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Tribologie und Schmierungstechnik EDITOR IN CHIEF MANFRED JUNGK 2 _ 25 VOLUME 72 Tribology—Lubrication Friction Wear An Official Journal of Gesellschaft für Tribologie An Official Journal of Österreichische Tribologische Gesellschaft An Official Journal of Swiss Tribology Issue 2 | 2025 Volume 72 Editor in chief: Dr. Manfred Jungk Tel.: +49 (0)177 1902330 eMail: jungk@verlag.expert www.mj-tribology.com Editorial director: Ulrich Sandten-Ma Tel.: +49 (0)7071 97 556 56 / eMail: sandten@verlag.expert Editor: Patrick Sorg Tel.: +49 (0)7071 97 556 57 / eMail: sorg@verlag.expert Dr. rer. nat. Erich Santner Tel.: +49 (0)2289 616136 / eMail: esantner@arcor.de Contributions marked with the author’s initials or full name represent the author’s opinion, not necessarily that of the editorial office. We take no responsibility for unsolicited contributions. The author is responsible for obtaining the rights to pictures. 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ISSN 0724-3472 ISBN 978-3-381-13781-7 Imprint Tribologie und Schmierungstechnik Tribology—Lubrication Friction Wear An Official Journal of Gesellschaft für Tribologie | An Official Journal of Österreichische Tribologische Gesellschaft | An Official Journal of Swiss Tribology Editorial 1 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TUS-2025-0006 “The Circularity Gap Report 2025” by Circle Economy (2025) Amsterdam states that more raw materials are still being extracted than recycled. Closed material cycles can significantly reduce the demand for raw materials and contribute to lowering greenhouse gas emissions. However, globally, the share of secondary raw materials used in production is declining. This fell from 7.2 to 6.9 percent compared to the previous report. Global exploitation of raw materials, namely Crops, Crop residues, Grazed biomass and fodder crops, Other biomass, Coal, Petroleum, Natural gas, Other fossil fuels, Iron ore concentrates and compounds, Copper ore concentrates and compounds, Gold ore concentrates and compounds, Other metal ores, Sand gravel and crushed rock for construction, Limestone, Structural clays and Other nonmetallic mineral has tripled over the past fifty years. Most recently, it rose to 100 billion tons. This could not be offset by secondary raw materials that are returned to the cycle from previous production, mostly Construction & demolition (49.6 %) and Industrial (44.0 %) waste. Their volume has also been growing for several years. A truly circular economy can be more resource-efficient, the report states. The potential for this is there: If all waste that is not currently recycled, 16 % is controlled and 57 % uncontrolled disposed, were made usable, the recycling rate used could increase from 6.9 to 25 percent, the experts calculated. The share of fossil fuels is declining from above 18 to under 16 % of all raw materials, according to data from Circle Economy. Nevertheless, the volume has increased from 6.1 billion tons at the beginning of the 1970s to 15.8 billion tons this decade. This corresponds to 13.3 percent of all materials used in the global economy. The transition to a solar-based economy will initially be material-intensive, but with intelligent processes, the use of materials can become more economical. Principles such as longer service life, reuse, and recycling earlier must be considered both in infrastructure investments and in the production of goods. By the middle of the century, the world population will grow by 2.5 billion people. This will require a significant expansion of infrastructure, for which countries should rely more than before on existing material stocks. The wealthier industrialized countries should ensure that they do not further expand existing material stocks, suggests the report. The report compares the global average lifetime in years of selected assets such as vehicles (17), buildings (54) and appliances (9) and remarks that there are no targets set to increase the lifetime or recycle rates. The “Gesellschaft für Tribologie e.V.” has published 3 Tribology in Germany studies, where the one named “Wear protection and sustainability as cross-sectional challenges” shows examples how to reduce the circularity gap, so remember Tribology is everywhere. Your editor in chief Manfred Jungk Circular Economy Events 2 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 Events We look forward to your contribution! The scientific journal Tribologie und Schmierungstechnik (TuS) is one of the leading publications for tribological research in Germany, Austria and Switzerland. As the official journal of the Society for Tribology (GfT) in Germany, the Austrian Tribological Society (ÖTG) and Swiss Tribology, the issues provide information on research from industry and science, current events and developments in the specialist community. Further information on the journal and publication: https: / / elibrary.narr.digital/ xibrary/ start.xav? zeitschriftid=tus&lang=en Date Place Event ► 02.09. - 04.09.25 Leeds, UK 50 th Leeds-Lyon Symposium on Tribology ► 20.09. - 25.09.25 Rio de Janeiro, Brasil 8 th WORLD TRIBOLOGY CONGRESS ► 29.09. - 01.10.25 Wernigerode, Germany 66. GfT Conference Tribology ► 14.10. - 16.10.25 Lemont, Illinois (USA) Argonne National Laboratory 2025 STLE Tribology Frontiers Conference ► 15.10. - 17.10.25 Stresa, Italy 2025 60 th UEIL Annual Congress ► 27.10. - 30.10.25 Amsterdam, Netherlands elgi Autumn Events ► 12.11. - 14.11.25 Dubrovnik, Croatia LUBRICANTS & BASE OILS SYMPOSIUM ► 19.11. - 21.11.25 Troy, Michigan (USA) 2025 STLE Tribology & Lubrication for E-Mobility Conference ► 27.01. - 29.01.26 Ostfildern, Germany 25 th International Colloquium Tribology Contents 3 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 Tribologie und Schmierungstechnik Tribology - Lubrication Friction Wear An Official Journal of Gesellschaft für Tribologie An Official Journal of Österreichische Tribologische Gesellschaft An Official Journal of Swiss Tribology Volume 72, Issue 2 August 2025 5 Christof Koplin, Bernadette Schlüter, Raimund Jaeger The sensitive dependence of sealing elastomers on lubricants 12 Alexander Roegnitz, Elias Merz, Hubert Mantz, Carsten Siemers, Andreas Haeger Tribological Investigation of the Novel Titanium Alloy TNTZ-O for Dental Implant Applications: Subsequent Results of a Comparative Study 20 Felix Farrenkopf, Thomas Lohner, Karsten Stahl, Kirsten Bobzin, Christian Kalscheuer, Max Philip Möbius, Marta Miranda Marti Dry Lubrication of Spur Gears Coated with CrAlN+Mo: W: S 30 Susanne Fritz Insight into large deformation contacts of soft polymers with molecular dynamics simulations 38 Sandra Kiese, Daniela Leistl, Jan Ulrich Michaelis, Stefan Hofmann, Thomas Lohner Cellulose in Motion: Enzymatically Modified Biopolymers and Glycerol in Tribological Interaction 1 Editorial Circular Economy 2 Events Science and Research 41 News Gesellschaft für Tribologie Columns Preface For authors Authors of scientific contributions are requested to submit their manuscripts directly to the editor, Dr. Jungk (see inside back cover for formatting guidelines). Anzeige 4 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 Gesundheit \ Romanistik \ Theologie \ Kulturwissenschaften \ Soziologie \ Theaterwissenschaft \ Geschichte \ Spracherwerb \ Philosophie \ Medien- und Kommunikationswiss chaft \ Linguistik \ Literaturgeschichte \ Anglistik \ Bauwesen \ Fremdsprachendidaktik \ DaF \ Germanistik \ Literaturwissenschaft \ Rechtswissenschaft \ Historische Sprachwiss chaft \ Slawistik \ Skandinavistik \ BWL \ Wirtschaft \ Tourismus \ VWL \ Maschinenbau \ Politikwissenschaft \ Elektrotechnik \ Mathematik & Statistik \ Management \ Altphilol Sport \ Gesundheit \ Romanistik \ Theologie \ Kulturwissenschaften \ Soziologie \ Theaterwissenschaft \ Geschichte \ Spracherwerb \ Philosophie \ Medien- und Kommunikatio issenschaft \ Linguistik \ Literaturgeschichte \ Anglistik \ Bauwesen \ Fremdsprachendidaktik \ DaF \ Germanistik \ Literaturwissenschaft \ Rechtswissenschaft \ Historische Spra issenschaft \ Slawistik \ Skandinavistik \ BWL \ Wirtschaft \ Tourismus \ VWL \ Maschinenbau \ Politikwissenschaft \ Elektrotechnik \ Mathematik & Statistik \ Management \ hilologie \ Sport \ Gesundheit \ Romanistik \ Theologie \ Kulturwissenschaften \ Soziologie \ Theaterwissenschaft \ Linguistik \ Literaturgeschichte \ Anglistik \ Bauwese remdsprachendidaktik \ DaF \ Germanistik \ Literaturwissenschaft \ Rechtswissenschaft \ Historische Sprachwissenschaft \ Slawistik \ Skandinavistik \ BWL \ Wirtschaft \ Touris VWL \ Maschinenbau \ Politikwissenschaft \ Elektrotechnik \ Mathematik & Statistik \ Management \ Altphilologie \ Sport \ Gesundheit \ Romanistik \ Theologie \ Kulturwiss chaften \ Soziologie \ Theaterwissenschaft \ Geschichte \ Spracherwerb \ Philosophie \ Medien- und Kommunikationswissenschaft \ Linguistik \ Literaturgeschichte \ Anglisti auwesen \ Fremdsprachendidaktik \ DaF \ Germanistik \ Literaturwissenschaft \ Rechtswissenschaft \ Historische Sprachwissenschaft \ Slawistik \ Skandinavistik \ BWL \ Wirtsc BUCHTIPP Nicole Dörr, Carsten Gachot, Max Marian, Katharina Völkel 24th International Colloquium Tribology Industrial and Automotive Lubrication Conference Proceedings 2024 1. Auflage 2024, 279 Seiten €[D] 148,00 ISBN 978-3-381-11831-1 eISBN 978-3-381-11832-8 expert verlag - Ein Unternehmen der Narr Francke Attempto Verlag GmbH + Co. KG Dischingerweg 5 \ 72070 Tübingen \ Germany Tel. +49 (0)7071 97 97 0 \ Fax +49 (0)7071 97 97 11 \ info@narr.de \ www.narr.de The conference provides an international exchange forum for the industry and the academia. Leading university researchers present their latest findings, and representatives of the industry inspire scientists to develop new solutions. Main Topics > Trends lubricants and additives > Automotive and transport industry > Industrial machine elements and wind turbine industry > Coatings, surfaces and underlying mechanisms > Test methodologies and measurement technologies > Digitalisation in tribology > Digital Tribological Services: i-TRIBOMAT > Sustainable lubrication Target Groups > Companies in the field of lubrication, additives and tribology > Research facilities Introduction The design of seals aims predominantly for service life and efficiency. The mineral oil-based and synthetic lubricants used are always associated with the risk of contamination, for example through leakage from gearboxes via seals, or via gear drives of conveyor systems, as well as through grease from bearings and pumps. Therefore, a minimum leakage and oil contamination is unavoidable. In addition, both dynamic and static seals for demanding fields of application are mostly based on fluoropolymers which will be affected by the possible ban of perfluorinated and polyfluorinated alkyl substances (PFAS). As with oil, wear of fluoropolymers can result in an undesirable contamination of the environment with wear particles. One solution being pursued is the use of water-based lubricated systems. However, due to the increased tendency of elastomers to build up a high level of adhesion without an oil-lubricated film, friction losses, e.g. on shafts and fatigue wear of the seal, must be avoided. For these applications, we investigate the influence of the choice of lubricant on the friction pairing in the form of friction changes and changes in the wear morphology and structure formation on the surface in model tests on sheet material. The wetting behavior of lubricants and base oils of greases is an essential information that can be used to correlate both the spreading behavior of the lubricant and the change in surface elasticity due to swelling by sorption. The surface elasticity and the spreading energy are the known characteristic values that determine the wave-like sliding and the tendency to form structure sizes of wear particles and crack structures. By investigating the tribological behavior of e.g. HNBR elastomers with various lubricating fluids, ranging from water-based mixtures to polyalkylene glycols and greases, the sensitive selection of advantageous systems and the mechanistically clear distinction between swelling and non-swelling systems becomes apparent. The adhesive friction of polymers arises from the continuous forming and breaking of adhesive bonds between the two gliding partners. This process was described by Schallamach as a thermally driven rate process. Schallamach’s model [1] for adhesive friction results in a coefficient of friction (COF) which initially grows monotonously with increasing gliding velocity. The COF reaches a maximum value at a specific gliding velocity and decreases if the gliding velocity is increased further. Since Schallamach’s approach is based on Science and Research 5 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 unikationswissenhe Sprachwissenent \ Altphilologie Kommunikationsistorische Sprachanagement \ Alttik \ Bauwesen \ schaft \ Tourismus ie \ Kulturwissenichte \ Anglistik \ \ BWL \ Wirtschaft The sensitive dependence of sealing elastomers on lubricants Christof Koplin, Bernadette Schlüter, Raimund Jaeger* submitted: 19.09.2024, accepted: 24.03.2025 (peer review) Presented at GfT Conference 2024 * Dr.-Ing. Christof Koplin Dr. Bernadette Schlüter Dr. Raimund Jaeger Fraunhofer-Institut für Werkstoffmechanik, Mikrotribologiezentrum µTC, Freiburg The design of elastomer seals for bearings and pumps requires the selection of the friction partner and lubricant. This study investigates the influence of swelling and spreading behavior on the tribological behavior of NBR systems. Interaction tendencies are described by the energies of interaction and spreading, which are determined by means of contact angle measurements. Intensive swelling was observed in elastomers with a high tendency to solve or interact. The elasticity and spreading energy are decisive for the formation of the elastic length of the adhesive contact. Highly deformable elastomers form an adhesive contact whose length influences the friction. If the elasticity of swollen elastomers is low, friction and wear increase. For elastomers with low lubricant absorption, the spreading tendency determines the friction. A characteristic speed shows at which a transition to forced wetting occurs. Polar lubricants as well as water-based lubricants show similar dependencies to dry systems, which still needs to be verified. For greases, the tendency to spreading and loosening was calculated for base oils, with non-spreading contacts predominating. In thermoplastic systems, the swelling tendency is dominant. Keywords seal, elastomer, lubricant, wear, friction, spreading, swelling, substitution, water-based Abstract amongst others, by Wu-Bavouzet et al. [8]. The effect of surface elasticity, interfacial energies, and viscosity on the formation of continuous or stick-slip dominated gliding was described. At low speeds, thermal fluctuations determine the formation of the adhesive bond and produce continuous gliding. At higher speeds, a minimum of the frictional force versus velocity occurs due to energetic equality of elastic strain and adhesion. High velocities lead to a cyclic tearing-off from the adhesively formed contact and a reduction of the pronounced contact surface. As a result, a slip movement is observed on the surface until a bonding or stick phase occurs again. The elastic length that is overcome until a slip phase occurs is described again by the ratio of interfacial energies and shear elasticity. The ratio of this length and the mean slip velocity is a characteristic time that is important for deriving further mechanisms. To do this, Wu-Bavouzet uses analytical approximations of the crack process of the adhesive interface. The main result is a model