usage of the same jetting parameters could be achieved.

Finally, the parameter set for the removal of single material layers should be chosen according to the following two criteria: 1. the complete layer should be removed, and 2. the subjacent layers should not be damaged.

Based on these criteria, a parameter set was chosen and this allowed certain tolerances for pressure fluctuations and standoff distance variations. The test specimens were prepared for repair using this parameter set.

Also, before the hardening, the prepreg repair patches were inserted. Base laminate and repair patches were hardened simultaneously. Tensile tests were made with the test material, the reference material and an undamaged CFRP laminate.

4 CONCLUSIONS Regarding the parameter test results, the manipulation direction has a bigger influence on the material removal than e.g. the nozzle standoff distance. If possible, this should be utilized. The parameter set should allow a way of tolerance for pressure variations.

With the chosen parameters, the selective removal of single layers and the manufacturing of a stepped peripheral zone were possible (Fig. 10).

3.1 Repair of CFRP Laminates

Fig. 10. Stepped peripheral zone (CFRP) after machining with the pure water jet

Subsequent to the successful machining of the stepped peripheral zone, first repair experiments were made. Two (0°/90°/90°/0°) 4-layer CFRP laminates with a layer thickness of 0.3 mm were prepared using the above method. The step widths were 10 mm (specimen 1) respectively 20 mm (specimen 2). Afterwards, they were repaired with fitted prepreg patches. The laminates were hardened using the hot press method with a platform load of 26.1 kN. A preferable homogenous load distribution was provided by a polysiloxane mat between laminate and press mould. At the beginning, the platforms were heated to 60 °C (base temperature). From this point, the laminates were heated up to the hardening temperature of 110 °C and later cooled down to 60 °C again. The hardening temperature was held for 60 minutes, while heating-up and cooling down required 30 minutes each. Along with the mentioned laminates, a reference specimen was manufactured to symbolize an idealized repair method. The stepped peripheral zone was cut manually into each prepreg layer of this reference laminate before the hardening process.

While the repaired reference laminates with the 10 mm-stepping had a residual tensile strength of 24% (244 MPa) in comparison to the undamaged material, the water jet machined specimen still had a residual strength of 20% (197 MPa). The values for the 20 mmstepping were a little higher, the strength amounted 29% (294 MPa, reference laminate) respectively 26% (263 MPa, water jet machined laminate). The undamaged CFRP laminate had a measured tensile strength of 1004 MPa. When comparing the mechanical values of the repaired laminates (idealized and water jet prepared) it becomes apparent that the water jet method allows an average of 86% remaining tensile strength compared to the idealized laminate. Through this, it can be shown that the water jet machining offers a lot of potential for the surface preparation of CFRP laminates. Compared to an undamaged laminate, the remaining tensile strength is significantly lower.

Certainly, this point can also be applied to the repaired laminates using the idealized method and is therefore not machining-dependent. It should also be considered

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that the repaired areas have not been supplied by external repair plies as has been done by other authors [2]. Since the geometrical properties of the repaired laminate should correspond with the original values, a solution without these external plies is preferable.

The selective layer removal with the pure water jet allows the machining of many different geometries and designs for the transition zone. This offers a lot of research potential for the future.

5 ACKNOWLEDGEMENT The authors are members of the Waterjet Laboratory Hannover (WLH) and of the German Working Group of Waterjet Technology (AWT).

The results presented here were acquired within a research project financed by the German Research Foundation (DFG). The authors would like to thank the DFG for their support.

6 REFERENCES [1] Ehrenstein, G.W. (2006). Faserverbund-Kunststoffe.

Werkstoffe – Verarbeitung – Eigenschaften. 2. Auflage.

Carl Hanser Verlag, München, Wien.

[2] Ahn, S.-H., Springer, G.-S. (1998). Repair of Composite Laminates. Journal of Composite Materials, vol. 32, p. 1035-1114.

[3] Ahn, S.H., Springer, G.S., Shyprykevic, P. (1996).

Composite Repair with Wet Lay-Up and Prepreg.

Composites ’96 Manufacturing and Tooling Conference, Society of Manufacturing Engineers, p.

211-226.

[4] Baker, A. (2006). Development of a Hard-Patch Approach for Scarf Repair of Composite Structure.

Defence Science and Technology Organisation, Canberra.

[5] Wang, C.-H., Gunnion, A.-J. (2008). On the design methodology of scarf repairs to composite laminates.

Composites Science and Technology, vol. 68, p. 35-46, DOI:10.1016/j.compscitech.2007.05.045.

[6] Öjmertz, C. (1997). A Study on Abrasive Waterjet Milling. Doctoral Thesis. Chalmers University of Technology, Gothenburg.

[7] Borkowski, P. (2010). Novel Technique for Spatial Objects Shaping With High-Pressure Abrasive Water Jet. Strojniški vestnik - Journal of Mechanical Engineering, vol. 56, no. 5, p. 287-294.

[8] Momber, A.W., Schulz, R.-R. (2006). Handbuch der Oberflächenbearbeitung Beton. Birkhäuser Verlag, Basel.

[9] Ligocki, A. (2005). Schneiden landwirtschaftlicher Güter mit dem Hochdruckwasserstrahl. Shaker Verlag, Aachen.

[10] Tönshoff, K., Kroos, F., Marzenell, C. (1997). Highpressure Water Peening – a New Mechanical SurfaceStrengthening Process. CIRP Annals – Manufacturing Technology, vol. 46, no. 1, p. 113-116.

[11] Momber, A.W., Kovacevic, R. (1998). Principles of Abrasive Water Jet Machining. Springer-Verlag Limited, London, DOI:10.1007/978-1-4471-1572-4.

Repair Preparation of Fiber-Reinforced Plastics by the Machining of a Stepped Peripheral Zone

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Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 578-586 DOI:10.5545/sv-jme.2012.538

Paper received: 2012-04-20, paper accepted: 2012-08-21 © 2012 Journal of Mechanical Engineering. All rights reserved.

An Improved Torque Method for Preload Control in Precision Assembly of Miniature Bolt Joints Zhang, X. – Wang, X. – Luo, Y. Xiwen Zhang – Xiaodong Wang* – Yi Luo

Key Laboratory for Micro/Nano Technology and System of Liaoning Province, Dalian University of Technology, China In this work, the improved torque method based on a mathematical model is proposed for preload control in the precision assembly of miniature bolt joints. The mathematical model was used for predicting the friction-compensated control torque. The predicted control torque is a function of preload and torque gradient which is determined by the friction of contact surfaces. Thus, the effect of friction on preload error was reduced, and the bolts preload scatter due to frictional variables of the bolts was decreased. Experiments were carried out to verify the proposed method using the bolts with a thread of M1.6, and the results show that the scatters of preload were less than ±13.0% in the elastic range, which is more consistent than the conventional torque method. The preload control error for most bolts was less than 10%.

Keyword: preload control, precision assembly, miniature bolt joints, torque gradient, friction-compensation

0 INTRODUCTION Threaded connection is a popular joining method in modern industries making up nearly 70% of all mechanical connections in industries worldwide [1].

The preload in threaded connections is an important factor affecting the joint performance and reliability.

Improper preload can degrade the behavior and life span of the joint and lead to joint problems, such as material failure, fastener loosening, joint separation, leakage, rattle, and fatigue failure [2] to [4]. It has been reported that up to 90% of bolted joint failures are caused by incorrect initial bolt tightening [5].

In the assembly of miniature precision equipment, performance degradation can be attributed to nonuniform stress, which is caused by the scatter of bolts preload. Thus, it is important to achieve an accurate preload during the assembly process for a specific bolt joint [6].

Several methods have been developed to control the preload of a bolt joint, for example, torque control tightening method, angle control tightening method, tension indicating method, ultrasonic method etc.

Among these methods, the most common way of tightening a bolt or a nut is the torque control method where the preload is predicted according to the torque-preload relationship. However, the torque-preload relationship is highly sensitive to the friction properties which may cause large scatter to about ±50% of the preload [7]. Even the most precise torque measurement can rarely predict the resulting bolt preload in a reliable way. The method of angle control tightening is supposed to be applicable to both sophisticated assembly tools and simple torque wrenches, the torque required to reach the yield point of bolt is very uncertain due to variations in geometry, bolt strength, and friction. However, 578

applying a specific angle after an initial torque leads to more consistent preload levels. Tightening over the yield point results in the preload less affected by friction than in the case of elastic tightening. The yield characteristic of bolt determines the preload and its scatter, which is often less than ±10%. Nevertheless, since the rotation angle in the elastic region is small, there is a risk of over-tightening the miniature bolts [8]. The tension indicating method includes the use of tension indicating devices to measure the preload indirectly, such as the load indicating washers. An obvious disadvantage of this method is poor precision, which is due to the bolt preload being estimated by measuring the gap between the bolt head and the washer. The ultrasonic method is used in industries where the bolted joints are critical. The method can successfully measure the bolts with large diameters under favorable conditions. However, when the bolts are smaller, environmental disturbances or even the operator induced error may exceed the instrument resolution [9] to [11].

Recently, some new tightening control methods have been developed to increase the preload accuracy. New approaches using shape memory alloy (SMA) [12] to [13] or electronic speckle pattern interferometry (ESPI) technique [14] have been applied to the monitoring of the preload. However, the former is costly and the latter demands very stringent environmental stability and hence cannot be employed in a manufacturing environment [15]. Optical digital imaging correlation (DIC) method [16] was proposed to measure compressive strain in a washer, and the bolt preload is subsequently determined from the measured strain. The strain measurement is limited to only one angular location on the washer, thus it is not suitable for small washers in miniature bolt joints. The digital speckle pattern interferometry

*Corr. Author’s Address: School of Mechanical Engineering, Dalian University of Technology, Dalian, China, xdwang@dlut.edu.cn

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(DSPI) system [17] was developed to measure and monitor the out-of-plane deformation on the surface of clamped joints in real-time. In the study, M1.5 miniature bolt joint applications were investigated, yet the control accuracy of bolt preload was not evaluated since load cells and strain-gauged bolts were not suitable for miniature bolts. As a final summary, most of the existing tightening methods are limited in the precise control of tightening miniature bolt joints. It is particularly challenging to find an effective tightening control method for miniature bolt joints due to their small size and low preload level.

In this paper, the improved torque method based on the mathematical model for preload precise control is proposed, and the effects of the friction under the head and in the threads were compensated in the model. The proposed method does not require any special equipment and is not restricted by the space, thus it is suitable for precision assembly of a miniature bolt joint. An introduction to the application of the improved method in M1.6 bolt tightening experiments is presented in detail. The effectiveness of the proposed model is demonstrated by the experiments.

1 PRINCIPLE FOR THE IMPROVED TORQUE METHOD 1.1 Theoretical Model One of the major problems with the use of bolt joints is the accuracy of preload. With the conventional torque control method, the control torque for the target preload is determined by the torque-preload relationship. The tightening torque has a significant dependence on the friction at contact surfaces and the majority of the torque is used to overcome friction (usually between 85 and 95% of the applied torque), so small variation in the frictional conditions can lead to large changes in the bolt preload [18]. However, if the effects of friction for a certain bolt are compensated, the preload scatter for these bolts will be decreased, and the tolerance of the bolt for critical application can be lowered.

When tightening bolt joints, as the nut rotation, part of the torque and preload is applied to close the gaps, and this situation is referred to as ‘snug torque’.

Then, the torque Tf and the preload Ff both increase in proportion to the nut rotation angle θ, the relationship between Tf /Ff and θ are linear and the gradient of Tf / Ff is almost constant. At last, the preload Ff stretching the bolt over its yield point until the bolt fails and the relationship between Tf /Ff and θ are nonlinear. In the elastic range, the equation relating preload Ff and nut rotation angle θ is expressed as follows [19]:

Pθ Cs

Ff =

2π

=

Pω tCs

, (1)

2π

where P is the thread pitch, ω and t are the rotation speed of nut and rotation time of nut, respectively. Cs is the system spring constant or system stiffness of bolt joint in the elastic range.

In the tightening procedure of the bolt joints, the relationship between the applied torque Tf on nut (or bolt head) and bolt preload Ff can be expressed as [20]: d  d T f = Ff d  2 tan ( ρ + β ) + n µ n  . (2)

2d   2d In Eq. (2), d2 and d are basic pitch diameter of the thread and nominal thread diameter, ρ and β are friction angles relating to friction coefficients and lead angle of threads, dn and μn represent the equivalent friction diameter and coefficient of friction at the bearing surface of nut, respectively. Usually, Eq. (2) is denoted by a simple form using torque coefficient K. T f = Ff dK . (3)

Taking the derivative of Tf and Ff with respect to nut rotation time t, the gradient of Tf and Ff can be expressed as:

dT f dFf dK dFf = = dK = K t . (4) dt dt dt dFf dt

=

d

Pω tCs 2π

dt

=

dC  Pω  CS + S  . (5) dt  2π 

Combining Eqs. (4) and (5), the torque coefficient K is deduced from the torque gradient Kt using the following relation:

K=

2π K t dC  Pd ω  CS + S dt 

  

. (6)

Substituting for torque coefficient K from Eq (6) into Eq (3) yields:

T f = Ff dK =

2π K t Ff dC   Pω  CS + S  dt  

. (7)

The system stiffness Cs in the elastic range is dC

S constant, which is mean = 0, thus the Eq. (7) can dt . be expressed as:

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T f = Ff dK =

2π K t Ff . (8) PωCS

From Eq. (8), the torque gradient Kt, system stiffness Cs and rotation speed ω are incorporated into the improved model for predicting the tightening torque Tf at target preload Ff. As shown in Eq. (6), the torque gradient Kt is linear related to torque coefficient K which is strongly affected by surface roughness and lubrication, so the control torque can be predicted by the model according to friction in the assembly processes of different bolts.

account in this paper. In the elastic range, the bolt joint is modeled as a set of linear springs together with a mass-less box, as shown in Fig. 1. The mass-less box has the function of providing displacement that is required in the actual system. Since the bolt joint model shows a linear behavior, the system stiffness Cs can be calculated using the following relations:

1.2 Model Parameters There are four specific parameters P, ω, Kt and Cs appearing in Eq. (8) related to the model in the tightening process. Among these parameters, the parameters Kt and Cs are determined by the bolt joint properties and friction coefficient, respectively. Hence it is critical to estimate the parameters Kt and Cs during bolt joint assembly.

1.2.1 Torque Gradient Kt In the linear elastic range, the torque Tf is linear related to the rotation time t, the torque gradient Kt is theoretically a constant. However, the torque fluctuation caused by the micro burrs or defects is inevitable, thus it is difficult to obtain the constant of Kt through directly taking the derivative of torque Tf with respect to nut rotation time t, as a result, it is difficult to determine if the bolt reaches the linear elastic range. In order to solve this problem, the smoothing filter and least squares linear fit method were employed. The torque Tf(i) at a specific time t(i) is obtained from the sensor data using the smoothing filter. Then the torque gradient Kt(i) at the time t(i) can be obtained by linearly fitting the torque data using linear least square fitting method. If the torque gradient between Kt(i-n) and Kt(i) remains stable around a constant, it is indicated that the bolt reach the linear elastic range, the average torque gradient Kt′ between t(i-n) and t(i) is considered as the torque gradient Kt during the linear increasing stage of torque.

Fig. 1. Stiffness model of bolt joint

Cs =

1 + K b . (9) ∑1 K c (i )

Kc(i) and Kb are stiffness of the clamped parts and bolt, respectively. Kc(i) and Kb are considered to be constant. The stiffness of clamped parts Kc(i) is unknown, thus experiments for measuring the stiffness are needed.

Bickford [21] provides the expression of the bolt stiffness for various bolt types and materials as described in Eq. (10):

L L 1 = be + se , K b EAB EAs

where, Lbe = Lb + 0.5 × Th, Lse = Lg – Lb + 0.5 × Tn, and As = 0.7854 × (D – 0.938 P)2.

As shown in Fig. 2. Lbe is the effective length of the bolts, and Lse is the effective length of the threads.

Th is the height of the bolt head and Tn is the height of the nut. Lb is the length of the unthreaded part of the bolt and Lg is the total gap between the bolt head and the nut. AB is the nominal cross sectional area, equal to the cross-sectional area based on the nominal diameter D. As is the effective stress area of the threads. E is the Young’s Modulus.

1.2.2 System Stiffness Cs The torsional stiffness is independent of axial force Ff, while the stiffness in the axial direction has a dominant effect on the relationship between Ff and Tf, therefore only the stiffness in the axial direction is taken into 580

(10)

Zhang, X. – Wang, X. – Luo, Y.

Fig. 2. Model of bolt joint

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2 EXPERIMENTS 2.1 Experimental Setup The schematic diagram of the experimental system is shown in Fig. 3a. The system is used for automatic assembly of the miniature bolts with the diameter less than 2.5 mm, which is widely employed in the precision miniature equipments. A torque sensor (Lorenz Messtechnik 0180, Germany) and a force sensor (CAAA BK-3A, China) were employed for an evaluation of the torque and preload during the bolt tightening. The resolution of the torque sensor is 0.01 Nmm with the accuracy of ±0.1% in the max range (range: ±200 Nmm), and the range of the force sensor is 200 N, with the resolution of 0.2 N and a maximum gain error of 0.1%. The torque loading is carried out with a step motor (SYNTRON 42BYG, China), which has the capacity of 230 Nmm. The rotary motion of the step motor is transmitted to a wrench through the torque sensor and flexible joints. The nut is tightened by the wrench, and as the wrench rotates, the linear motion stage drives the wrench to synchronous linear feed along the bolt, so there is no relative motion between the nut and wrench, and the effect of friction between the nut and the wrench is eliminated. The bolt head was fixed on the force sensor and sensor support, the rotation between bolt head and force sensor was restricted. Two washers were placed under the bolt head and the nut, respectively. The experimental setup

a)

used to investigate the improved torque method is shown in Fig. 3b.

In this study, a program has been developed in LabVIEW to calculate the torque gradient Kt in real time. When the constant torque gradient is determined, it is indicated that the bolt reached its linear elastic range, and then the friction-compensated control torque for target preload is predicted.

Bolt tightening experiments were performed using M1.6 bolts as the example. The property class of the bolts is 4.8, the geometric and material properties for the bolts are provided in Table 1. In the experiments, twenty-four bolts were divided into three groups for the target preload of 90, 120 and 150 N, respectively. The target preloads were selected to make sure that the bolts were in the elastic range, in which the relationship between the preload and nut rotation angle is linear. The maximum preload level is about 43% of the proof load of the bolts. In the tightening procedure, the bolts were tightened at fixed nut rotation speed of 5 rpm, which was selected based on preliminary tests. The total time to assembly a miniature bolt joint under the rotation speed was less than 20 seconds. The lower speed in tightening leads to low efficiency, but with a higher speed in tightening, the effect of hysteresis is not negligible, and the step motor is not able to stop at the prescribed target torque value. The experiments using the torque method were also carried out, providing a comparison with the proposed improved torque method. Another

b) Fig. 3. a) Schematic diagram of the experimental system, b) picture of the experimental setup An Improved Torque Method for Preload Control in Precision Assembly of Miniature Bolt Joints

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twenty-four bolts were divided into three groups for the experiments at target torque of 25, 35 and 45 Nmm, respectively.

Table 1. Geometrical and Material Properties of a M1.6 class 4.8 bolt Geometrical parameters Major Diameter [mm] 1.58 Minor diameter [mm] 1.20 Pitch [mm] 0.35 Length [mm] 8.00

Material properties Material low carbon steel Proof Load [MPa] 310 Yield Strength [MPa] 340 Tensile strength [MPa] 420

Kt(i) changed with the time t for miniature bolts were plotted in Fig. 5c.The target torque values were both 45 Nmm and the nut rotation speeds were 5 rpm. For bolt 1 and bolt 2, the preloads were 106.6 and 169.2 N, and the torque gradients were 9.43 and 6.04 Nmm/s, respectively. With the improved torque method, if the target preloads are both 150 N, the predicted torques calculated according to Eq. (8) should be 60.8 and 38.9 Nmm, respectively.

2.2 Model Parameters in Experiment 2.2.1 Stiffness of Bolt Joint In this paper, the stiffness of the bolt is calculated using Eq. (10), and the composited stiffness of the clamped parts 1/∑(1/Kc(i)) is determined experimentally by compressing the clamped parts and measuring the corresponding deformation. The clamped parts include washers, preload sensor and the support. Fig. 4 shows the calibration curve of deformation under a different load during the loading of clamped parts. The deformation is linearly related to force with a linearly dependent coefficient of 1/∑(1/Kc(i)) = 1.088 N/μm.

a) Torque versus time

b) Preload versus time

Fig. 4. Deformation versus force of clamped parts

2.2.2 Torque Gradient The torque gradient Kt is determined by the friction on contact faces, and is not calculated from the tightening torque in real time. As an example, Fig. 5 shows the curves obtained from the two bolts tightening experiments using torque method. The curves of torque changed with the time t are shown in the Fig. 5a and the curves of preload changed with the time t as is shown in the Fig. 5b. The trend of torque gradients 582

c) Torque gradient versus time Fig. 5. Experimental curves for tightening tests

As shown in Fig. 5, three stages can be defined. In the second stage, the Tf (t) and Ff (t) curves were linear, and the torque gradients were almost constant values.

Zhang, X. – Wang, X. – Luo, Y.

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As shown in Fig. 5a, the slopes of the Tf (t) curves were significantly different at the same target torque, and the corresponding Ff (t) curves have almost the same slopes and significantly different preload values.

These phenomena can be explained by Eqs. (6) and (7), the friction differences led to different preloads and a different slope of the Tf (t) curves. Moreover, almost the same slope of the Ff (t) linear curves means that the system stiffness of two bolt joints were both constant values and the stiffness difference was negligible.

A comparison of Fig. 5c with 5a, shows that torque gradients Kt(i) have a hysteresis less than 0.5 s to torques Tf, which is mainly caused by the time consumed in calculating and smoothing. The results indicate that as soon as the torque Tf was linearly related to time t, it is reliable for the proposed method to evaluate the torque gradient Kt.

a) Torque gradient for various target preload using improved torque method

3 RESULTS AND DISCUSSION 3.1 Preload Scatters of Improved Torque Method and Torque Method Fig. 6a shows the torque gradient results of every bolt during the tightening process. It is observed that the torque gradients at the target torque values of 90, 120 and 150 N range from 5.45 to 8.06 Nmm/s.

Fig. 6b and c show the predicted torque results and corresponding bolt preload results, respectively. The experimental results show that the scatters of torque gradients and predicted torques have a similar trend.

The predicted torques in the tightening process were 27.5±5.0, 37.1±6.9 and 44.4±8.9 Nmm, with an estimation scatters around the mean value of ±18.2, ±18.6 and ±20.0% for target preloads of 90, 120 and 150 N, respectively. As a result, the bolt preloads reached at the predicted torques value were 92.4±12.0, 120.4±12.1 and 149.8±12.3 N, with the scatters around the mean value of ±12.9, ±10.0 and ±8.9%, respectively. It is obvious that the scatters of predicted torques were larger than actual bolt preloads.

Furthermore, as the target preloads increase, the distributions of actual preloads increase slightly, while the scatter around the mean value of actual preload decrease significantly. The results demonstrate that the proposed improved torque method for preload precise control is reliable. For precision assembly of miniature bolt joints, it is feasible to predict the tightening torque in the elastic range using the proposed model. Furthermore, it validates that the higher the target preloads, the smaller the scatters around the mean value of bolts preload would be.

b) Predicted torque for various target preload using improved torque method

c) Actual preload for various target preload using improved torque method

d) Actual preload for various target torque using torque method Fig. 6. Results of tightening experiments

Fig. 6d represents the bolt preload results of the torque method, the bolt preloads reached at the target torque values were 84.6±20.1, 113.7±24.9 and 145.3±30.1 N, with scatters around the mean value of ±24.0, ±21.9 and ±20.7% for the corresponding

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target torques were 25, 35 and 45 Nmm, respectively.

It is clear that as the target torques increase, the distribution of the preloads increases significantly, while the scatters around the mean value of actual preload decrease.

3.2 The Improvements of the Improved Torque Method Table 2 shows the results of the comparison between the improved torque method and torque method. In contrast with the torque method, the improved torque method yields both smaller distribution of the preloads and smaller scatter around the mean value of bolts preload. The scatters around the mean value of bolts preload at the same level were decreased by at least 40.2% for which the reason may be that the effects of friction were compensated in the proposed improved mathematical model. Moreover, it is noticeable that in Fig. 7, the actual preload errors of most bolts were less than ±10.0%, and the maximum preload error in the experiments was 18.0%, with the actual preload and the target preload was 106.2 and 90 N, respectively.

As the target preloads increase, the percentage errors of actual preload decrease significantly.

Table 2. Comparison between improved torque method and torque method Preload control method Improved torque method Torque method

Target preload [N] 90 120 150 ∕ ∕ ∕

Tighten torque

Scatter of tighten torque [Nmm] [%] 27.5±5.0 ±18.2 37.1±6.9 ±18.6 44.4±8.9 ±20.0 25 ∕ 35 ∕ 45 ∕

Bolt preload

Scatter of preload

[N] 92.4±12.0 120.4±12.1 149.8±12.3 84.6±20.1 113.7±24.9 145.3±30.1

[%] ±12.9 ±10 ±8.9 ±24.0 ±21.9 ±20.7

Fig. 7. Percentage error of actual preload versus target preload

The preload errors observed in the experiment results can be explained by the effect of bolts stiffness error between the actual and theoretical value which is calculated by Eq. (10). It has been shown that 584

both friction coefficients and system stiffness play a significant role in determining the torque-preload relationship in accordance with Eq. (8). The effects of the friction were compensated in the experiments, yet the bolt dimensional tolerances which caused stiffness error of bolts were neglected. Another reason for the preload error was the accuracy of the torque gradient during the tightening process. The prediction accuracy was also affected by the calculation error, the micro burrs and defects of bolt threads.

The experimental results indicated that the proposed improved torque method is highly practical to realize a real-time control of preload; the method will result in a more controlled tightening of a miniature bolt joint.

3.3 Application of the Improved Torque Method M1.4, M1.6 and M2 bolts tightening experiments were performed to investigate the effect of bolt diameter on the preload scatter of the improved torque method. The results have demonstrated that, for the same experimental setup and the same preload level, as the bolts diameter increases from 1.4 to 2 mm, with the improved torque method, the scatters of bolts preload for the target preloads were ±8.6, ±10.0 and ±11.2%, respectively, and with the torque method, the scatters of preloads around the mean value were ±25.0, ±21.9 and ±22.7%, respectively. It is obvious that the diameter of the bolts influences the scatter of the improved torque method at the same preload level. It may induce larger preload scatters when the relatively lower tightened torques were applied on the bolt in larger diameters.

