B 16 569 ______________________________________________________________________________________________________________ Reduction of Bearing Load Capacity and Increase in Volume Flow Due to Wall Slip Maximilian M. G. Kuhr, M.Sc., Tobias Corneli, M.Sc., Univ.-Prof. Dr.-Ing. Peter F. Pelz, Institut für Fluidsystemtechnik (FST), Technische Universität Darmstadt, Germany 1 Introduction Since the beginning of the 20th century, hydraulic sealings and journal bearings are designed employing Reynolds lubrication theory /1–3/. The Reynolds lubrication the- ory presumes the no slip boundary condition at the liquid-solid interface. Recent studies conducted by the authors show, that the assumed no slip boundary condition at the liquid-solid interface is not valid for most fluid power applications; cf. /4/. This effects the prediction of leakage flow and frictional behaviour of sealing systems as well as bearing capacity of journal bearings. Thus, considering slip at the liquid-solid interface is important for the design of hydraulic components. The concept of wall slip was already discussed by Navier /5/ and Stokes /6/, when deriving the momen- tum equation for Newtonian fluids in the 19th century. Stokes favours the no slip boundary condition and justifies his hypothesis by a good agreement of the theory with experimental investigations of Poiseuille /7/. In contrast, Navier /5/ formulates the slip boundary condition, with the slip velocity being the product of the shear rate and the slip length. Regardless of this discussion, the no slip boundary condition is established over the centuries, based on the insufficient measurement techniques. However, in many technically important applications of fluid power technologies wall slip is not negligible. This is the case if the quotient of slip length and typical flow geometry is less than 10E-3. Thus, for typical hydraulic systems is reasonable to consider slip, if the gap geometries are of the order of magnitude of 10 µm. 2 Navier slip boundary condition and generalised Reynolds Equation Interesting is the train of thought, which Navier causes to formulate the slip boundary condition in 1822 /5/: Navier interprets the processes at the wall as a dynamic equi- librium between the shear force of the liquid at the wall � çR/çK and the wall-parallel adhesive forces (dynamic viscosity �). The adhesive forces are proportional to the slip velocity Ri, conforming to Stoke's law. Hence the balancing yields çR (1) Ri ∗ A � � çK . Today A is called the friction coefficient. Helmholtz /8/ interprets the constant in Na- vier's relationship by means of dimensional analysis as a length, the nowadays called slip length ÷ Reprinted with permission from [Kuhr, Maximilian; Corneli, Tobias; Pelz, Peter F. (2018): Reduction of Bearing Load Capacity and Increase in Volume Flow Due to Wall Slip. In: Proceedings of the 20th International Sealing Conference, S. 569-579, VDMA Fluidtechnik, ISBN 978-3-8163-0727-3]. Copyright 2018 VDMA Fluidtechnik - Licence: CC BY 4.0 / Creative Commons Attribution 4.0 International 570 20th ISC ______________________________________________________________________________________________________________ ÷: � �const . (2) Thus, (1) yields the purely kinematic form Ri � ÷ çRçK . (3) known today, no longer revealing Naviers original thought and dynamic nature of the boundary condition. Sometimes an anisotropic dynamic viscosity is introduced as a phenomenological model. I.e. the viscosity near the wall is considered to be smaller than in the bulk. This concept has to be judged critical from a scientific point of view. For a wall vis- cosity boundary layer the viscosity and boundary layer thickness has to be given, whereas for Naviers boundary condition only the slip length ÷ or alternatively the friction coefficient has to be quantified. The concept of anisotropic viscosity is there- fore no scientific progress, rather a drawback. Figure 1 illustrates the geometrical interpretation of the slip length Figure 1: Simple shearing flow for no slip (grey) and slip (black) by means of a simple shearing flow example. It shows the velocity profiles for no slip boundary condition (grey) and Navier slip boundary