for the coupling of wavelength, crack length and contact length, wave velocity, release frequency and sliding velocity to the energy release rate for a gliding step at the interface. The descriptions were obtained for linear elasticity, but Wu-Bavouzet assigns universal validity to the obtained dependencies. Several other authors as well described the separation of the interface by spreading and as a function of viscosity. The adhesion and friction of bio-inspired nanoand microtextured liquid silicone rubber (LSR) surfaces produced by injection molding were analyzed [9]. Different methods were used to create anisotropic structures. The structures created in this way on a surface of LSR have been proven to have a tribological effect. Nanostructures can increase the adhesive friction of surfaces during sliding, increase the wetting tendency of liquids and decrease adhesive friction. Meso-structures increase the movement amplitude of stick-slip and decrease dry friction, absorb shear deformation at low velocities, reduce film formation during lubrication and increase friction. Overall, the structures enabled frictional anisotropy. When elastomer substrates were laser microstructured and subsequently coated with DLC, the coated samples showed a significantly reduced run-in friction compared to the uncoated ones. On NBR samples, a “tile patterned” substrate could efficiently suppress any further cracking of the DLC-coating except in the predisposed trenches. The structuring in the coated as well as in the uncoated state helped to lower the CoF and reduce wear. This effect partly wore off over time and was dependent on the surface structure [10]. Our aim is to understand the effects or sensitivity of the lubrication pairing with these elastomers and to investigate the effects of structures and coatings in demonstrative cases. Science and Research 6 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 an energy-based description of the slip rate for adhesive processes between the frictional partners, their surfaceand interfacial energies can give some insight into the influence of the lubricant on the adhesive friction. A good correlation was found for wetting and spreading parameter of lubricated self-pairing contacts for a wide range of lubricants [2, 3]. Quantities like the work of spreading W spreading and the work of solving W solving can be calculated with surfaceand interfacial energies [4]. The determination of surface and interfacial energies of the frictional partners is typically based on contact angle measurements. The work of adhesion can be determined, for instance, with the Young-Dupree equation by measuring the contact angle between lubricant and polymer. An alternative approach described by Owens is to determine the work of adhesion using appropriate averages of the polar and dispersive components of the surface free energy for the polymer and the cohesive energy of the liquid. The latter is also determined with contact angle measurements, but possibly using several other liquid-solid combinations than the one for which the work of adhesion is calculated. If there is a strong interaction between the lubricant and the polymer, both approaches can yield different results. The ratio of interfacial energy determined by Owen’s approach and by a contact angle measurement with the pairing of lubricant and polymer (denoted as W solving) is called the “interaction parameter”. An interaction parameter greater than unity indicates a strong interaction between polymer and lubricant. To sum up, if a dry contact of two material is formed the work of adhesion has to overcome if a liquid invades, wets, and separates the contact faces. The change in energy of these states is the spreading energy. Pairings of semicrystalline thermoplastic polymers with lubricants were studied to differentiate between tribological effects due to sorption and plasticization and those due to spreading. For mild loading in mixed lubricated conditions against technical steel surface, friction and wear properties seem to be primarily determined by the hardness-dependence of abrasive contact and less by adhesion or hysteretic mechanisms [5]. When base oils are transformed to greases using Li as thickener, the increased spreading tendency of the base oils results in an increased stiction. A possible explanation for this behavior is that an increased tendency of spreading of the base oil increases the bleeding tendency of the grease. The bleeding and recovery behavior of the thickener-rich layer was described by Zhang et. al [6]. As a result, a thickener with a degraded rheology is obtained, thus an increase in grease lubricated stiction can be found [7]. The basic mechanisms of friction of elastomers in contact with smooth surfaces was systematically studied, Methods Materials Thermally pressed plates of 2 mm thickness of different NBR elastomer were investigated (Table 1). The listed lubricants were chosen (Table 2). Tribological Testing Ball-on-3-plates tests with a 1.3505 steel ball (diameter ½”, R a = 0.15 µm: R z = 1.4 µm, R ku = 3.85, R Pc = 37/ cm) gliding on 3 elastomer plates of 2 mm thickness were used to carry out oscillatory tribological experiments [11]. In these experiments, a polished steel ball loaded with 2 N is brought into spherical contact. The plates are separated by 120° around the rotating axis and have an inclination of 45° to this axis. Therefore the normal force on each plate is 2 0.5 / 3 times the tribometer load. The formed contact line on the ball has half diameter of the ball itself. Using the normal load and distance to the axis the coefficient of friction can then be calculated from the measured frictional moment. The contact pressure is limited and defined by hyperelastic behavior of the elastomer in the range of 0.1 to 1 MPa, which was verified by the diameter of contact marks. The amplitude of the velocity and deflections can be varied over a wide range, starting at very low velocities (1.3 ×10 -4 mm/ s - 3.3 mm/ s). To achieve stable and reproducible system conditions, a testing sequence was applied to form a run-in situation. Conducting the oscillatory experiments in this range of sliding speeds (oscillatory mode) makes it possible to record friction data in the static friction and boundary-lubrication regimes. Using a data acquisition rate of 10 data points per decade, the loading torque was increased on a logarithmic scale at an oscillation frequency of 1 Hz. 3-ball-abrasion test with 1.4125 steel balls (diameter ½”, R a = 0.15 µm: R z = 1.4 µm, R ku = 3.85, R Pc = 37/ cm) gliding on elastomer plates of 2 mm thickness dry and with lubricant. 3-ball-abrasion test were used to induce surface fatigue in elastomers. Three rigidly mounted steel balls were loaded with 30 N and rotated in a circular track at a velocity of 56 mm/ s on the elastomer surface. After a wear track becomes visible, the wear can be quantified by measuring the weight loss of the specimen. Measurements were done for 0.5 h at room temperature and stable run-in conditions were achieved for the last 10 min. The run-in friction and wear rate are calculated for the last 10 min. Ring-on-disc tests with axial bearing rings INA AS1528 made of 1.3505 of technical roughness (diameter of track 12.5 mm, R a = 0.1 µm, R z = 0.8 µm, R Pc = 64/ cm) were used for gliding contact of elastomer rings glued on INA AS0821 with 9.2 mm - 21 mm diameter and 2 mm height. A load of 0.3 MPa at 148 RPM (0,14 m/ s) was used. Additional Tests und Analysis Contact Angle Measurements The sessile drop method placing 10 times 2 µL droplets on the samples was used by a contact angle measurement device produced by Data Physics GmbH (figure 1). Three standard liquids (water, ethylene glycol and diiodomethane) were used to characterize the unknown surface and interfacial energies. The droplet profile is recognized and recorded by the software of the instrument, and the contact angle is calculated automatically. Science and Research 7 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 NBR-H-A-NT-68 Goorex Automotive GmbH & Co. KG Westerrönfeld, Germany HNBR-2513-80 Festo SE & Co. KG Esslingen, Germany HNBR-DX107-4 NBR-804506.1 Freudenberg Weinheim, Germany NBR-802607.2 pPG: polar polyalkylenglycole = 30 mm 2 / s at 40 °C Klüber Lubrication München SE & Co. KG uPG: non-polar polyalkylenglycole = 30 mm 2 / s at 40 °C Ester: ester oil = 23.5 mm 2 / s at 40 °C PAO: polyalphaolefin = 30 mm 2 / s at 40 °C pPG Ligrease uPG Ligrease Ester Ligrease PAO Ligrease EG: ethylenglycole = 17 mPa.s at 25 °C Omnilab (Bremen, Germany) Gly: glycerol = 1412 mPa.s at 20 °C Table 1 Table 2 in the lubricants and stored in ovens at the temperature of interest and weighed again (here: 100 °C). Results and Discussion It is common knowledge that for dry contact of elastomers with a frictional partner, the elastic length of the interface is proportional to the ratio of the work of adhesion and the elasticity of the elastomer. The work of adhesion can be calculated as W 12 by the surface energy of each partner. For simplicity, the elasticity is replaced by the shore hardness. As a result, the wear rate was found to be increased for a longer elastic length (figure 2). An increase in the extent of surface fatigue by an increase of localization of shear load by increased elastic length is the driving force for an increased wear rate. This assumption was confirmed for the tested elastomers. For lubricated elastomers in boundary lubrication, the work of adhesion must be replaced by the Work of spreading, W spreading . For the system with a higher spreading tendency, spreading of lubricant or bleeding of base lubricant for grease is likely to occur. The hypothesis is that oil or bled base oil is sorbed and does swell the elastomer. Science and Research 8 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 Measurement of Shore Hardness. Indentations were performed with a Bareis HPE II and the value shore A was obtained. Mass Uptake To measure the uptake of the lubricants by the elastomers, elastomeric specimens with a thickness l of 2 mm were first pre-dried at a temperature of 100 °C and a gas pressure less than 3 mbar until a weight equilibrium was reached. The specimens were then weighed, immersed Figure 1: Contact angle of a sessile drop of liquid on different elastomers Figure 2: correlation of wear at dry gliding contact of different NBR elastomer with elastic length Consequently, the swelling is followed by a decrease of the elastomer elasticity and an increase of elastic length and friction. In case of grease, a sorption of base oil into the elastomer also degrades the grease viscosity. In both cases, the friction or stiction (static friction) and wear rate would increase if there were a high tendency to spread (W spreading < = 0 mN/ m) (figure 2). When choosing the onset of gliding motion and measuring stiction there is no effect of viscosity and a comparison of systems with different types of lubrication is possible. For polar lubricants as ethylene glycole, the spreading behavior is reduced and consequently friction rises. It is defined in literature that for a reduction in spreading W spreading is increasing. Being highly polar, W spreading for water is highest, and since the viscosity is very low, it is squeezed out of the contact. A squeezed-out contact is formed by a pure elastomer steel interface. For a spheric contact a transition velocity was calculated [9] for a transition between squeeze-out and forced wetting to be v = (l/ R) 1/ 3 W spreading / viscosity (l: contact length, and R: contact radius). For ethylene glycole and water, friction decreases by forced wetting by gliding velocities beyond the characteristic velocity. This velocity-dependent transition behavior of polar lubricants will lead to high breakaway moments and massive occurrence of stick-slip phenomena in applications. For the lubricated ring on disc tests, only friction can be measured (figure 4) but no wear data were obtained. HNBR-2514-80 is non-spreading in oil which is reflected in the linearly increasing friction values for the different oils. In contrast, NBR-H-A-NT-68 and oil is a spreading system, but its friction values are higher due to the increased solving tendency. Of these 4 systems, the higher friction of NBR-H-A-NT-68 with three oils compared to the fourth oil (PAO) is a result of the fact that friction increases due to an increased tendency to solve. In summary, it can be said that friction is determined by its ten- Science and Research 9 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 Figure 4: Correlation of run-in friction at lubricated contact of different elastomer using different lubricant types with spreading and solving energy Figure 3: Correlation of stiction at lubricated contact of elastomer using different lubricant types with spreading energy high tendency to spread increases the tendency to bleed out base oil from the pressurized contact, it is sorbed into the elastomer. The swelling goes along with a reduction in shore hardness A. A mass uptake of 14 % for pPG was followed by a reduction in hardness by 28 %. The swelling of the elastomer together with the possible reduction of the lubrication capability of the thickener richened phase changes the friction of the contact. If for a system (e.g. HNBR-2513-80) low spreading and no swelling is observed, then friction is decreased with increasing spreading tendency. Conclusion The design of elastomer seals for bearings and pumps should include an appropriate selection of the frictional partner and lubricant. For the lubrication of elastomers, a significant influence of the swelling and spreading behavior on the tribological behavior was found for the investigated NBR systems. Depending on these interaction tendencies, changes in friction and wear were obtained. Science and Research 10 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 dency to dissolve and, if not present, by its tendency to spread. In the abrasion test with lubricated steel balls, only the friction can be measured (figure 5), as no wear was detected here either. In these experiments, a different steel type with a slightly different surface energy was used, hence solving and spreading energies of the oils differ. The loading, lubrication and contact pressure is as well different from the ring on disc test, but the dependence on energies follows similar rules. The system NBR-H- A-NT-68 with uPG is non-spreading in oil but it is solving, and as a result the friction is obtained higher than for the 3 other systems, both for oil and for grease. For the other 3 oils or greases the contact with NBR-H-A- NT-68 is spreading. In contrast to oil lubrication, the wear could be measured for grease lubrication (figure 6). Here, the friction and wear of elastomers is sensitive to the interaction with the base oils. The tribological behavior is determined by the spreadingor swelling-tendency of the systems. When a Figure 6: Correlation of run-in wear for lubricated gliding contact of elastomers using 4 different oil-based greases with spreading energy and the lubricant uptake when stored at 100 °C Figure 5: Correlation of run-in friction of lubricated contacts for different elastomers using different lubricant types ( uPG) with spreading and solving energy × The interaction tendencies can be calculated by the energies W solving and W spreading which can be efficiently determined by contact angle measurements. For elastomers with a high solving tendency, intensive swelling was found. The shore hardness of swollen elastomer was measured in order to quantify the effect of lubricant absorption. The elasticity and the spreading energy are the main properties that are relevant for the formation of the elastic length of adhesive