For miniature bolt applications of no more than M2, the improved torque method has an advantage over the torque method for preload control. However, if the bolt is larger than M2, other methods such as the angle control method, tension indicating method and even the ultrasonic method can also be used for better preload control in larger bolts application.

The bolts assembly setups have been implemented in an automatic precise assembly system for miniature products, as shown in Fig. 8, a specially designed gripper was used for picking up the nuts, and guided by a vision measurement unit to the position of the bolts. The assembly system with the proposed method accomplished the automatic assembly of the M1.4 bolts in a miniature products fabrication process [22].

Zhang, X. – Wang, X. – Luo, Y.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 578-586

Fig. 8. Automatic precise assembly system for miniature products

4 CONCLUSIONS An improved model based torque method was proposed for the real-time precise control of the preload of the miniature bolt joints. The torque gradient determined by friction effects was taken into account in the model. Moreover, the system stiffness and rotation speed were also incorporated into the model. Finally, miniature bolts tightening experiments which were conducted have demonstrated the validity of the proposed method, and the results indicated that the scatters of bolts preload for the target preloads of 90, 120 and 150 N were ±12.9, ±10.0 and ±8.9%, respectively. The magnitudes of the actual preload errors for most bolts were no less than 10%. However, if more accurate system stiffness can be obtained, better preload accuracy will be achieved. The method offers the benefit of convenience and does not require any special equipment, therefore, the proposed method is suitable for control the preload in real-time.

It is a promising method for the preload control in the precision assembly of miniature bolt joints.

5 ACKNOWLEDGEMENTS This research was supported by “the National Natural Science Foundation of China (No.51075058)” and by “the Fundamental Research Funds for the Central Universities (DUT10ZDG04)”.

6 REFERENCES [1] Roman, C. (2008). Optimal control of screwing speed in assembly with thread-forming screws.

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[2] Croccolo, D., Agostinis, M.D., Vincenzi, N. (2011). Failure analysis of bolted joints: Effect of friction coefficients in torque–preloading relationship. Engineering Failure Analysis, vol. 18, no. 1, p. 364-373, DOI:10.1016/j. engfailanal.2010.09.015.

[3] Zaletelj, H., Fajdiga, G., Nagode, M. (2011).

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[4] Stamenkovic, D., Maksimovic, K., NikolicStanojevic, V., Maksimovic, S., Stupar, S., Vasovic, I. (2010). Fatigue life estimation of notched structural components. Strojniški vestnik - Journal of Mechanical Engineering, vol. 56, no.

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[5] Arghavani, J., Derenne, M., Marchand, L. (2001).

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[6] Sawa, T., Omiya, Y., Takagi, Y., Torii, H. (2009).

Effects of scatter in axial bolt force on the sealing performance of pipe flange connections at elevated temperature. Proceedings of the ASME Pressure Vessels and Piping Conference, vol. 2, p.

139-147.

[7] Göran, R.T. (2003). Controlled tightening over the yield point of a screw: based on Taylor’s series expansions. Journal of Pressure Vessel Technology - Transactions of the ASME, vol. 125, no. 4, p. 460-466.

[8] Fukuoka, T., Takaki, T. (2004). Evaluations of the tightening process of bolted joint with elastic angle control method. Analysis of Bolted Joints. ASME/JSME Pressure Vessels and Piping Conference, vol. 478, p. 11-18.

[9] Jesse, R.M. (2010). Joint integrity monitoring using permanent ultrasonic bolt load transducers.

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[10] Jhang, K.Y., Quan, H.H., Ha, J., Kim, N. (2006). Estimation of clamping force in hightension bolts through ultrasonic velocity measurement. Ultrasonics, vol. 44, p. 1339-1342, DOI:10.1016/j.ultras.2006.05.190.

[11] Chaki, S., Corneloup, G., Lillamand, I., Walaszek, H. (2007). Combination of longitudinal and

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transverse ultrasonic waves for in situ control of the tightening of bolts. Journal of Pressure Vessel Technology - Transactions of the ASME, vol. 129, no. 3, p. 383-390.

[12] Hesse, T., Ghorashi, M., Inman, D.J. (2004).

Shape memory alloy in tension and compression and its application as clamping-force actuator in a bolted joint: Part 1 – Experimentation.

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[13] Ghorashi, M., Inman, D.J. (2004). Shape memory alloy in tension and compression and its application as clamping force actuator in a bolted joint: Part 2 – Modeling. Journal of Intelligent Material Systems and Structures, vol. 15, no. 8, p.

589-600, DOI:10.1177/1045389X04042791.

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61211-61215.

[18] Zou, Q., Sun, T.S., Nassar, S.A., Barber, G.C., Gumul, A.K. (2007). Effect of lubrication on friction and torque-tension relationship in threaded fasteners. Journal of Tribology Transactions of the ASME, vol. 50, no. 1, p.127136.

[19] Nassar, S.A., Yang, X.J. (2008). Torque-angle formulation of threaded fastener tightening.

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245011-245014, DOI:10.1115/1.2821388.

[20] Nassar, S.A., Matin, P.H., Barber, G.C. (2005).

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[21] Bickford, J.H. (1995). An Introduction to the Design and Behavior of Bolted Joints. 3rd ed.

Marcel Dekker, New York.

[22] Zhang, X.W., Wang, X.D., Luo, Y., Teng, L., Chen, L., Ma, T.M. (2011). A precise automatic assembly system for fabrication of complex miniature products.

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Zhang, X. – Wang, X. – Luo, Y.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 587-597 DOI:10.5545/sv-jme.2012.352

Paper received: 2012-02-13, paper accepted: 2012-06-29 © 2012 Journal of Mechanical Engineering. All rights reserved.

Determination of Surface Qualities on Inclined Surface Machining with Acoustic Sound Pressure Gok, A. – Gologlu, C. – Demirci, I.H. – Kurt, M. Arif Gok1 – Cevdet Gologlu2,* – Ibrahim H. Demirci2 – Mustafa Kurt3 1 Kastamonu

University, Vocational High School, Turkey University, Faculty of Engineering, Turkey 3 Marmara University, Technical Education Faculty, Turkey 2 Karabuk

Die parts used in automotive and aviation industry have complicated surfaces that require multiaxis machining. In machining of inclined surfaces with ball-end milling, the process is of great significance for its correctness and accuracy. In this study, Acoustic Sound Pressure (ASP) generated during the machining of a workpiece at a vertical machining centre has been measured. The experiments have been conducted in association with cutting velocity, feed rate and step over parameters determined by using different surface forms and different cutter path strategies. Therefore, the aim of this study is to understand the relationships between the generated sound signals and surface roughness in the machining of inclined concave and convex surfaces.

In the experiments, the workpiece material of EN X40CrMoV5-1 hot work tool steel, which is commonly used in the related die industry, has been chosen. The ball end mills with two indexable inserts with three different coatings of TiC, TiN, and TiAlN have been used. The results show that there is a rise at the value of surface roughness with a rise at the value of acoustic sound pressure and that surface roughness could be figured out with acoustic sound pressure level.

Key words: ball end milling, acoustic emission and sound pressure, linear regression, surface roughness

0 INTRODUCTION Machining is one of the most important methods in manufacturing [1]. To change the shape of a workpiece to a desired geometrical shape, it is necessary to use an appropriate machine tool and cutting tools to obtain the required dimensions and surface quality [2]. One of the most important factors to increase the quality in the manufacturing processes is to control surface quality of the product. Surface quality control is a costly process and a difficult task for the parts that are in the production line. Time devoted to quality control process and cost can be minimized by the help of prediction models and systems. Therefore, real-time model-based quality control is used by monitoring measurable processes [3]. The reason behind the monitoring of machining operations is to generally prevent undesired machining consequences such as chip formation and chip shape classification, tool wear, dimensional tolerances, surface texture (roughness and waviness), and tool deflection [4].

Researchers have been working on online monitoring with video based approaches, so that they can screen tool working conditions [5]. However, they are both difficult and costly, so that the implementation of these systems into industry is almost impossible.

Nevertheless, when the costs of these systems are reduced, it is likely to use them in many measurements such as tool wear, cutting force, Acoustic Emission (AE) and vibrations [6]. Ghosh et al. [7] focused on the prediction of tool wear in CNC milling using sensors in integration with neural networks. In their study, they found that the average flank wear of the

main cutting edge was predicted by the signals such as cutting force, cutting tool vibration, and sound pressure level obtained from the machining region.

Marinescu and Axinte [8] worked on the analysis of emission signals efficiency to determine damages on tool and workpiece in milling operations. At the end of the study, the damage on the tool has been determined with emission signals. Marinescu and Axinte [9] also focused on monitoring of time-acoustic emission frequency to describe surface defects on a workpiece in milling with more than one cutting edge. This was carried out with new methods for supervising cutting processes with multiple teeth cutting simultaneously, i.e. milling, by using of AE signals backed-up by force data. By means of this work, the researchers took signals into consideration for all simultaneous cutting edges. The results showed for the first time that identification of milling conditions (i.e. cutting with one tooth and two-three teeth) is possible using only AE signal in time–frequency domain. Additionally, surface deformations, related to the wearing of cutting edges, can be determined. Rivero et al. [10] worked on evaluation of the suitability of a tool wear monitoring system based on machine tool internal signals. The sensor data from internal signals were compared and analyzed, assessing the deviation in representative variables in time and frequency domains. As a result, tool wear has been estimated. Parallel to these studies, Wilcox et al. [11] worked on the use of cutting force and AE signals for the monitoring of tool insert geometry during rough face milling. In their studies, they simulated different forms of naturally occurring wear such as crater, notch and flank wear, local changes

*Corr. Author’s Address: Department of Mechanical Engineering, Faculty of Engineering, Karabuk University, 78050, Karabuk, Turkey, cgologlu@karabuk.edu.tr

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in rake angle and edge breakdown. Weingaertner et al.

[12] evaluated the influence of high speed end milling dynamic stability through audio signal measurements both experimentally and analytically. In the study, the stability evaluation was based on the workpiece surface finish and on the audio signals measured with a unidirectional microphone. The experimental and analytical results have been found very close to one another. Tekiner and Yeşilyurt [13] studied the cutting parameters depending on process sound during the turning machining process of AISI 304 austenitic stainless steel. In their study, the best cutting speed and feed rate values were determined according to the flank wear, built up edge, chip form, surface roughness of the machined samples and machine tool power consumption. In addition to the above mentioned studies, Salgado and Alonso [14] focused on Tool Condition Monitoring System (TCMS) for on-line tool wear monitoring in turning. In the study, the monitoring signals were feed motor current and sound signal. The tool wear has been found by TCMS. Ravindra et al. [15] worked on acoustic emissions for tool condition monitoring in metal cutting. Moreover, Haber et al. [16] studied tool-wear monitoring in a high-speed machining process on the basis of the analysis of different signals’ signatures in time and frequency domains. In their study, time and frequency domain analyses were confirmed wwith the relevance of cutting-force and vibration signals’ signatures for tool-wear monitoring in High-Speed Machining processes. Tool wear has been assessed as a result of their analysis. Quadro and Branco [17] carried out analysis of acoustic emission during drilling test.

They used AISI D3, drills of high speed steel with TiN coating. In their measurements, profilometry and light microscopy were used to characterize and quantify the wear on the drills’ cutting edges along the tests.

In another study, Guo and Ammula [18] developed a real-time acoustic emission monitoring system to investigate the sensitivity of broad AE signal parameters including RMS, frequency, amplitude, and count rate to white layer and corresponding surface finish and tool wear. In this way, tool wear has been observed as real time. Furthermore, Asilturk et al.

[19] conducted a study on modeling with regression of surface roughness dependent on cutting parameters, vibration and acoustic emission. In their model, the first degree, the second degree and logarithmic multiple regressions were used. In this way, it was found that the feed rate on the surface roughness parameter was the most effective and the better results were also attained at the second order regression model. Horvat et al. [20] studied on evaluation and 588

monitoring gas metal arc welding process by using an audible sound signal. In this way, the welding process has been assessed in terms of robustness and quality. In addition to this, a new algorithm based on the measured welding current has been established for the calculation of emitted sound during the welding process. The results of experimental and theoretical measurements were found to be in good agreement.

Kek and Grum [21] analysed acoustic emission (AE) signals obtained during laser cutting of a steel plate.

The acoustic emission signals in the plate have been measured with contact PZT sensors. During the laser cutting process, continuous AE signals have been captured by the action of a cutting gas. AE signals as a result of their research have demonstrated that an important indicator for quality of laser cut. The number of studies related to 3D machining [22] and tool life [23] were also studied by the researchers.

The previous studies focused on acoustic emission, tool wear and deformation, vibrations stability and best cutting parameters. In this study, the experiments were carried out in order to understand the relationships between the sound signals generated and surface roughness in the machining of inclined concave and convex surfaces.

1 EXPERIMENTAL STUDY 1.1 Cutter Path Styles It has been experimentally shown that the right choice on tool paths including different cutting movements affects production time, status of machining surfaces and cost [24]. Therefore, in these experimental studies, contouring and ramping tool paths styles are used to produce inclined surfaces. The tool at the machining of free form and inclined surfaces makes movements of ramping and contouring. Accordingly, for the implementation of up milling and down milling strategies at the machining of inclined surface, contouring and ramping are the inevitable choice of tool path styles. In contour operation, the tool moves parallel to the axis parts. As a consequence, the chip is easily disposed from the cutting zone. At the ramping operations, the effective diameter in ball end mill tools is easily visible and their effects to changes in the responses are easily observed. In the experiments, 40×30 mm islands on the 220×135×50 mm sized block were machined.

In contouring tool path styles, the cutter scans the inclined surface with the lines in parallel to surface radius (Fig. 1a). On the other hand, in ramping tool path styles, the cutter scans the inclined surface with

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the lines in vertical to surface radius (Fig. 1b). In Fig. 1, the feed rate and spindle speed are depicted by Vf and W, respectively.

Fig. 1. a) Contouring and, b) ramping cutter path styles for inclined concave and convex surfaces

In both tool path styles, step over values are constant. After machining each step, the cutter moves one step sideways in a position in which it returns back to the beginning level of that step and then processes the next step. Under these conditions, four tool path styles were generated as shown in Fig. 2. In Fig. 2, the form radius of workpiece, milling position angle, nominal depth of cut and step over are shown

by R, θ, a and fp, respectively.

After the machining operations are completed, the material is subjected to the heat process. As a cutter body, CoroMill (Sandvik Company) with an indexable (R216-16A20-045), Ø16 mm cylindrical shank, two fluted and 30° helix angle end mill were used.

Moreover, the ball end inserts (Sandvik Company) with TiC, TiN and TiAlN coated (R216-16 03 M-M H13A) were used. Every insert has a coating of 3 micron thickness. In addition to tool path styles, three different variable parameters were used. These were; cutting velocity (Vc), feed rate (Vf) and cutting step over (fp). Cutting tool step over affects the tracks on the surface made by the cutter, the load on the cutter and the processing time in a direct manner [24]. Step over value was chosen as 5% of the tool diameter and this value was set as the lower level of fp. Cutting velocity and feed rate values were picked up with reference to the catalogue values of Sandvik Company. These reference values were determined by carrying out a number of experiments for each tool coating and by taking the common use of the material for the industry into account (Table 2).

1.3 Experimental Apparatus and the Implementation of Experiment

Fig. 2. Cutter path styles- contouring: up milling (up step over) 1, down milling (up step over) 2, up milling (down step over) 3, down milling (down step over) 4; ramping: up milling (left step over) 5, down milling (left step over) 6, up milling (right step over) 7, down milling (right step over) 8

1.2 Material, Cutting Tool and Machining Parameters High machinability of the materials used in dies, automotive and space industry is of great importance for surface roughness of the workpiece production.

EN X40CrMoV5-1 (Böhler W302) hot work tool steel, which is commonly used in these industries, was preferred in the study. Chemical compositions of the material presented in Table 1. The material has 22 to 25 HRC hardness with yield strength of 1650 N/ mm2. When it is subjected to heat process under 1020 to 1080 °C for 15 to 30 minutes and cooled in oil, its hardness rises up to 50 to 54 HRC. The pre-hardening process is not implemented in its machining process.

Semi finishing operations were used in the experiments and coolant was not used due to constituting a layer between cutting edge and workpiece, and this layer causes shear in little slices on inclined forms. The experiments were carried out at a vertical machining center, John Ford VMC 550 CNC, having 12000 rev/min with a 12 kW engine power. The stages of the experimental system are depicted in Fig. 3. In pre-machining, tool wear did not occur in the machining of several blocks. Thus, every five blocks were machined by different cutting inserts.

After completing the pre-machining, a semi finishing operation was carried out using different cutter path styles and parameters on the workpiece. For the combinations of cutting parameters, L’16 standard orthogonal array was chosen as four different levels for each parameter appointed. For concave surface, a number of 48 experiments were conducted with TiC, TiN, and TiAlN coatings (16 experiments per each).

Similarly, a number of 48 experiments were made with TiC, TiN, and TiAlN coatings (16 experiments per each) for the convex surface.

Table 1. Chemical composition [wt%] of EN X40CrMoV5-1 %

C 0.39

Si 1.00

Mn 0.40

Cr 5.10

Determination of Surface Qualities on Inclined Surface Machining with Acoustic Sound Pressure

Mo 1.30

P 0.025

V 1.00

S 0.005

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analyzed in a time scale and its arihe metic mean was calculated. Then, signals were transformed into ASP values in dB. The transformations were made with the Eq. (1).

(

)

ASP = 20 log10 Voutput / Vo , (1)

where, Voutput is the arithmetic mean of the collected signals in Volt and Vo is the lowest recorded signals in Volt at the same experiment.

Some significant acoustic sound pressures for the related experiments are given below. In Fig. 4, experiment 3 has the lowest raw sound signal value of –0.5903 mV. On the other hand, the experiment 9 has –1 mV as the lowest raw sound signal value. The arithmetic mean of the collected signals for the related experiments is given in Table 3.

Table 3. Acoustic sound pressure values [mV] of sample experiments Min Max Arithmetic mean

Experiment #3 –0.5903 0.5293 1.477×10–5

Experiment #9 –1 0.9725 6.542×10–6

Fig. 3. The stages of the experimental system Table 2. Assignment of the levels to factors Factors Cutting velocity, Vc [m/min] Feed rate, Vf [mm/rev] Step over, fp [mm]

TiC TiN TiAlN TiC TiN TiAlN

Level 1 70 100 110 223 318 350 0.8

Level 2 80 110 120 255 350 382 1

Level 3 90 120 130 286 382 414 1.5

Level 4 100 130 140 318 414 445 2

1.4 Measurement of Machining Sound Pressure Level Tool defects can be observed by analyzing the sounds generated in machining. In the study, the sound sensor was placed to the closest position to the cutting tool.

The acoustic sound pressures were collected by a sound sensor (microphone) at the sampling of 100 ms in mV units. An algorithm written in MATLAB is used for digitizing and collecting the sound pressures from the sound sensor. For calibration of the sound sensor, the same sound values were measured at the same time with a sound measurement device of CEMDT 8850. The real sound pressure values were determined by evaluating the differences found via the algorithm. The signals that were recorded in mV were 590

Fig. 4. Samples of the measured acoustic sound

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In the experiments, MahrSurf PS1 surface roughness measurement device was used. The measurements were performed at a direction normal to cutting tool paths with an angle of 45° to standing position in order to take the effective cutting diameter into consideration in every sample. Average surface roughness (Ra) is represented as in Eq. (2).

Ra =

an increase with the increase of feed rate and step over. The effects of feed rate and step over values on obtained ASP and Ra values are given in Fig. 6.

L

1 Y ( x ) dx , (2) L ∫0

where, Ra represents deviation from the average line, Y represents the ordinate of profile curve and L represents the measurement length. In the experimental measurements, every measurement was carried out three times and their average values have been taken into account.

a)

2 EXPERIMENTAL RESULTS One of the most important criteria for determining surface quality in cutting is surface roughness. The acoustic emission can be employed for predicting surface roughness of machining surfaces [25]. In other words, the relation between surface roughness values and variation of ASP levels can be established.

The changes in sound pressure levels have been seen available to determine an average surface roughness.

For concave surface forms, the ASP [dB] and surface roughness [µm] values which are acquired from the experiments for the related coatings are shown in Fig. 5. When the acquired ASP and Ra graphics for TiC, TiN and TiAIN coatings were examined, it wasobserved that surface roughness values increase with the increase in acoustic sound pressure values. However, surface roughness values decrease with the decrease of acoustic sound pressure values.

According to Fig. 5, for the cutter with TiC coating, the largest ASP of 105.2 dB causes the largest Ra value of 5.09 µm at the experiment 4. On the other hand, the lowest ASP of 83.1 dB is observed at the Ra value of 1.51 µm at the experiment 1. For the cutter with TiN coating, the largest ASP of 104.3 dB resulted in the largest Ra value of 5.69 µm at the experiment 13. The lowest ASP of 86.1 dB is observed at the Ra value of 1.50 µm at the experiment 1. Lastly, for the cutter with TiAlN coating, the largest ASP of 101.3 dB resulted in the largest Ra value of 5.66 µm at the experiment 4. On the other hand, the lowest ASP of 74.1 dB is observed at the Ra value of 1.68 µm at the experiment 1. When the results are evaluated in terms of parameters (Vf and fp), ASP and Ra values show

b)

c) Fig. 5. Ra versus ASP for; a) TiC, b)TiN, and c) TiAlN coatings in concave surface

According to Fig. 6a, for the TiC coating, the largest values of Ra and ASP have been observed at the feed rate of 318 mm/rev and the step over of 2 mm.

The lowest values of Ra and ASP have been observed at the feed rate of 223 mm/rev and the step over of

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a)

b)

c)

Fig. 6. Ra versus ASP for; a) TiC, b) TiN , and c) TiAlN as to feed rate and step over levels in concave surface

0.8 mm. In Fig. 6b, for the TiN coating, the largest values of Ra and ASP have been observed at the feed rate of 414 mm/rev and the step over of 2 mm. The lowest values of Ra and ASP have been observed at the feed rate of 318 mm/rev and the step over of 0.8 mm. Finally in Fig. 6c, for the TiAlN coating, the largest values of Ra and ASP have been detected at the feed rate of 445 mm/rev and the step over of 2 mm.

The lowest values of Ra and ASP have been observed at the feed rate of 350 mm/rev and the step over of 0.8 mm.

592

In this study, linear regression analysis was carried out to determine the relationship between the sound pressure level and surface roughness average.

Linear regression is represented as [26]:

S = β 0 + β1 x1 + .... + β i xi + β n xn + ε , (3)

where, S represents the response and βi represents a regression factor of ith and residual. ε is a random defect term and is presumed to show normal distribution having average zero with σ2 variance.

The standardized residual is the residual divided by the standard deviation, where the residual is the

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a) a)

b) b)

c)

c) Fig. 8. Ra versus ASP for; a) TiC, b) TiN, and c) TiAlN coatings in convex surface Fig. 7. ASP as a function of Ra for; a) TiC, b) TiN, and c) TiAlN coatings in concave surface

difference between the data response and the fitted response. In other words, it is residual standardized to have standard deviation 1 [26]. According to linear regression analysis, it was found that the relationship between sound pressure level and surface roughness is positive, linear and statistically significant

(respectively R2 =0.875; 0.822 and 0.873) for TiC, TiN, and TiAIN (Fig. 7). When all cutting parameters are taken into consideration, it is determined that the correlation coefficient between sound pressure level and surface roughness is better than R2 = 0.8. This shows that it is beneficial to adopt the sound pressure level during manufacturing.

Similarly, for convex surface forms, the ASP [dB] and surface roughness [µm] values which are

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a)

b)

c)

Fig. 9. Ra versus ASP for; a) TiC, b) TiN, and c) TiAlN as to feed rate and step over levels in convex surface

acquired from the experiments for the related coatings are shown in Fig. 8. When it is examined the acquired ASP and Ra graphics for TiC, TiN, and TiAIN coatings in Fig. 8, it is observed that surface roughness values increase with the increase of acoustic sound pressure as in concave surface type.

According to Fig. 8, for the cutter with TiC coating, the largest ASP of 99.9 dB causes the largest Ra value of 4.95 µm at the experiment 13. On the other hand, the lowest ASP of 85.3 dB is observed at the Ra value of 1.25 µm at the experiment 16. For the cutter with TiN coating, the largest ASP of 113.4 dB resulted in the largest Ra value of 5.01 µm at the experiment 13. The lowest ASP of 84 dB is observed at the Ra value of 1.59 µm at the experiment 16. Lastly, for the 594

cutter with TiAlN coating, the largest ASP of 104.1 dB resulted in the largest Ra value of 5.16 µm at the experiment 4. On the other hand, the lowest ASP of 83.7 dB is observed at the Ra value of 1.51 µm at the experiment 16.

Upon the evaluation of parameters (Vf and fp), ASP and Ra values show an increase with the increase of feed rate and step over as happened in concave surface type. The effects of feed rate and step over values on acquired ASP and Ra values are given in Fig. 9. As it is seen in Fig. 9a, for the TiC coating, the largest values of Ra and ASP have been observed at the feed rate of 318 mm/rev and the step over of 2 mm. The lowest values of Ra and ASP have been seen at the feed rate of 223 mm/rev and the step over

Gok, A. – Gologlu, C. – Demirci, I.H. – Kurt, M.

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coating, the largest values of Ra and ASP have been detected at the feed rate of 445 mm/rev and the step over of 2 mm. The lowest values of Ra and ASP have been observed at the feed rate of 350 mm/rev and the step over of 0.8 mm.