condition (black). The lower sur- face is fixed while the upper surface moves with constant velocity ø at a distance >. Figure 1 also illustrates the geometrical interpretation of the slip length by means of a simple shearing flow example. It shows the velocity profiles for no slip boundary condition (grey) and Navier slip boundary condition (black). In the case of no-slip, the velocity of the liquid molecules at the wall is identical to the wall velocity. In the case of slip, there is a relative velocity between the wall near molecules of the liquid and the wall. The relative velocity at the fixed wall Riá is greater than zero and the relative velocity at the moving wall RiQ reduces the liquid velocity relative to the wall velocity ø. Extrapolating the velocity profile down to zero and up to the surface ve- locity ø yields the slip length as the perpendicular distance from the surface to the boundaries of the extrapolated velocity profile. Due to this, Helmholtz’s [8] geometric B 16 571 ______________________________________________________________________________________________________________ interpretation of an apparent gap opening due to wall slip is deduced. Slip velocity and shear rate are proportional and the proportionality constant is the slip length. So far, the Reynolds equation is derived and applied only assuming no-slip boundary condition. Since wall slip affects both the leakage flow of seals and the bearing ca- pacity of plain bearings, it is necessary to solve the Reynolds equation while taking the dynamic slip boundary condition into account. The generalized Reynolds equa- tion incorporating wall slip can be derived in five steps. Starting from the well-known assumptions of the Reynolds equation (cf. /9/), i.e. %?� ≪ 1, the Navier Stokes equation simplify to the linear differential equations ç¶ç�3 � � ç'R3ç��' ; ç¶ç�' � � ç'R'ç��' ; ç¶ç�� � 0. (4) By integrating the linear differential equation twice, the general solution of the veloc- ity distribution in the lubrication gap is derived. Applying Navier's slip boundary condition for the velocities R at the fixed ��� � 0 and the moving wall ��� � > R���� � 0 � ÷µ çR� ç�� ���Ûg, (5) R���� � > � ø� − ÷T çR�ç�����Û���� , (6) yields the special solution of the velocity distribution for < � 1, 2. Substituting the in- tegration constants using the boundary conditions from Equation (5) and (6) provides the solution to the boundary value problem, with > � >��� , � . The volume flow ± per unit depth in the ��-direction, is obtained by integrating the velocity distribution across the lubrication gap height. ±��� � ø�>2 1 + 2÷µ/>1 + ÷µ/> + ÷T/> − >�12� ç¶ç�� 1 + 4÷µ/> + 4÷T/> + 12÷µ÷T/>'1 + ÷µ/> + ÷T/> . (7) Inserting (7) into the integral form of the continuity equation yields the generalized Reynolds equation for plane flows. New is the now included wall slip. ∂ç�� � >�� d¶ d�� ¹1 + 4÷µ/> + 4÷T/> + 12÷µ÷T/>'1 + ÷µ/> + ÷T/> »� � � ∂ ∂�� 6ø�> ! 1 + 2÷µ/>1 + ÷µ/> + ÷T/>"# + ç>ç� . (8) 572 20th ISC ______________________________________________________________________________________________________________ 2.1 Application to Slider Bearings The pressure distribution in the plane lubrication gap of a slider bearing as well as the increased volume flow is investigated, taking Navier's slip boundary condition into account; cf. Figure 2. The upper wall is inclined by the angle % with respect to the � - axis and moves at the constant velocity ø. The liquid is pulled into the nar- rowing gap, leading to a pressure increase. So far, the load-bearing capacity for this arrangement is only examined for the no-slip boundary condition. In the following, the influence of wall slip is considered as well. Figure 2: Slider bearing Deriving the bearing capacity of the slider bearing, it is assumed for simplicity that both the fixed and moving wall are of the same material and Temperature, thus the slip length is identical at both solid-liquid interfaces: ÷µ � ÷T � ÷. Under this assump- tion and for quasi-stationary flow, the modified Reynolds equation (8) reads d d� �>�� d¶ d� ¹1 + 8÷'/> + 12÷'/>'1 + 2÷/> »� � d d� �6ø> . (9) Integration of (9) leads to >�� d¶ d� ¹1 + 8÷'/> + 12÷'/>'1 + 2÷/> » � 6ø> + �3 . (10) The integration constant �3 is determined at the location � � �̅ of maximum pressure, where d¶/d� � 0, giving the pressure gradient depending on the gap height >�� and the gap height at maximum pressure >$ d¶ d� � 6�ø 1 − >/>>�6÷ + > . (11) With the gap height >�� � >3 − %� and substituting d� � −d>/%, the pressure dis- tribution yields B 16 573 ______________________________________________________________________________________________________________ ¶�>�� � − 6�ø% � 1 − >/>>�6÷ + > d> (12) Applying the boundary condition ¶�� � 0 � ¶�>3 � 0 the integration constant can be determined. Determining the position of the maximum pressure >$ at ¶�� � t �¶�>' � 0 resulting in the pressure distribution and the position of the maximum pres- sure. ¶�> � �ø6%÷' %>$ + 6÷& ln ! > + 6÷>3 + 6÷" − ln ! >>3"# + 6>$÷ !> − >3>>3 "#, (13) > � 6÷ (ln �>' + 6÷>3 + 6÷� + ln �>3>'�) ln �>3 + 6÷>' + 6÷� + ln �>'>3� + 6÷ �>3 − >'>'>3 �. (14) With the made assumptions, inserting (13) into (7) leads to the volume flow per unit depth of the slider bearing including slip ± � 3ø÷ (ln �>'>3� + ln �>3 + 6÷>' + 6÷�) 6÷ �>' − >3>'>3 � + ln �>3>'� + ln �>' + 6÷>3 + 6÷�. (12) 3 Darmstadt Slip Length Tribometer The Darmstadt Slip length Tribometer (DSLT) /4, 10/ is an indirect measuring method for quantifying wall slip. The slip length is determined by the measurement of an integral quantity and a suitable model. The integral measured quantity is hereby the friction torque between rotating and the stationary disk depending on the gap height. As a suitable model, the Reynolds equation (8) is used. The DSLT is a classical plate-plate tribometer (cf. Figure 3) measuring the friction torque transmitted from the rotating disk through the liquid film of height h to the stationary disk. The torque is measured with a reaction torque sensor at the station- ary disk. The distance is measured by means of integrated capacitive distance sensors with a resolution of 4 nm. In order to allow a cardanic self-levelling of the two disks relative to each other, the lower one is supported by a jewel bearing. The adjustment of the lubrication gap is achieved by the axial spring stiffness and the feed pressure of the test liquid. 574 20th ISC ______________________________________________________________________________________________________________ Figure 3: Principle sketch of the slip length tribometer with a disk diameter of 64 mm Figure 4: Inverse friction torque depending on the gap height for slip and no slip boundary condition In the lubrication gap, the pressure flow in radial direction and the drag flow in cir- cumferential direction are superimposed. For small gap heights, the Reynolds number is of the order of magnitude of 0.1 and the tilt angle of the disks is smaller than 0.001°. Thus the Reynolds equation (8) for wall slip can be used, giving the Couette velocity profile R�B, º � ΩB�º + > /�> + ÷3 + ÷' . With the velocity profile in the circumferential direction, the friction torque is determined by integrating the shear stresses. �â3 � > + ÷3 + ÷'�ΩÉ� . (13) B 16 575 ______________________________________________________________________________________________________________ The inverse of the friction torque is a linear equation with the polar moment of area É+ � , B'd 0 , the sum of the slip lengths at the stationary and rotating disk can be obtained by determining the �-axis intercept for the curve of the equation. Figure 4 illustrates schematically the relationship of equation (13) for Newtonian flu- ids. The inverse friction torque depends linearly on the gap height. As the no-slip boundary condition holds true, the linear curve of the inverse friction torque intersects the coordinate origin. As the rotational frequency of the rotating disks increases, the slope of the straight-line equation decreases. If wall slip occurs at the liquid-solid interface, the curve intersects the �-axis in the negative. This distance is equal to the sum of the two slip lengths at the stationary and the rotating disk. 4 Results 4.1 Slip length measurement Figure 5 shows the measurement of the inverse torque as a function of varying gap heights, showing the �â3- > - curve for an