contact. Highly deformable elastomers form an adhesive contact with a specific length in the test before a slip process occurs. This can be described in a first approximation with an energetically based elastic length, for which the deformation energy and adhesion energy are in thermodynamic equilibrium. Accordingly, the proportion of friction increases with the dissipated adhesion energy and the elastic length. For the lower elasticity of swollen elastomers, the measured friction was immensely increased. Consequently, wear was also increased. If the elastomer absorbs little lubricant, the tendency to spread determines the friction. A high spreading tendency would result in a reduction in elastic length and therefore, a reduced friction and wear was found. Since the spreading energy is calculated for a static situation the description must be completed for a dynamic situation and the effect of lubricant viscosity. A characteristic velocity can be calculated for which a transition to forced wetting or squeeze-out occurs. One simple rule is that a constantly high friction occurs for non-spreading systems at velocities below the characteristic velocity. Polar lubricants are highly non-spreading. It can be assumed that the same dependence of friction and wear on the elastic length exists for these systems as for dry systems, but this still needs to be verified. For greases, the tendency to spread and solve was calculated for their base oils. For elastomers and grease lubrication, the measurable dependence on the spreading tendency is visible, as the base oils predominantly produce non-spreading contact. For thermoplastic systems, the tendency to swell is usually dominant, as these systems produce a predominantly spreading contact. References [1] Schallamach, A. A theory of dynamic rubber friction. Wear 1963, 6, 375-382, doi: 10.1016/ 0043-1648(63)90206-0. [2] Schertzer, M.; Iglesias, P. Meta-Analysis Comparing Wettability Parameters and the Effect of Wettability on Friction Coefficient in Lubrication. Lubricants 2018, 6, 70, doi: 10.3390/ lubricants6030070. [3] Kalin, M.; Polajnar, M. The Effect of Wetting and Surface Energy on the Friction and Slip in Oil-Lubricated Contacts. Tribol Lett 2013, 52, 185-194, doi: 10.1007/ s11249-013-0194-y. [4] Israelachvili j.n. Intermolecular and Surface Forces; Elsevier, 2011, ISBN 9780123751829. [5] Koplin, C.; Oehler, H.; Praß, O.; Schlüter, B.; Alig, I.; Jaeger, R. Wear and the Transition from Static to Mixed Lubricated Friction of Sorption or Spreading Dominated Metal-Thermoplastic Contacts. Lubricants 2022, 10, 93, doi: 10.3390/ lubricants10050093. [6] Zhang, S.; Klinghart, B.; Georg Jacobs, G.; von Goeldel, S.; König, F. Prediction of bleeding behavior and film thickness evolution in grease lubricated rolling contacts, Tribology International, Volume 193, 2024, doi: 10.1016/ j.triboint.2024.109369. [7] Koplin, C.; unpublished results [8] Wu-Bavouzet, F.; Clain-Burckbuchler, J.; Buguin, A.; Gennes†, P.-G. de; Brochard-Wyart, F. Stick-Slip: Wet Versus Dry. The Journal of Adhesion 2007, 83, 761-784, doi: 10.1080/ 00218460701586178. [9] Koplin, C.; Weißer, D.F.; Fromm, A.; Deckert, M.H. Stiction and Friction of Nanoand Microtextured Liquid Silicon Rubber Surface Formed by Injection Molding. Appl. Mech. 2022, 3, 1270-1287. https: / / doi.org/ 10.3390/ applmech3040073 [10] Vogel, S.; Brenner, A.; Schlüter, B.; Blug, B.; Kirsch, F.; Roo, T.v. Laser Structuring and DLC Coating of Elastomers for High Performance Applications. Materials 2022, 15, 3271. https: / / doi.org/ 10.3390/ ma15093271 [11] Koplin, C.; Abdel-Wahed, S.A.; Jaeger, R.; Scherge, M. The Transition from Static to Dynamic Boundary Friction of a Lubricated Spreading and a Non-Spreading Adhesive Contact by Macroscopic Oscillatory Tribometry. Lubricants 2019, 7, 6, doi: 10.3390/ lubricants7010006. Science and Research 11 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0007 Friction and wear are systemic properties that extend beyond specific material characteristics, playing a key role in determining the longevity, durability, and overall functionality of dental implants. The interplay of these factors can affect the ability of the implant to withstand mechanical stress, resist degradation, and maintain its structural integrity over time. Therefore, a comprehensive understanding of these interactions is essential for optimizing implant design, improving material choices, and ensuring the long-term success and reliability of dental implants. Figure 1 illustrates the various tribological contacts present in dental implants, including those between the abutment and the restoration (1), between the implant and the abutment (2), within the structure of the implant itself (3), and at the interface between the implant and the surrounding bone tissue (4). In this study, the tribo- Science and Research 12 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 1 Introduction Dental implants play a crucial role in restoring oral function and must meet specific requirements to ensure their effectiveness and longevity. To successfully replace missing teeth, the materials used for dental implants must possess high strength in combination with a moderate Young’s modulus, ideally ranging from 70 to 80 GPa. This balance helps prevent the overloading of newly formed bone cells while simultaneously minimizing stress shielding over the long term, which could otherwise compromise the stability of the implant. Furthermore, dental implant materials must provide biocompatibility and corrosion resistance to reduce the risk of implant rejection. In addition, the materials must exhibit sufficient wear resistance to prevent metal particle formation and the accumulation of wear debris, which in turn can result in material degradation and inflammatory reactions in adjacent tissues. This raises the risk of complications such as peri-implantitis and bone loss, potentially resulting in complete implant loss.[1] Understanding tribological interactions, particularly friction and wear, is crucial for evaluating the performance of dental implants within the dynamic and complex conditions of the oral environment. These interactions are significantly influenced by various factors such as material pairings, loading conditions, environmental variables as well as the presence and type of lubrication. Tribological Investigation of the Novel Titanium Alloy TNTZ-O for Dental Implant Applications: Subsequent Results of a Comparative Study Alexander Roegnitz, Elias Merz, Hubert Mantz, Carsten Siemers, Andreas Haeger* submitted: 27.09.2024 accepted: 16.04.2025 (peer review) Presented at GfT Conference 2024 To date, the titanium alloy Ti-36Nb-2Ta-3Zr-0.3O (TNTZ-O) has not been used in class III medical products, despite its favorable properties for implant applications, including high strength, low stiffness, high elastic strain, and good biocompatibility. To assess the performance and longevity of dental implants within the dynamic oral environment, it is vital to understand its tribological interactions. This study examines the frictional behavior and wear characteristics of TNTZ-O in artificial saliva against 100Cr6 and compares it to established titanium alloys, specifically CP-Ti Grade 4 and Ti-6Al-4V ELI. The findings indicate that TNTZ-O exhibits higher wear volume and frictional instability, indicating the need for further optimization for dental implant applications. Keywords TNTZ-O, Tribology, Dental implants, Wear, Biomaterials, Titanium alloys Abstract * Alexander Roegnitz, M. Eng. 1 (corresponding author) Elias Merz, M. Sc. 2 Prof. Dr. rer. nat. Hubert Mantz 1 Carsten Siemers 2 Prof. Dr.-Ing. Andreas Haeger 1 1 Ulm University of Applied Sciences, Institute for Manufacturing Technology and Materials Testing, 89075 Ulm, Germany 2 Technische Universität Braunschweig, Institute for Materials Science, 38106 Braunschweig, Germany logical aspects related to the contact area between the implant and the abutment (2) are investigated. For the last three decades, “Commercially Pure Titanium” (CP-Ti) has been extensively used in dental implants due to their favorable corrosion resistance and good mechanical properties. In particular, CP-Ti Grade 4 has been preferred due to its superior strength and reliability. However, the mechanical properties of CP-Ti are still limited and its relatively poor tribological performance highlights the need for more advanced alternatives. [2] To address these limitations, various alloying elements have been incorporated into titanium-based materials to improve their mechanical and tribological properties. For instance, Ti-6Al-4V ELI (Titanium Grade 23) is highly valued for its enhanced strength and corrosion resistance. However, the potential cytotoxic effects associated with aluminum and vanadium ions in this alloy raise concerns about its long-term safety and suitability for implants [3]. Recent advancements have led to the development of low-modulus Ti-based alloys with higher contents of β-stabilizing elements, free of aluminum and vanadium, such as Ti-13Nb-13Zr (TNZ), Ti-15Mo and Ti-36Nb-2Ta-3Zr-0.3O (TNTZ-O). These alloys show considerable promise due to their high strength, low Young’s modulus, and absence of cytotoxic elements, making them an attractive alternative for dental applications [4]. Among these so-called second and third generation medical titanium alloys, the latter exhibits the lowest Young’s modulus as well as high strength. While these properties are promising, further research is necessary to fully evaluate the tribological behavior and overall suitability of TNTZ-O for dental implants and other medical applications. This study investigates the frictional behavior and wear characteristics of TNTZ-O and compares it to established titanium alloys, specifically CP Ti Grade 4 and Ti-6Al-4V ELI. 2 Experimental methods 2.1 Materials TNTZ-O is classified as a third-generation titanium alloy specifically designed for class II medical applications, i.e. orthopedic wires. Depending on its specific chemical composition, the microstructure can comprise a range of phases, including α-, β-, α"and ω-phases [4, 5]. Currently, this alloy is manufactured by powder metallurgy and sintering, with subsequent rotary swaging, and wire drawing. This manufacturing process restricts its applications to wires with diameters less than 2 mm. To enable future dental implant applications as a Class III medical product, a conventional manufacturing route involving ingot production followed by deformation processes is currently under development by one of our project partners. The TNTZ-O material used in the present study was obtained in the “as received” (AR) state from Toyota Tsusho Material Incorporated, where it had been roll-leveled from a 7.3 mm diameter coil. The material exhibits a β-phase microstructure, which is a critical factor in its performance characteristics. For comparative Science and Research 13 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 Figure 1: Relevant tribological contact areas in dental implants Material Ti Nb Ta Zr O Al V Fe CP-Ti Grade 4 bal. - - - 0,30 - - 0,08 Ti-6Al-4V ELI bal. - - - 0,11 6,08 3,99 0,14 TNTZ-O (AR) bal. 36,2 ± 0,1 1,9 ± 0,1 3,0 ± 0,1 0,21 ± 0,01 - - - Table 1: Chemical composition of the tested materials (in wt.%) Material Tensile strength (MPa) Young’s Modulus (GPa) Vickers Hardness (HV5) Phases TNTZ-O (AR) 896 ± 23 65 ± 3 253 ± 7 β CP-Ti Grade 4 835 ± 8 105 ± 7 255 ± 7 α Ti-6Al-4V ELI 1057 ± 9 110 ± 6 312 ± 9 α + β Table 2: Mechanical properties and phases of the tested materials β-phase regions concentrated at the grain boundaries. The microstructural features are illustrated in Figure 2. 2.2 Tribological testing and characterization The tribological behavior of the materials was evaluated utilizing an oscillating ball-on-disc SRV ® III-Tribometer from Optimol, conducting tests in artificial saliva at a controlled temperature of 37 °C to simulate the conditions in the oral environment. The Aldiamed mouth spray, produced by Certmedica Int. GmbH was selected as the saliva substitute due to its viscosity and lubricating properties, which closely resemble those of natural saliva. The tribological samples were initially sectioned into disks with a height of 3 mm. These disks underwent a preparation process involving grinding and polishing to ensure comparable and uniform surface properties. The final surface finishing was conducted using a 50 nm alkaline colloidal silica suspension from Cloeren Technology GmbH, which provided a mirror-like surface and minimized the impact of surface roughness (Ra: 0,013 - 0,049 µm) on the measurements of friction and wear. A hardened 100Cr6 - G20 steel ball (HRC: 62) with a diameter of 10 mm was selected as the counter body for the tests. The specific load parameters and experimental conditions that were used during the tests are summarized in Table 3. Science and Research 14 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 analysis, the chemical composition and mechanical properties of TNTZ-O, CP-Ti Grade 4, and Ti-6Al-4V ELI are detailed in Tables 1 and 2. The CP-Ti Grade 4 and Ti-6Al-4V ELI alloys, used as reference materials in this study, were supplied by from Klein SA. These reference alloys have a diameter of 8 mm and conform to ASTM F-67-13 and ASTM F-136 standards, respectively [6, 7]. The microstructures of TNTZ-O, CP-Ti Grade 4, and Ti-6Al-4V ELI were analyzed using optical microscopy (OM) and scanning electron microscopy (SEM). Prior to imaging, the samples were ground and polished to obtain smooth surfaces. TNTZ-O and CP-Ti Grade 4 were etched with a custom solution composed of 86 ml H 2 O, 12 ml H 2 O 2 , 5 ml HF and 4,5 ml HNO 3 while Ti-6Al-4V ELI was etched using Kroll’s reagent. After etching, the samples were examined with a Zeiss Axio Imager.M2m optical microscope and analyzed using AxioVision 4.8 software. Additionally, the Hitachi TM3000 desktop SEM was used to acquire images with backscattered electron (BSE) contrast. TNTZ-O exhibited large, globular β-phase grains, with apparent deformation features, presumably twinning. CP-Ti Grade 4 is characterized by a uniform structure consisting mostly of equiaxed α-phase grains. In contrast, Ti-6Al-4V ELI revealed a fine-grained, globular microstructure, with a duplex arrangement of equiaxed α-phase grains and Figure 2: Microstructural images of TNTZ-O, CP-Ti Grade 4, and Ti-6Al-4V ELI. The top row includes optical microscopy (OM) images: TNTZ-O (A), CP-Ti Grade 4 (B), and Ti-6Al-4V ELI (C). The bottom row displays scanning electron microscopy (REM) images with backscattered electron (BSE) contrast: TNTZ-O (A'), CP-Ti Grade 4 (B'), and Ti-6Al-4V ELI (C'). Load (N) Duration (s) Frequency (Hz) Stroke (mm) Temperature (°C) 20 2520 1 2 37 Table 3: Load collective parameters for tribological testing The coefficient of friction (COF) was recorded, providing a detailed analysis of frictional behavior and tribological stability during reciprocating sliding in an artificial oral environment. For each material, three comparable test runs were conducted to ensure reliable and consistent results. Post-test examination of the wear tracks was performed using the μsurf #1018 confocal microscope from NanoFokus, equipped with an Olympus LMPLFL20x/ 0.40 objective. The captured images were then processed and analyzed using Mountain 8 software from DigitalSurf to determine the wear volumes. Further investigations of the present wear mechanisms involved SEM using a Zeiss Sigma 300 VP. Elemental analysis of the wear tracks was conducted using Energy Dispersive X-ray Spectroscopy (EDX) with a Bruker XFlash 6|60 detector, identifying the elemental composition and distribution within the wear tracks. 