According to linear regression analysis, it has been found that the relationship between the sound pressure level for convex surface type and surface roughness is positive, linear and statistically significant (respectively R2 =0.888; 0.899 and 0.916) for TiC, TiN, and TiAIN (Fig. 10). Depending on all cutting parameters, it was determined that the correlation coefficient between surface roughness and sound pressure level is better than R2 = 0.8. Again, this shows that it is beneficial to adopt the sound pressure level during manufacturing.

a)

3 CONCLUSION

b)

c) Fig. 10. ASP as a function of surface roughness for; a) TiC, b) TiN, and c) TiAlN coating in convex surface

of 0.8 mm. On the other hand, in Fig. 9b, for the TiN coating, the largest values of Ra and ASP have been seen at the feed rate of 414 mm/rev and the step over of 2 mm. The lowest values of Ra and ASP have been observed at the feed rate of 318 mm/rev and the step over of 0.8 mm. Finally, in Fig. 9c, for the TiAlN

In this study, contouring and ramping cutting path styles set on down milling and up milling strategies were generated in concave and convex surfaces in semi finishing machining operations. In the experiments, the different tools with different coatings were employed. In addition, the relationship between surface roughness and cutting sound pressure level was observed with different levels of cutting velocity, feed rate and step over values. The indexable insert of ball end mills with TiC, TiN, and TiAlN coatings were used for machining the formed inclined surfaces. As a result: • The previous studies were focused on estimation of tool wear, vibrations stability and best cutting parameters by means of acoustic emission. In this study, the experiments on EN X40CrMoV5-1 material were carried out in order to understand the relationships between the sound signals generated and surface roughness in the machining of inclined concave and convex surfaces. The previous studies do not investigate the studied experimental cases of the inclined surfaces.

• The obtaining smaller ASP levels in concave surface shape machining (Fig. 5) in comparison to convex surface shape machining (Fig 8) might be explained by the surface form in which convex surface allows the microphone to gather the sounds with less obstruction.

• In convex surface shape machining when the all coatings are considered, the greater ASP levels of 111.9 and 113.4 dB in the TiN coating have been seen at the experiments 10 and 13 as they are depicted in Fig. 8, respectively. The reason for that is that the ball end mill has less contact

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•

•

•

•

•

• •

with the workpiece during the machining and it produces chatter mechanism. The surface roughness values belonging to those experiments are greater than of the other coatings as seen in Fig. 8.

Acoustic sound pressure level decreased in the concave surface because the tool worked inside the workpiece and contacted the workpiece with the larger cutting edge in inner surface.

After the evaluation of the sound pressure and surface roughness in terms of feed rate and step over parameters, the increase in the parameters raised ASP and Ra values without a dependency of tool coatings and surface shape.

As for the effects of tool coatings on the values of Ra; in the experiments, TiAlN coating for low cutting velocity displayed a good performance, and the increase in feed rate and step over values raised Ra values independent of tool coatings.

The greater ASP and Ra values are formed for concave surface type in up milling strategy.

The reason for that can be explained by chatter mechanism. The chatter occurred because of the type of surface, longer cutting tool and cutting tool’s movement from less chip volume to more chip volume.

ASP and Ra values of convex surface type in down milling strategy are higher than up milling.

The overlap of cutting edges of the cutting tool on the workpiece increased both, sound pressure and surface roughness because up milling and the machined part were convex.

It has been observed that in contouring, the ASP, which influences Ra, decreases with the increase in milling position angle.

In ramping, the ASP, which again influences Ra, is hardly affected by the milling position angle.

4 ACKNOWLEDGMENT

The authors wish to thank Karabük University (Project code: KBÜ-BAP-C-11-D-004) for providing financial support to conduct this study.

5 REFERENCES [1] Koksal, S. (2000). Face Milling of Nickel-Based Super Alloys with Coated and Uncoated Carbide Tools. PhD thesis, Coventry University, Coventry.

[2] Shaw, M.C. (2005). Metal Cutting Principles. Oxford University Press, Oxford.

[3] Wolter, B., Dobmann, G., Boller, C. (2011). NDT based process monitoring and control. Strojniški vestnik -

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[4] Bagci, E., (2011). Monitoring and analysis of MRRbased feedrate optimization approach and effects of cutting conditions using acoustic sound pressure level in free-form surface milling. Scientific Research and Essays, vol. 6, no. 2, p. 256-277.

[5] Inasaki, I. (1998). Application of acoustic emission sensor for monitoring machining processes.

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[6] Jemielnaik, K., Otman, O. (1998). Tool failure detection based on analysis of acoustic emission signals. Journal of Materials Processing Technology, vol. 76, p. 192197, DOI:10.1016/S0924-0136(97)00379-8.

[7] Ghosh, N., Ravi, Y.B., Patra, A., Mukhopadhyay, S., Paul, S., Mohanty, A.R., Chattopadhyay, A.B. (2007).

Estimation of tool wear during CNC milling using neural network-based sensor fusion. Mechanical Systems and Signal Processing, vol. 21, p. 466-479, DOI:10.1016/j.ymssp.2005.10.010.

[8] Marinescu, I., Axinte, D.A. (2008). A critical analysis of effectiveness of acoustic emission signals to detect tool and workpiece malfunctions in milling operations. International Journal of Machine Tools & Manufacture, vol. 48, p. 1148-1160, DOI:10.1016/j. ijmachtools.2008.01.011.

[9] Marinescu, I., Axinte, D.A. (2009). A time–frequency acoustic emission-based monitoring technique to identify workpiece surface malfunctions in milling with multiple teeth cutting simultaneously. International Journal of Machine Tools & Manufacture, vol. 49, p.

53-65, DOI:10.1016/j.ijmachtools.2008.08.002.

[10] Rivero, A., Lopez de Lacelle, L.N., Penalva, L.M. (2008). Tool wear detection in dry high-speed milling based upon the analysis of machine internal signals.

Mechatronics, vol. 18, p. 627-633, DOI:10.1016/j. mechatronics.2008.06.008.

[11] Wilcox S.J., Reuben R.L., Souquet P. (1997). The use of cutting force and acoustic emission signals for the monitoring of tool insert geometry during rough face milling. International Journal of Machine Tools and Manufacture, vol. 37, p. 481-494, DOI:10.1016/S08906955(96)00069-7.

[12] Weingaertner, L.W., Schroeter, B.R., Polli, L.M., Gomes O.J., (2006). Evaluation of high-speed end-milling dynamic stability through audio signal measurements.

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[13] Tekiner, Z., Yeşilyurt, S. (2004). Investigation of the cutting parameters depending on process sound during turning of AISI 304 austenitic stainless steel. Materials and Design, vol. 25, p. 507-513, DOI:10.1016/j. matdes.2003.12.011.

[14] Salgado, R.D., Alonso, J.F., (2007). An approach based on current and sound signals for in-process tool wear monitoring. International Journal of Machine Tools &

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Manufacture, vol. 47, p. 2140-2152, DOI:10.1016/j. ijmachtools.2007.04.013.

[15] Ravindra, V.H., Srinivasa, G.Y., Krishnamurthy, R. (1997). Acoustic emission for tool condition monitoring in metal cutting. Wear, vol. 212, p. 78-84, DOI:10.1016/S0043-1648(97)00137-3.

[16] Haber, E.R., Jimenez, E.J., Peres, R.C., Alique, R.J. (2004). An investigation of tool-wear monitoring in a high-speed machining process. Sensors and Actuators, vol. 116, p. 539-545, DOI:10.1016/j.sna.2004.05.017.

[17] Quadro, L.A., Branco, J.R.T. (1997). Analysis of the acoustic emission during drilling test. Surface & Coatings Technology, vol. 94-95, p. 691-695, DOI:10.1016/S0257-8972(97)00509-4.

[18] Guo, Y.B., Ammula, S.C. (2005). Real-time acoustic emission monitoring for surface damage in hard machining. International Journal of Machine Tools & Manufacture, vol. 45, p. 1622-1627, DOI:10.1016/j. ijmachtools.2005.02.007.

[19] Asilturk, İ., Akkus, H., Demirci, T.M. (2011). Surface roughness modelling based on vibration, acoustic and cutting parameters via regression analysis. VI. Machine Design and Manufacturing Technologies Congress, p.

47-52.

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Monitoring gas metal arc welding process by using audible sound signal. Strojniški vestnik - Journal of

Mechanical Engineering, vol. 57, no 3, p. 267-278, DOI:10.5545/sv-jme.2010.181.

[21] Kek, T., Grum, J. (2009). AE signals as laser cutting quality indicators. Insight, vol. 51, no. 3, p. 124-128, DOI:10.1784/insi.2009.51.3.124.

[22] Chaari, R., Abdennadher, M., Louati, J., Haddar, M. (2011). Modelling of the 3D machining geometric defects accounting for workpiece vibratory behaviour.

International Journal of Simulation Modelling, vol. 10, no. 2, p. 66-77, DOI:10.2507/IJSIMM10(2)2.173.

[23] Roy, S.S. (2010). Modelling of tool life, torque and thrust force in drilling: a neuro-fuzzy approach.

International Journal of Simulation Modelling, vol. 9, no. 2, p. 74-85, DOI:10.2507/IJSIMM09(2)2.149.

[24] Gologlu, C., Sakarya, N. (2008). The effects of cutter path strategies on surface roughness of pocket milling of 1.2738 steel based on Taguchi method. Journal of Materials Processing Technology, vol. 206, p. 7-15, DOI:10.1016/j.jmatprotec.2007.11.300.

[25] Singh, S.K., Srinivasan, K., Chakraborty, D. (2004). Acoustic characterization and prediction of surface roughness. Journal of Materials Processing Technology, vol. 152 p. 127-130, DOI:10.1016/j. jmatprotec.2004.03.023.

[26] Montgomery, D.C., Peck, E.A., Vining G.G. (2006).

Introduction to Linear Regression Analysis. Wiley, Hoboken.

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Paper received: 2012-04-20, paper accepted: 2012-08-21 © 2012 Journal of Mechanical Engineering. All rights reserved.

Vacuum Insulation Panels - An Assessment of the Impact of Accelerated Ageing on Service Life Kunič, R. Roman Kunič

University of Ljubljana, Faculty of Civil and Geodetic Engineering, Slovenia Vacuum insulation panels (VIP) are the most effective thermal insulation and are primarily used where not much space is available and extremely high thermal insulation properties are required. Panels are already widely used in industry (insulation of the front doors and side walls of refrigerators and freezers for domestic appliances), for logistic purposes (transport containers, storage tanks, campers, refrigeration and other vehicles), for the insulation of laboratory and professional devices and insulation of various construction elements, such as elements for doors, furniture.

Since the entry of atmospheric air gases into VIP panels represents the most harmful influence on their excellent thermal insulation properties, our research was focused on the analysis of accelerated ageing and determination of service life caused by this influence. The primary objective of the research was to define the theoretical base supporting measurements of accelerated testing and the connection between ageing and the service life of VIP panels. Arrhenius law of accelerated testing was used as a theoretical support to experimental testing of VIP panels at high temperatures and different exposure times. These results are useful in order to control quality, resistance to ageing and determining the life expectancy of VIP panel products, the characteristics of which can be determined in a significantly shorter time during accelerated testing than during real service life conditions.

The accelerated ageing laboratory test measurements and simulation analysis confirm that VIP panels are a durable and high quality product.

Keywords: service life, ageing, accelerated testing, Arrhenius law, activation energy, vacuum insulation panel (VIP)

0 INTRODUCTION Most envelope constructions applied due to new demands for energy savings, greatly increase the thickness of thermal insulation, whereby a thickness of 30 or even 50 centimetres is not unusual. Extremely high energy efficient buildings, such as Near-0-Energy (N-0-E) houses, require thermal insulation with a U-value even lower than 0.10 W/(m2·K), which has resulted in layers of conventional thermal insulation of at least 30 cm. This imposes not only technical constraints but also has negative aesthetic effects.

The same situation is also encountered in the industry of refrigerators and other domestic and professional appliances, transport containers and vehicles, campers and laboratory equipment, where there is a space limitation.

It should be noted that the optimal thickness of the heat insulating layer is always obtained at the expense of the efficiency of the available volume, since the net yield of the useful volume of the envelope is often much smaller. On the other hand, it is also true that designers want to gain space.

Vacuum insulation panels are substantially more efficient (up to ten times) than conventional insulating materials [1]. The thermal conductivity may be lower than 0.0030 W/(m·K) and the thermal resistance of a panel only 20 mm thick can even reach 6.66 (m2·K)/W. This means that high heat resistance with an extremely small thickness of insulation panels can be achieved with the help of VIP. The thermal 598

insulation envelope shrinks, becoming up to five times thinner.

Growing interest has been noted recently in replacing standard insulation with the VIP type, leading to new issues that need to be addressed and clarified. One of the most important aspects of VIP is assessment of their service life and their reliability during their life cycle. This has a direct impact on reducing the consumption of raw materials and environmental protection aspects, such as recycling, planning and management of waste resulting from its removal. Accelerated ageing and determination of the life expectancy of thermal insulation materials play an important part in this.

The experimental methods that were used in this work aimed at evaluating the rate of degradation of VIP panels, i.e., to determining the values of thermal conductivity as a function of thermal loads. By taking into account Arrhenius law of accelerated ageing [2], the ageing mechanism of vacuum insulation panels was determined. Other loads, such as humidity and thermal cycling can be used and their effect on the service life of VIP can be assessed. However, we believe that temperature loads have the primary impact on VIP failure and thus deserve a detailed investigation.

1 VIP PANELS THEORY VIP panels are extremely suitable for cases in which a large thickness of insulating layers represents a

*Corr. Author’s Address: University of Ljubljana, Faculty of Civil and Geodetic Engineering, Jamova c. 2, SI-1000 Ljubljana, Slovenia, roman.kunic@fragmat.si

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physical barrier, where using VIP panels achieves high thermal insulation without significantly increasing the thickness of the complete insulation layer.

The panels are already used for insulation in various construction elements, such as elements for doors, furniture), logistics (transport and packaging containers, special containers and storage tanks, refrigeration vehicles), insulation of domestic appliances (refrigerator doors, sides of refrigerators and freezers) and for insulation of various laboratory and professional appliances and devices. The Slovenian manufacturer of domestic appliances thus already uses VIP panels in selected models of refrigerators and freezers for more demanding customers.

A VIP panel - non-homogenous composite thermal insulation material - is made from different components: first, a special film, which is vacuum tight and enables the insulating core to be sealed against air and moisture diffusion: second, vacuumated thermal insulation without air or moisture and, third, special gas and/or moisture absorbers, which are usually inserted inside the VIP panel based on special needs.

Panels can be manufactured in any size. In contrast to other insulation materials, after their manufacture, VIP cannot be cut to specific dimensions or disintegrated, and no piercing or mechanical fasteners are allowed.

1.1 Thermal Insulated Core Material Thermal insulation of the core at normal atmospheric pressure is also important because after the VIP panel service time is reached or the vacuum is lost due to puncturing, deformation of the outer foils or damage during transport, the residual thermal insulation still provides a certain insulation effect. In the study, we used core material made of specially processed glass fibres, without any binders, with a thermal conductivity value of about 32 mW/(m·K), as well as more efficient core materials such as micro-porous based silica powder structures. Unfortunately, the latter are considerably more expensive (such materials have a thermal conductivity of about 22 mW/(m·K).

Core material densities are usually in the range of 170 to 210 kg/m3.

The core material must have an open cell structure. It can be in powder form or have a fibrous structure, both allowing the air to be pumped out almost completely, thus enabling the creation of near vacuum conditions, estimated to be up to 1/1000’s of atmospheric pressure (1 mbar) or even 100×smaller (i.e., just 1 Pa). The core material must be compact enough to withstand high compression stresses caused by the air pressure without any additional load on the panel (represents a pressure of 1 bar).

1.2 Multilayer Envelope Foil as a Protection Film Special multilayer films (foils) should ensure air tightness against all atmospheric gases and a total vapour barrier for preventing the diffusion of moisture and air. One of the most crucial properties of VIP is foil resistance to ageing, which is related to its chemical composition and overall quality, and its ability to be bonded. The film structure is multi-layer (aluminium foil, PT, PET, HDPE, PP, nylon, etc.), while bonding is achieved by using appropriate adhesives. High quality foils have recently entered the market and better bonding systems have been developed [3].

Fig. 1. Vacuum insulation panel

The most important issue is to prevent air from entering from outside, which can happen due to mechanical damage or an imperfect weld bonding the two halves of the outer foils. It should be born in mind that even a small amount of air entering a VIP has an enormous effect on increasing its thermal conductivity.

1.3 Gas and Moisture Absorbers In order to reduce the influence of penetration of moisture and gases, special absorbers can be added to panels during production. They absorb gases, moisture or a combination of both. In our basic study, we focused mainly on the impact of accelerated ageing without additional absorbers, and a core made of silica powder is already a good moisture absorber.

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1.4 Basic Instructions for Using Vacuum Insulation

Table 1. Values of activation energy for some materials

VIP panels should be well protected during transport, storage and handling and any mechanical damage must be avoided. Practice shows that most damage occurs before and during the installation of panels.

Sharp objects or surfaces should not be in contact with VIP panels.

A larger panel performs better because the initial thermal conductivity is lower and resistance to ageing is better. In other words, thermal conduction through the edges of a panel is quite low compared to thermal conduction through the whole surface of a VIP panel.

At the same time, a relatively shorter length of edges compared to the whole panel area, means the vacuum inside VIP panels is better maintained.

An additional vapour barrier is usually needed in cases in which VIP panels are used on the inner or warmer and usually higher vapour pressure side of the thermal barrier.

Material Polyvinylchloride – PVC [9] Cement, contaminated with soil [10] Soil [11] Polycarbonate – UV stabilized [12] Batteries – empty (used) [13] Polypropylene [14] Natural oil sand bitumen [15] Rubber, vulcanized with sulphur [9] Rubber, vulcanized with sulphur and radiation [12] Plastic tiles for cars – durable [16] Batteries – new (not used) [13] Bitumen [17] Concrete [11] Bituminous membrane – oxidized bitumen [4] Elastomeric butyl [14] Polyurethane [14] Cement [18] Bituminous membrane – APP modified bitumen [4] EPDM [14] Epoxy resin reinforcement with glass fibbers [19] Polychloroprene rubber [20] Neoprene rubber [14] Bituminous membrane – SBS modified bitumen [4] Polyethylene for cable insulation [5] TISS – Thickness Insensitive Spectrally Sensitive coating, Polyurethane substrate, type B [7] Polyvinylchloride – PVC for cable insulation [6] Casts made of Al alloy [9] EP rubber [21] TISS – Thickness Insensitive Spectrally Sensitive coating, Polyurethane substrate, type A [7], [8] Ethylene Vinyl Acetate (EVA) [22] Alpha Olyphine Co-Polymer [21] EP Graphite Co-Polymer [21] Atactic Polypropylene – APP [21] Isotactic Polypropylene – IPP [21]

2 METHODOLOGY 2.1 History of Accelerated Ageing Historically, accelerated ageing has been used in very rapid testing under simulated field conditions in the laboratory. In many cases, this leads to excessive failures of such materials or products, which would not have occurred under normal conditions of use during normal service life. The conditions of accelerated ageing of various materials or products must therefore be adjusted to the conditions of normal use as far as possible. Some references to the results of relevant successful studies, tests and analysis are shown in Table 1; bituminous products with different additives which are used to extend service life [4], analysis of electrical cable insulation against ageing [5] and [6], ageing studies of different Thickness Insensitive Spectrally Sensitive (TISS) coatings for solar collectors [7] and [8].

2.2 Activation Energy The threshold for the reaction is the minimum energy - activation energy - that leads to a reaction.

In chemistry and biology, it is the threshold energy, or the energy required to produce a chemical reaction.

In other words, this is the minimum energy required for the initiation of certain chemical changes and it describes the material’s characteristics and allows indirect prediction of its degradation or ageing under the influence of various external factors. In order to 600

Activation energy Ea [kJ/mol] from to 24.5 37.5 25.0 58.0 25.0 71.0 26.0 28.5 31.0 34.0 35.0 50.0 36.9 46.7 37.0 40.0 45.0

49.0

48.0 50.0 53.9 55.0 58.4 60.0 65.0 71.0

69.0 55.0 57.9 70.0 58.4 100.0 119.0 73.0

75.7

75.7

78.0

127.0

80.0

135.0

82.0 84.0

96.0 88.0

85.2

85.2

88.0

107.0

96.0

96.0

98.0 135.0 160.0

99.0 145.0 160.0

160.0

166.2

176.0 180.0 200.0 600.0 1,300.0

184.0 180.0 200.0 600.0 1,300.0

exceed the reaction level of the activation energy, we need to achieve a sufficiently high temperature and correct orientation and energy of molecules.

Activation energy is usually expressed as the energy [J] required for the reaction of one mole of a substance or reactant, i.e., J/mol. For substances with high activation energy (more than 170 kJ/mol), Kunič, R.

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reactions are only detectable at temperatures higher than 400 °C. Table 1 shows the activation energy for various materials. If cement (Ea is from 71 to 73 kJ/mol) is contaminated with soil, the activation energy is only one-third to one-quarter of the original value (from 24.5 to 37.5 kJ/mol). Polymers that are used for mixing with other materials in order to improve resistance to low and high temperatures and the UV spectrum of solar radiation, i.e., to increase the resistance to ageing of the majority of products, have extremely high levels of activation energies; atactic polypropylene (APP) has an activation energy of 600 kJ/mol and isotactic polypropylene (IPP) of 1300 kJ/mol.

2.3 The Arrhenius Equation The Arrhenius equation describes the dependence of the speed of chemical reactions on various physical parameters, such as temperature, humidity, pressure etc. The magnitude of changes affects the ability of the material to resist those factors that provide activation energy. The first steps of the theory were proposed by the Dutch chemist, J. H. van’t Hoff, in 1884. Five years later, the Swedish chemist and later Nobel Prize winner, Svante Arrhenius published a physical interpretation of the equation. Arrhenius claimed that, in order to achieve a reaction that changes the material, a sufficient amount of energy needs to be supplied.

The Arrhenius equation describes the logarithmic relationship (based on a natural logarithm) between temperature and activation energy. Despite the fact that the defaults simplify many effects as being temperature independent, several studies have clearly confirmed that the process of degradation of a material can be sufficiently accurately described by a simple Arrhenius equation [2].

In thermal balance at absolute temperature (T), the parts of molecules that have a kinetic energy of more than (Ea) can be calculated by MaxwellBoltzmann’s distribution of statistical mechanics and are in accordance with the literature [2] in relation to:

k = A⋅e

−

Ea R ⋅T

, (1)

where Ea is activation energy [J/mol], R ideal gas constant, T temperature [K], k specific rate of reaction, reaction constant factor or reaction speed [1/s], and A pre-exponent factor or frequency factor of a certain reaction [1/s]. By using the natural logarithm, the Arrhenius equation can be expressed as:

ln k = −

Ea 1 ⋅ + ln A, (2) R T

If the specific rates of reactions (k1 and k2) at two different temperatures (T1 and T2) are known, the activation energy can be determined [16]:

ln

E T −T k2 = ln aT = a ⋅ 2 1 , (3) k1 R T1 ⋅ T2

where αT is coefficient of acceleration of ageing due to degradation at elevated temperatures [-].

In accordance with Arrhenius law (Eq. 3), in order to determine the activation energy (Ea), at least two accelerated ageing testing procedures at two different temperatures are needed. Changes of temperature from T1 to T2 lead to an acceleration of ageing (aT), which depends on the value of activation energy Ea (Eq. 3). In other words, if after a time period t1 at a temperature level T1 the same degradation is caused as after a time period t2 at temperature level T2 , then the acceleration of degradation processes caused by changes in temperature can also be discovered. This feature is called the activation energy (Ea).

Since the relationship between the temperature and the reaction is constant and follows Arrhenius law, then the graph representing ln k as a function of 1/T is a straight line (Fig. 2). The slope of the graph is in interdependence with the activation energy Ea , more precisely with –Ea/R. The higher the activation energy, the greater the slope of the curve.

Fig. 2. According to Arrhenius law, the graph represents the relation between ln k and 1/T

Arrhenius equation thus provides a qualitative basis for the relation between activation energy and the rate at which a certain reaction takes place.

This process is the basis for experimental chemical kinetics, which is used for the determination of activation energy for individual reactions.

In the present analysis, Arrhenius equation (Eq. 3) was used as a suitable empirical expression to describe temperature dependence of VIP panels’ degradation process.

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Since most films are unstable or could be permanently damaged above temperatures around 105 to 110 °C, we selected the following accelerated ageing temperatures: 100, 90, 80, 70 and 60 °C, with time exposures from half a day up to 3 months or, in some cases, even longer.

Accelerated ageing test was performed on VIP panels with dimensions approx. 25×25 cm (total of 112 samples).

2.4 Simulation Analyses The main advantages of simulation analyses in comparison with testing in a real, natural environment are rapid implementation and rapid acquisition of results and thus lower prices of testing procedures [12]. The procedure of accelerated ageing in which the level of a particular load or several different loads is increased above the load level in normal use, is an alternative to observation under natural conditions of use. Testing time should thus be significantly shorter [12].

Service life does not usually mean a period without any deterioration. It is also a function of the manner of use and maintenance. Flaws and damage are a matter of subjective assessment but are used as the basis for determining the life expectancy of a particular material, product or component in the system. Properties often have more than two states (airproof against air penetration, waterproof against water leakage, resistance to instability, etc.) but the determination of slow degradation due to a continuous ageing process is much more difficult. In the case of a dramatic or even catastrophic drop in quality (also called ‘sudden death’), the determination of service life is very simple. More often, degradation is a slow, continuous reduction of quality and usefulness and other characteristics of the product. In such a case, we need to determine the maximum or acceptable change in the product quality and its properties, i.e., we need to define the maximum allowable degradation of a product that is still acceptable to users, as well as the appearance, functionality and all other relevant parameters of products at the end of their service life [23].

3.1 Thermal Conductivity We can assume that the primary function and the most important requirement for VIP panels is excellent thermal conductivity, which also depends on the quality of application, installation, details, thermal bridges, design connections of constructional complexes and base surfaces. The thermal conductivity measurements were taken with a thermal conductivity meter (Fig. 4).

The goal is to maintain thermal conductivity as low as possible. Thermal conductivity measurements at the UAE Institute in Bavaria have confirmed values even lower than 3 mW/(m·K).

Fig. 4. Laboratory measurement of thermal conductivity

Because of high hygroscopic properties of the tested micro-porous-based silica powder core material in VIP panels, the further aspect of accelerated tests focused on thermal loads and not the influence of humidity. Thus, the assumption with only temperature dependence of the degradation process was accepted for indoor environment or refrigerator use of VIP panels. However, if the same insulation was used in high humidity and at elevated or high temperatures (up to around 100 °C), both, thermal and humidity dependence of VIP panels’ degradation process should be taken into account.

3 EXPERIMENTAL Laboratory measurements and numerical analyses were carried out on VIP panels exposed to stress under high temperature in an oven (Fig. 3).