alpha-olefin 6 at constant temperature of 29.9°C and constant rotational frequency of 2 Hz. The symbols mark the individual measurement points at which the torque was measured depending on the gap height. The figure shows 20 measurement series. The best fit linear regression curve is determined individually for each measurement series and the slip length is deter- mined from the intersection with the �-axis. Figure 5: Slip length measurements for POA 6 Figure 6: Detailed view of the intersection Figure 6 gives a detailed view at the intersection of the regression curves with the �- axis. The measured slip length for an alpha-olefin at 29.9°C averages to ÷ � 540 nm. The measurements can be repeated with a standard deviation of _ � 50 nm. Due to the 20 repetitions, the statistical uncertainty is reduced by the factor �/√K (student- 576 20th ISC ______________________________________________________________________________________________________________ t-distribution). The statistical uncertainty of the measured value with K � 20 repeti- tions and a confidence interval of 95% is less than 24 nm and thus less than 5% of the measured mean value. 4.2 Bearing capacity and volume flow of a slider bearing Figure 7 gives the pressure distribution in a lubrication gap for the slip as well as for the no slip boundary condition applying Equation (13). Figure 7: Pressure distribution in the lubri- cated gap Figure 8: Load capacity of the slider bearing The gap height >3 is 20 μm, the tilt angle % of the upper plate is 0.286° and the plate is moved at constant velocity ø of 0.3 m/s. The utilized fluid is an alpha-olefin with a dynamic viscosity � of 0.039 Pa·s and a slip length ÷ of 540 nm at 29.9 °C. The quotient of slip length and typical flow geometry >3 is thus 0.027. Figure 7 exhibits, that the peak pressure is reduced due to slip by approximately 20 %. Integrating the pressure along the horizontal directions yields the load capacity per unit depth, which is also reduced by approximately 20 %. Hydraulic applications operate in a wide temperature range. Thus the thermal be- haviour of the load capacity and leakage flow is of major interest. The temperature dependent dynamic viscosity and the temperature dependent slip length are ob- tained by means of Arrhenius relations; cf. /4/. Using these Arrhenius relations, Figure 8 gives the temperature depending load capacity of the above mentioned slider bearing. The presented results clearly show that difference in load capacity per unit depth for slip and no slip decreases with increasing temperature. This is reasonable since the activation energy for wall slip is smaller than the activation en- ergy for shearing, cf. /4/. Figure 9 gives the temperature depending volume flow per unit depth of the slider bearing using equation (12) and the Arrhenius relation for the temperature depend- ence of the slip length. At 29.9 °C the volume flow including slip increases in contrast B 16 577 ______________________________________________________________________________________________________________ to the no-slip solution by approximately 0.75 %. With increasing system temperature, the volume flow including slip starts to approach the no-slip values. Figure 9: Volume flow ±� per unit depth 5 Summary and Conclusion The presented paper investigates the temperature dependence of the Navier slip boundary condition and the related reduction of load capacity of a bearing and the increased volume flow. First, the Navier slip boundary condition is discussed and a modified Reynolds equation, including slip, is derived. Second, the modified Reyn- olds equation, the pressure distribution and the load capacity of a slider bearing as well as the corresponding volume flow are deduced. Third, the Darmstadt Slip Length Tribometer is presented, utilized for measuring the slip length of technical rough sur- faces. Finally, the paper closes with the comparison of the temperature dependent results of the slip length measurements and the effect on the load capacity of the slider bearing as well as the increased volume flow in comparison to the common no slip boundary condition. In future further efforts have to be made in understanding the effects on the temperature-dependent sliding length. In a first step the molecular structure of the oil has to be varied. Afterwards the influence of surface roughness on the slip length needs further investigation. In a final step, the influence of different material combinations should be addressed. 