3 Results and discussion 3.1 Friction behavior The coefficient of friction graph, as depicted in Figure 3, illustrates the distinct tribological behavior of TNTZ-O, CP-Ti Grade 4 and Ti-6Al-4V ELI throughout the 42-minute test duration, revealing their varying frictional responses and wear characteristics. After an initial run-in phase, the friction values of TNTZ-O stabilize within the range of 0,25 to 0,28, reaching COF comparable to those observed in the reference materials. However, these stabilized values are intermittently disrupted by significant friction peaks, with COF values occasionally exceeding 0,8, indicating periods of instability during sliding. Although the material’s friction behavior intersects with that of the other reference materials at various points, it ultimately diverges, displaying a slight upward trend toward the end likely due to the accumulating friction fluctuations. In contrast, CP-Ti Grade 4 and Ti-6Al-4V ELI exhibited stable COF trends throughout the test. Both materials settled quickly after a brief run-in phase of approximately 180 to 240 seconds, with Grade 4 maintaining values around 0.21 to 0.23. Meanwhile, the COF of Ti-6Al-4V ELI gradually decreased, ultimately finishing slightly lower than Grade 4 by the end of testing. The low standard deviation (SD COF ) values in Table 4 demonstrate the consistent and stable frictional performance of CP-Ti Grade 4 and Ti-6Al-4V ELI, especially when compared to the fluctuating and unstable frictional behavior observed in TNTZ-O. 3.2 Wear behavior TNTZ-O exhibited the highest wear volume at 0,098 mm 3 after the full test duration of 42 minutes, which corresponds with significant fluctuations in its COF. This indicates increased adhesive wear and delamination of the forming oxide layer. The observed peaks in the COF suggest stick-slip behavior, where temporary adhesion or interfacial locking contributes to the wear mechanisms. An example of a wear track for TNTZ-O is shown in Figure 4A, featuring an oval-shaped pattern that illustrates the observed wear. The increased wear standard deviation (SD wear ) for TNTZ-O of 0,0164 mm 3 , reflects the unstable friction behavior of TNTZ-O. These fluctuations in friction lead to inconsistent sliding con- Science and Research 15 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 Figure 3: Exemplary COF curves for TNTZ-O, CP-Ti Grade 4 and Ti-6Al-4V ELI tested in reciprocating ball-on-disc tribometer for 42 min in artificial saliva against hardened 100Cr6. within the contact area, as illustrated in Figure 5A, supports this observation. The oxidation of the surface and subsequent delamination of the oxide layer seems to be the primary causes for the COF fluctuations and the pronounced friction peaks observed during testing. This adhesive wear mechanism is further verified by the EDS measurements of the 100Cr6 steel balls, as depicted in Figure 6B. These analyses reveal the presence of titanium (Ti) and niobium (Nb) from the TNTZ-O alloy and provide evidence of notable material transfer between the disc and the steel balls. While the precise wear volume of TNTZ-O on the ball surface could not be accurately quantified, the detected elements strongly suggest substantial material detachment and re-adhesion during sliding. In contrast, the SEM images of CP-Ti Grade 4 and Ti-6Al-4V ELI, presented in Figure 5B and Figure 5C respectively, reveal wear tracks with fine grooves and ridges. These features are indicative of abrasive wear, suggesting that the wear process for these alloys primarily involves micro-cutting or micro-plowing of the surface material, with significantly less material transfer and delamination compared to TNTZ-O. This results in a more uniform and consistent friction behavior, characterized by fewer fluctuations or peaks in friction and minimal adhesive wear. Consequently, these alloys exhibit stable tribological performance, highlighting their robustness and reliability under the tested conditions. This stable performance is consistent with existing literature, Science and Research 16 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 ditions, resulting in increased variability in wear volume. Conversely, CP-Ti Grade 4 and Ti-6Al-4V ELI demonstrated considerably lower wear volumes, recorded at 0,012 mm 3 and 0,017 mm 3 , respectively. Despite Ti-6Al-4V ELI having a lower final COF than CP-Ti Grade 4, it exhibited a slightly higher wear volume. Nevertheless, both materials displayed overall comparable tribological performance. Table 4 presents a summary of the tribological data, including the average COF after a 10-minute running-in phase, the standard deviation (SD) of COF, minimum (Min COF ) and maximum (Max COF ) COF values, and the corresponding wear volumes. To further investigate the observed wear behavior, a detailed SEM analysis of the wear tracks was conducted, as illustrated in Figure 5. This investigation revealed distinct wear mechanisms for each alloy, providing additional insight into their tribological performance. In the case of TNTZ-O, the SEM images (Figure 5A) illustrate a wear track characterized by larger flakes of material compared to those observed in CP-Ti Grade 4 (Figure 5B) and Ti-6Al-4V ELI (Figure 5C). This indicates a dominant adhesive wear mechanism, where material is torn from the surface, re-adheres, or undergoes delamination. The presence of a heavily oxidized region Figure 4: Exemplary wear track on TNTZ-O disc using confocal microscopy (A). Average wear volumes for TNTZ-O, CP-Ti Grade 4 and Ti-6Al-4V ELI (B) tested in reciprocating ball-on-disc tribometer for 42 min in artificial saliva against hardened 100Cr6. Material COF (Avg. after 10 min) SD COF Min COF Max COF Wear volume (mm³) SD wear (mm³) CP-Ti Grade 4 0,23 0,01 0,22 0,28 0,012 0,0003 Ti-6Al-4V ELI 0,23 0,01 0,20 0,27 0,017 0,0011 TNTZ-O (AR) 0,32 0,07 0,25 0,68 0,098 0,0164 Table 4: Friction and wear characteristics of the tested materials which suggests that under low load and temperature conditions a transition from adhesive wear to a combination of abrasion and oxidative wear occurs [8]. Additionally, the presence of artificial saliva provides lubricating effects that further reduce the primarily oxidative and abrasive wear mechanisms affecting Ti-6Al-4V ELI and CP-Ti [9, 10]. Although EDS results for the counter-bodies of these alloys are not presented, preliminary observations suggest reduced material transfer to the steel balls due to the smaller contact area and wear scar. A homogeneously distributed titanium signal is detected in both CP-Ti Grade 4 and Ti-6Al-4V ELI counter-bodies, with a reduced presence of aluminum and vanadium specifically identified on the Ti-6Al-4V ELI counter-body. In contrast, the TNTZ-O counter-body exhibits distinct areas of material detachment, characterized by pronounced saturated regions. 4 Conclusion and outlook TNTZ-O exhibits the highest wear volume and significant fluctuations in the COF among the tested materials. This increased wear and friction instability can be attributed to its β-phase body-centered cubic microstructure, which exhibits lower hardness and a reduced Young’s modulus compared to CP-Ti Grade 4 and Ti-6Al-4V ELI. The inherent properties of the β-phase microstructure contribute to greater elastic deformation and adhesive wear, which lead to higher wear volumes and inconsistent friction behavior. In contrast, CP-Ti Grade 4 and Ti-6Al-4V ELI, with their hexagonal close-packed α and α+β microstructures respectively, demonstrate more stable frictional behavior and lower wear volumes. The increased stiffness and hardness of these materials contribute to their superior performance in terms of wear resistance and friction stability. To enhance the wear behavior of TNTZ-O, future studies will focus on optimizing its microstructure through advanced thermo-mechanical processing. Specifically, promoting the transformation of the β-phase to the αor ω-phases could potentially improve its tribological performance by increasing hardness and reducing elastic deformation, thus mitigating wear and friction instability. Further research is required to gain a comprehensive understanding of the wear mechanisms of TNTZ-O. In order to provide a more detailed analysis of the wear processes and material interactions, it would be beneficial to complement the existing detailed SEM and EDS analyses of the wear tracks with a cross-sectional examination. Additionally, utilizing more application-specific Science and Research 17 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 Figure 5: Exemplary wear track of TNTZ-O (A), CP-Ti Grade 4 (B) and Ti-6Al-4V ELI (C) tested in reciprocating ball-on-disc tribometer for 42 min in artificial saliva against hardened 100Cr6. Images are captured in SE contrast at 1000: 1 magnification. Figure 6: Exemplary SEM images with subsequent EDS analysis of a wear track on a TNTZ-O disc A, and the corresponding 100Cr6 ball (B) tested in reciprocating ball-on-disc tribometer for 42 min in artificial saliva. Images are captured at 130: 1 magnification. [5] Zhang J-L, Tasan CC, Lai MJ, et al. Partial recrystallization of gum metal to achieve enhanced strength and ductility. Acta Mater 2017; 135: 400 - 410. [6] American Society for Testing and Materials (ASTM). Specification for Unalloyed Titanium, for Surgical Implant Applications (ASTM F-67-13; UNS R50250, UNS R50400, UNS R50550, UNS R50700). [7] American Society for Testing and Materials (ASTM). Specification for Wrought Titanium-6Aluminum-4Vanadium ELI (Extra Low Interstitial) Alloy for Surgical Implant Applications (ASTM. F-136; UNS R56401). [8] Mao YS, Wang L, Chen KM, et al. Tribo-layer and its role in dry sliding wear of Ti-6Al-4V alloy. Wear 2013; 297: 1032-1039. [9] Lee Y-S, Niinomi M, Nakai M, et al. Differences in Wear Behaviors at Sliding Contacts for β-Type and (α + β)- Type Titanium Alloys in Ringer’s Solution and Air. Mater Trans 2015; 56: 317-326. [10] Cvijović-Alagić I, Cvijović Z, Bajat J, et al. Electrochemical behaviour of Ti-6Al-4V alloy with different microstructures in a simulated bio-environment. Mater Corros 2016; 67: 1075 -1087. [11] Lee Y-S, Niinomi M, Nakai M, et al. Predominant factor determining wear properties of β-type and (α + β)-type titanium alloys in metal-to-metal contact for biomedical applications. J Mech Behav Biomed Mater 2015; 41: 208 -220. Science and Research 18 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0008 counter body materials in testing could offer a more accurate evaluation of TNTZ-O’s wear resistance in practical dental applications. Despite its promising mechanical properties for dental implants, TNTZ-O’s current tribological performance indicates challenges that must be addressed. Enhancing its wear behavior is crucial to ensure long-term durability and reliability in dental applications, which highlights the need for continued development and optimization. References [1] Abd-Elaziem W, Darwish MA, Hamada A, et al. Titanium-Based alloys and composites for orthopedic implants Applications: A comprehensive review. Mater Des 2024; 112850. [2] Breme J, Eisenbarth E, Biehl V. Titanium and its Alloys for Medical Applications. In: Titanium and Titanium Alloys. John Wiley & Sons, Ltd, pp. 423-451. [3] Willis J, Li S, Crean SJ, et al. Is titanium alloy Ti-6Al-4 V cytotoxic to gingival fibroblasts - A systematic review. Clin Exp Dent Res 2021; 7: 1037-1044. [4] Gordin DM, Ion R, Vasilescu C, et al. Potentiality of the ‘Gum Metal’ titanium-based alloy for biomedical applications. Mater Sci Eng C Mater Biol Appl 2014; 44: 362- 370. Science and Research 19 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 BOOK RECOMMENDATION expert verlag - Ein Unternehmen der Narr Francke Attempto Verlag GmbH + Co. KG Dischingerweg 5 \ 72070 Tübingen \ Germany \ Tel. +49 (0)7071 97 97 0 \ info@narr.de \ www.narr.de In addition to the indisputably necessary electrification of the transport sector, which is currently being ramped up, internal combustion engines will still be urgently needed in the future. Otherwise, the demand for mobility in the on-road, off-road and non-road sectors cannot be met. There is no doubt that these internal combustion engines will have to be improved regarding efficiency plus lower emissions and nowadays more and more important upgraded for zero and low carbon fuels. Even though Spark Ignition (SI) engines have been around for more than a century, there is still a lot of room for improvement, particularly in terms of power density, ignition, combustion control, and preventing uncontrolled combustion. To offer all interested developers an inspiring exchange platform for the latest developments, IAV established two exciting conferences more than two decades ago, which are now held under the heading “Two Conferences - One Goal”. This volume brings together the contributions to this conference. Content Ignition and inflammation of conventional and alternative fuels such as hydrogen, ammonia, methanol etc. - Combustion processes for alternative fuels - Prevention of irregular combustion phenomena when using conventional and alternative fuels - Methods for measurement and analysis of irregular combustion phenomena - Modern virtual development methods - Control, regulation and latest function algorithms Marc Sens (ED.) 6th International Conference on Ignition Systems for SI Engines 7th International Conference on Knocking in SI Engines 1st edition 2024, 386 p. €[D] 189,00 ISBN 978-3-381-12991-1 eISBN 978-3-381-12992-8 the base pitch p et based on ISO/ TR 14179-2 [4]. To reduce the load-dependent gear power loss P LGP under dry lubrication, it is essential to reduce the coefficient of friction µ. Solid lubricants are used to achieve this. These are typically transition metal dichalcogenides (TMD), such as molybdenum disulfide (MoS 2 ) and tungsten disulfide (WS 2 ), as reviewed in [5]. The solid lubrication mechanism involves the formation of a lubricating layer when a tribological stress is applied. This lubricating tribofilm then transfers and redistributes itself across the contact surfaces by means of shearing the solid lubricant, with the objective of spreading it uniformly. The transfer mechanism of solid lubricant coatings has been investigated on gears, revealing that the wear resistance of coated gears is influenced by both the coating material and the specific gear being coated [6]. Studies on the application of WC/ C coatings on the gear wheel results in improved tribological properties, with the uncoated pinion undergoing a polishing process due to interaction with the coating. The transfer Science and Research 20 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 1 Introduction Lubricants in tribological contacts of machine elements, such as gears, are intended to reduce friction and wear and to remove frictional heat generated in the contact. Oil and grease lubrication cannot be used in applications with extreme ambient conditions such as in a vacuum, investigated in [1], with low weight or extreme thermal requirements, reviewed in [2], and in hygienic applications such as systems in the food or pharmaceutical industries, investigated in [3]. Dry lubrication results in a limited heat transfer by convection. Therefore, a low load-dependent gear power loss P LGP is essential to stay below critical temperature regimes for durable and safe operation. Equation (1) shows the load-dependent gear power loss P LGP , (1) with the coefficient of friction µ, the line load f N and the sliding velocity v g across the area of tooth contact, and = 1 ⋅ ( ̅ , ) ⋅ ( ̅ , ) ⋅ ( ̅ , ) ̅ ̅ Dry Lubrication of Spur Gears Coated with CrAlN+Mo: W: S Felix Farrenkopf, Thomas Lohner, Karsten Stahl, Kirsten Bobzin, Christian Kalscheuer, Max Philip Möbius, Marta Miranda Marti* submitted: 1.08.2024 Accepted: 25.04.2025 Presented at GfT-Conference 2024 Cylindrical gears are coated with CrAlN+Mo: W: S using High Power Pulsed Magnetron Sputtering and direct current Magnetron Sputtering. Their power loss, bulk temperature, and wear behavior are analyzed on the FZG gear efficiency test rig under dry lubrication. Coated and uncoated gears are examined using laser microscopy, energy-dispersive spectroscopy, and Raman spectroscopy. For operating points with low gear power loss, a tribofilm containing MoS 2 as well as Mo and W oxides is detected on the pinion and wheel tooth flank, resulting in a stable tribosystem for several thousands load cycles. Cylindrical gear pairs with CrAlN+Mo: W: S coated wheels exhibit reduced power loss and lower bulk temperatures compared to uncoated gears. Keywords PVD, CrAlN+Mo: W: S, Solid Lubricants, Gears, Efficiency, Friction, Wear Abstract * Felix Farrenkopf, M.Sc. 1 (corresponding author) Orcid-ID: https: / / orcid.org/ 0000-0002-8675-3040 Dr.