3.2 Definition of Service Life of VIP Panels, Mechanism of Degradation and Its Functional Capability The main problems encountered in applying the recommended accelerated test procedure for VIP

Fig. 3. Oven for accelerated ageing temperature stress

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panels are, firstly, the selection and application of suitable thermal loads and secondly, to observe degradation over different periods of time. Due to the non-reversible deformation of VIP panels, overexposure to excessively high thermal loads must be avoided during the procedure of accelerated ageing, since they could cause thermal instability and disintegration of the material (mainly of multilayer protection foils).

It is difficult to determine the deterioration of the product or the whole system in percentage terms. The thermal conductivity value of VIP panels with higher than double value of initial thermal conductivity, thus worse insulation property, do not perform their important function of resistance to ageing well. We set the assumption of service life as the time at which the value of thermal conduction doubled.

In order to predict the service life according to Arrhenius law, it is necessary to know the activation energy (Ea), i.e., the threshold energy required to produce a chemical reaction. In general, the activation energy can be determined from the measured changes of any of the material’s properties that can serve as a degradation indicator. In order to obtain the activation energy connected to studying the degradation process of VIP panels, we used the determination of thermal conductivity value.

In addition to mechanical damage or puncture, the entry of gases (i.e., air or other atmospheric gases) or moisture (primarily through diffusion) significantly reduces thermal insulation. Even small amounts of gases or moisture cause a large increase in thermal conductivity [24].

The assumption for service life span in this study was the time at which double of the initial thermal conductivity of VIP panel is reached (i.e. increase of thermal conductivity by 100%).

4 RESULTS AND DISCUSSION The most important loads on a VIP panel, if mechanical damage is excluded, are elevated temperature and humidity. These two loads individually or simultaneously and at higher intensity, lead to an increase in the passage of gases and moisture into a VIP panel. This process significantly increases the thermal conductivity, which is limited by the value of the thermal insulation of the core material. Our task was to establish the time at which double the initial thermal conductivity is reached.

During service life period over years or decades, gases and moisture slowly and constantly penetrate into VIP panels, primarily through sealants of multilayer

foils, thus increasing vapour pressure, resulting in higher thermal conductivity. The measurements and analysis in this study were focused on thermal load of VIP panels during accelerating ageing tests. It is possible to neglect the effects of moisture dependence because of high hygroscopic properties of VIP core material (micro-porous silica powder) and because of relative low ambient temperature during the entire service life (indoor or domestic appliance refrigerator use). In addition, temperature and humidity have possible influence on service life only when mechanical deformation of VIP panels occurs, and not if they are used properly; without any deleterious chemical, physical or mechanical influence.

4.1 Assessment of Accelerated Ageing Degradation at elevated temperature is determined as follows [12]: E

aT =

aT =

k2 k1

1

1

a ( − ) k2 τ 2 R ⋅ lnaT = = e R T1 T2 ⇒ Ea = , (4) 1 1 k1 τ 1 ( − ) T1 T2

=

τ1 τ2

=e

−

Ea  1 1   −  R  T2 T1 

=e

Ea  1 1   −  R  T2 T1 

=e

Ea  T2 −T1    R  T1 ⋅T2 

, (5)

where k1 is specific reaction rate under accelerated test conditions [1/s], k2 specific reaction rate under normal or working conditions [1/s], τ1 time to degradation during accelerated ageing test temperature [hour, day, week, month, year], τ2 time to degradation at operating temperature, therefore during normal use [month, year, decade], T1 testing temperature [K], and T2 operating or working temperature [K].

We assume that the limit of functionality of a VIP panel (i.e., complete degradation) occurs when it reaches the thermal conductivity of the core under normal atmospheric pressure, which in our case is 32 mW/(m·K) for glass fibres and 22 mW/ (m·K) for silica powder. In accordance with Eq. (4), the procedure of accelerated ageing at two or more different temperatures and, as a consequence, at two or more different exposure times, enables the calculation of the activation energy. The proposed temperatures for accelerated ageing testing procedures (from 60 to 100 °C) were confirmed as the confidence interval by means of the results of partly degraded samples.

Above such temperatures, too high or even totally degraded samples can be expected, i.e., no reliable comparisons can be made and the results represent non-Arrhenius behaviour.

From all the obtained results, the value of the activation energy for VIP panel products (Ea = 66.35

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kJ/mol, Table 2) can be determined. The value of activation energy is used to evaluate the behaviour of the VIP panel products exposed to different temperatures, by interpolation within the testing temperature interval at which accelerated ageing is performed or, more commonly, by extrapolation outside the testing temperature interval (Tables 3 and 4). In doing so, it is necessary to be careful and critical of the results and to be aware that extrapolation to temperatures far outside the testing interval cannot be accurate since, in such cases, the influence of the neglected temperature dependent pre-exponent factor in the Arrhenius equation becomes more important and the results are increasingly inaccurate.

possible to transform the thermal load (which represents degradation kinetics at the specific level of temperature) in any of the other thermal load curves, only in cases in which the same activation energy (Ea) is used for all transformations [8]. In other words, the results obtained from experimentally accelerated ageing, from the activation energy obtained by Arrhenius law and, consequently, on the basis of the calculated times and temperatures of ageing (Tables 3 and 4), can be graphically presented in Figs. 5 and 6, showing the expected ageing period and the degree of degradation as a function of the ageing temperature.

If, however, we move along a curve, the same level of ageing or degradation is achieved.

100

Table 2. Activation energy of vacuum insulation panels Activation energy Ea [kJ/mol]

Ageing [log years]'

Ageing at lower Ageing at higher temperatures temperatures No. Temp.

Ageing Temp.

Ageing [°C] [days] [°C] [days] 1 60 51 80 14 2 60 58 80 14 3 60 148 80 21 4 60 178 80 51 5 60 51 100 5 6 60 58 100 6 7 60 148 100 6 8 60 178 100 14 9 80 148 100 58 10 80 148 100 51 11 80 178 100 51 12 80 178 100 58 Average value of activation energy (Ea): Median value of activation energy (Ea): Standard deviation of activation energy (Ea):

63.234 69.525 95.513 61.140 60.015 58.627 82.835 65.709 51.322 58.368 68.480 61.434 66.350 62.334 12.014

10

1

0 0

10

20

30

40

50

60

70

80

Temperature [oC]

Fig. 5. Ageing of vacuum insulation panels (VIP); expected time when thermal conductivity reaches double the initial value in relation to service life temperature

Prognosis of behaviour when exposed to lower temperatures during the lifetime of the product, according to Arrhenius law, is carried out by extrapolation to longer time periods, taking into account the activation energy Ea = 66.35 kJ/mol (Table 4).

Ageing in accordance with the times and temperatures listed in Tables 3 and 4, is such that complete disintegration of the product, or the end of its expected service life, does not occur except in cases of

4.2 Determination of Service Life It can be concluded that the thermal load to which we exposed the VIP panels caused a degradation process that followed Arrhenius law. It is therefore Table 3. Extrapolation of accelerated ageing over shorter periods of time

Accelerated time [days] 240 120 90 60 30 21 Exposure temperature of VIP panel, when its thermal 72.6 83.3 88.0 94.7 106.8* 113.4* conductivity reached double its initial value [°C] * because of irreversible damage (chemical decomposition), VIP panels should not be exposed to such high temperatures.

14

7

121.2*

135.1*

Table 4. Extrapolation of accelerated ageing over longer periods of time Expected lifetime [years] 1 Exposure temperature of VIP panel, when its thermal 66.4 conductivity reached double its initial value [°C] * accuracy of extrapolation to such a long service life is questionable.

604

2

3

5

10*

15*

20*

30*

56.7

51.3

44.7

36.1

31.3

28.0

23.5

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extremely high temperatures, when such degradation occurs and results in non-reversible degradation or leads to chemical degradation.

Ageing [log years]'

100

10

1

periodic loads up to 80 °C, without the presence of mechanical damage, such as piercing, strong abrasion, bending or other deformations.

On the basis of our assumptions, service life is the time until the thermal conductivity value is double the initial value of the insulation panel. The results of our accelerated tests show that the thermal conductivity of a VIP panel exposed to constant ambient temperature (25 °C for indoor conditions) doubles, i.e., thermal insulation property of VIP panel falls by 50%, after a period of 26.2 years and a thermal conduction value of 12 mW/(m·K) is reached after 48.1 years.

6 NOMENCLATURES

0 0

10

20

30

40

50

60

70

80

Temperature [oC]

Fig. 6. Ageing of vacuum insulation panels (VIP): expected time when thermal conductivity reach value of 12 mW/(m·K) in relation with service life temperature

5 CONCLUSIONS As a result of accelerated ageing laboratory test measurements, calculation simulation analysis and other scientific investigations, it can be concluded that VIP panels are a durable and high quality product, mainly in terms of maintaining exceptional thermal insulation, which was confirmed during testing, measurements and analysis taking into account Arrhenius law.

VIP panels can be used for insulating barriers, both in the industry (household appliances, laboratory equipment, medical devices, equipment for various manufacturing and other processes), transport and trade (refrigeration trucks, refrigeration containers, refrigerated boxes, food machines and refrigerators, mobile homes, campers and other vehicles), as well as for construction purposes (outer walls, terraces, special panels, containers, for insulation of alternative systems for cooling and ventilation [25] and in general where space is very valuable). Due to their high price, the use of VIP panels is justified in particular when there is a lack of space for thermal insulation. VIP panels provide, with minimum thicknesses, much higher thermal insulation than in standard, widely used construction insulation materials.

Caution should be exercised during transport, storage and manipulation with VIP panels. Packaging of the finished product is recommended in well protected boxes. Mechanical damage during the installation and use of VIP panels should be prevented at all costs. Exposure of VIP panels is recommended in dry, not excessive humid conditions, with thermal

A pre-exponent factor or frequency factor of a

particular reaction [1/s] Ea activation energy [J/mol] k specific rate of reaction, reaction constant

factor or reaction speed [1/s] k1 specific reaction rate under test conditions [1/s] k2 specific reaction rate under normal or

working conditions [1/s] R ideal gas constant, R=8.314472 J/(K·mol) T temperature [K] T1 test temperature [K] T2 operating or working temperature [K] αT coefficient of acceleration of ageing due to

degradation at elevated temperatures [-] τ1 time to degradation at an accelerated ageing

test temperature [hour, day, week, month, year] τ2 time to degradation at operating temperature,

therefore during normal use [month, year, decade].

6 REFERENCES [1] Tenpierik, M.J., Cauberg, J.J.M. (2007). VIP integrated facade designs: the advantage of combining high thermal performance with limited construction thickness. Proceedings of the 24th International - PLEA Conference on Passive and Low Energy Architecture, PLEA/NUS/RBP, p. 303-310.

[2] Glasstone, S. (1951). Textbook of Physical Chemistry.

Macmillan & Co. Ltd., London, p. 1193-1210.

[3] Carmi, Y. (2011). Developing 60 year lifetime metallized VIP laminates – Challenges and solutions.

10th International Vacuum Insulation Symposium, Vacuum Insulation Panels: Advances in Applications, p. 167-170.

[4] Kunič, R., Orel, B., Krainer, A. (2011). An assessment of the impact of accelerated ageing on the service life of

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bituminous waterproofing sheets. Journal of Materials in Civil Engineering, vol. 23, no. 12, p. 1746-1754, DOI:10.1061/(ASCE)MT.1943-5533.0000326.

[5] Gillen, K.T., Bernstein, R., Celina, M. (2004). Non – arrhenius behavior for oxidative degradation of chlorosulfonated polyethylene materials. Polymer Degradation and Stability, vol. 87, no. 2, p. 335-346, DOI:10.1016/j.polymdegradstab.2004.09.004.

[6] Jabukowitcz, I., Nazdaneh, Y., Geart, T. (1999). Effects of accelerated and natural ageing on Plasticized Polyvinyl Chloride (PVC). Polymer Degradation and Stability, vol. 66, no. 3, p. 415-421, DOI:10.1016/ S0141-3910(99)00094-4.

[7] Kunič, R., Mihelčič, M., Orel, B., Slemenik Perše, L., Bizjak, M., Kovač, J., Brunold, S. (2011). Life expectancy prediction and application properties of novel polyurethane based thickness sensitive and thickness insensitive spectrally selective paint coatings for solar absorbers. Solar Energy Materials And Solar Cells, vol. 95, no. 11, p. 2965-2975, DOI:10.1016/j. bbr.2011.03.031.

[8] Kunič, R., Koželj, M., Orel, B., Šurca Vuk, A., Vilčnik, A., Slemenik Perše, L., Merlini, D., Brunold, S. (2009). Adhesion and thermal stability of thickness insensitive spectrally selective (TISS) polyurethanebased paint coatings on copper substrates. Solar Energy Materials And Solar Cells, vol. 93, no. 5, p. 630-640, DOI:10.1016/j.solmat.2008.12.026.

[9] Jorgensen, G. (2003). A phenomenological approach to obtaining correlations between accelerated and outdoor exposure test results of organic materials.

New Directions in Coatings Performance Technology, ASTM STP 1435, ASTM International, p. 12-22.

[10] Chitambira, B., AL-Tabbaa, A., Perera, A.S.R., Yu, X.D. (2007). The activation energy of stabilized/ solidified contami-nated soils. Journal of Hazardous Materials, vol. 141, no. 2, p. 422-429, DOI:10.1016/j. jhazmat.2006.05.080.

[11] Jorgensen, G., Brunold, S., Carlsson, B., Möller, K., Heck, M., Köhl, M. (2003). Durability of polymeric glazing materials for solar applications. First European Wethering Symposyum, p. 1-15.

[12] Köhl, M., Carlsson, B., Jorgensen, G., Czanderna, A.W. (2004). Performance and durability assessment: Optical Materials for Solar Thermal Systems, Elsevier, Oxford.

[13] Liaw, B.Y., Roth, E.P., Jungst, R.G., Nagasubramanian, G., Case, H.L., Doughy, D.H. (2003). Correlation of Arrhenius behaviours in power and capacity fades with cell impedance and heat generation in cylindrical lithium-ion cells. Journal of Power Sources, vol. 119121, p. 874-886, DOI:10.1016/S0378-7753(03)001964.

[14] Celina, M., Gillen, K.T., Assink, R.A. (2005).

Accelerated ageing and lifetime prediction: Review of non-Arrhenius behaviour due to two competing processes. Polymer Degradation and

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Stability, vol. 90, no. 3, p. 395-404, DOI:10.1016/j. polymdegradstab.2005.05.004.

[15] Sonibare, O.O., Egashira, R., Adedosu, T.A. (2003).

Thermo-oxidative reactions of Nigerian oil sand bitumen. Thermochimica Acta, vol. 405, no. 2, p. 195205, DOI:10.1016/S0040-6031(03)00192-8.

[16] Nohara, M. (1997). Study of prediction method for thermo-oxidative life of plastics by thermal analysis.

JSAE Review, vol. 19, no. 3, p. 263-268, DOI:10.1016/ S0389-4304(98)00009-5.

[17] Owens, J.W. (1997). Life cycle assessment: constraints on moving from inventory to impact assessment.

Journal of Industrial Ecology, vol. 1, no. 1, p. 37-49, DOI:10.1162/jiec.1997.1.1.37.

[18] Bochen, J., Gil, S., Szwabowski, J. (2005). Influence of ageing process on porosity changes of the external plasters. Cement and concrete composites, vol. 27, no. 7-8, p. 769-775, DOI:10.1016/j. cemconcomp.2005.01.003.

[19] Budrugeac, P. (2001). Thermal degradation of glass reinforcement epoxy resin and polychloroprene rubber: the correlation of kinetic parameters of isothermal accelerater ageing with those obtained from nonisothermal data. Polymer Degradation and Stability, vol. 74, no. 1, p. 125-132, DOI:10.1016/S01413910(01)00112-4.

[20] Gillen, K.T., Bernstein, R., Derzon, D.K. (2004).

Evidence of non-Arrhenius behaviour from laboratory ageing and 24-year field ageing of polychloroprene rubber materials. Polymer Degradation and Stability, vol. 87, no. 1, p. 57-67, DOI:10.1016/j. polymdegradstab.2004.06.010.

[21] Fawcett, A.H., Mcnally, T. (1999). Blends of bitumen with various polyolefins. Polymer, vol. 40, no. 23, p.

6337-6349, DOI:10.1016/S0032-3861(98)00779-4.

[22] Czarderna, A.W., Pern, F.J. (1995). Encapsulation of PV modules using ethylene vinyl acetate (EVA) copolimer as a pottant: A critic review. Solar Energy Materials and Solar Cells, vol. 43, no. 2, p. 101-181, DOI:10.1016/0927-0248(95)00150-6.

[23] Brunold, S., Frei, U., Carlsson, B., Möller, K., Köhl, M. (2000). Accelerated life testing of solar absorber coatings: Testing procedure and results. Solar Energy, vol. 68, no. 4, p. 313-323, DOI:10.1016/S0038092X(00)00034-7.

[24] Mukhopadhyaya, P., Kumaran, M.K., Sherrer, G., van Reenen, D. (2011). An investigation on longterm thermal performance of vacuum insulation panels (VIPs). 10th International Vacuum Insulation Symposium, Vacuum Insulation Panels: Advances in Applications, p. 143-148.

[25] Stritih, U., Butala, V. (2011). Energy savings in building with a PCM free cooling system. Strojniški vestnik – Journal of Mechanical Engineering, vol. 57, no. 2, p. 125-134, DOI:10.5545/sv-jme.2010.066.

Kunič, R.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 607-613 DOI:10.5545/sv-jme.2010.265

Paper received: 2010-12-28, paper accepted: 2012-04-04 © 2012 Journal of Mechanical Engineering. All rights reserved.

Nautical Risk Assessment for LNG Operations at the Port of Koper

Perkovic, M – Gucma, L. – Przywarty, M. – Gucma, M. – Petelin, S. – Vidmar, P. Marko Perkovic1,* – Lucjan Gucma2 – Marcin Przywarty2 – Maciej Gucma2 – Stojan Petelin1 – Peter Vidmar1 1University

of Ljubljana, Faculty of Maritime Studies and Transport, Slovenia 2Maritime University of Szczecin, Poland

Construction of an LNG (Liquid Natural Gas) terminal by TGE (TGE Gas Engineering GmbH) is currently planned to be located along the Slovenian coast near the city of Koper. Two other LNG terminals are also plannedl: one in the Trieste port and the other off shore in Trieste Bay in Italy. Focusing on nautical operations, the purpose of this paper is to identify potential risks and to assess their levels as consequences of increased LNG activities. The ports in the area are host to a variety of vessels, including containers, tankers and chemical carriers, general cargo vessels, passenger ferries, bulk carriers, ro-ro carriers, etc.; and a large number of recreational and fishing vessels can be located on the navigational line towards ports located in Trieste bay. There are around 2500 vessels calling at the Port of Koper a year and approximately the same number at the Port of Trieste as well as a few hundred more nearby at the Port of Monfalcone. Using a quantitative approach, collision and grounding risk assessment will be analyzed for the shipping situation in the area, obtained through AIS (Automatic Identification System).

Keywords: LNG terminal, shipping, accidents, risk

0 INTRODUCTION On the Slovenian coast near the city of Koper an LNG (Liquefied Natural Gas) terminal is planned to be built

by TGE. This LNG terminal will have a throughput of 5 billion Nm³/year. To enable this throughput a jetty will be constructed at the end of mole II between basin 2 and basin 3 (Fig. 1) for the import of LNG using specialised LNG carriers [1].

be gasified and exported to the external gas pipeline distribution system. A similar layout is envisioned for Trieste by the German utility group E.ON. A special off shore terminal - though not floating - is planned to be located in the middle of Trieste bay [2].

To identify and assess the nautical level of risk of the LNG activities a risk study has been performed focusing on the nautical operations of the foreseen LNG activities in the bay and the Port of Koper [2] to [4]. This risk assessment compares the risks in the Port of Koper and in the bay following potential accidents/collisions with an LNG carrier.

1 SETTING THE SCENE 1.1 Activities in the Port of Koper

Fig. 1. Location of proposed LNG terminal

At the north side of this jetty, LNG carriers (LNGC) with capacities varying from 70,000 m³ to 220,000 m³ will berth and unload LNG. From the jetty an above-ground pipeline will transport the LNG up to the western end of mole II. Here, the pipeline will go underground and transport the LNG further to the LNG terminal where the LNG will be stored in LNG tanks. From these tanks liquid natural gas will

The Port of Koper, as the lone port of a sovereign nation is as it must be a multipurpose port, equipped and prepared for handling and warehousing all types of goods. The basic port activities are carried out at specialized terminals, which are technologically and organizationally suitable for handling and warehousing specific cargo groups. Fig. 2 gives information about the throughput of various types of cargo, by not actually providing an overall picture of the port, its capacities, complications and specific circumstances.

Just as the Adriatic is a shallow finger of the Mediterranean, the bay of Koper is shallow – Koper is a shallow port with many traffic constraints and challenges. The region also combines typical northern Mediterranean weather and a micro-climate that includes the burja (bora) wind as the chief obstacle

*Corr. Author’s Address: University of Ljubljana, Faculty of Maritime Studies and Transport, Pot pomorscakov 4, 6320 Portoroz, Slovenia, marko.perkovic@fpp.uni-lj.si

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to easy ingress and egress, as well as local variants on the typical storms and winds of the Mediterranean. As the sole port for Slovenia, it must handle every type of cargo that is traded by ship, those listed or implied in Fig. 2, along with refrigerated consumables, timber, specific bulk cargos like coal and iron, innumerable minerals and foodstuffs, sandy alumina (with a 20,000 ton capacity silo), as well as livestock.

3 45°35.81’N; 13°39.53’E (center of outbound traffic lane from Koper) 4 45°36.71’N; 13°36.42’E (in precaution area, approximate center of prolonged traffic lane to Trieste) 5 45°37.86’N; 13°33.07’E (center of SW bound traffic lane) 6 45°35.88’N; 13°26.38’E (center of SW bound traffic lane) xkp 45°37.68’N; 13°35.59’E (center of precaution area)

Fig. 3. Actual AIS tracks in the area Fig. 2. Maritime cargo throughput in tons

The Port of Koper is therefore a relatively small and shallow bay with terminals for virtually all of the country’s imports and exports with a notable exception of LNG. Traffic includes a variety of ship sizes, for the most part limited on its larger side only by the depth of the Adriatic and its ports in general – Koper is just a single peninsula southwest of the port of Trieste, and in fact Trieste is often used as the point of orientation for port workers such as pilots and tug captains. Given the variety of cargo types, there is also a tremendous variety of ship types, converted vessels, some of the oldest seagoing vessels, as well as the newest. Being so close to the waters of two different countries, and especially so close to Trieste, traffic considerations do not begin or end at the port itself, but must be considered beyond the bay as well [5].

Positions of AIS gates and modeled routes are depicted in Fig. 4, where traffic is analyzed for; A Incoming vessels with destination: - Koper crossing gates no. 1 and 2, - Trieste or Monfalcone crossing gates no. 1

and 4, B Outgoing vessels sailing from: - Koper crossing gates no. 3, 5 and 6, - Trieste or Monfalcone crossing gates no. 5

and 6, C Vessels from Koper to Trieste or Monfalcone crossing only gate no. 3, D Vessels from Trieste/Monfalcone to Koper crossing gates no. xkp and 2.

1.2 Traffic in the Area Traffic have been analyzed on the basis of AIS research (Fig. 3). Vessel passages are checked in 6 gates; 1 45°33.14’N; 13°28.21’E (center of NE bound traffic lane) 2 45°34.48’N; 13°39.60’E (center of inbound traffic lane to Koper) 608

Fig. 4. AIS Observation gates and modeling routs

Perkovic, M – Gucma, L. – Przywarty, M. – Gucma, M. – Petelin, S. – Vidmar, P.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 607-613

For each vessel course, speed and distance to check point at the time of crossing the perpendicular line to the direction of traffic flow have been measured (Fig. 5). On the basis of AIS measurements ship routes have been determined. After an analysis due to different behavior of navigators on different sizes of ships, routes have been analyzed for three different ship sizes separately (up to 120 m, 120 to 180 m and over 180 m in length). The AIS research enabled the evaluation of two probabilistic parameters of ships routes: mean and standard deviation of Way Points in given routes. Traffic on the routes has been modeled with the use of the Poisson distribution and Way Points have been modeled by 2-dimensional normal distribution with mean and standard deviation assessed from AIS data.

The graphs (Fig. 5) show the number and type of vessels entering and leaving the Port of Koper for each basin.

For the purpose of grounding in the port vicinity and collision risk assessment with LNG jetty several additional parameters are considered such as: vessel type, draught, pilotage and tug escorting status, approaching course and speed and anchorage occupancy. Fig. 6 presents the yearly period during which vessels are grouped by size and type for basin 1. The same data collection approach was used for basin 2 and 3. Fig. 7 presents control lines and sample data for one approaching vessel.

Fig. 6. Vessels by size and type entering and leaving Port of Koper basin 1

a)

b) Fig. 7. Vessels by size and type entering and leaving Port of Koper basin 1

2 METHODOLOGIES

c) Fig. 5. Histogram function of ships variability from route on AIS Gate no. 1 for three different sizes; a) up to 120 m, b) 120 to 180 m, c) over 180 m in length

One of the most appropriate approaches to assessing the safety of complex marine traffic engineering systems is the use of stochastic simulation models [6]. The model can be used for almost all navigational accident assessments, such as collisions between vessels, groundings, collisions with fixed objects, indirect accidents such as those involving anchors or

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accidents caused by ship generated waves [7]. The model could comprise several modules responsible for different navigational accidents.

The design of the collision model was divided into 4 stages. The main goal of this model is to calculate the probability of collision for a given type of encounter. Because there are no statistical data about collisions in the Gulfs of Trieste and Koper the probability calculated for the Baltic Sea was used [8].

Stage I: Division of the encounters into characteristic types • Head-on encounter – difference of headings 170 to 190° and distance less than critical.

• Overtaking – difference of headings more than 350° or less than 10° and distance less than critical.

• Crossing – difference of headings (rest of cases) and distance less than critical.

The critical distances where navigators perform anti-collision manoeuvres was assumed on the basis of expert opinions separately for each type of encounter presented by Table 1 where Lmax correspond to the length of bigger ship.

Table 1. Critical distances where navigators perform anti-collision manoeuvres

Head on (port/port-side) Head on (starboard/strb-side) Overtaking Crossing

Good visibility 2.5 Lmax 5 Lmax 2.5 Lmax 5 Lmax

Restricted visibility 5 Lmax 10 Lmax 5 Lmax 10 Lmax

Stage II: Calculation of the number of encounters of each type.

For the southern part of the Baltic Sea - the overall number of encounters estimated by the simulation model is around 140,000 per year, 30% of them headon situations, 40% crossing and 30% overtaking. This could be done for the Gulf of Trieste, but due to the lack of collision data it would be pointless.