6 Acknowledgements The authors thank the Research Association for Fluid Power of the German Engi- neering Federation VDMA for its financial support. Special gratitude is expressed to the participating companies and their representatives in the accompanying industrial committee for their advisory and technical support. 578 20th ISC ______________________________________________________________________________________________________________ 7 Nomenclature Variable Description Unit C frequency T-1 > height L > height at maximum pressure L � torque L2 M T-2 ¶ pressure L-1 M T-2 ± volume flow L3 T-1 ø velocity of the slider bearing L T-1 R velocity L T-1 % angle * Θ temperature Θ ÷ slip length L � dynamic viscosity L-1 M T-1 Ω angular velocity T-1 8 References /1/ Müller, H.K.: Abdichtung bewegter Maschinenteile. Funktion - Gestaltung - Be- rechnung - Anwendung. Müller, Waiblingen (1990) /2. Petrov, N.P., Reynolds, O., Sommerfeld, A., Michell, A.G.M., Hopf, L. (eds.): Abhandlungen über die hydrodynamische Theorie der Schmiermittelreibung. Ostwald's Klassiker der exakten Wissenschaften, vol. 218. Akad. Verl.-Ges, Leipzig (1927) /3/ Trutnovsky, K.: Die Dichtung bewegter Maschinenteile. In: Berührungs- dichtungen /4/ Corneli, T., Pelz, P.F.: The Activation Energy for Wall Slip. submitted to Phyis- cal Review Letters, 2017 /5/ Navier, M.: Mémoires de l'Académie des Sciences de l'Institut Impérial de France. Firmin Didot (1827) /6/ Stokes, G.G. (ed.): Mathematical and Physical Papers. Cambridge Library Col- lection - Mathematics. Cambridge University Press, Cambridge (2009) B 16 579 ______________________________________________________________________________________________________________ /7/ Poiseuille, J.L.: Recherches experimentales sur le mouvement des liquides dans les tubes de tres-petits diametres. Imprimerie Royale (1844) /8/ Helmholtz, H.: Über Reibung tropfbarer Flüssigkeiten. Von H. Helmholtz und G. v. Piotrowski. (Mit 2 Taff.) (Aus d. XL. Bd. S. 607. 1860. der Sitzgsber. der math-nat. Cl. der k. Ak. der Wiss. bes. dbg.). Hof- & Stts.-Druck (1860) /9/ Spurk, J.H., Aksel, N.: Strömungslehre. Einführung in die Theorie der Strömun- gen, 8th edn. Springer-Lehrbuch. Springer-Verlag Berlin Heidelberg, Berlin, Heidelberg (2010) /10/ Corneli, T., Pelz, P.F., Ludwig, G.: Slip length in narrow sealing gaps an expe- rimental approach. In: 18th International Sealing Conference, Stuttgart (2014) Englisch Summaries Group I Session 1: Introduction Lectures I 01 (DE): A Few Ideas About Squeaking and Whistling by Mechanical Seals I 02 (DE): Causes of Failure of Elastomeric Seals – Evaluation of More than 2000 Analyses Group A Session 2: Rotary Shaft Seals A 01 (DE): Modern Visual Methods for Wear Analysis at Radial Shaft Seals A 02 (EN): Functional Behaviour of Different Back Structures for PTFE Shaft Seals A 03 (DE): Adaptive PTFE Rotary Shaft Seals Session 3: Mechanical Seals A 04 (DE): Hybrid Seal Faces Produced with Laser Beam Melting A 05 (DE): Dimensional Stability of Seal Faces Made of Carbon Ceramics A 06 (EN): Realization of Ultra-High Speed, Zero-Leakage and Low-Friction Textured Mechanical Seals by combining Liquid and Gas Lubrications – Gas Liquid Hybrid Face A 07 (EN): Realisation of Zero-Leakage and Low-Friction Surface Textured Mechanical Seals for Tidal Turbine Session 4: Tribology A 08 (DE): Influence of Operating Parameters on the Visual Indication of the Sealing Zone of Radial Shaft Seals under Constant Power Loss A 09 (DE): Experimental Determination and Comparison of the Contact Temperature of Radial Shaft Seals and its Derived Tribological System A 11 (DE): Influences of System Parameters in Practice on the Pumping Rate Behavior of Radial Lip Seals Session 5: Rotary Shaft Seals A 12 (EN): Deployment of Fast Tools to Predict Performance of Rotary Seals A 13 (DE): A New Interpretation of Elastomer Contact Seal (ROS) with Improved Performance