-Ing. Thomas Lohner 1 Orcid-ID: https: / / orcid.org/ 0000-0002-6067-9399 Prof. Dr.-Ing. Karsten Stahl 1 Orcid-ID: https: / / orcid.org/ 0000-0001-7177-5207 Prof. Dr.-Ing. Kirsten Bobzin 2 Orcid-ID: https: / / orcid.org/ 0000-0003-3797-8347 Dr.-Ing. Christian Kalscheuer 2 Orcid-ID: https: / / orcid.org/ 0000-0002-1606-5090 Max Philip Möbius, M.Sc. 2 Orcid-ID: https: / / orcid.org/ 0000-0002-2982-8720 Marta Miranda Marti, M.Sc.(corresponding author) Orcid-ID: https: / / orcid.org/ 0000-0002-2048-7042 1 Gear Research Center (FZG), Department of Mechanical Engineering, School of Engineering and Design, Technical University of Munich, Boltzmannstraße 15, 85748 Garching near Munich, Germany 2 Surface Engineering Institute (IOT), RWTH Aachen University, Kackertstraße 15, 52072 Aachen, Germany mechanisms of solid lubricant coatings such as W-S-C [7], h-BN [8], Ti-WS 2 [9], Mo 2 N/ MoS 2 / Ag [10] have been extensively investigated in various studies but primarily on flat samples. A promising approach for low friction and high wear resistance is by applying a CrAlN coating enriched with solid lubricants components such as Mo, W and S. Previous studies explored the use of graded CrAlN+Mo: W: S deposited by means of physical vapor deposition (PVD), demonstrating a significant reduction in friction compared to uncoated substrates [11]. These studies also investigated the tribofilm and transfer mechanism on flat samples [12]. However, these coatings have not been analyzed in machine elements such as gears tested under stationary operating conditions, and the transfer mechanisms of these triboactive coatings remain not fully understood. Therefore, the subject of the present study is the investigation of the triboactive CrAlN+Mo: W: S coating in the gear contact and the analysis of the tribofilm and transfer film formation mechanism. 2 Methods and Materials This section describes the FZG efficiency test rig, the CrAlN+Mo: W: S coating and application process used, the test procedure for the coated gears in the FZG efficiency test rig, and the gear analysis procedure. 2.1 Test Rig and Gears The considered FZG efficiency test rig is a modified FZG back-to-back test rig with a center distance of a = 91.5 mm according to DIN ISO 14635-1 [13]. It is based on the principle of power circulation in which the load torque is applied by the load clutch in-between the test and transmission gearbox. The test rig is driven by an electric engine that provides the loss torque of the power circle. The load torque is measured in the power circle, the power loss is measured between the power circle and the electric engine. Further information can be found in [14,15]. The gear contact in the test gearbox is dry-lubricated and bearings of the type 6406 are used, which are both sided sealed. In the transmission gearbox, bearings of the type NU406 are used and injection lubrication with 2 l/ min of an ISO VG 100 mineral oil at ϑ oil = 60 °C is applied. The test gears used in the test and transmission gearbox are of FZG type C mod (cf. Figure 1) made of case-hardened 16MnCr5 steel with ground and superfinished gear flanks. A tip relief C a of 35 µm is applied. According to [16] and using the loaded tooth contact analysis (LTCA) software RIKOR [17], a local geometric tooth power loss factor H VL of 0.166 is derived for T 1 = 94.1 Nm. The bulk temperature ϑ M is measured by a Pt100 temperature sensor in the middle of one tooth. The main gear geometry data is given in Table 1. Science and Research 21 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 (a) (b) Figure 1: FZG test gear of type C mod (a) and transverse section of the gear mesh (b) Parameter Value Center distance, a in mm 91.5 Common tooth width, b in mm Normal module, m n in mm 4.5 Pressure angle, α n in ° 20 Number of teeth, z 1 | z 2 2 1 b | 16 | 24 14 | 14 Tip relief, C a in μm 35 Table 1: Main geometry data of the FZG test gear of type C mod Pinion (1) Wheel (2) Pinion (1) Wheel (2) d 1 d 2 1.2 0.9 0.6 0.3 0A B C D E 10 8 6 4 2 0 p H in GPa R red in mm p H R red 4.5 3 1.5 0A B C D E v in m/ s v Σ |v g | (a) (b) To achieve a homogeneous coating of the gears, a substrate holder with a spit geometry is employed for a twofold rotation of the gear. Prior to coating, a chemical cleaning process is conducted. 2.3 Test Procedure The operating point considered is selected from the method FVA 345 [21]. A moderate load of T 1 = 94.1 Nm and medium circumferential speed of v t,C = 5 m/ s at ambient temperature and atmosphere are applied. Figure 2 shows the Hertzian pressure p H , the radius of curvature R red , the sliding velocity v g and the sum velocity v ∑ along the path of contact for the FZG test gear of type C mod at T 1 = 94.1 Nm and v t,C = 5 m/ s. Due to the test gears’ tip relief C a of 35 µm (cf. Table 1) designed for T 1 = 302 Nm, there is no tooth contact close to A and E for the considered operating condition. While there is pure rolling |v g | = 0 m/ s at the pitch point C, the sliding portion increases with increasing distance to the pitch point C. Science and Research 22 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 The superfinished gear flanks have a roughness of Ra = 0.07…0.12 µm. Only the wheels are coated, resulting in a roughness of Ra = 0.06…0.11 µm. The surface roughness was measured according to DIN EN ISO 4287 [18] and DIN EN ISO 4288 [19] with a limiting wavelength of λ c = 0.8 mm on a length of L t = 4.8 mm in tooth profile direction. Due to the manufacturing and coating process (cf. section 2.2), the coated wheels have an average lead crowning of up to C β = 2.4 µm. 2.2 Coating The CrAlN+Mo: W: S coating is deposited using an industrial CC800/ 9 coating unit, CemeCon AG, Würselen, Germany. This coating unit is equipped with four direct current Magnetron Sputtering (dcMS) cathodes and two High-Power Pulsed Magnetron Sputtering (HPPMS) cathodes. The cathodes operate simultaneously to achieve a hybrid dcMS/ HPPMS process, resulting in a coating thickness of d = 3 µm. Table 2 displays the deposition process parameters. Further coating parameters can be found in [20]. Process phase Process parameters Unit Value Interlayer and Pressure, p mPa 490 - 560 ramps Bias voltage, U B V -150 to (-75) Argon flow, j Ar sccm 200 Nitrogen flow, j N sccm pressure controlled Pulse frequency f Hz 500 Pulse duration, t on μs 40 Speed of table rotation, v rot rpm 3 Toplayer Pressure, p mPa 590 Bias voltage, U B V -75 Argon flow, j Ar sccm pressure controlled Nitrogen flow, j N sccm 100 - 140 Table 2: Process parameters of coating deposition Figure 2: Hertzian pressure p H and radius of curvature R red (a), and sliding velocity v g and sum velocity v ∑ (b) along the line of action of the FZG test gear of type C mod (T 1 = 94.1 Nm, v t,C = 5 m/ s) For the coated gear tests, only the wheel is coated while the pinion is uncoated. For reference tests, an uncoated pinion and wheel are used. Upon reaching the maximum annealing temperature of ϑ M = 200 °C, as defined by DIN EN ISO 683-3 [22], the tests are aborted. A calibrated transmission gearbox is used for all tests. For calibration, bearings of type NU406 and uncoated gears are mounted in the test gearbox, to ensure the same set-up as in the transmission gearbox. Calibration tests are carried out under no-load and load conditions with both gearboxes are injection-lubricated. Since the test and transmission gearbox are equipped with FZG test gears of type C mod , featuring a symmetrical gear mesh behavior, the measured power loss can be approximately halved to derive the no-load and load-dependent power loss of the transmission gearbox, P L0,cal and P LP,cal . With the calibrated transmission gearbox, the bearings of type 6406 and test gears are mounted in the test gearbox. Before each test under dry-lubrication, the circumferential speed is held for t = 15 min under no-load to determine the no-load power loss of the test gearbox P L0,dry : (2) In the subsequent test under load, the load-dependent power loss P LP,dry is determined by subtracting the power loss of the transmission gearbox (P L0,cal + P LP,cal ) and the no-load power loss of the test gearbox P L0,dry . This results in the load-dependent loss of the test gearbox P LP,dry : (3) The SKF method in [23] method is used to determine the load-dependent bearing power loss P LBP of the test gearbox, which can also be subtracted to finally derive the load-dependent gear power loss P LGP,dry of the test gearbox: (4) The load-dependent gear power loss P LGP acc. to equation (1) can be rewritten by the assumption of a mean , = − , ! , = − , − ( , ! + , ! ) , = , − " coefficient of gear friction µ mz acc. to ISO/ TR 14179-2 [4]: (5) Subsequently, the mean coefficient of friction can be determined from the load-dependent gear power loss P LGP,dry using P in and H VL . 2.4 Analysis of Tribofilm and Transfer film To analyze the tooth flank of the considered test gears, a single tooth is separated due to geometric limitations of the characterization techniques. This is achieved by cutting the gears between two teeth. This is done fluid-free with an angle grinder to prevent removing any tribofilm or transfer film. Compressed air cooled the gear, and temperature is monitored during cutting. Subsequently, a surface analysis after mechanical testing of the coated gear flanks and uncoated pinion flank is conducted by means of confocal laser scanning microscopy (CLSM), VK-X210, Keyence, NeuIsenburg, Germany. For the characterization of the chemical composition of tribofilm on wheel and pinion flanks after the efficiency test, an energy-dispersive X-ray spectroscopy (EDX) detector integrated into a Phenom XL G2 desktop scanning electron microscope (SEM) by ThermoFischer Scientific, Dreieich, Germany, is used. The employed accelerating voltage is E V = 15 kV with a working distance of W D = 8 mm, and the mapping resolution range is set to 480 x 480 pixel with a pixel dwell time range of t = 10 ms/ pixel. A systematic chemical analysis of tribofilm on the coated gear and uncoated pinion is conducted by means of Raman spectroscopy employing a Renishaw InVia Reflex, Renishaw GmbH, Pliezhausen, Germany. The parameters for analysis are shown in Table 3. The laser is calibrated before the measurement by using a silicon reference sample. Representative spectra are chosen for evaluation and fitted by OriginPro Software using a Lorentz fitting filter as recommended in [24]. = #$ ⋅ 1 ⋅ ( ̅ , ) ⋅ ( ̅ , ) ̅ ̅ %&&&&&&&&&&&'&&&&&&&&&&&* -. ⋅/ 02 Science and Research 23 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 Parameter Value Laser wavelength λ in nm 532 Laser power P in mW 2.6 Accumulations n in - 5 Grating in 1/ mm 1,800 Exposure time t in s 15 Objective lens 5x Table 3: Raman acquisition parameters ϑ M = 200 °C already at N 1 = 5,273 and N 1 = 6,295. One CrAlN+Mo: W: S/ uncoated test reaches the abort criteria of ϑ M = 200 °C at N 1 = 19,212, while another CrAlN+Mo: W: S/ uncoated test is stopped at N 1 = 5,284 to analyze the gear flanks at the corresponding state. A further repetition test shows a similar trend and is stopped at N 1 = 41,680 due to no significant changes in the operating behavior. The load-dependent gear power loss P LGP , dry in Figure 3a and mean coefficient of gear friction µ mz in Figure 3b shows a steep increase from the very beginning for the tests uncoated/ uncoated. A plateau is finally reached at a high level of power loss and mean coefficient of gear friction. Consequently, this results in a steep and Science and Research 24 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 3 Results and Discussion This section presents the results of the tests conducted on the FZG efficiency test rig and an analysis of the gears. 3.1 Experimental Results Figure 3 shows the measured load-dependent gear power loss P LGP,dry , mean coefficient of gear friction µ mz and gear bulk temperature ϑ M of the pinion and wheel over the load cycles at the pinion N 1 . Test results are shown for uncoated/ uncoated and CrAlN+Mo: W: S/ uncoated configurations. The two uncoated/ uncoated tests show similar behavior and reach the abort criteria of Figure 3: Load-dependent gear power loss P LGP (a), mean coefficient of gear friction µ mz (b) and gear bulk temperature ϑ M (c) over the load cycles at the pinion for tests with uncoated/ uncoated and coated/ uncoated gears of type C mod (T 1 = 94.1 Nm, v t,C = 5 m/ s) 1.5 1 0.5 0 0 2 4 6 8 10 12 14 16 18 20 Load cycles at pinion N 1 in 1,000 test uncoated/ uncoated 1 test uncoated/ uncoated 2 test CrAlN+Mo: W: S/ uncoated (N 1 =5,284) test CrAlN+Mo: W: S/ uncoated (N 1 =19,212) test CrAlN+Mo: W: S/ uncoated (N 1 =41,680)* 0.5 0.3 0.2 0 0 2 4 6 8 10 12 14 16 18 20 0.4 0.1 Load cycles at pinion N 1 in 1,000 200 150 100 20 0 2 4 6 8 10 12 14 16 18 20 50 Pinion Wheel Load cycles at pinion N 1 in 1,000 *stopped intentionally (a) (b) (c) continuous increase in the bulk temperatures of the pinion and wheel (cf. Figure 3c). Both bulk temperatures are very close to each other. Once the abort criteria of ϑ M = 200 °C is reached at N 1 = 6,295 for the test uncoated/ uncoated 1 and at N 1 = 5,273 for the test uncoated/ uncoated 2, the tests are aborted. In contrast, the tests CrAlN+Mo: W: S/ uncoated in Figure 3 show a decrease of the load-dependent gear power loss P LGP,dry and the mean coefficient of gear friction µ mz immediately after the beginning. A minimum mean coefficient of gear friction µ mz of 0.048 and 0.064 and minimum load-dependent gear power loss P LGP,dry of 0.105 and 0.144 kW is reached. This means a reduction of 84 and 88 % compared to the test uncoated/ uncoated 1, which is achieved by the CrAlN+Mo: W: S coating applied. The low mean coefficients of gear friction indicate a functioning friction-reducing tribofilm and thus the formation of the solid lubricant. Observing the CrAlN+Mo: W: S/ uncoated (N 1 = 19,212) test, there is a steep increase in the mean coefficient of gear friction μ mz and load-dependent gear power loss P LGP,dry at around 10,000 load cycles (cf. Figure 3a and b). From 12,000 load cycles on, this increase becomes more moderate but continues to increase until the abort criterion is reached. This indicates that the effectiveness of the tribofilm is no longer sufficient at the point of steep increase. The second test CrAlN+Mo: W: S/ uncoated (N 1 = 41,680) shows a similar trend from beginning until it reaches a second plateau from 14,000 load cycles. To analyze the coating and transfer film in the minimum plateau, the test CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) is stopped intentionally. As illustrated in Figure 3c, the bulk temperatures of the tests CrAlN+Mo: W: S/ uncoated exhibits a characteristic pattern. A slight increase in bulk temperature ϑ M is observed in the low friction plateau region. As friction and power loss increase, the bulk temperature ϑ M continues to rise as well. The test CrAlN+Mo: W: S/ uncoated (N 1 = 19,212) reaches the abort criterion while the test CrAlN+Mo: W: S/ uncoated (N 1 = 41,680) stabilizes due to the second friction plateau after 14,000 load cycles. It is notable that there is a considerable difference in the bulk temperatures ϑ M of the uncoated pinion and the coated wheel. The findings of [25] indicate that a coating can function as a thermally insulating layer, which results in diminished heat dissipation into the substrate. 