Stage III: Study of the statistical data and evaluation of the intensity of collisions.

For the southern part of the Baltic Sea the mean intensity of collision accidents equals 2.2 per year.

Only open sea area accidents were considered.

Stage IV: Calculation of the probability of collision for a given type of situation.

To simplify the calculations it was assumed that the intensity of a collision is equal in all considered situations. The existing databases of real accident scenarios justifies this assumption: 610

head-on encounter: 0.73 collisions per year, overtaking: 0.73 collisions per year, crossing: 0.73 collisions per year.

The probability of collision for a given type of encounter can be calculated by using the Eq. (1): C Pc = I , (1)

nES where Pc is the probability of collision for a given type of encounter, CI intensity of collision for given type of encounter and nES number of encounters of a given type.

The number of ships navigating on a given route, in cases when vessels have freedom of selecting their speed, route and time of departure can be described as a stationary Poisson distributed stochastic process, where probability of appearance of k ships in one step of simulation equals: λ k e−λ P( X = k ) = , (2)

k! where k is the number of ships navigating on a given route in one step of simulation, and λ the expected number of ships that occur during that one step of simulation.

The probability that in one step no ship will appear on a given route equals:

P ( X = 0 ) = e − λ . (3)

Distance Type of encounter

• • •

If the assumption of freedom of traffic cannot be accepted, the non-homogeneous Poisson process should be used. Let us assume that on the given route in a period T of time I ships appear. Period T is divided into sections (t1, t2], (t2, t3]…(tn-1, tn] with the same length Δt where t1=0 and tn=T. Since the n1, n2, …, nm is the number of ships navigating in a particular section, the total number of ships equals: m

na = ∑ ni . (4)

i =1

Let us assume that the ship traffic flow process derives from NHPP. The rate of number of ships λ(t) is considered as constant in a given section. The average intensity function in a given section is the rate of the number of ships per unit of period, the maximal likelihood estimator is the average rate of number of ships in a section normalized to the length of the section:

λˆ (t ) =

ni , ti −1 < t ≤ ti , for i = 1, 2, ..., m. (5) I (ti − ti −1 )

The associated cumulative distribution is:

Perkovic, M – Gucma, L. – Przywarty, M. – Gucma, M. – Petelin, S. – Vidmar, P.

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 i −1 n j Λˆ (t ) =  ∑  j =1 I

 ni (t − ti −1 ) + . (6)  I (t − t ) , ti −1 < t ≤ ti i i −1 

It should be noted that if there are no accidents in a given interval the cumulative function estimate is constant. The non-parametric confidence interval for such cumulative intensity function can be expressed by the following expression:

ˆ (t ) ˆ (t ) Λ Λ , (7) ˆ (t ) − z ˆ (t ) + z Λ < Λ (t ) < Λ α /2 α /2 I I

for 0<t<T, where zα/2 is the 1-α/2 fractal of the standard normal distribution.

Spatial distribution is one of the main parameters describing traffic flow. It describes the ship’s hull position relative to the axis of the route.

The information about the position of the vessel’s center of gravity, the shape of the waterline and the course are used to define the distribution. It should be noted that different types of distributions are used depending on the type and shape of the waterway like: normal, logarithmic, logistic or triangular distribution.

Here normal distribution with PDF (probability density function) was used:

Following are the results for the traffic analyzed for the complete area (Trieste bay) presented in Fig. 8 and further analyzed for the inbound (green) traffic lane calling at the Port of Koper and the outbound traffic lane (orange) corresponding to the departure passage.

dl ( y ) =

1

σ l 2π

e

−

( y − ml ) 2σ l2

2

, (8)

where y is distance to the axis, m average of ship’s distance to the waterway axis, σ standard deviation of ship’s distance to the waterway axis..

Coordinates of Way Points are modeled using two-dimensional normal distribution. Each coordinate of each Way Point for a ship on a given route is generated separately with the use of the following formula: ( x − µ )2 − 1 2 f ( x) = e 2σ , (9)

σ 2π

Fig. 8. AIS base traffic and two locations for offshore terminal

3.1 Existing Traffic in the Bay Existing traffic for the Port of Koper (which includes traffic destined to Trieste and departing from Trieste) was analyzed by performing 40,000 simulations (800 per year over 5 years). The results are presented in Fig. 9 and Table 2. As illustrated, the departure area is more risky due to the fact that in this direction vessels cross the precautionary area.

where μ is the mean coordinate (latitude, longitude) of Way Point, and σ standard deviation.

Fig. 9. Simulated places of collisions during (800×) 5 years period

3 RESULTS The AIS tracks and LNG locations for the foreseen offshore terminal located at the separation zone are presented in Fig. 8. The figure shows the initial position located inside the precautionary area and then the second location chosen after it was determined that the first was inappropriate.

3.2 Offshore Location The location of an LNG terminal in the center of the existing TSS was also analyzed (Fig. 10 and Table 3). For that purpose it was necessary to widen the TSS. The model demonstrates that the mean time

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between collisions has increased (from 80 to 112.68 years) meaning that expanding the TSS can contribute towards safety at sea.

Table 2. Mean number of accidents, probability of accidents and mean time between accidents Number of Total Expected number Mean time between simulations number of of collisions collisions per year collision [year-1] Complete Area Arrival Departure

800

50

0.0125

80

800 800

13 27

0.00325 0.00675

308 148

instances of Main Engine failures. Statistics for ME failures were used from MEHRA [11]. Fig. 12 shows that some ME failures can result in collisions with the terminal. The LNG terminal is presented as a dot in the separation area (Fig. 11). ME failures inside the safety domain of the LNG terminal have potential for collisions. In that case considering a safety domain with a diameter of 3 NM suggests that only around 15% of total expected failures in the analyzed area can result in a collision. Complete results are presented in Table 3.

3.3 The Koper Area On the basis of AIS measurements the routes of ships have been determined before Koper. Due to a very low number of simulated collisions the risk of collision can be described by the number of encounter situations (Fig. 12). For the Port of Koper the model of grounding was applied with the assumption that the probability of an accident depends on the distance to the given safety contour (or shore).

Fig. 10. Simulated places of collisions in scenario 1 during 1600×5 years period with modified TSS and offshore LNG terminal as an obstruction

Fig. 12. Simulated encounter situations [5 years]

Fig. 11. Simulated places of ME failures in scenario 1 (no increase of traffic) during 1600×5 years period

Further, the location of the LNG terminal was analyzed as an obstruction [9] and [10] relevant to

Numerical results are presented in Table 3.

It is evident that the collision period is seriously lower than for the offshore location. At the same time vessel speeds are slower in the port vicinity, so the consequences can not be the same as for the offshore location where traffic moves at “sea speed”.

By implementing this severity factor the mean time

Table 3. Mean number of accidents, probability of accidents and mean time between accidents

Existing traffic

612

Number of 5 year simulation periods 1600

Total number of collisions 71

Expected number of collisions per year 0.008875

Mean time between collisions [year] 112.68

Total number of ME failures 1343

Perkovic, M – Gucma, L. – Przywarty, M. – Gucma, M. – Petelin, S. – Vidmar, P.

Expected number of ME failures per year 0.167875

Mean time between ME failures 5.96

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between collisions rises to 265 years, compared to offshore at 80 years or 71 collisions if the traffic separation area is extended in width.

Table 4. Estimated collisions and grounding factors Name Probability of collision Collision intensity per year Collision period Severe collision factor Severe collision period Grounding intensity per year Grounding period Severe grounding factor Severe grounding period

Value 0.00002 0.0378 26.5 0.1 265 0.11 9.1 0.1 90

unit

[1/year] [years]

[years] [1/year] [years]

[years]

4 CONCLUSIONS The results presented in Figs. 8 to 10 and Table 1 to 3 provide a general overview of collision risk spots in the analyzed area. The most risk-affected place in the analyzed area is the precaution area near AIS gate xkp. This area should be covered and protected by future VTS with special care. The ships in this area significantly change courses, forcing other ships to predict their maneuvers. New routing measures could be considered in this place in the future to increase navigational safety (e.g., roundabout traffic scheme). But this is standard traffic risk analysis with no bearing on the question of any specific type of vessel unless we consider abnormalities of size and maneuverability.

For the present purposes it is significant is that traffic increases in the ports of Koper and Trieste expected to result from the installation of LNG facilities is approximately 80 LNG vessels per year and therefore no affect on collision analysis is perceptible.

On the contrary, it may even be considered that LNG vessels require high safety standards, and along with that we must consider the introduction of VTS, enhanced traffic lanes arrangements and advanced berthing facilities, extended reporting, improved escorting procedures – so that, almost paradoxically, the introduction of LNG facilities and ships should, in the end, actually increase the safety level in the Gulf of Trieste. On the other hand, although we may feel confident in our results specific to LNG ships and port activities, even the most cursory analysis of traffic in the Gulf of Trieste, including both ports, makes it clear that traffic control is a pressing issue and that each year with the increases in traffic the danger

of a variety of accidents increases. At some point then, although the chances of an LNG carrier being involved in an accident may be slight, a worst case scenario involving this type of ship must be analyzed.

Regarding an offshore terminal located inside the TSS, similar results are obtained in that, logically, a widening of the TSS results in an expected decrease in the chance of collisions; yet the addition of a stable structure in the TSS virtually invents a new possibility for disaster in that main engine failure can lead to a ship colliding with the structure.

5 REFERENCES [1] Vanem, E., Antao, P., Ostvik, I., Del Castillo de Comas, F. (2008). Analyzing the risk of LNG carrier operations.

Reliability Engineering & System Safety, vol. 93, no. 9, p. 1328-1344, DOI:10.1016/j.ress.2007.07.007.

[2] Perkovič, M., David, M., Gucma, L., Przywarty, M. (2008). LNG terminals – cross-border influence, University of Ljubljana, Portorož. (In Slovenian) [3] Risk Assessment LNG import Koper. (2009). Nautical and unloading operations. Report no/DNV Reg No.: /124UI0A-4, Rev 1, Det Norske Veritas, Hovik.

[4] Gucma, L., Perkovic, M., Przywarty, M. (2009).

Assessment of influence of traffic intensity increase on collision probability in the gulf of Trieste. Annual of Navigation, no.15, p. 41-48 [5] Perkovic, M. (2002). AIS and global safety.

Proceedings of 10th International Symposium on Electronics in Transport, Electro-technical Society of Slovenia, Ljubljana.

[6] Gucma, L. (2003). Models of ship’s traffic flow for the safety of marine engineering structures evaluation, Proceedings of European Safety & Reliability Conference, p. 713-718.

[7] Gucma, L., Zalewski, P. (2003). Damage probability of offshore pipelines due to anchoring ships. Polish Maritime Research, vol. 10, no. 4, p. 6-12.

[8] Helcom Report (2006). Report on shipping accidents in the Baltic Sea area for the year 2005. Helsinki Commission Baltic Marine Environment Protection Commission, Helsinki.

[9] MAIB (2005). UK registered ships accident database, Maritime Accident Investigation Branch, Southampton.

[10] Sandia Report (2004). Guidance on Risk Analysis and Safety Implications of a Large Liquefied Natural Gas (LNG) Spill on Water, SAND 6258, Sandia National Laboratories, Albuquerque, New Mexico and Livermore.

(1999).

Establishment of Marine [11] MEHRA Environmental High Risk Areas, Factors Influencing Vessel Risks in UK Waters, app. 6, Maritime and Coastguard Agency, London.

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Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 614-620 DOI:10.5545/sv-jme.2012.696

Paper received: 2012-07-06, paper accepted: 2012-08-28 © 2012 Journal of Mechanical Engineering. All rights reserved.

Influence of Laser Surface Remelting on Al-Si Alloy Properties Sušnik, J. – Šturm, R. – Grum, J. Janez Sušnik – Roman Šturm – Janez Grum*

University of Ljubljana, Faculty of Mechanical Engineering, Slovenia The paper treats Al-Si alloy after laser surface remelting by the changes in microstructure and micro-hardness of a modified surface layer.

Laser remelting of the thin surface layer was carried out with different energy inputs into the as-cast specimen surface. The remelting conditions were varied by application of different laser beam pulse duration. After solidification of surface remelted layer a fine-grained microstructure is formed. Such a microstructure increases micro-hardness of Al-Si alloy by about 60 to 80%. The variation and size of mainly tensile residual stresses in surface remelted layer greatly depends on the cooling rates i.e. laser pulse duration. Precipitation annealing after the process of laser surface remelting additionally influences the reduction of residual stresses.

Keywords: Al-Si alloy, laser surface remelting, residual stresses, microhardeness, precipitation annealing

0 INTRODUCTION Hardening of thin surface layers of aluminium alloys by laser remelting has become valued due to the economy of the process and its applicability, in comparison with other processes, to both small- and large-series manufacture [1]. With a selected laser beam power, the energy input may be adapted to individual needs by changing the degree of defocusing and the laser beam travel rate. In the engineering practice the aluminium alloys are often used for parts which are to be built in, after machining, into functional assemblies; therefore, it is often required that their surfaces and surface layers are of good quality, which is defined as surface integrity. Laser remelting of the thin surface layer is one of the surfacehardening processes which produce higher material hardness levels, and consequently improve wear resistance of machine parts [2] to [4]. With numerous alloys, which show no phase transformations up to the melting temperature, this is the only process for alloy hardening. By selecting a suitable energy input, rapid local heating of the material over the melting temperature could be achieved, followed by rapid cooling and solidification. Thus, a thin remelted layer consists of a solid phase finely distributed in the basic matrix. Properties of the remelted layer are dependent on the chemical composition of the alloy prior to remelting, partly also on the microstructure of alloys prior to the remelting process and on the magnitude of energy input into the sample surface causing thermal stress field [5].

In the technical literature results of numerous investigations on hardening of the thin surface layer by remelting of different aluminium alloys with silicon and other alloying elements may be found. Increasing the hardness of surface could also be achieved by dispersion of hard particles into the material surface as was studied by Anandkumar et al. [6].

614

It has been proved that the efficiency of hardening depends on cooling rates at the liquid/solid boundary and concentrations of alloying elements in the alloy [2] and [5].

Luft et al. [7] also studied aluminium alloys with particular regard to microstructure, i.e. the size and distribution of individual phases. Vollmer and Hornbogen [8] investigated various Al-Si alloys, i.e., two hypoeutectic alloys, one eutectic and one hypereutectic alloy. With different energy inputs, which were ensured by different power densities and different interaction times, hardening of the solid solution of aluminium with individual alloying elements was determined. Conquerelle et al. [4] also investigated Al-Si alloys in terms of the size of crystal grains of the solid solution with different energy inputs with regard to constant power density and different laser beam travel speeds. Wear resistance of these alloys was determined in terms of the size distribution of silicon particles in the solid solution of aluminium and silicon as well as of different degrees of overlapping of laser traces. Wear mechanisms in terms of micro plastic deformation of the soft base and catastrophically failure of silicon crystals were explained by means of micro structural and micro chemical analyses. Rapid heating and cooling, i.e. rapid solidification, result in formation of metastable microstructures which are carriers of important technological properties such as vastly improved material hardness, improved wear resistance and often also improved corrosion resistance [9] to [11].

Rapid solidification is achieved by heat removal into the remaining cold part of the material, which has a much greater mass. The influence of graphite absorber coating thickness during laser surface hardening on a modified layer depth was analyzed by Kek et al. [12].

The optimum coating thickness was determined in a range around 0.03 mm. In their earlier research, Šturm

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

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 614-620

et al. [13] described the influence of alloying elements in Al-Si alloys on the CO2 laser remelting process.

The aim of this study was to define changes in the microstructure of the remelted surface layer after Nd:YAG laser remelting followed by precipitation annealing.

1 EXPERIMENTAL PROCEDURE 1.1 Selection of Al-Si Alloy Table 1 shows the selected Al-Si alloy EN AC48000-K-T6. The microstructure was analyzed in as-delivered state, i.e. as-cast state, and after laser remelting of the thin surface layer.

Table 1. Chemical composition and mechanical properties of EN AC-48000-K-T6 Al-Si alloy a) Chemical composition of AlSi12CuNiMg alloy [wt. %], Al balance Si 12

Fe 0.01

Mn 0.01

Mg 1.04

Cu 0.93

Ni 0.9

Ti 0.01

b) Mechanical properties of AlSi12CuNiMg alloy HV0.2 [MPa] 100

Rm [MPa] 280

Rp0.2 [MPa] 240

A5 [%] 1

The selected alloy AlSi12CuNiMg is hypoeutectic and contains inter-dendritic network of eutectic silicon with different intermetallic compounds which improve alloy micro-hardness. Intermetallic compounds are Mg2Si (black script), Al6Cu3Ni (light-gray script) and Al3Ni (dark-gray script).

The alloy AlSi12CuNiMg shows micro-hardness of around 100 HV0.2. The selected hypoeutectic alloy AlSi12CuNiMg shows tensile strength of around 280 MPa. The microstructure is shown in Fig. 1.

Fig. 1. Microstructure of as-cast AlSi12CuNiMg

1.2 Selection of Laser Remelting Conditions Tests of laser hardening by remelting of the thin surface layer were carried out with a Nd:YAG laser system OR-LASER with laser source maximum power P = 80 W. Based on the previous research, the following laser parameters were selected: Laser source power P = 40 W Mode structure multimode – top hat Pulse duration tb = 4, 6 and 8 ms Pulse frequency ν = 7 Hz Laser beam travel speed vb = 150 mm/min Beam diameter on the specimen surface Db = 1.4 mm Beam overlapping 0.9 mm, 40% of Db 1.3 Preparation of Specimen Surfaces by an Absorbent Absorptivity of the aluminium or its alloys for laser light of Nd:YAG radiation (λ = 1.06µm) is poor (0.07) [14]; therefore, the surface of the aluminium alloys should be suitably prepared prior to laser remelting.

This can be done by surface treatment or by deposition of a suitable absorbent [9] to [13]. Absorptivity is defined as the relationship between the absorbed and the incident energy flows. It also increases with an increase in the energy at the specimen material surface [15]. For our experiment of laser surface remelting of Al-Si alloy colloidal graphite as an absorbent with very high absorbtion coefficient in the range of 60 to 80% [16] was used.

2 EXPERIMENTAL RESULTS 2.1 Size of the Remelted Layer When applying laser hardening by remelting the thin surface layer of the material some basic conditions should be satisfied if a machine part is to operate in a machine or installation efficiently. They are as follows: • Such a high energy input should be ensured that melting of the thin surface layer is guaranteed in the required depth, min. of 0.2 mm.

• A significant size of the remelted surface area is obtained by overlapping of laser remelted traces.

The size of the remelted layer was defined in the cross-section of the remelted specimen. For metallographic analysis, the specimen was cut in the transverse direction, ground, polished and etched for observation with an optical measuring microscope.

The depth of the remelted trace was then measured.

The results of surface modified layer dimensions are

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shown in Table 2 as a function of laser pulse duration.

The greatest depth of the remelted layer hRL, is obtained with the laser pulse duration ti = 8 ms and it amounts of 0.34 mm.

Table 2. Laser modified layer dimensions Laser pulse duration ti [ms] 4 6 8

Depth of laser remelted layer hRL [mm] 233 to 259 241 to 316 274 to 343

Depth of laser modified layer hML [mm] 270 to 334 281 to 378 310 to 419

2.2 Macro and Micro Analysis of the Modified Layer The metallographic preparation of samples and subsequent analysis of the microstructure of the modified surface layer of the studied AlSi12CuNiMg alloy helped us to assess the micro structural changes and to quantify them by measurements.

The as-cast hypo-eutectoid AlSi12CuNiMg consists of different intermetallic compounds, as Al6Cu3Ni and Al3Ni were detected in the microstructure. These intermetallic compounds improve alloy micro-hardness to a level of 100 HV0.2.

Measurements of intermetallic compounds were made according to form factor defined in Equation 1, where parameters of AREA, Dmax and Dmin are defined in Fig. 2.

FORM ⋅ FACTOR =

AREA . (1) π ⋅ ⋅ 4 D max D min

a very fine, homogeneous microstructure. By varying the laser pulse duration, changes of the melting rate are produced, which further produces changes in the morphology of the eutectic and thus also changes in the mechanical and technological properties of given alloys.

Laser remelting of the thin surface layer on the studied AlSi12CuNiMg alloy, using the chosen remelting conditions, creates a very fine and homogeneous microstructure with increased hardness or micro-hardness.

Based on the microstructure analysis, modified layer consists of two zones with different microstructure, i.e. remelted zone, and transition zone between the remelted zone and the base material.

a)

b)

Fig. 3. Dependance of Form factor and cross-section area of intermetallic compounds; a) Al6Cu3Ni, b) Al3Ni

2.2.1 Remelted Zone

Fig. 2. Definition of Form factor parameters

The results of intermetallic compounds measurements are shown in Fig. 3. In both cases it can be seen that the majority of intermetallic compounds were small and round and that only few were long and massive. Laser heating indirectly provokes micro-structural changes in the specimen material. High heating and cooling rates lead to resolidification of the melt in the melt pool, resulting in 616

By laser remelting of the thin surface layer it is possible to achieve local changes in the microstructure of the AlSi12CuNiMg alloy induced by the high heating and cooling rates. High cooling rates of the homogeneous melt gives us an important effect of formation of fine dendrite microstructure. Thus, formation of very fine and homogeneous microstructure of the solution crystals of silicon can be explained by thermo kinetic processes caused by high cooling rates (Fig. 4a).

2.2.2 Transition Zone between the Remelted Zone and Base Material Fig. 4b shows the remelted area and the transition of the remelted zone into the basic microstructure. In the transition zone of the remelted layer into the base material a-dendrites of aluminium and silicon can be

Sušnik, J. – Šturm, R. – Grum, J.

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noted. In the transition of the modified layer extending into the base material, a narrow, non-uniform area has formed with a partially homogeneous microstructure presenting a solid solution rich with silicon.

•

a)

Fig. 5. Micro-hardness measurements through the laser modified surface layer

b)

Fig. 4. Microstructure of the laser surface modified layer of AlSi12CuNiMg alloy; a) remelted zone, b) transition zone

2.3 Micro-hardness in the Remelted Layer Micro-hardness was measured in accordance with the Vickers method, i.e. diamond pyramid hardness measurement, through the depth of the remelted layer. Micro-hardness was measured at a load of 2 N. Fig. 5 shows micro-hardness values through the remelted layer for different laser pulse durations.

From the results shown in Fig. 5 the following may be concluded: • Through-thickness micro-hardness variations in the remelted layers are very similar. The depth with increased micro-hardness is dependent on the energy input; therefore, with longer laser pulse duration a greater depth of the remelted layer is achieved, and consequently also a greater depth of uniform and increased micro-hardness.

•

After remelting the surface layer of alloy AlSi12CuNiMg an average micro-hardness level of around 160 to 180 HV0.2 was obtained.

Micro-hardness gradually reduces in the transition zone from the remelted zone to the base metal.

Hardening of the thin surface layer may be influenced by [3], [10] and [11]: • Finest possible distribution of a-dendrites in the oversaturated solid solution of aluminium and silicon with other alloying elements.

• Finest possible and uniform distribution of intermetallic compounds such as Al6Cu3Ni, Al2Cu and Ni3Al. Micro-hardness was measured in dependence of laser pulse duration. It can be noticed that laser pulse duration has almost no influence on micro-hardness achieved in the remelted zone. In the case of laser remelting with a pulse duration of 8 ms micro-hardness in the surface remelted region is a bit lower than in the case of pulse duration of 4 and 6 ms. Micro-hardness results confirm the difference measured in the depth of the remelted layers when laser operates with 4, 6 or 8 ms long pulses. The longer the laser pulse duration, the greater depth of the remelted zone is achieved.

2.4 Measurements of Residual Stresses Strain measurements and calculations of residual stresses in the surfaced layer were based on the relaxation hole-drilling method in accordance with ASTM standards (ASTM Int., E 837-01, 1995 and ASTM Int., E 837-08, 2008) and were performed by using the CEA-06-062-UM measuring resistance rosettes and the RS-200 Milling Guide device, a product of Vishay Group. The residual-stress variation in the laser-surfaced specimens was obtained by implementing the integral method and the H-drill program [17]. The measurements of residual stresses

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were carried out on all specimens. Due to a relatively wavy and uneven surface of the surfaced layer, the resistance rosettes were glued by applying a twocomponent epoxy resin (Vishay, M-Bond, GA-2).

The graph in Fig. 6 presents the maximum residual stress profiles versus the depth of the modified layer of AlSi12CuNiMg alloy as a function of different laser pulse durations, defined by measurements of strain in a given direction and by calculating the directions of the main axes. From the results in Fig. 6 the following can be concluded: • In all cases of different laser pulse durations residual stresses have a very similar profile differing only in absolute values. In the surface remelted layer, tensile residual stresses were found in a range between 45 and 67 MPa. • The maximum tensile residual stresses were found in the lower part of the remelted layer, or in the transition zone between remelted zone and base metal.

• The maximum tensile residual stresses were achieved in the case of pulse duration of 6 ms, the minimum in the case of pulse duration of 4 ms.

Fig. 6. Residual stresses versus depth of the modified layer after laser surface remelting

Residual stresses are a result of temperature and micro-structural stresses occurring in the specimen material directly after the process of remelting a thin surface layer. During the process of rapid cooling, when the process of solidification goes on, the volume of the sample contracts, resulting in temperature stresses. However, the variation and size of residual stresses in the remelted layer depends also on the composition and homogeneity of the melt and condition of cooling. The cooling conditions are very important since, at higher cooling rates, it is possible to achieve a fine distribution of a-dendrites in the matrix, i.e. in the solid solution of aluminium and silicon.

618

Regarding the nature of the laser remelting process, in which there is a melt pool that mixes due to hydrodynamic and electromagnetic forces, around the laser beam, a rather homogeneous melt, and, after a rather rapid cooling, quite uniform micro-hardness in the remelted specimen layer may be established. The nature of the residual stresses in the surface remelted layer is tensile. The maximum value is around 70 MPa, which is still low enough to select the process of laser surface remelting as apropriate procedure for surface hardness improvement.