A 14 (EN): Analysis of Large Diameter Lip Seal Behaviour for Hydraulic Dam Turbine Bearing A 15 (EN): Evaluating Leakage from Radial Lip Seals Affected by BearingArea of Shaft Topography Session 6: Rotary Shaft Seals A 16 (DE): Further Development and Calculation of Contactless Shaft Seals A 17 (DE): Impact of Imperfections on the Sealing Function of Radial-Lip-Seals A 18 (DE): Research on the Pumping Rate of Sealing Counterfaces with Macro-Lead A 19 (EN): Contact Investigation of Rotary Shaft Seals Session 7: Application in Practice A 20 (DE): New Generation of Omni-Directional Large Diameter RSS for Grease Lubricated Main Bearings A 21 (EN): Abrasive Applications and their Challenges for Dynamic Seals in Food&Pharma Environment A 22 (EN): Towards Cognitive Sealing: Artificial Intelligence based Seal Condition Monitoring Group B Session 2: Reciprocating Seals B 01 (EN): Analysis of Particle Currents of a Wiper Seal at a Cylinder Rod B 02 (DE): Seal Gap and Guiding, a General Conflict Situation? B 03 (EN): Enhanced Performance of Hydraulic Rod Seals by Use of 3DPrinted Reinforcement Session 3: Static Seals B 04 (EN): Unavoidable Minimum Leakage of PTFE Gaskets Due to Diffusion B 05 (DE): Fibre Gaskets – Latest Development B 06 (DE): Next Level of PTFE Gasketing Material – Higher Safety at Installation and During Operation B 07 (EN): Universal Aspects on Functionality and Performance of HNBR FS Seals Session 4: Simulation B 08 (EN): Analysis of Continuously Film Coating Conditions On Seal Interface of Plunger Reciprocating Pump for Water-jet Cutting Use B 09 (DE): Friction Reduction of Dynamic Elastomeric Seals by Simulation-Based Optimisation of the Microstructured Sealing Surface B 10 (EN): Simulation of the Stick-Slip Effect Observed in Hydraulic Sealing Systems Made of Polyurethanes B 11 (EN): On the Modeling of Ageing on Rubber Seals Session 5: Materials and Surfaces B 12 (EN): Sliding Friction Behavior of Polymer Seals – Effect of Polymer Structure and Modifications B 13 (EN): New Grade Thermoplastic Polyurethane with High Thermal Conductivity and Low Coefficient of Friction B 14 (DE): Conductive Elastomers for Diaphragm Seals B 15 (DE): Leakage Newly Defined: Electro-Magnetic Radiation and How to Seal It Session 6: Reciprocating Seals B 16 (EN): Reduction of Bearing Load Capacity and Increase in Volume Flow Due to Wall Slip B 17 (DE): Extended Approach to Specify Counter Surfaces for Reciprocating Seals B 18 (DE): Can a Hydraulic Rod Seal Wipe Out Surface Structures on the Rod? – A Viscoelastic FE-Analysis of a PU-U Cup Seal on Different Surfaces B 19 (EN): Surface Energy Wetting Effects on Hydrodynamic Lubrication ofHydraulic Rod Seals Session 7: Static Seals B 20 (DE): A Question of Hygiene - What to Consider When Choosing a Gasket B 21 (DE): The 3D Seal - A New Sealing Solution for Steel Enamel Applications with Large Deviations B 22 (DE): Large Format Gaskets for Battery Housings Group C Session 8: Closing Lectures C 01 (EN): Influence of Hydraulic Fluid Properties on a Tribological System Containing a Reciprocating Sealing Element. C 02 (DE):Transient Dynamic Multiphysic Simulations of PTFE Shaft Seals Considering Friction, Heat Generation and Wear Deutsch Zusammenfassungen Gruppe I Session 1: Einführungsvorträge I 01 (DE): Ein paar Ideen zum Quietschen und Pfeifen durch Gleitringdichtungen I 02 (DE): Ausfallstatistik von Elastomerdichtungen Gruppe A Session 2: Wellendichtungen A 01 (DE): Moderne visuelle Untersuchungsmethoden für die Verschleißanalyse am Beispiel Radial-Wellendichtring A 02 (EN): Funktionsverhalten von unterschiedlichen Rückenstrukturen für PTFE Wellendichtungen A 03 (DE): Adaptive PTFE Wellendichtungen Session 3: Gleitringdichtungen A 04 (DE): Herstellung von hybriden Gleitringen mittels additivem Verfahren A 05 (DE): Maßstabilität von Gleitringen aus Kohlenstoffkeramiken A 06 (EN): Realisierung von reibungsarmen mechanischen Gleitringdichtungen mit extrem hohen Geschwindigkeiten, Null-Leckage und Oberflächentextur durch die Kombination von Flüssigkeits- und Gasschmierung - Gas Liquid Hybrid Face A 07 (EN): Realisierung von reibungsarmen strukturierten Gleitringdichtungen mit Null-Leckage für Gezeiten-Turbinen