3.2 Analysis of Tribofilm and Transfer Film Formation To gain an overview of the gear tooth flanks and identify the critical wear areas, CLSM overview images are obtained, as illustrated in Figure 4. To accurately determine the presence of the coating, an EDX mapping is conducted along the coated wheel tooth flank. In this mapping, the elements Mo, W and S are represented in yellow, with no difference among them due to the overlapping of the characteristic Mo and S EDX-peaks [26], whereas interlayer elements such as Al and Cr are depicted in blue. The substrate material Fe is shown in green, to provide a clear contrast between the coating and the substrate. As illustrated in Figure 4a, the uncoated wheel from the test uncoated/ uncoated 1 exhibits significant wear along its tooth flank in the areas of high sliding, with a maximum wear depth of z = 20 µm. Negligible wear is observed in the vicinity of the pitch point C, where rolling friction is predominated. The overall high wear value on the uncoated gear tooth flanks correlates with the rapid increase in power loss and bulk temperature observed from the beginning of the test. The tooth flank root and tip area does not contact the pinion surface (cf. Figure 2), thus the coating remains unaffected. For the coated wheel from the test CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) in Figure 4b), less substrate is exposed when compared to the wheel after completion, and the elements of the interlayer are minimally present. Wear along the tooth flank is barely noticeable with a maximum wear depth of z = 3 µm. However, the CrAlN+Mo: W: S coating with a coating thickness of d = 3 µm is already partly removed from the surface. Due to the lead crowning C β , a wear concentration in the midsection can be observed. For the coated wheel from the test CrAlN+Mo: W: S/ uncoated (N 1 = 19,212), the delamination of the coating’s solid lubrication elements on the upper part of the tooth flank becomes more evident in Figure 4c, indicated by increased measurements of Al and Cr content. Moderate wear can be observed in the upper part of the tooth flank, while wear in the lower part of the tooth flank is much more pronounced. This result suggests that the toplayer and the graded function layer containing the solid lubricant elements is consumed to achieve lubrication along the tooth flank, creating MoS 2 and reducing the wear. As a result, the power loss is reduced compared to uncoated gears, and a reduction in wear is achieved, as determined by the profile measurement. In order to determine the triboactive transformation of the coating on the wheel of the test CrAlN+Mo: W: S/ uncoated (N 1 = 5,284), Raman measurements on the remaining coating are conducted. The results of an exemplary spectrum found on the analyzed wheel tooth flank are presented in Figure 5. The characteristic double peak of MoS 2 is clearly observed at ω E2g = 385 cm -1 and ω A1g = 405 cm -1 corresponding to the E 2g and the A 1g band respectively, as measured in [27]. Tillmann et al. report similar results Science and Research 25 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 and W oxides are also found on the tooth flank surface, which might contribute to the lubricating effect. To analyze the formation of a transfer film on the pinion, the same characterization techniques are employed. The formation of a transfer film on the pinion tooth flank surface is analyzed using EDX. However, due to the scattering of triboactive elements along the pinion flank, no Raman measurement of the triboactive elements is pos- Science and Research 26 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 where formation of a lubricating tribofilm contributes to low friction for non-synchronized, dry running twinscrew machines [28]. Due to this mechanical activation of the solid-lubricant elements Mo, W and S, a MoS 2 tribofilm is obtained. The characteristic WS 2 peak is observable at ω E2g = 350 cm -1 , while the second characteristic WS 2 peak is reported in [29] to overlap with the MoS 2 peak at around ω A1g = 405 cm -1 . This suggests the presence of WS 2 in the resulting tribofilm. Various Mo Figure 4: CLSM images, EDX mapping and profile measurements of wheel tooth flanks of the tests (a) uncoated/ uncoated 1, (b) CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) and (c) CrAlN+Mo: W: S/ uncoated (N 1 = 19,212) (a) (b) (c) Science and Research 27 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 Figure 5: Raman spectra on wheel of test CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) Figure 6: CLSM images, EDX mapping and flank profile from the tests (a) uncoated/ uncoated 1, (b) CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) and (c) CrAlN+Mo: W: S/ uncoated (N 1 = 19,212) (a) (b) (c) 3. On the coated wheel, from a test stopped after N 1 = 5,284 pinion load cycles, EDX mapping measured remaining triboactive elements. Raman spectroscopy identified them as solid lubricants MoS 2 and WS 2 . Furthermore, various Moand W-oxides were found on the tooth flanks. This demonstrates that solid lubricants are formed in situ under the tribological conditions found in gear contacts. 4. The presence of a transfer film on the pinion tooth flanks, as observed by EDX analysis, confirmed the formation of a transfer film, which may have contributed to the measured reduction in friction. Considering these findings, the next step should be a comprehensive parameter study to evaluate the solid lubricant and its lifetime under different operating conditions. This further study should identify limits and potential areas of application. Acknowledgement The authors acknowledge the financial support of the German Research Foundation, Deutsche Forschungsgemeinschaft (DFG), within the priority program SPP 2074 and the research project “Fluid-less lubrication systems with high mechanical load” (No. BO 1979/ 66-2 / STA 1198/ 17-2). References [1] C. Donnet, Advanced solid lubricant coatings for high vacuum environments, Surface and Coatings Technology 80 (1996) 151-156. https: / / doi.org/ 10.1016/ 0257-8972(95)02702-5. [2] I.M. Allam, Solid lubricants for applications at elevated temperatures, J Mater Sci 26 (1991) 3977-3984. https: / / doi.org/ 10.1007/ BF00553478. [3] C. Donnet, A. Erdemir, Solid Lubricant Coatings: Recent Developments and Future Trends, Tribology Letters 17 (2004) 389-397. https: / / doi.org/ 10.1023/ B: TRIL.0000044487.32514.1d. [4] ISO-International Organization for Standardization, Gears - Thermal capacity -Part 2: Thermal load-carrying capacity: ISO/ TR 14179-2: 2001, ISO-International Organization for Standardization, 2001. [5] T.W. Scharf, Transition Metal Dichalcogenide-Based (MoS 2 , WS 2 ) Coatings, in: G.E. Totten (Ed.), Friction, Lubrication, and Wear Technology, ASM International, 2017, pp. 583-596. [6] R. Michalczewski, M. Kalbarczyk, W. Piekoszewski, M. Szczerek, W. Tuszyński, The rolling contact fatigue of WC/ C-coated spur gears, Proceedings of the Institution of Mechanical Engineers, Part J: Journal of Engineering Tribology 227 (2013) 850-860. https: / / doi.org/ 10.1177/ 1350650113478179. Science and Research 28 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 sible. Figure 6 presents a summary of the results, comparing the uncoated pinion flanks from the tests uncoated/ uncoated 1, CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) and CrAlN+Mo: W: S/ uncoated (N 1 = 19,212). Due to the presence of Cr in both the coating interlayer and the 16MnCr5 substrate, a mapping including this element cannot differentiate whether the interlayer has been transferred from the coated wheel or originates from the pinion substrate. Consequently, the Cr content is excluded from this analysis. At the uncoated pinion from the test uncoated/ uncoated 1 in Figure 6a, a high wear of up to z = 21 µm is noticeable. On the tooth flank surface of the pinion from the test CrAlN+Mo: W: S/ uncoated (N 1 = 5,284) in Figure 6b, adhesions containing solid lubricant elements can be observed. This confirms the transfer of the tribofilm from wheel to pinion. After the test CrAlN+Mo: W: S/ uncoated (N 1 = 19,212) (cf. Figure 6c), no transfer film can be measured, only elements prevenient of the interface like Al. Higher wear of up to z = 11 µm can be measured, especially in the region of the tooth flank root area. This reinforces the theory that the tribofilm on the wheel is consumed with increasing load cycles to lubricate the dry contact. This process reduces power loss and controls the bulk gear temperature. Due to the consumption of the triboelements after higher pinion load cycles, the coating failure increases the friction, power loss and bulk temperature. 4 Conclusion This study investigated the tribological behavior of CrAlN+Mo: W: S coated wheels paired with uncoated pinions under dry lubrication. To investigate the relationship between tribofilm formation and pinion load cycles, gear tests were carried out at a pinion torque T 1 = 94.1 Nm and circumferential speed v t,C = 5 m/ s. The following conclusions can be drawn: 1. The deposition of a CrAlN+Mo: W: S coating on the wheel has been shown to result in a significant reduction in the mean coefficient of gear friction µ mz when compared to uncoated tests. The reduction was approximately 84 % to 88 %, resulting in values ranging from µ mz = 0.048 to µ mz = 0.064. This result validates the friction reduction with CrAlN+Mo: W: S coatings in gear contacts. 2. The detection of interlayer elements such as Al and Cr on the wheel tooth flanks after reaching the abortion criteria ϑ M = 200 °C confirms the partial consumption of the CrAlN+Mo: W: S-toplayer and the subsequent wear of the functional coating. This observation highlights the coating’s active contribution in the tribological contact and demonstrates a correlation between coating wear and operational lifetime. [7] J.V. Pimentel, T. Polcar, M. Evaristo, A. Cavaleiro, Examination of the tribolayer formation of a self-lubricant W-S-C sputtered coating, Tribology International 47 (2012) 188-193. https: / / doi.org/ 10.1016/ j.triboint.2011.10.021. [8] C. Zishan, L. Hejun, F. Qiangang, Q. Xinfa, Tribological behaviors of SiC/ h-BN composite coating at elevated temperatures, Tribology International 56 (2012) 58-65. https: / / doi.org/ 10.1016/ j.triboint.2012.06.026. [9] T.W. Scharf, A. Rajendran, R. Banerjee, F. Sequeda, Growth, structure and friction behavior of titanium doped tungsten disulphide (Ti-WS2) nanocomposite thin films, Thin Solid Films 517 (2009) 5666-5675. https: / / doi.org/ 10.1016/ j.tsf.2009.02.103. [10] S.M. Aouadi, Y. Paudel, W.J. Simonson, Q. Ge, P. Kohli, C. Muratore, A.A. Voevodin, Tribological investigation of adaptive Mo 2 N/ MoS 2 / Ag coatings with high sulfur content, Surface and Coatings Technology 203 (2009) 1304-1309. https: / / doi.org/ 10.1016/ j.surfcoat.2008.10.040. [11] K. Bobzin, T. Brögelmann, C. Kalscheuer, M. Thiex, Selflubricating triboactive (Cr,Al)N+Mo: S coatings for fluidfree applications, J Mater Sci 56 (2021) 15040-15060. https: / / doi.org/ 10.1007/ s10853-021-06255-9. [12] K. Bobzin, C. Kalscheuer, M. Carlet, C. Schulze, 2023. Influence of the etching process on the coating performance in dry tribological contacts. Journal of Vacuum Science & Technology A 41, 033104. https: / / doi.org/ 10.1116/ 6.0002363. [13] DIN ISO 14635-1, Zahnräder - FZG-Prüfverfahren - Teil 1: FZG-Prüfverfahren A/ 8,3/ 90 zur Bestimmung der relativen Fresstragfähigkeit von Schmierölen, 2023. [14] A. Doleschel, Wirkungsgradberechnung von Zahnradgetrieben in Abhängigkeit vom Schmierstoff: Efficiency calculation of gear drives depending on the lubricant. Dissertation, 2003. [15] T. Lohner, Berechnung von TEHD Kontakten und Einlaufverhalten von Verzahnungen: Calculation of TEHL contacts and run-in behaviour of gears. Dissertation, 2016. [16] A. Wimmer, Lastverluste von Stirnradverzahnungen: Load dependent power loss of spur gears. Dissertation, 2006. [17] U. Weinberger, M.K. Otto, K. Stahl, 2020. Closed-Form Calculation of Lead Flank Modification Proposal for Spur and Helical Gear Stages. Journal of Mechanical Design 142, 031106. https: / / doi.org/ 10.1115/ 1.4045396. [18] DIN EN ISO 4287: 2010-07, Geometrische Produktspezifikation (GPS) - Oberflächenbeschaffenheit: Tastschnittverfahren - Benennungen, Definitionen und Kenngrößen der Oberflächenbeschaffenheit, 2010. [19] DIN EN ISO 4288: 1998-04, Geometrische Produktspezifikation (GPS) - Oberflächenbeschaffenheit: Tastschnittverfahren - Regeln und Verfahren für die Beurteilung der Oberflächenbeschaffenheit, 1997. [20] K. Bobzin, C. Kalscheuer, M.P. Möbius, C. Schulze, M. Miranda Marti, 2024. High power pulsed magnetron sputtering tailored low temperature CrAlN + Mo: W: S coatings for dry tribological contacts. Journal of Vacuum Science & Technology A 42, 013404. https: / / doi.org/ 10.1116/ 6.0003259. [21] A. Wimmer (Ed.), Vergleich der Wirkungsgradmessmethoden nach VW und FVA und die Übertragbarkeit der Ergebnisse auf Praxisgetriebe. Comparison of the Efficiency Measurement Methods according to VW and FVA and the Transferability of th eResults to Practical Gearboxes, Frankfurt am Main, Germany, 2004. [22] DIN EN ISO 683-3: 2022-06, Für eine Wärmebehandlung bestimmte Stähle, legierte Stähle und Automatenstähle - Teil 3: Einsatzstähle, 2022. [23] SKF Group, Roller bearings, 4000/ IV T, Schweinfurt, 1994. [24] Xueyin Yuan, Robert A. Mayanovic, An Empirical Study on Raman Peak Fitting and Its Application to Raman Quantitative Research, Appl. Spectrosc. 71 (2017) 2325- 2338. [25] M. Ebner, A. Ziegltrum, T. Lohner, K. Michaelis, K. Stahl, Measurement of EHL temperature by thin film sensors - Thermal insulation effects, Tribology International 149 (2020) 105515. https: / / doi.org/ 10.1016/ j.triboint.2018.12.015. [26] K.E. Ramohlola, E.I. Iwuoha, M.J. Hato, K.D. Modibane, Instrumental Techniques for Characterization of Molybdenum Disulphide Nanostructures, J. Anal. Methods Chem. 2020 (2020) 8896698. https: / / doi.org/ 10.1155/ 2020/ 8896698. [27] H. Li, Q. Zhang, C.C.R. Yap, B.K. Tay, T.H.T. Edwin, A. Olivier, D. Baillargeat, From Bulk to Monolayer MoS 2 Evolution of Raman Scattering, Adv Funct Materials 22 (2012) 1385-1390. https: / / doi.org/ 10.1002/ adfm.201102111. [28] W. Tillmann, A. Wittig, D. Stangier, H. Moldenhauer, C.