2.5 Precipitation Annealing After the process of laser surface remelting, additional heat treatment procedure was carried out. Precipitation annealing of laser surface remelted layer was carried out at a temperature of 150°C for 6 and 12 hours. In Fig. 6 it was found that the lowest residual stresses in the remelted surface layer were obtained in the case of the shortest laser pulse duration of 4 ms. The influence of precipitation annealing on residual stresses in laser surface remelted layer is shown in Fig. 7. It can be noticed that after 6 hours of annealing tensile residual stresses at the surface increased from 64 to 114 MPa. After 12 hours of precipitation annealing, tensile residual stresses at the surface of the remelted layer decrease extensively to a level of 36 MPa. We assume that the quantity and largeness of precipitates, as Al6Cu3Ni, Al2Cu and Ni3Al, in the laser surface remelted layer directly influences on presence of residual stresses. The influence of laser pulse duration, i.e. the remelting time, after 12 hours of precipitation annealing on residual stresses is shown in Fig. 8. In the remelted surface area tensile residual stresses decrease, when laser pulse duration is increased. In the case of pulse duration of 8 ms residual stresses at the surface get even compressive character of –9 MPa. Maximum tensile residual stresses of 81 MPa were observed in the transition zone in the depth of 0.4 mm. Such a big difference from compressive to tensile residual stresses in a very small dimension of 0.4 mm is not favorable course of residual stresses through thin surface remelted layer.

In addition to residual stress measurements the precipitation process was evaluated also with microhardness measurement through the surface remelted layer (Fig. 9). In the case of laser pulse duration of 4 ms micro-hardness in the remelted layer reduced for about 10 HV0.2 when we performed precipitation annealing for 6 hours. In the case of precipitation annealing for 12 hours micro-hardness increased with

Sušnik, J. – Šturm, R. – Grum, J.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 614-620

respect to initial remelted microstructure in a range of 5 to 20 HV0.2.

Fig. 7. Residual stresses versus depth of the modified layer after laser surface remelting and additional precipitation annealing at 150 °C

hours at 150 °C. The main difference is in comparison to initial remelted microstructure. The new annealed microstructure indicated higher micro-hardness as it was in the initial remelted microstructure, for 20 HV0.2, and reached level of 180 to 200 HV0.2.

Fig. 10. Micro-hardness versus depth of the modified layer after laser surface remelting and additional precipitation annealing at 150 °C at different laser pulse durations

3 CONCLUSIONS

Fig. 8. Residual stresses versus depth of the modified layer after laser surface remelting and additional precipitation annealing at 150 °C at different laser pulse durations

Fig. 9. Micro-hardness versus depth of the modified layer after laser surface remelting and additional precipitation annealing at 150 °C

The influence of laser pulse duration, i.e. the remelting time, after 12 hours of precipitation annealing on micro-hardness is shown in Fig. 10.

It can be noticed that laser pulse duration had no significant effect on micro-hardness of laser surface remelted layer after precipitation annealing of 12

Traveling of laser light across the sample surface produces rapid heating and, consequently, melting and further solidification of the material. In addition to material hardening due to rapid solidification, there are several other advantages of the process, i.e. it is a clean, simple, practical and cheap process in comparison to the other processes of thermal surface treatment. The following findings confirm the efficiency of laser surface remelting process of AlSi12CuNiMg alloy: • After solidification of surface remelted layer a fine-grained microstructure was formed.

• With the given remelting conditions the depth of the remelted layer hRL varied between 0.27 and 0.42 mm.

• The process of laser remelting caused improvement of micro-hardness in the remelted surface layer. High quantities of Si in the alloy which were additionally modified with other alloying elements causing the micro-hardness increase from 100 HV0.2 in base alloy to 160 to 180 HV0.2 in laser surface remelted layer. Such a micro-hardness increase of the remelted layer represents 60 to 80% increase in micro-hardness regarding to the base alloy.

• Measurements of residual stresses were made on the specimen on which laser beam was led at a 40 % overlapping of the remelted layer. After laser surface remelting, tensile residual stresses of a magnitude of 45 to 105 MPa were identified

Influence of Laser Surface Remelting on Al-Si Alloy Properties

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Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, 614-620

into a depth of 0.25 mm. In the depth greater than 0.25 mm from the sample surface, tensile residual stresses gradually decreased. The variation and size of residual stresses depends on the cooling rates or time necessary for the remelted layer to cool down to the ambient temperature.

• High cooling down rates in the remelted material caused formation of small a-dendrites, which influenced formation of higher internal stresses of tensile character.

• Additional precipitation annealing of 12 hours at 150 °C caused a reduction of tensile residual stresses at the surface of the remelted layer to level around zero. At the same time an increase of micro-hardness of 20 HV0.2 occurred. The reason for this lies in a precipitation growth of intermetallic compounds in the remelted microstructure.

In the future, a detailed analysis of the precipitated intermetallic compounds in the laser surface remelted layer shall be undertaken to characterize the dependence between quantity, distribution and largeness of precipitates, as Al6Cu3Ni, Al2Cu and Ni3Al, on residual stresses and micro-hardness.

4 REFERENCES [1] Antona, P.L., Appiano, S., Moschini, G. (1987). Laser surface remelting and alloying of aluminium alloys.

Laser Treatment of Materials, Mordike, B.L. (ed.), DEM Informationsgeseelschaff Verlag, Oberursel, p.

133-145.

[2] Pinto, M.A., Cheung, N., Ierardi, M.C.F., Garcia, A. (2003). Microstructural and Hardness Investigation of an Aluminum-copper Alloy Processed by Laser Surface Melting. Materials Characterization, vol. 50, p. 249253, DOI:10.1016/S1044-5803(03)00091-3.

[3] Tomida, S., Nakata, K., Shibata, S., Zenkouji, I., Saji, S. (2003). Improvement in wear resistance of hypereutectic Al-Si cast alloy by laser surface remelting.

Surface and Coatings Technology, vol. 167-170, p.

468-471, DOI:10.1016/S0257-8972(03)00100-2.

[4] Conquerelle, G., Fachinetti, J.L. (1987). Friction and Wear of Laser Treated Aluminium-Silicon alloys. Laser Treatment of Materials, Mordike, B.L. (ed.), DEM Informationsgeseelschaff Verlag, Oberursel, p. 171178.

[5] Yilbas, B.S., Arif, A.F.M., Karats, C., Raza, K. (2009).

Laser treatment of aluminum surface: Analysis of thermal stress field in the irradiated region. Journal of Materials Processing Technology, vol. 209, p. 77-88, DOI:10.1016/j.jmatprotec.2008.01.047.

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[6] Anandkumar, R., Almeida, A., Colaço, R., Vilar, R., Ocelik, V., De Hosson, J. Th. M. (2007). Microstructure and wear studies of laser clad Al-Si/SiC(p) composite coatings. Surface & Coatings Technology, vol. 201, p.

9497-9505, DOI:10.1016/j.surfcoat.2007.04.003.

[7] Luft, U., Bergmann, W., Mordike, G.L. (1987).

Laser Surface melting of aluminium alloys. Laser Treatment of Materials, Mordike, B.L. (ed.), DEM Informationsgeseelschaff Verlag, Oberursel, p. 147163.

[8] Vollmer, H., Hornbogen, E. (1987). Microstructure of laser treated Al-Si alloys. Laser Treatment of Materials, Mordike, B.L. (ed.), DEM Informationsgeseelschaff Verlag, Oberursel, p. 164-170.

[9] Wong, T.T., Liang, G.Y. (1997). Effect of Laser Melting Treatment on the Structure and Corrosion Behaviour of Aluminium and Al-Si Alloys. Journal of Materials Processing Technology, vol. 63, p. 930-934, DOI:10.1016/S0924-0136(96)00098-2.

[10] Chong, P.H., Liu, Z., Skeldon, P., Thompson, G.E. (2003). Large area surface treatment of aluminium alloys for pitting corrosion protection. Applied Surface Science, vol. 208-209, p. 399-404, DOI:10.1016/ S0169-4332(02)01418-6.

[11] Osorio, W.R., Cheung, N., Spinelli, J.E., Cruz, K.S., Garcia, A. (2008). Microstructural modification by laser surface remelting and its effect on the corrosion resistance of an Al-9 wt%Si casting alloy.

Applied Surface Science, vol. 254, p. 2763-2770, DOI:10.1016/j.apsusc.2007.10.013.

[12] Kek, T., Grum, J. (2010). Influence of the graphite absorber during laser surface hardening. Strojniški vestnik - Journal of Mechanical Engineering, vol. 56, no. 2, p. 150-157.

[13] Šturm, R., Grum, J., Božič, S. (2012). Influence of the alloying elements in Al-Si alloys on the laser remelting process, Lasers in Engineering, vol. 22, no. 1-2, p. 4761.

[14] Anandkumar, R., Almeida, A., Vilar, R., Ocelik, V., De Hosson, J. Th. M. (2009). Influence of powder particle injection velocity on the microstructure of Al–12Si/SiC(p) coatings produced by laser cladding.

Surface & Coatings Technology, vol. 204, p. 285-290, DOI:10.1016/j.surfcoat.2009.07.025.

[15] Pierron, N., Sallamand, P., Mattei, S. (2007). Study of Magnesium and Aluminum Alloys Absorption Coefficient during Nd:YAG laser interaction.

Applied Surface Science, vol. 253, p. 3208-3214, DOI:10.1016/j.apsusc.2006.07.035.

[16] Schuöcker, D. (1998), Handbook of the Eurolaser Academy, Chapman & Hall, London.

[17] Schajer, G.S. (1988). Measurement of non-uniform residual stresses using the hole drilling method. Journal of Engineering Materials and Technology, vol. 110, no.

4, p. 338-349, DOI:10.1115/1.3226059.

Sušnik, J. – Šturm, R. – Grum, J.

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10 Vsebina

Vsebina Strojniški vestnik - Journal of Mechanical Engineering letnik 58, (2012), številka 10 Ljubljana, oktober 2012 ISSN 0039-2480 Izhaja mesečno

Razširjeni povzetki člankov Matej Rajh, Srečko Glodež, Jože Flašker, Karl Gotlih: Gibljivost haptičnega mehanizma v cilindričnem prostoru MR-skenerja David Zaremba, Christian Biskup, Thomas Heber, Nico Weckend, Werner Hufenbach, Frank Adam, Friedrich-Wilhelm Bach, Thomas Hassel: Priprava z ogljikovimi vlakni ojačene plastike za popravila z obdelavo stopničastega perifernega območja Xiwen Zhang, Xiaodong Wang, Yi Luo: Izboljšana momentna metoda za nadzor predobremenitve pri natančni montaži miniaturnih vijačnih zvez Arif Gok, Cevdet Gologlu, Ibrahim H. Demirci, Mustafa Kurt: Opredelitev kakovosti obdelanih neravnih površin s pomočjo zvočnega tlaka Roman Kunič: Vrednotenje vpliva pospešenega staranja na življenjsko dobo vakuumskoizolacijskih panelov Marko Perkovic, Lucjan Gucma, Marcin Przywarty, Maciej Gucma, Stojan Petelin, Peter Vidmar: Ocena navtičnih tveganj za operacije UZP v koprskem pristanišču Janez Sušnik, Roman Šturm, Janez Grum: Vpliv parametrov laserskega površinskega pretaljevanja na lastnosti Al-Si zlitine Osebne vesti Diplomske naloge

SI 117 SI 118 SI 119 SI 120 SI 121 SI 122 SI 123 SI 124

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 117

Prejeto: 2009-07-13, sprejeto: 2012-07-12 © 2012 Strojniški vestnik. Vse pravice pridržane.

Gibljivost haptičnega mehanizma v cilindričnem prostoru MR-skenerja Rajh, M. – Glodež, S. – Flašker, J. – Gotlih, K. Matej Rajh1,* – Srečko Glodež2 – Jože Flašker3 – Karl Gotlih3

Rajh Plus d.o.o., Slovenija v Mariboru, Fakulteta za naravoslovje in matematiko, Slovenija 3 Univerza v Mariboru, Fakulteta za strojništvo, Slovenija 1

2 Univerza

Haptični vmesnik je naprava, ki s povratno silo uporabniku omogoča vključenost v navidezno ali oddaljeno okolje.

Med razvojem MR-združljivega haptičnega mehanizma smo zaznali nekatere probleme in pomembne lastnosti, skupne vsem robotskim napravam, ki delujejo v MR-tunelu. Namen prispevka je prikaz razvoja in kinematične optimizacije haptičnega mehanizma s tremi prostostnimi stopnjami ter predstavitev 3D-vizualizacijske metode za analizo gibljivosti mehanizmov v omejenem prostoru. Glavni cilj dela je, na osnovi izsledkov raziskave, konstruiranje in izdelava novega fMRI-združljivega haptičnega mehanizma s tremi prostostnimi stopnjami, gnanega z elektromotorji, ki so kot aktivni mehanski elementi nameščeni zunaj kritičnega območja glede na gostoto magnetnega polja.

Haptična naprava mora imeti odlično gibljivost v vseh smereh za izpolnjevanje operaterjevih zahtev. Ta zahteva je še posebej pomembna v omejenem prostoru, kot je MR-tunel. Zato smo oblikovali kinematični model, ki omogoča reševanje zaznanih problemov in posledično izdelavo mehanizma z enakomerno porazdeljenimi silami in hitrostmi. Rešitev omogoča 3D-grafični prikaz za vsako točko delovnega prostora mehanizma, z vnaprej izbrano ločljivostjo. Presek med omejenim območjem in celotnim delovnim prostorom prikaže zmožnost izvajanja gibov. S tem modelom je mogoče predstaviti in izboljšati gibljivostne lastnosti mehanizmov v tunelu MR-skenerja.

Za optimalno določanje pozicije robota ni dovolj, če poznamo samo delovni prostor robota, saj obstajajo tudi mehanske omejitve. Te izvirajo iz kinematike mehanizma oz. matematične posebnosti singularnosti Jacobijeve matrike, ki se izraža v zmanjšani gibljivosti mehanizma oziroma kinematični singularnosti, ko mehanizem zaradi mehanskih lastnosti ne more v celoti izpolnjevati naloge. Zato je pomembno, da se singularnostim izognemo že pri načrtovanju mehanizma.

Gibljivost je definirana kot merilo delovanja robotskega mehanizma in opisuje pogoje transformacije vektorja hitrosti iz notranjih v zunanje koordinate. Kinematična obravnava mehanizma je izvedena s pomočjo DenavitHartenbergovih parametrov in Jacobijeve matrike, ki je osnova za analizo gibljivosti z razvitim računalniškim modelom. V modelu je delovni prostor mehanizma diskretiziran in ga predstavimo z ekvidistantno mrežo točk, ki definirajo gibljivost mehanizma na osnovi najkrajše osi hitrostnega elipsoida. Za predstavitev rezultatov je uporabljen volumetrični prikaz indeksov gibljivosti. Pri iskanju najboljše pozicije za izvajanje naloge znotraj delovnega prostora algoritem premika prostor naloge skozi skupino diskretnih točk v delovnem prostoru. Izmed vseh možnih položajev izbere v prostoru naloge tistega, ki ima najvišje število točk z ustrezno gibljivostjo.

Haptične naprave za delo v MR-okolju imajo zaradi magnetnega polja in oblike skenerja določene zahteve glede zgradbe in delovanja. Nekateri viri obravnavajo omenjeno problematiko, vendar ne vključujejo gibljivosti MR-mehanizma kot ene izmed ključnih lastnosti za kakovosten prenos sile v haptičnih napravah, kjer je treba upoštevati tudi omejitev prostora. Prav to pa lahko pomembno vpliva na konstruiranje mehanizma, saj bi delovanje v bližini singularne točke povzročalo resne obratovalne probleme.

Vsaka naloga gibanja človeške roke in s tem vrha mehanizma ima svoje karakteristike in zahteva drugačen položaj v tunelu. Razviti program omogoča prikazovanje delovnega prostora mehanizma in pripadajočih indeksov gibljivosti v odprtem prostoru ali v vnaprej definiranem omejenem prostoru. Na osnovi 3D-grafične predstavitve je možno ocenjevati lastnosti gibljivosti v različnih položajih. S tem je uporabniku omogočena prilagoditev položaja baze mehanizma skladno z zahtevami po višjem indeksu gibljivosti. Rezultat tega sta lažje premikanje mehanizma in boljši prenos mehanske energije od vrha mehanizma, ki je v stiku z uporabnikom, do motorjev, kar izboljša haptičnost in vključenost v navidezno okolje. Ta program pa je uporaben tudi kot konstrukcijski pripomoček za off-line načrtovanje robotskih proizvodnih celic in haptičnih robotov, saj omogoča boljši vpogled v uporabni delovni prostor in posledično boljši izkoristek.

Ključne besede: gibljivost, Jacobijeva matrika, delovni prostor, haptični mehanizem, MR-združljivost, omejen prostor *Naslov avtorja za dopisovanje: Rajh Plus d.o.o., Črešnjevec 143, 2310 Slovenska Bistrica, Slovenija, matej.rajh@rajh.eu

SI 117

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 118

Prejeto: 2012-01-05, sprejeto: 2012-08-23 © 2012 Strojniški vestnik. Vse pravice pridržane.

Priprava z ogljikovimi vlakni ojačene plastike za popravila z obdelavo stopničastega perifernega območja

Zaremba, D. – Biskup, C. – Heber, T. – Weckend, N. – Hufenbach, W. – Adam, F. – Bach, Fr.-W. – Hassel, T. David Zaremba1,* – Christian Biskup1 – Thomas Heber2 – Nico Weckend2 – Werner Hufenbach2 – Frank Adam2 – Friedrich-Wilhelm Bach1 – Thomas Hassel1 1 Leibnizova 2 Tehniška

univerza v Hannovru, Nemčija univerza v Dresdnu, Nemčija

Današnji standardni postopki za popravilo plastike, ojačene z vlakni, niso optimizirani z ozirom na strukturno porazdelitev lastnosti izvirnega materiala. Glavni razlog za to je v dejstvu, da so posamezne mehanske lastnosti takih materialov odvisne od velikosti in usmeritve materiala v vlaknih, ki ga je treba obnoviti. Razen stopničastega prekrivnega spoja in krpanja je obravnavana tudi nova metoda za ponovljivo in samodejno pripravo materiala za popravila.

Metoda je zasnovana na ponovljivi izdelavi stopničastega perifernega območja v materialu, ki ustreza strukturnim lastnostim osnovnega laminata. Del tega postopka vključuje tudi inkrementalno izdelavo stopničastega dela in razkrivanje vlaken v prehodnem območju med osnovnim laminatom in krpo, ki zahteva primerno tehniko odstranjevanja materiala. Odstranjevanje materiala je bilo opravljeno z različnimi postopki, ki dajejo odlične možnosti uporabe pri obdelavi površin: s suhim ledom, s snegom in s čistim vodnim curkom. Referenčni material je bila plastika, ojačena z ogljikovimi vlakni, ki je primerljiva z laminati v letalski industriji.

V eksperimentalnem okolju je bilo ugotovljeno, da obdelava s snegom in suhim ledom ni primerna za pripravo materiala za popravila, medtem ko pa omogoča obdelava z vodnim curkom natančen nadzor nad odstranjevanjem materiala. Kombinacija majhnega delovnega območja vodnega curka in natančnega pozicioniranja vodilne naprave omogoča natančen nadzor nad odstranjevanjem materiala. Postopek zaradi svoje stabilnosti omogoča tudi razmeroma zvezno obliko curka. Posebna prednost je tudi v visoki selektivnosti procesa, ki omogoča odstranjevanje posameznih plasti kompozitnega materiala z ortogonalno konfiguracijo. Na osnovi teh obetavnih rezultatov je bila opravljena parametrična študija za ugotavljanje parametrov priprave površine plastike, ojačene z ogljikovimi vlakni.

Med obravnavanimi parametri so bili oddaljenost šobe, tlak vode in hitrost podajanja orodja. Vpliv pomembnih procesnih parametrov je bil preučen predvsem z ozirom na dva glavna kriterija – globino odnašanja in poškodbe. Končno je bil določen tudi nabor parametrov za odstranjevanje posameznih plasti materiala. Po uspešni obdelavi stopničastega perifernega območja so bila opravljena tudi prva preizkusna popravila. Stopničasto periferno območje je bilo v ta namen napolnjeno s krpo iz usmerjenega mehkega preprega. Za utrjevanje je bil uporabljen postopek vročega prešanja. Skupaj z omenjenimi laminati so bili pripravljeni tudi referenčni vzorci, ki predstavljajo idealiziran popravljeni material. Stopničasta periferna cona je bila pred procesom utrjevanja ročno vrezana v vsako plast preprega teh laminatov. Pred utrjevanjem so bile vstavljene tudi krpe preprega za popravilo.

Utrjevanje osnovnega laminata in krpe za popravilo je potekalo sočasno. Nato so bili izdelani vzorci za natezni preizkus iz laminata, popravljenega s pomočjo vodnega curka, in iz idealiziranega laminata.

Natezni preizkusi so pokazali uporabnost čistega vodnega curka za pripravo plastike, ojačene z ogljikovimi vlakni, za popravila.

Ključne besede: priprava površin, z ogljikovimi vlakni ojačena plastika, CFRP, vodni curek, obdelava s suhim ledom, obdelava s snegom, popravilo z ogljikovimi vlakni ojačene plastike, stopničasto periferno območje

SI 118

*Naslov avtorja za dopisovanje: Leibnizova univerza v Hannovru, An der Universität 2, D-30823 Garbsen, Nemčija, zaremba@iw.uni-hannover.de

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 119

Prejeto: 2012-04-20, sprejeto: 2012-08-21 © 2012 Strojniški vestnik. Vse pravice pridržane.

Izboljšana momentna metoda za nadzor predobremenitve pri natančni montaži miniaturnih vijačnih zvez Zhang, X. – Wang, X. – Luo, Y. Xiwen Zhang – Xiaodong Wang* – Yi Luo

Laboratorij za mikro/nanotehnologije in sisteme province Liaoning, Tehniška univerza v Dalianu, Kitajska

Pri montaži miniaturnih preciznih komponent lahko pride do poslabšanja zmogljivosti zaradi neenakomernih napetosti, nastalih zaradi raztrosa predobremenitve miniaturnih vijačnih zvez. Pri montažnem procesu je zato pomembna natančnost predobremenitve, ki zagotavlja kakovost in zanesljivost vijačnih zvez. Večina obstoječih metod pa ni primerna za natančen nadzor nad zategovanjem miniaturnih vijačnih zvez zaradi njihove majhne velikosti in nizke ravni predobremenitev.

Najpogostejši način nadzora nad zategovanjem vijakov je z določanjem vrednosti momenta. Razmerje med momentom in predobremenitvijo pri miniaturnih navojnih zvezah pa je močno odvisno od trenja kontaktnih površin, in celo zmerna variabilnost ali netočnost pri določanju zateznega momenta pomembno vpliva na predobremenitev.

Zato je podan predlog izboljšane momentne metode na osnovi matematičnega modela za nadzor predobremenitve pri precizni montaži miniaturnih vijačnih zvez. Model odvisnosti med momentom in predobremenitvijo je bil razvit za napovedovanje zateznega momenta ter upošteva togost sistema, kot vrtenja matice, predobremenitev in momentni gradient. Pri zategovanju vijačnih zvez obstaja linearna povezava med kontrolnim momentom in momentnim gradientom. Momentni gradient ima linearen potek s koeficientom momenta, na katerega močno vplivata površinska hrapavost in mazanje ter se izračunava v realnem času iz zateznega momenta. Predlagani matematični model je zato mogoče uporabiti za napovedovanje s trenjem kompenziranega kontrolnega momenta in izboljšana momentna metoda bi zmanjšala vpliv trenja na predobremenitev ter raztros predobremenitve vijakov zaradi tornih spremenljivk vijakov.

Razvit je bil eksperimentalni sistem za samodejno montažo miniaturnih vijakov premera pod 2,5 mm.

Sistem je bil razvit posebej za dinamičen nadzor nad procesom zategovanja s trajnim napovedovanjem zateznega momenta iz izračunanega momentnega gradienta. Za vrednotenje momenta in predobremenitve med zategovanjem vijakov sta bila uporabljena senzorja momenta in sile. Obremenjevanje je bilo izvedeno s koračnim motorjem, za linearno podajanje pa je bila uporabljena linearna pogonska stopnja. Ključ zateguje matico in linearna pogonska stopnja izvaja med obračanjem ključa sinhrono linearno gibanje po osi vijaka. Za izračunavanje momentnega gradienta in krmiljenje procesa montaže vijačnih zvez je bila razvita programska oprema v okolju LabVIEW. Togost sestavljenih delov je bila ugotovljena eksperimentalno.

Poskusi montaže vijačnih zvez so bili izvedeni z eksperimentalnim sistemom. Izvedene so bile tri vrste poskusov montaže miniaturnih vijakov dimenzije M1,4, M1,6 in M2. Pri poskusnih montažah sta bili uporabljeni izboljšana momentna metoda in momentna metoda. Zategovanje vijakov je potekalo s fiksno hitrostjo vrtenja matic, ciljne predobremenitve pa so bile izbrane tako, da so bili miniaturni vijaki v elastičnem območju, kjer je odvisnost med predobremenitvijo in vrtilnim kotom matice linearna. Rezultati poskusov kažejo, da je raztros napovedanega momenta večji od dejanske predobremenitve vijakov, izboljšana momentna metoda pa daje manjšo porazdelitev predobremenitev in manjši raztros okrog srednje vrednosti predobremenitve vijakov. Rezultati tudi kažejo, da premer vijakov vpliva na raztros izboljšane momentne metode. Pri enaki ravni predobremenitve je ob večanju premera vijakov predobremenitev v primerjavi z zateznim momentom razmeroma nižja in lahko povzroči večji raztros predobremenitve. Raztros predobremenitve je bil manjši od ±13,00 % v elastičnem območju, kar je doslednejši rezultat kot pri običajni momentni metodi. Napaka nadzora predobremenitve je bila pri večini vijakov manjša od 10 %.

Izboljšana momentna metoda je priročna in ne zahteva posebne opreme, zato se izkaže kot zelo praktična za nadzor predobremenitve pri precizni montaži miniaturnih navojnih zvez. Če pa dimenzija vijaka presega M2, je za boljši nadzor predobremenitve mogoče uporabiti tudi druge postopke, kot je postopek nadzora kota, postopek indikacije napetosti ali celo ultrazvočni postopek.

Ključne besede: nadzor predobremenitve, precizna montaža, miniaturne vijačne zveze, momentni gradient, elastično območje, kompenzacija trenja

*Naslov avtorja za dopisovanje: Laboratorij za mikro/nanotehnologije in sisteme province Liaoning, Tehniška univerza v Dalianu, Dalian, Kitajska, xdwang@dlut.edu.cn

SI 119

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 120

Prejeto: 2012-02-13, sprejeto: 2012-06-29 © 2012 Strojniški vestnik. Vse pravice pridržane.