Session 4: Tribologie A 08 (DE): Zum Einfluss der Betriebsparameter auf den optischen Befund einer RWDR-Dichtzone bei konstanter Verlustleistung A 09 (DE): Vergleichende Untersuchungen der Kontakttemperatur im Real- und Ersatzsystem von Radialwellendichtringen A 11 (DE): Einflüsse praxisrelevanter Systemparameter auf das Förderverhalten von Radial-Wellendichtungen Session 5: Wellendichtungen A 12 (EN): Bereitstellung schneller Werkzeuge zur Vorhersage der Leistung von Rotationsdichtungen A 13 (DE): Eine neue Interpretation der elastomeren Kontaktdichtungen (RWDR) zur Steigerung der Leistungsfähigkeit A 14 (EN): Analyse des Lippendichtungsverhaltens großer Durchmesser für hydraulische Turbinenlager A 15 (EN): Wellenoberflächentopographie als ein Faktor, der die radiale Lippendichtungsleistung beeinflusst Session 6: Wellendichtungen A 16 (DE): Weiterentwicklung von berührungslosen Wellendichtungen A 17 (DE): Förderwirkung von Fehlstellen im Dichtsystem Radial-Wellendichtung A 18 (DE): Untersuchung des Förderverhaltens makrodrallbehafteter Dichtungsgegenlaufflächen A 19 (EN): Kontaktuntersuchung von Radial-Wellendichtringen Session 7: Anwendungsthemen A 20 (DE): Eine neue Generation von omni-direktional wirkenden RWDR für fettgeschmierte Hauptlager in Windkraftanlagen A 21 (EN): Abrasive Anwendungen und deren Herausforderungen für dynamische Dichtungen im Bereich Food&Pharma A 22 (EN): Cognitive Sealing: Zustandserkennung von Dichtungen basierend auf künstlicher Intelligenz Gruppe B Session 2: Translat. Dichtungen B 01 (EN): Analyse der Partikelströme über einen Abstreifer an einer Zylinderkolbenstange B 02 (DE): Dichtspalt und Führung, ein grundsätzlicher Konflikt? B 03 (EN): Verbesserte Leistung der Hydraulischen Stangendichtungen durch Verwendung von 3D-Druckverstärkungen Session 3: Statische Dichtungen B 04 (EN): Unvermeidbare Leckage bei PTFE Dichtungen aufgrund von Diffusionsvorgängen B 05 (DE): Faserstoff Reloaded – Die Entwicklung geht weiter B 06 (DE): Die nächste Ebene von PTFE Dichtungen - Höhere Sicherheit bei der Montage und im Prozess bei gleichzeitiger Lagerkostenoptimierung B 07 (EN): Generelle Aspekte hinsichtlich Funktionalität und Performance von HNBR FS-Dichtungen Session 4: Simulation B 08 (EN): Kontinuierliche Flüssigkeitsfilmbildungszustandsanalyse unter Abdichtung von Wasserschneiden im Plungerpumpenbereich B 09 (DE): Reibungsreduzierung dynamischer Elastomerdichtungen durch die simulationsgestützte Optimierung der mikrostrukturierten Dichtfläche B 10 (EN): Simulation des Stick-Slip Effekts bei Hydraulikdichtungen aus Polyurethan B 11 (EN): Ein Alterungungs-Modell für elastomere Dichtungen Session 5: Werkstoffe und Oberflächen B 12 (EN): Gleitreibungsverhalten von Dichtungen – Einfluss der Polymermorphologie und von Modifikatoren B 13 (EN): Neue Sorte thermoplastisches Polyurethan mit hoher Wärmeleitfähigkeit und niedrigem Reibungskoeffizienten B 14 (DE): Elektrisch leitfähige Elastomere für Membranventildichtungen B 15 (DE): Leckage neu definiert: elektro-magnetische Strahlung und wie man sie abdichten kann Session 6: Translat. Dichtungen B 16 (EN): Reduzierung der Lagertragfähigkeit und Erhöhung des Volumenstroms durch Wandgleiten B 17 (DE): Erweiterter Ansatz zur Spezifikation von Gegenlaufflächen für translatorische Dichtungen B 18 (DE): Kann ein Hydraulikdichtring Strukturen auf der Stangenoberfläche auswischen? – Viskoelastische FE-Untersuchungen an einem PU-Nutring auf unterschiedlichen Stangenoberflächen B 19 (EN): Einfluss der Oberflächenenergie auf die Schmierfilmbildung bei Hydraulikstangendichtungen Session 7: Statische Dichtungen B 20 (DE): Alles eine Frage der Hygiene – Was bei der Dichtungsauswahl zu beachten ist B 21 (DE): Die 3D-Dichtung – Eine neue Dichtungslösung für Stahl-Email-Anwendungen mit großen Ebenheitsabweichungen B 22 (DE): Großformatige Dichtungen für Batteriegehäuse Gruppe C Session 8: Abschlussvorträge C 01 (EN): Einfluss von Eigenschaften der Hydraulikflüssigkeiten auf ein tribologisches System mit einem translatorischen Dichtungselement C 02 (DE): Transiente Mehrfeldsimulation der Dynamik vonPTFE-Wellendichtungen unter der Berücksichtigungvon Reibung, Wärmeentwicklung und Verschleiß