- A. Thomann, J. Debus, D. Aurich, A. Bruemmer, Temperature-dependent tribological behavior of MoS x thin films synthesized by HiPIMS, Tribology International 153 (2021) 106655. https: / / doi.org/ 10.1016/ j.triboint.2020.106655. [29] M. Viršek, A. Jesih, I. Milošević, M. Damnjanović, M. Remškar, Raman scattering of the MoS 2 and WS 2 single nanotubes, Surface Science 601 (2007) 2868-2872. https: / / doi.org/ 10.1016/ j.susc.2006.12.050. Science and Research 29 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0009 nents of composites, visual and tactile perception, thermal conductivity and even electrical conductivity in the case of some new smart materials and sensors. However, the quantification and modelling of the contact and tribological properties of soft polymers remains a significant challenge, necessitating a trial-and-error approach and a substantial number of measurements for the optimisation of product properties. Since an analytical model that would take into account all the special properties of soft polymers would be so complicated as to prevent its application to practical problems, it is desirable to find an applicable approximate model that adequately addresses the main issues of contacts with soft polymers. As a first step towards such an approximate model for soft polymer contacts, it is logi- Science and Research 30 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 Introduction The real contact area between two materials is typically orders of magnitude smaller than the apparent contact area, due to the ubiquitous roughness of all surfaces. However, it is the real contact area that determines the tribological phenomena occurring at the interface [1]. The present study considers the phenomenon of large deformation contacts, which are of particular practical relevance in the context of soft polymers. Here, the term ‘soft polymers’ shall refer to polymeric materials that are used above their glass transition temperature. In this so-called rubbery state, the Young’s modulus can be orders of magnitude lower than in the hard, glassy state below the glass transition temperature. Due to the small forces required for deformation, typical application forces can easily lead to large deformations. The present study does not focus on soft polymers such as silicones or rubber, for which the deformation behaviour can be adequately described by elastic or hyperelastic models and the contact properties have been extensively examined by other researchers [2, 3, 4]. The study instead focuses on soft polymers with non-ideal elastic deformation behaviour, including polyethylene (PE), plasticised polyvinyl chloride (p-PVC) and polyurethanes (PUR). Applied in foils, paints, varnishes and coatings on a wide range of substrates, these materials determine the surface properties of a vast variety of everyday products, including contact properties in contacts to other surfaces. These contact properties, in turn, influence a number of crucial product characteristics, including blocking and sliding behaviour, grip, the generation of disturbing noises (such as squeaking and creaking), durability and susceptibility to wear, adhesion between different compo- Insight into large deformation contacts of soft polymers with molecular dynamics simulations Susanne Fritz* submitted: 20.09.2024 accepted: 9.05.2025 (peer review) Presented at GfT-Conference 2024 The size of the real contact area between mating materials represents a crucial quantity in the context of all tribological problems. Consequently, it is a topic that has been the subject of intensive investigation. Nevertheless, the subject of contact formation with soft polymers remains largely uninvestigated. Soft polymers are frequently employed as foils or coatings, thereby influencing the surface characteristics of a vast variety of everyday products. Employed at temperatures above their glass transition point, soft polymers with low values of the Young’s modulus, permit significant deformation during contact formation. To assess the impact of such large deformations on the size of the real contact area, atomistic molecular dynamics (MD) simulations of nanoindentation tests have been conducted. Depending on the material properties and loading conditions, the formation of the contact area was observed across the entire range of indentation depths. Keywords contact, contact mechanics, polymers, nanoindentation, MD simulations, Hertz-model Abstract * Dr. rer. nat. Susanne Fritz Orcid-ID: https: / / orcid.org/ 0000-0002-5006-8826 Department Surfaces, FILK Freiberg Institute gGmbH Meißner Ring 1-5, 09599 Freiberg, Germany cal to begin with the Hertz model, given that the majority of models for the real contact area in contact mechanics are based on the Hertz model. Heinrich Hertz [5] derived an analytical solution to the simple contact problem between a rigid sphere and an elastic half-space. A numerous number of other authors proceeded to expand his theory to include the effects of adhesion [6], friction [7, 8], viscoelasticity [9, 10] or roughness [11, 12], for example. However, since Hertz built his equations upon a certain set of fundamental assumptions, all the other theories are also only valid under these assumptions, which are not always met by contacts with soft polymers. These assumptions are in particular: • homogeneous, isotropic solids that can be considered as a continuum, • linear elastic behaviour, and • deformations that are so small, that the shape of the surface in the contact region can be reasonably well approximated by a second-degree polynomial. In contrast, the behaviour of soft polymers is typically characterised by a pronounced viscoelastic-viscoplastic response and can even exhibit viscous flow to some extent. At the length scale of small surface asperities (micro-roughness), a polymeric material made of large chain molecules and maybe containing crystallites cannot simply be assumed to behave as a homogenous continuum and usually deforms anisotropically. But the main difference surely are the large deformations, that can easily occur with soft polymers, especially at small asperities. The question of high deformation Hertz contacts has already been addressed by other authors [13, 14, 15, 16, 17, 18, 19], with results ranging from the good applicability of the Hertz model to the introduction of correction terms and the development of new empirical equations. Nevertheless, as the actual contact area is not readily accessible through direct experimentation, all the referenced papers have employed analytical or finite element analysis (FEA). This involves defining a mathematical equation to describe the deformation behaviour of the material, along with the appropriate material properties, as the input for the model. This approach is effective when the deformation model accurately represents the material’s behaviour, but it can also introduce unpredictable errors when the deformation behaviour is not straightforward to describe mathematically, as is the case with soft polymers. For this reason, molecular dynamics (MD) simulations [20, 21] were used in the present study instead of FEA. In MD simulations, the material is described at the atomic level, and the behaviour of the material is calculated based on the known interactions between the atoms. As a consequence, only small time and length scales are accessible, but no models or material properties are needed to describe, for example, the deformation behaviour, and nanoscopic properties that are difficult to measure can be accessed. The use of MD simulations for the analysis of contact problems has already been demonstrated in other studies [22, 23, 24, 25, 26, 27]. The present study should be regarded as an empirical investigation utilising MD simulations as an alternative to experiments, assuming that the simulated material behaviour is identical to the real behaviour of soft polymers. The simulations were designed based on the ideal Hertz contact between a non-deformable spherical indenter and a surface of a soft polymer. Nanoindentation simulations were conducted under various conditions, with the calculated properties subsequently compared to the equations of the Hertz model. The objective was to evaluate the applicability of the Hertz equations for approximate predictions of the real contact area for soft polymers where the basic assumptions are not met. Methods MD simulations were conducted using the GROMACS software package [28, 29, 30], which was run on NVIDIA Quadro P2000 graphics cards. The leap-frog algorithm [31] was employed for the integration of the equation of motion with a time step of 3 fs (or 1 fs for equilibration runs). The application of three-dimensional periodic boundary conditions enabled the simulation of condensed matter at the nanoscale. Nonbonded interactions were cut off at a distance of 1 nm with the Verlet algorithm for neighbour searching [32]. Temperature and pressure control were realised by the velocity-rescaling thermostat [33], and the anisotropic Berendsen barostat [34], respectively. Interaction energies were calculated using the GROMOS 53a6 force field [35, 36, 37]. Polyethylene (PE) was selected as a representative of soft polymers for a number of reasons, which were beneficial for the investigations: Its glass transition point lies considerably below ambient conditions, it is chemically relatively simple, and, as a semi-crystalline polymer, it offers the possibility of considering a range of deformation properties without affecting the chemical properties of the material. Starting from linear, unbranched polyethylene chains comprising 2000 carbon atoms, bulk simulation boxes of dense PE with edge lengths of at least 50 nm were generated via a subsequent heating and cooling process, whereby the different polymer chains underwent coiling, mixing and entanglement. By varying the cooling rate between 100 and 0.04 K/ ns, the degree of crystallisation could be adjusted between 0 and 50 %. A higher degree of crystallisation of the PE is associated with a higher density, a higher Young’s modulus and a lower Poisson’s Science and Research 31 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 distances in the xy direction. These atoms were not subject to mutual interaction but were constrained to maintain their respective distances. With regard to the interaction with the PE surface, the indenter atoms were treated as PE atoms. For the nanoindentation setup the indenter was positioned above the PE surface, as illustrated in Figure 2b. The bottom atoms of the PE layer model were fixed in space to prevent the model as a whole from moving through the application of external forces. Distance-controlled indentation simulations were conducted using the slow-growth functionality of GROMACS [39] to move the indenter towards the surface by a minimal defined distance every time step, determining the indentation velocity υ, which was varied in 4 stages (0.5, 1, 10, and 20 nm/ ns). The indentation process was simulated up to the full penetration of the spherical tip (p = R), but to a maximum of p = 15 nm to ensure, that the deformation was not affected by the limited size of the PE surface layer. For each parameter variant (c, R, v), a minimum of three determinations were conducted with the tip position varying along the surface to ensure statistical reliability. The normal forces were calculated by summing all the Science and Research 32 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 ratio (Figure 1). A more detailed description of the box generation method can be found in [38]. For the nanoindentation simulations PE with degrees of crystallization c of 0, 12, 30, 40 and 50 % was used. In order to form a 2D surface model and still utilise the faster 3D periodic routines, a vacuum slice with a thickness of 150 nm was introduced by enlarging the simulation box in the z-direction without modifying the atomic positions. The vacuum slice is of sufficient size to prevent interactions between the periodic PE slices, ensuring that the z-direction periodicity does not affect the surface. The simulation boxes were equilibrated for 3 ns to allow recombination on the freshly cleaved surfaces. Due to the presence of stable crystallites within the surface plane, the obtained surface models exhibited a certain nanoscale roughness (S a ≈ 1 nm). For the nanoindentation simulations, non-deformable indenters with spherical tips were produced. The initial basis geometry was a cone with an opening angle of 25°, the apex of which was rounded with a curvature radius R (3, 8, 10, 12, 14, and 18 nm). The indenter was composed of atoms positioned according to a mathematical calculation and arranged in a regular pattern with fixed Figure 2: a) Illustration of the symbols used to characterize the Hertz contact, b) setup of the nanoindentation simulations, illustrated by a detail of a cross-section through the simulation box of a polyethylene surface with c = 0.5 and an indentation radius of R = 8 nm, and c) exemplary snapshot of tip atoms in contact with PE atoms projected onto the xy-plane, together with the according contact circle of area A and radius a Figure 1: Relation between the simulated values of crystallinity c, cooling time τ c , density p, Young’s modulus E and Poisson ration υ for the semi-crystalline PE model polymer forces acting on the PE atoms in the z-direction. In order to quantify the size of the real contact area, the number of tip atoms that met a specific distance criterion (0.65 nm) to the PE atoms was counted for each time step. As a result of the uniform distribution of the tip atoms, geometric considerations could be employed to calculate the penetration depth of the tip p, and both area A and radius a of the contact circle parallel to the xy plane from the determined contact number (Figure 2c). Results and discussion In order to ensure the greatest possible comparability between the Hertz model and the simulations, the present investigation focuses on the short-term deformation behaviour. Given the applied velocities, it can be reasonably assumed that slow processes, such as material deformation caused by viscoelasticity or adhesion, can be disregarded during the indentation part and are only of relevance for the retraction part of the nanoindentation, which shall not be discussed here. In the absence of viscoelasticity and adhesion, the dimensions of the contact area can be assumed to be independent of previous states, and only dependent on the current position of the indenter. This assumption was tested through a series of simulations, which were repeated identically, apart from the adhesive force between the indenter and the surface (realised by the variation of the interaction potential between the tip and the PE atoms by ±50 % relative to the original values). The results showed no effects on the formation of the real contact area or the acting forces during the indentation part of the simulation (while significant effects were observed during the retraction part, which will be discussed elsewhere). Thereby, In the following, the term ‘data set’ is used to refer to the corresponding values of p, d, A and a, resulting from an indentation process with specific values of c, R, and v after time t in determination x. The analysis of the present study is based on a total of 23.400 of such data sets. In comparison to experiments, simulations facilitate a straightforward quantification of the real contact area. Nevertheless, the indentation depth is less clearly defined on the nanoscale for a nanoscopically rough surface. In the present study, two distinct definitions were employed for the indentation depth. In the first definition, d c represents the distance in the z-direction between the indenter positions at the specified time step and the initial atomic contact time step. This definition allows for a unique determination of the indentation depth within the simulation, is reasonably comparable to the Hertz model and will be used in the first part of the presented analysis. Nevertheless, since the atomic contact is not easily accessible within experiments, experimental setups typically employ an increasing force to determine the initial position of the surface. This definition was also applied in the presented simulations, resulting in d f , which was then used in the second part of the comparison with the Hertz model. Figure 3 shows a scatter plot of corresponding values of d c and d f for all data sets. Despite the inherent uncertainty in d f , stemming from the conjunction of high velocities and the statistical noise associated with the calculated normal force, d c and d f exhibit a constant shift of