Opredelitev kakovosti obdelanih neravnih površin s pomočjo zvočnega tlaka Gok, A. – Gologlu, C. – Demirci, I.H. – Kurt, M. Arif Gok1 – Cevdet Gologlu2,* – Ibrahim H. Demirci2 – Mustafa Kurt3 1 Univerza

Kastamonu, Visoka poklicna šola, Turčija Karabuk, Fakulteta za strojništvo, Turčija 3 Univerza Marmara, Tehniška fakulteta, Turčija

2 Univerza

Namen te študije je preučitev odvisnosti med ustvarjenimi zvočnimi signali ter površinsko hrapavostjo po obdelavi konveksnih in konkavnih neravnih površin.

Kontrola kakovosti površine delov na proizvodni liniji je draga in težavna naloga. V študiji so bile za operacije srednje fine obdelave uporabljene strategije istosmernega in protismernega rezkanja s konturno in radialnoaksialno potjo orodja. Opazovane so bile odvisnosti med površinsko hrapavostjo in ravnjo zvočnega tlaka pri odrezavanju z različnimi rezalnimi hitrostmi, podajanjem na obrat in korakom podajanja. Za obdelavo oblikovanih neravnih površin so bile uporabljeni steblasti rezkarji z obračalnimi ploščicami s prevlekami TiC, TiN in TiAlN. Za kombinacijo rezalnih parametrov je bilo izbrano standardno ortogonalno polje L'16 s po štirimi različnimi ravnmi za vsak parameter. Pri eksperimentih je bil uporabljen merilni instrument MahrSurf PS1 v smeri pravokotno na pot orodja pod kotom 45° glede na stojni položaj, s čimer je bil pri vsakem vzorcu upoštevan efektivni premer odrezavanja. Za digitalizacijo in zajem zvočnega tlaka iz zvočnega senzorja je bil uporabljen algoritem, napisan v MATLAB-u. Realne vrednosti zvočnega tlaka so bile določene z vrednotenjem razlik, ki jih najde algoritem.

Za ugotavljanje odvisnosti med ravnjo zvočnega tlaka in povprečno površinsko hrapavostjo je bila uporabljena linearna regresijska analiza.

Dejstvo, da so bile pri obdelavi konkavnih površin dobljene manjše vrednosti ravni zvočnega tlaka kot pri obdelavi konveksnih površin, je mogoče razložiti s tem, da konveksna oblika površine omogoča mikrofonu manj oviran zajem zvoka. Pri obdelavi konveksnih površin z različnimi prevlekami sta bili doseženi višji ravni zvočnega tlaka 111,9 dB in 113,4 dB. Razlog je v tem, da ima steblasti rezkar med obdelavo manj stika z obdelovancem in da se pojavi mehanizem vibracij.

Raven zvočnega tlaka je bila manjša pri konkavni površini, ker je orodje delovalo znotraj obdelovanca in je bilo z njim v stiku z daljšim rezalnim robom. Vrednotenje zvočnega tlaka in površinske hrapavosti z ozirom na parametra podajanje na obrat in širina podajanja je pokazalo povečanje vrednosti zvočnega tlaka in Ra pri povišanih rezalnih parametrih, ne glede na prevleko orodja in obliko površine. Eksperimentalna raziskava vpliva orodnih prevlek na vrednosti Ra je pokazala, da daje prevleka TiAlN zelo dobre rezultate pri majhni rezalni hitrosti.

Večje vrednosti zvočnega tlaka in Ra so bile ugotovljene pri protismernem rezkanju konkavne površine, kar je mogoče pojasniti z vplivom vibracij. Vrednosti zvočnega tlaka in Ra za konveksne površine so pri istosmernem rezkanju večje kot pri protismernem rezkanju. Prekrivanje rezalnih robov orodja na obdelovancu povečuje zvočni tlak in površinsko hrapavost zaradi protismernega rezkanja in konveksne oblike obdelovanca.

Linearna regresijska analiza je pokazala pozitivno, linearno in statistično signifikantno odvisnost med ravnjo zvočnega tlaka in površinsko hrapavostjo pri konkavnih in konveksnih površinah za TiC, TiN in TiAlN (R2 = 0,875, 0,822 in 0,873 pri konkavnih površinah; oz. R2 = 0,888; 0,899 in 0,916 pri konveksnih površinah). Ob upoštevanju vseh rezalnih parametrov je bilo ugotovljeno, da je koeficient korelacije med ravnjo zvočnega tlaka in površinsko hrapavostjo boljši od R2 = 0,8. Iz tega sledi, da je smiselno spremljati raven zvočnega tlaka med obdelavo.

Prejšnje študije so bile osredotočene na ocenjevanje obrabe orodja, vibracijske stabilnosti in najboljših rezalnih parametrov s pomočjo akustične emisije. V tej študiji so bili izvedeni eksperimenti z materialom EN X40CrMoV5-1, ki poglabljajo razumevanje odvisnosti med ustvarjenim zvočnim signalom in površinsko hrapavostjo pri obdelavi neravnih konveksnih in konkavnih površin.

Ključne besede: steblasto rezkanje, akustična emisija in zvočni tlak, površinska hrapavost, linearna regresija

SI 120

*Naslov avtorja za dopisovanje: Univerza Karabuk, Fakulteta za strojništvo, 78050, Karabuk, Turčija, cgologlu@karabuk.edu.tr

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 121

Prejeto: 2012-04-20, sprejeto: 2012-08-21 © 2012 Strojniški vestnik. Vse pravice pridržane.

Vrednotenje vpliva pospešenega staranja na življenjsko dobo vakuumskoizolacijskih panelov Kunič, R. Roman Kunič

Univerza v Ljubljani, Fakulteta za gradbeništvo in geodezijo, Slovenija

Vakuumskoizolacijski paneli (VIP) so toplotnoizolacijski paneli, ki so bistveno učinkovitejši – tudi do desetkrat – kot do sedaj poznani in razširjeni toplotnoizolacijski materiali. Toplotna prevodnost lahko znaša samo 0,003 W/(m·K) ali 3 mW/(m·K), toplotna upornost samo 20 mm debelega panela pa celo 6,66 (m2·K)/W. S pomočjo napredne in inovativne tehnologije tako dosegamo izjemo toplotno izolativnost, ne da bi bistveno povečali debelino plasti, ki ovira prehod toplote.

VIP-paneli se že s pridom uporabljajo za ustvarjanje vsakršnih toplotnih ovir, tako v industriji (bela tehnika, laboratorijska in medicinska oprema, avtomati za pijače in hrano, naprave za razne proizvodne in druge procese), v transportu in trgovini (tovornjaki hladilniki, hladilni kontejnerji, hladilne torbe za zdravila ali hrano, mobilni domovi, prikolice in druga vozila), kot tudi v gradbeništvu (zunanje stene, terase, ravne strehe, posebne panelne plošče, izolacija sistemov za hlajenje in prezračevanje), ter povsod tam, kjer je prostor dragocen.

Namen pospešenega staranja je pridobiti podatke o tistih materialih, ki bi jih uporabili zato, da bi dosegli pričakovano življenjsko dobo izdelka ali sistema. Čas testiranja v primeru pospešenega staranja je v primerjavi z izpostavitvijo normalnim pogojem uporabe znatno krajši. Metoda je posebej primerna za določanje življenjske dobe izdelkov v primerih, ko so podatki o pričakovanih življenjskih dobah nedostopni ali se jih ne izplača pridobiti s pomočjo testiranja. Razlog je enostaven: čas nam ne dopušča, da bi merili celotno življenjsko dobo.

Najpomembnejši obremenitvi VIP-panela, če seveda izvzamemo mehanske poškodbe, sta povišana temperatura in vlaga. Ti dve obremenitvi, posamič ali hkratno, povzročata ter ob višji intenzivnosti povečujeta prehod plinov in vlage v sam VIP-panel. Ob tem procesu se povečuje toplotna prevodnost, ki limitira k vrednosti za samo jedro.

Glavni cilj raziskave je bil opredeliti podporo teoretičnim osnovam Arrheniuosovega zakona pospešenega staranja v povezavi z življenjsko dobo VIP-panelov. Zanimajo nas vrednosti toplotne prevodnosti v odvisnosti od temperature in časa obremenitve. S pomočjo Arrheniusovega zakona smo določili mehanizem staranja VIPpanelov. Določitev ustrezne temperaturne obremenitve, časa trajanja in obremenitve predstavlja velik izziv, saj zlahka presežemo temperaturo in čas testiranja, ko reverzibilni postopki niso več mogoči zaradi trajnih poškodb vzorca. Ker je večina zaščitnih folij nestabilna pri temperaturah med 105 do 110 °C, ko nastopijo tudi trajne poškodbe, smo izbrali naslednje temperaturne obremenitve: 100, 90, 80, 70 in 60 °C, ter čas izpostavljenosti od najmanj pol dneva do treh mesecev, v nekaterih primerih pa celo dlje.

Upravičeno lahko menimo, da so VIP-paneli trajen in kakovosten izdelek, tudi z ozirom na trajnost ohranjanja izredne toplotne izolativnosti, kar potrjujemo tudi z laboratorijskimi meritvami, lastnimi znanstvenimi dognanji in računskimi analizami ob upoštevanju Arrheniusovega zakona. Kot visokokakovosten izdelek se zlasti zaradi svoje izredne toplotne izolativnosti uporabljajo tam, kjer nastopa pomanjkanje prostora, ali pa je prostor zelo dragocen in ga na ta način izrabimo bistveno učinkoviteje.

Na osnovi sprejetih predpostavk je življenjska doba definirana kot čas, v katerem VIP-panel doseže dvojno vrednost toplotne prevodnosti glede na začetno stanje. Na osnovi opravljenih meritev pri različnih temperaturnih obremenitvah in po različnih časih trajanja, ter ob uporabi Arrheniusovega zakona pospešenega staranja, je življenjska doba VIP-panelov v primeru trajne izpostavitve sobni temperaturi (25 °C) 26,2 leta, vrednost toplotne prevodnosti 12 mW/(m·K) pa je pod istimi pogoji dosežena v času 48,1 leta.

Ključne besede: življenjska doba, pospešeno staranje, Arrheniusov zakon, aktivacijska energija, toplotnoizolacijski materiali, vakuumskoizolacijski panel (VIP)

*Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za gradbeništvo in geodezijo, Jamova 2, Ljubljana, Slovenija, roman.kunic@fragmat.si

SI 121

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 122

Prejeto: 2010-12-28, sprejeto: 2012-04-04 © 2012 Strojniški vestnik. Vse pravice pridržane.

Ocena navtičnih tveganj za operacije UZP v koprskem pristanišču

Perkovic, M – Gucma, L. – Przywarty, M. – Gucma, M. – Petelin, S. – Vidmar, P. Marko Perkovic1,* – Lucjan Gucma2 – Marcin Przywarty2 – Maciej Gucma2 – Stojan Petelin1 – Peter Vidmar1 1 Univerza

v Ljubljani, Fakulteta za pomorstvo in promet, Slovenija 2Univerza za pomorstvo v Szczecinu, Poljska

Na slovenski obali je v okolici mesta Koper v fazi idejne zasnove postavitev terminala za utekočinjen zemeljski plin (UZP) graditelja TGE (TGE Gas Engineering GmbH). Na območju Tržaškega zaliva pa sta načrtovana še dva terminala UZP: terminal v Tržaškem pristanišču in morski terminal, ki bi se nahajal znotraj separacijske plovne cone. Namen članka je predstaviti navtični vidik vpliva postavitve novih terminalov v tržaškem zalivu, z namenom identifikacije možnih tveganj in njihovega obsega zaradi povečanega ladijskega prometa tankerjev za UZP in pretovornih operacij. Pristanišče v tržaškem zalivu obiskujejo različne vrste ladij: kontejnerske, tankerji za prevoz nafte, naftnih derivatov in kemikalij, ladje za prevoz generalnih tovorov, ladje za prevoz razsutih tovorov, ro-ro ladje, potniške ladje idr. Veliko pomorskega prometa pa se odvija tudi s plovili za šport in razvedrilo ter z ribiškimi čolni, in sicer prav na območju plovnih poti, ki vodijo v pristanišča Tržaškega zaliva. Na letni ravni pride v koprsko pristanišče okoli 2500 ladij, prav toliko jih pride tudi v tržaško pristanišče, v pristanišče Tržič pa približno 3000 ladij. Že zaradi tega so v sosednji Italiji že vzpostavili sistem nadzora ladijskega prometa VTS, ki bo v naslednjih dveh letih postavljen tudi v Sloveniji.

V članku je predstavljena ocena tveganja trčenja ali nasedanja s statističnim modelom, uporabljenim na realnih ladijskih trajektorijah. Z uporabo kvantitativnih metod je izračunano tveganje za primere trčenj na morju in nasedanja ob upoštevanju trenutnega prometa, pri čemer so podatki o plovnih poteh ladij pridobljeni iz sistema AIS (Automatic Identification System). Poleg samega števila prehodov se lahko s podatki AIS analizira tudi “vedenje” posameznih ladij, razvrščenih v skupine po velikosti. Skupine ladij so analizirane na trajektorijah (kurzih).

Tako je s pomočjo AIS-a ocenjena povprečna pozicija med-točk, kakor tudi standardni odklon trajektorij okoli te pozicije. Iz teh podatkov so bili v nadaljevanju s Poissonovo distribucijo modelirani kurzi med posameznimi med-točkami. Za modeliranje verjetnosti trčenja na morju je bil uporabljen poenostavljen statistični model. Model pravzaprav zanemarja številne elemente in njihove povezave/odvisnosti, ker je enostavno zasnovan na statističnih podatkih, pridobljenih iz opazovanja realnega ladijskega prometa. Največja neznanka (neznani parameter) za tovrstno modeliranje je natančno število bližnjih srečanj. Nekaj se jih da razbrati iz arhiva AIS, vendar pa je treba omeniti, da je teh srečanj več, saj je prav v področju pred Koprom križišče ladijskih poti, kjer se pogosto nahajajo tudi ribiške ladje in druge ladje, ki niso bile obravnavane v arhivu prometa. Zato je edina možnost za določanje parametra bližnjih srečanj v takšnih kompleksnih prometnih režimih modeliranje prometnih tokov v daljših časovnih obdobjih.

Z obdelavo podatkov je mogoče prikazati, da ladje ne plujejo samo v predpisanih področjih, temveč tudi v področjih separacijske cone, po plovnih pasovih v nasprotni smeri, kot tudi v področjih obalnega lokalnega prometa. Zaradi pomanjkljivosti sistema za nadzor (ki sloni samo na podatkih AIS) se da razbrati le neustrezno ravnanje plovil, opremljenih z napravo AIS, ki pa je lahko le posledica kršenja pravil s strani plovil (predvsem čolnov) brez AIS-a. Model je podal največjo gostoto možnih trčenj prav v področju med separacijami – v tako imenovanem področju s povišano pozornostjo. Na osnovi izračunanih podatkov je mogoče določiti raven navtičnega tveganja in koliko bi le-ta vplivala na obratovanje terminala za UZP. Ključne besede: terminal za UZP, ladijski promet, nesreče, tveganje

SI 122

*Naslov avtorja za dopisovanje: Univerza v Ljubljani, Fakulteta za pomorstvo in promet, Pot pomorščakov 4, 6320 Portorož, Slovenija, marko.perkovic@fpp.uni-lj.si

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 123

Prejeto: 2012-07-06, sprejeto: 2012-08-28 © 2012 Strojniški vestnik. Vse pravice pridržane.

Vpliv parametrov laserskega površinskega pretaljevanja na lastnosti Al-Si zlitine Sušnik, J. – Šturm, R. – Grum, J. Janez Sušnik – Roman Šturm – Janez Grum*

Univerza v Ljubljani, Fakulteta za strojništvo, Slovenija

Članek obravnava pretaljevanje aluminijeve zlitine Al-Si z Nd:YAG laserjem. Zlitina AlSi12CuNiMg z oznako EN AC-48000-K-T6, ki je bila v litem stanju in katere trdota se giblje okoli 100 HV0,2, ima dobro korozijsko odpornost in izredne mehanske lastnosti glede na njeno specifično gostoto. Zlitina se uporablja za izdelavo batov v motorjih z notranjim zgorevanjem, kjer je želja doseči višjo trdoto na izpostavljenih predelih. Sama zlitina je bila najprej metalografsko pregledana. Določene so bile velikosti in oblike intermetalnih faz Al6Cu3Ni in Al3Ni. Izločene intermetalne faze so predvsem majhnih in okroglih oblik, le redki izločki pa so podolgovati in veliki. Za povečanje trdote površinskega sloja smo uporabili postopek laserskega pretaljevanja, ki nam omogoča lokalno modifikacijo površine.

Za optimalno lasersko pretaljevanje je bilo treba nastaviti takšne parametre, da se zlitina pod laserskim snopom raztali, vendar se ne upari. Izbrani parametri so bili tako določeni s preliminarnimi testi, kjer smo dobili optimalno globino in širino pretalitve. Zaradi slabe absorptivnosti laserske svetlobe v aluminijevo zlitino smo vzorce predhodno modificirali z grafitnim nanosom, ki ima absorptivnost med 60 in 80%. Tako smo dosegli boljše rezultate z manjšimi močmi laserskega snopa, pri tem pa sam nanos ni vplival na spremembo sestave površinskega modificiranega sloja. Pri pretaljevanju je bil uporabljen zaščitni plin argon, ki preprečuje oksidacijo taline. Z laserskim pretaljevanjem se je močno spremenila mikrostruktura pretaljenega sloja, saj gre za hitro taljenje, ki mu sledi hitro ohlajanje oz. samogašenje.

Po laserskem pretaljevanju smo izvedli metalografsko analizo. V pretaljenem površinskem sloju smo dobili fino in homogeno mikrostrukturo, sestavljeno iz faz α-Al, Mg2Si, Al6Cu3Ni in Al3Ni. Spreminjanje dolžine laserskega pulza vpliva tako na globino pretaljenega sloja kot tudi na trdoto modificiranega sloja. Sprememba mikrostrukture se močno odraža na mikrotrdoti, saj je le-ta v pretaljenem delu narasla do 180 HV0,2. Z globino pretalitve trdota počasi pada, na prehodu med pretaljenim in osnovnim materialom pa se trdota izenači z osnovnim materialom. Spremenjena mikrostruktura vpliva na nastanek zaostalih napetosti.

Meritve zaostalih napetosti so pokazale, da so v pretaljenem sloju dosežene v glavnem natezne napetosti velikosti od 45 do 105 MPa. Na večjih globinah, kjer ni prišlo do pretalitve, zaostale napetosti močno padejo.

Velikostni razred zaostalih napetosti je močno odvisen od dolžine pulza laserskega snopa. Na globini 0,25 mm pri pulzu 6 ms narastejo zaostale napetosti do 105 MPa, pri pulzu 4 ms pa le do 60 MPa, pulz dolžine 8 ms pa doseže natezne napetosti 90 MPa. Neugodne natezne napetosti na površini, ki so nastale z laserskim pretaljevanjem, smo želeli odpraviti. Z naknadnim izločevalnim utrjevanjem smo odpravili visoke zaostale natezne napetosti, obenem pa se je še povečala trdota pretaljenega sloja. Po 12-urnem izločevalnem utrjevanju pri temperaturi 150 °C, ki je sledilo laserskemu pretaljevanju, so se sprostile natezne napetosti na površini. Trdota po izločevalnem utrjevanju se je povzpela za 20 HV0,2 in tako smo na površini dosegli trdoto 200 HV0,2. Razlog za povečanje trdote je v rasti intermetalnih precipitatov v strukturi pretaljenega sloja, zaostale napetosti po izločevalnem utrjevanju lasersko pretaljene površine so padle na 36 MPa. Pri pulzih dolžine 8 ms so natezne napetosti tik pod površino prešle celo v tlačne napetosti in sicer na –9 MPa. Pri nadaljnjem preučevanju laserskega pretaljevanja utrjene aluminijeve zlitine je treba preučiti, kako oblike in velikosti intermetalnih faz vplivajo na potek zaostalih napetosti.

Ključne besede: Al-Si zlitina, lasersko površinsko pretaljevanje, zaostale napetosti, mikrotrdota, precipitacijsko izločanje

*Naslov avtorje za dopisovanje: Univerza v Ljubljani, Fakulteta za strojništvo, Aškerčeva 6, 1000 Ljubljana, Slovenija, janez.grum@fs.uni-lj.si

SI 123

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Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 3. septembra 2012: Matija GOMIŠČEK z naslovom: »Krčni nased statorja in ohišja elektromotorja« (mentor: prof. dr.

Matija Fajdiga); Andraž STRLE z naslovom: »Napetostno deformacijska analiza izvleka dolivka« (mentor: izr. prof. dr. Jernej Klemenc); dne 4. septembra 2012: Gregor GRILC z naslovom: »Vrednotenje potenciala sončnega obsevanja v urbanem okolju« (mentor: prof. dr. Sašo Medved, somentor: doc. dr.

Ciril Arkar); Uroš JANČIČ z naslovom: »Varjenje nerjavnih dupleks jekel« (mentor: prof. dr. Janez Tušek); Simon KONDA z naslovom: »Batni hidravlični akumulator za vodno pogonsko-krmilno hidravliko« (mentor: doc. dr. Jožef Pezdirnik); Matej BOGATAJ z naslovom: »Določitev nosilnosti obremenitvene zveze kost-vsadek v predelu komolca na nadlahtnici« (mentor: prof. dr. Franc Kosel, somentor: doc. dr. Miha Brojan); Urban JERNEJČIČ z naslovom: »Uporaba metod vzvratnega inženiringa pri rekonstrukciji dojk« (mentor: izr. prof. dr. Peter Butala, somentor: doc. dr.

Drago Bračun); Denis KLEP z naslovom: »Pogonska enota simulatorja človeškega želodca« (mentor: prof. dr.

Janez Diaci); Grega RETELJ z naslovom: »Merjenje razpršenosti visokohitrostnega vodnega curka« (mentor: doc. dr. Joško Valentinčič); dne 26. septembra 2012: Goran GAVRANIĆ z naslovom: »Preiskava tlačnih posod z akustično emisijo« (mentor: prof. dr.

Janez Grum, somentor: asist. dr. Tomaž Kek); Uroš ILNIKAR z naslovom: »Optimizacija sočasnega osvajanja izdelka« (mentor: prof. dr. Marko Starbek, somentor: izr. prof. dr. Janez Kušar); Urban POTOČNIK z naslovom: »Nadgradnja tehničnega informacijskega sistema za delo v dobaviteljski verigi« (mentor: izr. prof. dr. Jože Tavčar, somentor: prof. dr. Jožef Duhovnik); Miha PRAŠNIKAR z naslovom: »Optimizacija zaporedja in časov trajanja operacij« (mentor: prof. dr.

Marko Starbek, somentor: izr. prof. dr. Janez Kušar); SI 124

dne 27. septembra 2012: Matija OBLAK z naslovom: »Študija preoblikovanja tanke kovinske pločevine s teflonskim nanosom« (mentor: prof. dr. Karl Kuzman); Miha PREMRL z naslovom: »Zasnova in trdnostni preračun hangarske konstrukcije« (mentor: prof. dr. Boris Štok); Urban ŽIGANTE z naslovom: »Mehanska analiza povozne vkopane cisterne« (mentor: prof. dr.

Boris Štok, somentor: doc. dr. Nikolaj Mole); dne 28. oktobra 2012: Aleš ERŽEN z naslovom: »Vzdržljivost spojev pri vibracijskih obremenitvah sesalnih enot« (mentor: prof. dr. Miha Boltežar); Iztok LONČAR z naslovom: »Antiresonančna strukturna optimizacija« (mentor: prof. dr. Miha Boltežar, somentor: doc. dr. Janko Slavič); Blaž ROŽMAN z naslovom: »Nosilna konstrukcija podesta za filter lesovine« (mentor: doc. dr. Boris Jerman); David KOZINC z naslovom: »Določanje kinematskih parametrov objektov na osnovi strojnega vida« (mentor: prof. dr. Janez Diaci); Damjan LOKNAR z naslovom: »Razvoj sistema za sledenje gibanja objektov na osnovi stereo vida« (mentor: prof. dr. Alojzij Sluga, somentor: doc. dr.

Drago Bračun); Luka PETERNEL z naslovom: »Optimizacija hidravličnega delilnika/združevalnika toka« (mentor: doc. dr. Jožef Pezdirnik); Jernej ROŠER z naslovom: »Zasnova tribometra in analiza tornih razmer pri odrezovalnih procesih« (mentor: doc. dr. Franci Pušavec, somentor: prof. dr.

Janez Kopač); Valter SMRDEL z naslovom: »Ustreznost simulacijske napovedi delovanja hidravličnih sistemov« (mentor: doc. dr. Jožef Pezdirnik). * Na Fakulteti za strojništvo Univerze v Ljubljani sta pridobila naziv magister inženir strojništva: dne 3. septembra 2012: Jesus Alejandro Moreno LOPEZ z naslovom: »Razvoj merilne naprave za analizo vpliva velikosti granuliranih mikro prahov iz nerjavečega jekla 316LW na njihovo pretočnost« (mentor: prof. dr. Igor Emri); Jamal UMER z naslovom: »Določitev ekvivalentnosti vpliva temperature in tlaka ter

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 124-129

preverba veljavnosti koncepta prostega volumna za poliamid 6« (mentor: prof. dr. Igor Emri). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv univerzitetni diplomirani inženir strojništva: dne 27. september 2012: Jurij KOŠIR z naslovom: »Predelava starih mlinov in žag na vodo v učinkovite proizvajalce električne energije« (mentor: prof. dr. Zoran Ren, somentor: prof. dr. Aleš Hribernik); Jani PROŠT z naslovom: »Razvoj varilnovpenjalne naprave za robotsko varjenje Higienskega stroja« (mentor: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Tomaž Vuherer); Kristijan TOMAŠ z naslovom: »Mestni skiro s pomožnim motorjem« (mentorica: doc. dr. Aleš Belšak, somentorica: izr. prof. dr. Miran Ulbin); Boštjan TROBENTAR z naslovom: »Razvoj ohišja diferenciala za formula student dirkalnik« (mentor: prof. dr. Jože Flašker). * Na Fakulteti za strojništvo Univerze v Mariboru sta pridobila naziv univerzitetni diplomirani gospodarski inženir: dne 21. septembra 2012: Jurij PUŠNIK z naslovom: »Optimizacija linije za proizvodnjo metlic in gnetilcev« (mentor: izr. prof. dr. Borut Buchmeister, somentorica: prof. dr. Anton Hauc); dne 27. septembra 2012: Marko SOVIČ z naslovom: »Koncept in zasnova karoserije električnega avtomobila« (mentor: izr. prof. dr. Stanislav Pehan, somentor: doc. dr. Karin Širec). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv magister inženir strojništva: dne 12. septembra 2012: Nika SAJKO z naslovom: »Karakterizacija mehanskih lastnosti visokozmogljivih materialov za letalske aplikacije pri udarnih obremenitvah« (mentor: prof. dr. Zoran Ren, somentor: doc. dr. Nikica Petrinić); dne 19. septembra 2012: Uroš JEKE z naslovom: »Modeliranje kondenzacije in uparjanja vlage na trdnih površinah z računalniško dinamiko tekočin« (mentor: prof. dr.

Matjaž Hriberšek); Rolando KOREN z naslovom: »Projekt vzpostavitve preizkuševališča opreme v

Premogovniku Velenje« (mentor: doc. dr. Iztok Palčič, somentor: izr. prof. dr. Borut Buchmeister) Matej ČONTALA z naslovom: »Optimizacija prenosnikov toplote v toplotnih črpalkah: teorija in meritve« (mentor: doc. dr. Matjaž Ramšak, somentor: prof. dr. Aleš Hribernik) * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva (UN): dne 13. septembra 2012: Peter GOLOB z naslovom: »Izdelava diagrama nosilnosti mobilnega žerjava v odvisnosti od dolžine izteka in nagiba teleskopa« (mentor: prof. dr. Iztok Potrč, somentor:doc. dr. Tone Lerher; Matjaž HRŽIČ z naslovom: »Načrtovanje in postavitev mostnega žerjava« (mentor: prof. dr. Iztok Potrč, somentor:doc. dr. Tone Lerher); Niko SRT z naslovom: »Računalniško podprto načrtovanje mostnih žerjavov« (mentor: prof. dr. Iztok Potrč, somentor:doc. dr. Tone Lerher); Vid ŠARMAN z naslovom: »Izdelava komponente stroja iz jekel različnih kvalitet« (mentor: izr. prof. dr. Vladimir Gliha); Jaka ŠTEFANČIČ z naslovom: »Preračun tračnega transporterja za transport gline« (mentor: prof. dr. Iztok Potrč, somentorja: izr. prof. dr. Ivan Pahole, doc. dr. Tone Lerher); Tadej TOPLAK z naslovom: »Tehnika žičniških naprav za prevoz oseb« (mentor: prof. dr. Iztok Potrč, somentor: doc. dr. Tone Lerher); dne 20. septembra 2012: Renato BRODAR z naslovom: »Parametrična analiza pretoka skozi aortno zaklopko« (mentor: izr. prof. dr. Jure Marn, somentor: asist. Jurij Iljaž); Benjamin CERKVENIK z naslovom: »Sušenje modrega bakra« (mentor: prof. dr. Matjaž Hriberšek, somentor: prof. dr. Zoran Ren); Primož OBAL z naslovom: »Ogrevanje enostanovanjske hiše z enoto za soproizvodnjo električne energije in toplote« (mentor: prof. dr. Aleš Hribernik); Jure ŠANTL z naslovom: »Analiza in izboljšava vzmetenja gorskega kolesa« (mentor: izr. prof. dr.

Marko Kegl, somentor: doc. dr. Boštjan Harl); Jan TIBAUT z naslovom: »Računska analiza obtekanja lopatice lopatične rešetke« (mentor: prof. dr. Aleš Hribernik); dne 21. septembra 2012: Sandi KEŠPRET z naslovom: »Računalniško podprta izdelava časovnih normativov« (mentor: doc. dr. Marjan Leber, somentorica: doc. dr. Nataša Vujica Herzog); SI 125

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dne 25 septembra 2012: Gregor DOBNIK z naslovom: »Vpliv biogoriv na karakteristike dizelskega motorja« (mentor: prof. dr.

Breda Kegl, somentor: asist. Blaž Vajda); Gregor GRAČNER z naslovom: »Primerjava reaktorske tlačne posode II. in III. generacije« (mentor: izr. prof. dr. Jure Marn); Mitja IVANČIČ z naslovom: »Numerična simulacija ohlajanja ulitka« (mentor: doc. dr. Jure Ravnik, somentor: prof. dr. Leopold Škerget); Anže JAVORNIK z naslovom: »Ogrevanje proizvodne hale z lesno biomaso« (mentor: prof. dr.

Aleš Hribernik); Blaž OREŠNIK z naslovom: »Postavitev male vetrne turbine v urbanem okolju« (mentor: prof. dr.

Aleš Hribernik, somentor: doc. dr. Matjaž Ramšak); Sašo PUŠNIK z naslovom: »Postavitev merilne proge za testiranje turbokompresorjev« (mentor: prof. dr. Aleš Hribernik somentor: dr. Gorazd Bombek); Tomaž ROBIČ z naslovom: »Ogrevanje otroškega vrtca z lesno biomaso« (mentor: prof. dr.

Aleš Hribernik, somentor: doc. dr. Matjaž Ramšak); Matevž ŠTUMBERGER z naslovom: »Numerična simulacija procesa zgorevanja v dizelskem motorju ob uporabi različnih goriv« (mentorica: prof. dr. Breda Kegl, somentorja: Blaž Vajda, Luka Lešnik); Sandi TRPLAN z naslovom: »Vpliv biogoriv na karakteristike procesa vbrizgavanja dizelskega motorja« (mentorica: prof. dr. Breda Kegl, somentor: prof. dr. Aleš Hribernik); dne 27. septembra 2012: Tadej CRNJAC z naslovom: »Integracija fotovoltaičnih sistemov v človekov bivalni, delovni in naravni prostor« (mentor: izr. prof. dr. Bojan Dolšak, somentorica: Urška Sancin); Iztok DERŽANIČ z naslovom: »Dimenzioniranje nosilne konstrukcije žerjavne proge« (mentor: prof. dr.

Srečko Glodež, somentor: doc. dr. Janez Kramberger); Marko HRIBERŠEK z naslovom: »Konstruiranje naprave za vzvojno preizkušanje« (mentor: prof. dr.

Nenad Gubeljak, somentor: izr. prof. dr. Jožef Predan); Nebojša ILIĆ z naslovom: »Fizikalno modeliranje menjalnika z dvojno sklopko« (mentor: izr. prof. dr.

Bojan Dolšak, somentorica: Urška Sancin); Luka JERMAN z naslovom: »Razvoj kompozitnega kolesa za dirkalnik Formula S GPE12« (mentor: izr. prof. dr. Jožef Predan, somentor: prof. dr.

Nenad Gubeljak); Matej KAPUN z naslovom: »Fizikalno modeliranje regeneracijskih zavor« (mentor:izr. prof. dr. Bojan Dolšak, somentorica: Urška Sancin); Mitja KRIŽNIK z naslovom: »Izdelava prototipa ohišja uparjalnika s postopkom litja z izgubljenim SI 126

jedrom« (mentor: izr. prof. dr. Igor Drstvenšek, somentor: mag. Tomaž Brajlih); Luka KUTNJAK z naslovom: »Obdelovalni stroji za orodjarstvo« (mentor: doc. dr. Mirko Ficko); Matic PLAZL z naslovom: »Računalniško podprta izdelava orodij za brizganje avtomobilskih vzglavnikov« (mentor: doc. dr. Mirko Ficko); Matjaž ŠTANER z naslovom: »Računalniška simulacija tlačnega litja dela zavornega mehanizma« (mentor: prof. dr. Zoran Ren, somentor: doc. dr.

Gorazd Lojen); Primož ŠTEFANE z naslovom: »Določitev preizkuševalne obremenitve za os veterne elektrarne« (mentor: prof. dr. Nenad Gubeljak, somentor: izr. prof. dr. Jožef Predan); Matevž ŠTERN z naslovom: »Konstruiranje zavornega pedalnega sklopa dirkalnika Formula Student GPE-12« (mentor: prof. dr. Zoran Ren, somentor: dr. Matej Borovinšek); Luka VODIŠEK z naslovom: »Parametrično modeliranje zobnih vsadkov« (mentor: izr. prof. dr.

Miran Ulbin, somentor: Franci Gačnik); Tomas ZADRAVEC z naslovom: »Naprava za transportiranje lesnih pelet« (mentor: prof. dr. Srečko Glodež, somentor: prof. dr. Iztok Potrč); Peter ZOBEC z naslovom: »Vležajenje bobna pralnega stroja« (mentor: prof. dr. Srečko Glodež); dne 28. septembra 2012: Tadej PREAC z naslovom: »Preskus sposobnosti postopka laserskega sintranja z uporabo trikoordinatne merilne tehnike« (mentor: prof. dr. Bojan Ačko). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir mehatronike (UN): dne 2. septembra 2012: Tadej PUŠNIK z naslovom: »Navigacija robota v delovnem prostoru« (mentor: izr. prof. dr. Karl Gotlih, somentor: prof. dr. Riko Šafarič); dne 20. septembra 2012: Tadej BERNHARD z naslovom: »NI FPGA sistem za zajemanje podatkov preizkuševališča pralnega stroja« (mentor: izr. prof. dr. Karl Gotlih, somentor: izr. prof. dr. Aleš Hace); Sabina MUMINOVIĆ z naslovom: »Uporabniški vmesnik ni sistema za zajemanje podatkov preizkuševališča pralnega stroja« (mentor: izr. prof. dr. Karl Gotlih, somentor: izr. prof. dr. Aleš Hace); David ROŽMARIN z naslovom: »Modifikacija mehanizma robotskega stereo vida« (mentor: izr. prof. dr. Karl Gotlih, somentor: prof. dr. Riko Šafarič); David SIMONČIČ z naslovom: »Krmilnik za univerzalni motor pralnega stroja« (mentor: izr. prof. dr. Karl Gotlih, somentor: izr. prof. dr. Aleš Hace);

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dne 27. septembra 2012: Domen BELE z naslovom: »Elektronski sistem za vodenje izmeničnih motorjev - vmesniško vezje za mikrokrmilnik TMS320F28335« (mentor: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Miran Rodič); Gašper TRSTENJAK z naslovom: »Stereo vid z meritvijo razdalje do objekta« (mentor: izr. prof. dr.

Karl Gotlih, somentorja: doc. dr. Suzana Uran, prof. dr. Riko Šafarič). * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani gospodarski inženir (UN): dne 13. septembra 2012: Tadej PUNGARTNIK z naslovom: »Analiza linijske razmestitve opreme v proizvodnem podjetju« (mentor: doc. dr. Iztok Palčič, somentor: prof. dr.

Vojko Potočan); dne 20. septembra 2012: Matic DOBRAJNŠČAK z naslovom: »Reorganizacija razmestitve proizvodne opreme v podjetju Farmtech d.o.o.« (mentor: doc. dr. Iztok Palčič, somentor: dr. Vojko Potočan); dne 21. septembra 2012: Jernej FIŠTRAVEC z naslovom: »Vrednostni menedžment v storitveni dejavnosti« (mentor: doc. dr.

Marjan Leber, somentorica: prof. dr. Majda Bastič); Benjamin KLAUŽAR z naslovom: »Model snovanja novega izdelka od ideje do koncepta« (mentor: doc. dr. Marjan Leber, somentor: prof. dr.

Anton Hauc); Jaka PAPIĆ z naslovom: »Inovacijski potencial Mariborske livarne Maribor« (mentor: doc. dr. Marjan Leber, somentorica: izr. prof. dr. Zdenka Ženko); dne 27. septembra 2012: Franci GOLEŽ z naslovom: »Inovativni pristop k razvoju izdelka s poudarkom na tržni funkciji« (mentor: doc. dr. Marjan Leber, somentorica: izr. prof. dr. Zdenka Ženko); Tine KOCBEK z naslovom: »Zbiranje in selekcija idej za inovacije s pomočjo podpornega orodja« (mentor:doc. dr. Marjan Leber, somentorica: izr. prof. dr. Zdenka Ženko); Domen PLANINŠIČ z naslovom: »Inovativne obdelave z elektroerozijo« (mentor: prof. dr. Miran Brezočnik, somentorica: izr. prof. dr. Zdenka Ženko); dne 28. septembra 2012: Rok BRAČIČ z naslovom: »Vrednostni menedžment v proizvodni dejavnosti« (mentor: doc. dr. Marjan Leber, somentorja: prof. dr. Majda Bastič).

* Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva: dne 11. septembra 2012: Klemen KALTNEKAR z naslovom: »Radionavigacijske naprave in instrumentne procedure na Letališču Portorož« (mentor: izr. prof. dr. Tadej Kosel, somentor: pred. mag. Andrej Grebenšek); Tim PRESKAR z naslovom: »Vzpostavitev enotnega evropskega neba s funkcionalnimi bloki zračnega prostora« (mentor: izr. prof. dr. Tadej Kosel, somentor: pred. mag. Andrej Grebenšek); Jasmin ŠEHOVIĆ z naslovom: »Vzdrževanje letalskega motorja CFM56-5A« (mentor: izr. prof. dr.

Tadej Kosel); Rok ŠKERJANC z naslovom: »Organizacija za vodenje neprekinjene letalnosti po EASA (Part M)« (mentor: izr. prof. dr. Tadej Kosel); dne 12. septembra 2012: Klemen ENIKO z naslovom: »Izboljšan hidravlični valj za vodno-pogonsko krmilno hidravliko« (mentor: doc. dr. Jožef Pezdirnik); Anže JELOVČAN z naslovom: »Rekonstrukcija hidravlične naprave za odvijanje tkanine« (mentor: doc. dr. Jožef Pezdirnik); Andraž PAVLICA z naslovom: »Ekonomska upravičenost investicijskega projekta« (mentor: izr. prof. dr. Janez Kušar, somentor: prof. dr. Marko Starbek); Aljaž PODOBNIK z naslovom: »Analiza uvedbe alternativne tehnologije razreza posebnih materialov v kamnoseškem podjetju« (mentor: doc. dr. Henri Orbanić, somentor: prof. dr. Mihael Junkar); Gašper VERBIČ z naslovom: »Analiza toka vrednosti« (mentor: izr. prof. dr. Janez Kušar, somentor: prof. dr. Marko Starbek). dne 13. septembra 2012: Luka STRAŽAR z naslovom: »Obvladovanje kakovosti injekcijskega brizganja leče avtomobilskega žarometa« (mentor: prof. dr. Karl Kuzman); dne 17. septembra 2012: Borut JERMAN z naslovom: »Vpliv vhodnih parametrov brušenja na cilindričnost in premer kolenastih gredi« (mentor: prof. dr. Mirko Soković); Anton KIRN z naslovom: »Tehnični in zakonski vidiki uporabe tahografov v tovornih vozilih in avtobusih« (mentor: doc. dr. Samo Zupan, somentor: prof. dr. Ivan Prebil); Marko KRAJNC z naslovom: »Pravice letalskih potnikov« (mentor: viš. pred. mag. Aleksander Čičerov, somentor: izr. prof. dr. Tadej Kosel); SI 127

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 124-129

Cene RANT z naslovom: »Nadgradnja analize FMEA z uporabo nestatističnih metod menedžmenta kakovosti« (mentor: prof. dr. Mirko Soković); dne 18. septembra 2012: Ana MARUŠIČ z naslovom: »Popis termodinamičnih lastnosti plinske zmesi« (mentor: prof. dr. Iztok Golobič); Martin VALIČ z naslovom: »Določitev optimalnih rezalnih orodij za obdelavo jekel Toolox« (mentor: prof. dr. Janez Kopač); Matej VIDIC z naslovom: »Vpliv dinamike energijskih tokov na izračun letne rabe energije stavbe z računalniškimi programi« (mentor: prof. dr. Vincenc Butala, somentor: doc. dr. Uroš Stritih); dne 19. septembra 2012: Klemen JELEN z naslovom: »Eksperimentalna analiza piezo-električnih aktuatorjev« (mentor: izr. prof. dr. Niko Herakovič); Miroslav PIRKOVIČ z naslovom: »Sodelovanje robotov pri montaži protipovratnega ventila« (mentor: izr. prof. dr. Niko Herakovič); Ina ŠPAREMBLEK z naslovom: »Optimizacija montažnih in tehnoloških procesov pri izdelavi visokonapetostnih varovalk« (mentor: izr. prof. dr.

Niko Herakovič); Matic ŠPELKO z naslovom: »Izvedbeni razredi jeklenih objektov po standardu SIST EN 1090« (mentor: doc. dr. Boris Jerman). * Na Fakulteti za strojništvo Univerze v Ljubljani so pridobili naziv diplomirani inženir strojništva (VS): dne 13. septembra 2012: Jaka BENEDIK z naslovom: »Vodno-hidravlična roka nakladalnika« (mentor: doc. dr. Jožef Pezdirnik); Alen LJOKI z naslovom: »Analiza hidravličnega motorja tipa gerotor« (mentor: doc. dr. Jožef Pezdirnik); Jaka TRČEK z naslovom: »Analiza delovanja toplotne črpalke, podprte s fotonapetostnim toplotnim sprejemnikom sončne energije« (mentor: doc. dr.

Andrej Kitanovski); dne 17. septembra 2012: Matic SMUK z naslovom: »Konstrukcija strežne naprave pri brizganju gumijastega izdelka« (mentor: izr. prof. dr. Jože Tavčar, somentor: prof. dr. Jožef Duhovnik); dne 18. septembra 2012: Jure POLJANŠEK z naslovom: »Mašenje ultrafiltracijskih membran« (mentor: prof. dr. Iztok Golobič); Edvard ŠTEFANIČ z naslovom: »Priprava procesne vode v farmacevtski industriji« (mentor: prof. dr. Iztok Golobič); SI 128

dne 19. septembra 2012: Anže GAŠPIRC z naslovom: »Neporušno testiranje kompozitnih materialov s termografijo« (mentor: prof. dr. Janez Grum, somentor: doc. dr.

Tomaž Kek); Luka PODOBNIK z naslovom: »Primerjava linearnih in nelinearnih materialnih modelov za opis vedenja vratnih ligamentov med vretenci C3 in C4« (mentor: doc. dr. Robert Kunc, somentor: prof. dr.

Ivan Prebil); * Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva: dne 20. septembra 2012: Gregor AČKO z naslovom: »Konstruiranje manipulatorja za prenos aluminijevih bram« (mentor: doc. dr. Janez Kramberger); Gregor BLAGOTINŠEK z naslovom: »Konstruiranje naprave za merjenje lastnosti sipkega materiala« (mentor: doc. dr. Janez Kramberger, somentor: prof. dr. Srečko Glodež); dne 21. septembra 2012: Mitja PLAVEC z naslovom: »Povečanje produktivnosti zalivanja anod v podjetju TALUM Aluminij d.o.o.« (mentor: doc. dr. Marjan Leber, somentor: izr. prof. dr. Borut Buchmeister); Primož ŠTINGL z naslovom: »Racionalizacija stroškov proizvodnje v podjetju VAR d.o.o.« (mentor: doc. dr. Marjan Leber); dne 27. septembra 2012: Martin AMON z naslovom: »Numerična analiza ojnice hibridne sestave« (mentor: prof. dr. Zoran Ren, somentor: dr. Matej Borovinšek); Blaž CVIKL z naslovom: »Konstruiranje pogonov diskaste kosilnice« (mentor: doc. dr. Samo Ulaga); Rok KUMAR z naslovom: »Konstruiranje in izdelava hidravličnega cepilnika drv« (mentor: doc. dr. Samo Ulaga, somentor: doc. dr. Darko Lovrec); Uroš MUŠIČ z naslovom: »Vzdrževanje strojev in naprav v sistemu DARS« (mentor: prof. dr. Boris Aberšek); Luka PELKO z naslovom: »Preračun tračnega transporterja za podporo separaciji kamnitih agregatov« (mentor: prof. dr. Iztok Potrč, somentorja: izr. prof. dr. Ivan Pahole, doc. dr. Tone Lerher); Jernej STANKO z naslovom: »Razvoj in izdelava kotla na lesne sekance« (mentor: izr. prof. dr. Vladimir Gliha, somentor: doc. dr. Tomaž Vuherer); dne 28. septembra 2012: Rok KRAMAR z naslovom: »Nadzor procesa izdelave kavnega aparata TK4« (mentor: prof. dr.

Bojan Ačko, somentor: doc. dr. Andrej Godina);

Strojniški vestnik - Journal of Mechanical Engineering 58(2012)10, SI 124-129

* Na Fakulteti za strojništvo Univerze v Mariboru so pridobili naziv diplomirani inženir strojništva (VS): dne 13. septembra 2012: Jure LESKOVAR z naslovom: »Virtualna robotizirana celica« (mentor: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Tomaž Vuherer); Dejan RUKAV z naslovom: »Modeliranje proizvodnih sistemov z Robot studio ABB« (mentor: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Tomaž Vuherer); dne 20. septembra 2012: Erik BAUMAN z naslovom: »Merjenje rezalnih pogojev pri brušenju« (mentor: prof. dr. Franci Čuš, somentor: izr. prof. dr. Ivan Pahole); Sandi KOLETNIK z naslovom: »Numerična analiza zavornega valja« (mentor: izr. prof. dr. Miran Ulbin, somentor: doc. dr. Aleš Belšak); Milan LESKOVAR z naslovom: »Optimiranje izdelave orodja za utopno kovanje« (mentor: doc. dr.

Aleš Belšak, somentor: izr. prof. dr. Miran Ulbin); Martin NOVAK z naslovom: »Numerična analiza modela poroznega materiala« (mentor: izr. prof. dr.

Miran Ulbin, somentor: prof. dr. Zoran Ren);

Matej PAL z naslovom: »Koncipiranje in zasnova avtobusnih vrat« (mentor: izr. prof. dr. Stanislav Pehan); dne 21. septembra 2012: Mihael VITKO z naslovom: »Primerjava časovnih analiz z metodo REFA in WORK FACTOR« (mentor: doc. dr. Marjan Leber, somentorica: doc. dr.

Nataša Vujica Herzog); dne 27. septembra 2012: Tadej JURIČ z naslovom: »Valjanje in vtiskovanje navojev« (mentor: izr. prof. dr. Ivan Pahole, somentor: doc. dr. Mirko Ficko); Aljaž OŽIR z naslovom: »Racionalizacija montaže bobnastih kosilnic« (mentor: prof. dr. Miran Brezočnik, somentor: Simon Klančnik); Tomaž RAJH z naslovom: »Avtomatizacija delovnega mesta za predmontažo bovdenov« (mentor: prof. dr. Miran Brezočnik, somentor: dr. Simon Brezovnik); Sašo ŠTRAJHAR z naslovom: »Manipulacija z objekti v programu Robot studio ABB« (mentor: izr. prof. dr. Karl Gotlih, somentor: doc. dr. Tomaž Vuherer); Kristijan VUGRINEC z naslovom: »Oblikovanje študentskega dirkalnega vozila« (mentor: izr. prof.

Vojmir Pogačar).

SI 129

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

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

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

Editor in Chief Vincenc Butala University of Ljubljana Faculty of Mechanical Engineering, Slovenia Technical Editor Pika Škraba University of Ljubljana Faculty of Mechanical Engineering, Slovenia Editorial Office University of Ljubljana (UL) Faculty of Mechanical Engineering SV-JME Aškerčeva 6, SI-1000 Ljubljana, Slovenia Phone: 386-(0)1-4771 137 Fax: 386-(0)1-2518 567 E-mail: info@sv-jme.eu, http://www.sv-jme.eu Print Tiskarna Knjigoveznica Radovljica, printed in 480 copies Founders and Publishers University of Ljubljana (UL) Faculty of Mechanical Engineering, Slovenia University of Maribor (UM) Faculty of Mechanical Engineering, Slovenia Association of Mechanical Engineers of Slovenia

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58 (2012) 10

Chamber of Commerce and Industry of Slovenia Metal Processing Industry Association

Since 1955

Strojniški vestnik Journal of Mechanical Engineering

Journal of Mechanical Engineering - Strojniški vestnik

F. Tasič, Marjeta Terčelj-Zorman, Matej Zajc: ronchoscopy Image Processing in the Selected Colour

year

no. 10 2012 58

volume

Cover: Cover page shows special shape of developed 3DOF MR compatible haptic mechanism and the kinematic test of the mechanism in MR environment. Figure also presents limited working space within the tunnel of MR scanner. This is the basis for manipulability analysis and development of visualization method for representation of manipulability indices in a limited space. The analysis and optimization of mechanisms shape is performed in order to achieve high quality force transmission.

Image Courtesy: Laboratory for Computer Aided Design, Faculty of Mechanical Engineering, University of Maribor

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

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[1] Hackenschmidt, R., Alber-Laukant, B., Rieg, F. (2010). Simulating nonlinear materials under centrifugal forces by using intelligent crosslinked simulations. Strojniški vestnik - Journal of Mechanical Engineering, vol. 57, no. 7-8, p. 531-538, DOI:10.5545/sv-jme.2011.013.

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[6] Rockwell Automation. Arena, from http://www.arenasimulation.com, accessed on 2009-09-07.

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http://www.sv-jme.eu

58 (2012) 10

Since 1955

Contents Papers

Matej Rajh, Srečko Glodež, Jože Flašker, Karl Gotlih: 563 Manipulability of a Haptic Mechanism within the Cylindrical Space of an MR Scanner

David Zaremba, Christian Biskup, Thomas Heber, Nico Weckend,

Werner Hufenbach, Frank Adam, Friedrich-Wilhelm Bach, Thomas Hassel: 571 Repair Preparation of Fiber-Reinforced Plastics by the Machining of a Stepped

Peripheral Zone

Xiwen Zhang, Xiaodong Wang, Yi Luo: 578 An Improved Torque Method for Preload Control in Precision Assembly of

Miniature Bolt Joints

Arif Gok, Cevdet Gologlu, Ibrahim H. Demirci, Mustafa Kurt: 587 Determination of Surface Qualities on Inclined Surface Machining with Acoustic

Sound Pressure

Roman Kunič: 598 Vacuum Insulation Panels - An Assessment of the Impact of Accelerated

Ageing on Service Life

Marko Perkovic, Lucjan Gucma, Marcin Przywarty, Maciej Gucma, Stojan Petelin,

Peter Vidmar: 607 Nautical Risk Assessment for LNG Operations at the Port of Koper

Janez Sušnik, Roman Šturm, Janez Grum: 614 Influence of Laser Surface Remelting on Al-Si Alloy Properties

Journal of Mechanical Engineering - Strojniški vestnik

10 year 2012 volume 58 no.

Strojniški vestnik Journal of Mechanical Engineering