approximately 1 nm. This shift is observed to be independent of the indenter radius and the crystallinity of the PE. The shift between d c and d f can be attributed to adhesion forces. The attractive interactions between the atoms of the indenter and the PE result in a negative normal force at the initial atomic contact. It is only when repulsive forces due to material displacement exceed the attractive forces that the normal force increases, which is interpreted as reaching the surface in terms of d f . In comparison to harder materials, the necessary penetration depth to overcome the attractive forces is significantly greater for soft polymers, due to the lower modulus of deformation. This is the reason for the discrepancy between d c and d f , which should have minimal impact at the macro level but significant effects at the nanoscale. According to the Hertz model [7], the indentation depth d of a spherical indenter with radius R results in a displacement u z of the elastic surface within the contact area, that depends on the distance r in xy-direction to the con- Science and Research 33 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 Figure 3: Relationship between the differently defined values for the indentation depth for all datasets; the values are shifted by a constant of ≈ 1 nm independently of indenter radius R or crystallinity c (6) In order to facilitate a comparison between these theoretical considerations and the conducted simulations, the proportionality a 2 = Rd, as defined in Equation (2), was examined for all data sets, with d c representing the indentation depth. At the nanoscale, the material does not exhibit the characteristics of a homogeneous, isotropic continuum. Instead, it deforms anisotropically due to the semi-crystalline structure, which gives rise to considerable fluctuations in the calculated quantities. Nevertheless, when considering the mean behaviour across all data sets, it can be seen from Figure 4a that a 2 = Rd seems to be a very good approximation for the whole data range. This means, the approximation is not exclusive to small deformations, as given by the Hertz model, but can also be reasonably well applied to large deformations extending up to the indenter radius, despite the violation of all crucial assumptions. This is an unexpected result, given that for large deformations, the linear relationship between a 2 and p is no longer valid and must be replaced by the non-linear relation described in Equation (6) (Figure 4b). Concurrently, however, an increasing deviation from the relationship p = d/ 2 (Equation (5)) is also observed with increasing indentation depth (Figure 4c). This appears to compensate for the non-linear relationship between a 2 and p, resulting in a seemingly linear relationship between a 2 and d. Thus, it can be seen that the violation of the small deformation condition results in significant discrepancies from the Equations (1) and (5), yet it nonetheless produces a seemingly identical relationship to Equation (2). As previously stated, this is not a consequence of material deformation caused by adhesion or viscoelasticity. Given that Equation (2) offers a satisfactory approximation of the simulated contacts, even in the context of significant deformations, Equation (4) was similarly considered by examining the proportionalities F ~ a 3 , F ~ 1⁄R, and F ~ E * . and . In conducting this analysis, only those simulations with a velocity of 20 nm/ ns were considered, as this velocity was also applied for the cal- 3 4 5 67 8 4 9 Science and Research 34 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 tact point at r = 0 (Figure 2a). Equation (1) is derived from the premise that a spherical surface can be approximated with a high degree of accuracy by a seconddegree polynomial in the vicinity of the contact point. This implies that the approximation is only applicable to small deformations, where the displacement around the contact point is the primary concern. (1) By identifying a pressure distribution that causes the displacement described in Equation (1) and integrating it over the contact area, the Equations (2) and (3) for the contact radius a and the normal force F, respectively, can be derived [7]. (2) (3) Equation (2) and (3) can be combined to give Equation (4), which relates the normal force with the contact radius. (4) By employing Equation (1) to ascertain the displacement u z at the boundary of the contact area r = a with the aid of Equation (2), one arrives at Equation (5), which asserts that the penetration depth p is precisely half of the indentation depth d. (5) Obviously, Equations (1) to (5) are only valid for values of d that are small compared to R. Considering the simple case of full penetration (p = R), it can be seen from Equation (2) that the approximation already causes an error of 50 % for the real contact area (a 2 = 2R 2 compared to the exact geometrical solution of a 2 = R 2 ). For geometric reasons, the precise relation between the contact radius a and the penetration depth p for sphere segment is given by Equation (6). : $ 5 8 ; 4 67 9 3 4 5 7 9 < 5 > ? @ A 8 B 4 C 7 D 9 < 5 > ? @ EA 8 B 4 F 3 D 7 5 > ? @ G 7 3 D 9 5 6 9 Figure 4: Observed relations between the contact radius a, the indentation depth compared to first atomic contact d c and the penetration depth p for all simulated data sets culation of the bulk deformation properties in Figure 1c, which are markedly velocity-dependent. As the example simulation results in Figure 5a show, there is a perfect correlation between F and a 3 not only for small deformations as proposed by Equation (4), but over the entire range of indentation depths up to the indenter radius. According to Equation (4), the slope of the regression line m should be proportional to the inverse of the indenter radius 1/ R. This hypothesis was tested and confirmed in Figure 5b for simulations with the same degree of crystallisation but different indenter radii. According to Equation ( 4 ), the slope of the regression line should now be solely dependent on the deformation characteristics of the surface E * = E ⁄ (1 - υ 2 ). E * was calculated in this way for all the considered PE surfaces with different degree of crystallisation. Despite the excellent reflection of the relationships expressed in Equation (4) in the simulation results, Figure 5c demonstrates a significant discrepancy between the calculated E *values from indentation simulations and the E * -values determined from bulk simulations (see Figure 1c). While a similar tendency is evident, a systematic deviation is also present. The discrepancy between the deformation properties determined from indentation and bulk simulations was also confirmed qualitatively by experiments, which will not be illustrated in detail here. Creep experiments were conducted on a variety of different samples of PE, p-PVC, PUR, and silicones, using spherical indenters with radii between 0.8 and 20 mm and constant loads ranging from 20 to 80 mN. In order to obtain the Young’s modulus, the measured creep curves were fitted with the equation provided by [40], which accounts for the viscoelasticity of the material. Despite the fact that the fits resampled the measured curves with great precision and the small deformation condition was fulfilled, the obtained Young’s moduli exhibited a notable and systematic discrepancy from the values obtained through standardised tensile tests. This finding aligns with the simulation results. The findings presented thus far have utilised the indentation depth defined by the initial atomic contact, d c . For d f values, which are more appropriate for experimental setups but exhibit a constant shift in comparison to d c (see Figure 3), there is a notable discrepancy between the analytical equations and the simulation results. Ne- Science and Research 35 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 Figure 5: Evaluation of the relationship in equation (4) for the conducted simulations by testing the proportionality between a) F and a 3 , b) F and 1/ R, and c) the resulting E * from indentation simulations and the calculated E * from bulk simulations Figure 6: Utilizing the definition of the indentation depth based on the force, d f , to get approximate values of the real contact area: a) approximation of d c via equation (7), b) and c) approximation of p and A via equation (8) Acknowledgement The research project “Determination of the real contact area of soft polymers” (49VF190053) was financially supported by German Federal Ministry for Economic Affairs and Climate Action (BMWK). References [1] Jacobs T D B and Martini A 2017 Measuring and Understanding Contact Area at the Nanoscale: A Review Appl. Mech. Rev. 69 060802. [2] Persson B N J, Albohr O, Tartaglino U, Volokitin A I and Tosatti E 2004 On the nature of surface roughness with application to contact mechanics, sealing, rubber friction and adhesion J. Phys. Condens. Matt 17 R1. [3] Persson B N J 2006 Contact mechanics for randomly rough surfaces Surf. Sci. Rep. 61 201-227. [4] Klüppel M and Heinrich G 2000 Rubber Friction on selfaffine Road Tracks Rubber Chem. Technol. 73(4) 578- 606. [5] Hertz H 1881 Über die Berührung fester elastischer Körper J. F. reine u. angew. Math. 92 159-171. [6] Johnson K L, Kendall K and Roberts A D 1971 Surface Energy and the Contact of Elastic Solids Proc. R. Soc. Lond. A Math. Phys. Sci. 324(1558) 301-313. [7] Popov V L 2017 Contact Mechanics and Friction Springer Heidelberg Dordrecht London New York 2 nd edn.. [8] Johnson K L 1955 Surface interaction between elastically loaded bodies under tangential forces Proc. Roy. Soc. A320 531-548. [9] Banks H T, Hu S and Kenz Z R 2011 A Brief Review of Elasticity and Viscoelasticity for Solids Adv. Appl. Math. Mech. 3 1-51. [10] Chen W W, Wang Q J, Huan Z and Luo X 2011 Semi- Analytical Viscoelastic Contact Modeling of Polymer- Based Materials J. Tribol. 133 041404. [11] Greenwood J and Williamson J 1966 Contact of Nominally Flat Surfaces Proc. R. Soc. A 295(1442) 300-319. [12] Bush A W and Gibson R D 1975 The Elastic Contact of a Rough Surface Wear 35(1) 87-111. [13] Shull K R 2002 Contact mechanics and adhesion of soft solids Mater. Sci. Eng. R36 1-45. [14] Wu C-E, Lin K-H and, Juang J-Y 2016 Hertzian load-displacement relation holds for spherical indentation on soft elastic solids undergoing large deformations Tribol. Int. 97 71-76. [15] Dintwa E, Tijskens E and Ramon H 2008 On the accuracy of the Hertz model to describe the normal contact of soft elastic spheres Granul. Matter 10 209-221. [16] Lin Y-Y and Chen H-Y 2006 Effect of Large Deformation and Material Nonlinearity on the JKR (Johnson-Kendall- Roberts) Test of Soft Elastic Materials J. Polym. Sci. B44(19) 2912-2922. [17] Feng Z Q, Peyraut F and Labed N 2003 Solution of large deformation contact problems with friction between Blatz-Ko hyperelastic bodies Int. J. Eng. Sci. 41(19) 2213-2225. Science and Research 36 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 vertheless, there are two distinct approaches that can be employed to obtain approximate values for the real contact area based on d f . Since these approximations are based on empirical evidence, it is yet to be determined whether they can be applied to other systems and, in particular, to different length scales. The most evident possibility is to calculate d c from d f via Equation (7), that can be derived from Figure 3a. This produces a satisfactory agreement between the simulation results and Equation (2), as illustrated by the correlation in Figure 6a. (7) Nevertheless, a somewhat stronger agreement between the A-values calculated from the simulations and approximated from d f (Figure 6c) can be achieved by employing the observed empirical relationship p ≈ 0.83 d f (Figure 6b) in conjunction with the precise analytical relationship between p and a 2 (Equation (8)). (8) Conclusion The empirical simulations demonstrate that, under the specified conditions of investigation (nanometre level, PE, high velocities and short-term deformation behaviour), despite significant deviations from the fundamental assumptions (i.e., small deformations and ideal elastic behaviour), the Hertz Equations (2) and (4) provide a reasonable approximation for the real contact area. Nevertheless, it should be noted that the Young’s modulus in Equation (4) is not identical to the usual Young’s modulus that can be obtained from experimental tensile tests or bulk simulations. Furthermore, the indentation depth d is defined from the position of the first atomic contact, which is not easily accessible in experiments. Further approximations have to be utilized to calculate the real contact area from more customary values for the indentation depth d f , defined from increasing forces. Based on findings of these initial investigations, future work will address remaining questions, including: a) How can a suitable Young’s modulus be determined for predictions? b) Are the findings also transferable to other soft polymers and conditions (velocity, length scales)? c) To what extent do alternative models based on the Hertz model (including, but not limited to, time dependency, adhesion, rough surfaces, friction, and so forth) apply to soft polymers? H 5 I3 4 5 I E 67 8 4 F J IKALMM7 N 8 OLMP N 4 Q W J N X A9 RSTU999 J OLV? N 9 [18] Dimitri R, De Lorenzis L, Scott M A, Wriggers P, Taylor R L and Zavarise G 2014 Isogeometric large deformation frictionless contact using T-splines Comput. Methods Appl. Mech. Eng. 269 394-414. [19] Lin D C, Shreiber D I, Dimitriadis E K and Horkay F 2009 Spherical indentation of soft matter beyond the Hertzian regime: numerical and experimental validation of hyperelastic models Biomech. Model. Mechanobiol. 8 345-358. [20] Frenkel D and Smit B 2023 Understanding molecular simulation: from algorithms to applications Academic Press London California 3 rd edn.. [21] Zhang J, Wang Z, Yan Y and Sun T 2016 Concise Review: Recent Advances in Molecular Dynamics Simulation of Nanomachining of Metals Curr. Nanosci. 12(6) 653-665. [22] Yang C and Persson B N J 2008 Contact mechanics: contact area and interfacial separation from small contact to full contact J. Phys.: Condens. Matter 20 215214. [23] Solhjoo S and Vakis A I 2015 Definition and Detection of Contact in Atomistic Simulations Comput. Mater. Sci. 109 172-182. [24] Vishnubhotla S B, Chen R, Khanal S, Hu X, Martini A and Jacobs T D B 2019 Matching Atomistic Simulations and In Situ Experiments to Investigate the Mechanics of Nanoscale Contacts Tribol. Lett. 67 97. [25] Li S, Li Q, Carpick R W, Gumbsch P, Liu X Z, Ding X, Sun J and Li J 2016 The Evolving Quality of Frictional Contact with Graphene Nature 539(7630) 541-545. [26] Spijker P, Anciaux G and Molinari J F 2012 The Effect of Loading on Surface Roughness at the Atomistic Level Comput. Mech. 50(3) 273-283. [27] Luan B and Robbins M O 2005 The Breakdown of Continuum Models for Mechanical Contacts Nature 435(7044) 929-932. [28] Abraham M J, Murtola T, Schulz R, Páll S, Smith J C, Hess B and Lindahl E 2015 GROMACS: High performance molecular simulations through multi-level parallelism from laptops to supercomputers SoftwareX 1-2 19- 25. [29] Berendsen H J C, van der Spoel D and van Drunen R 1995 GROMACS: A message-passing parallel molecular dynamics implementation Comput. Phys. Commun. 91(1- 3) 43-56. [30] Gromacs. Fast. Flexible. Free 2024 https: / / www.gromacs.org/ . [31] Hockney R W, Goel S P and Eastwood J 1974 Quiet High Resolution Computer Models of a Plasma J. Comput. Phys. 14(2) 148-158. [32] Páll S and Hess B 2013 A flexible algorithm for calculation pair interactions on SIMD architectures Comput. Phys. Commun. 184(12) 2641-2650. [33] Bussi G, Donadio D and Parrinello M 2007 Canonical sampling through velocity rescaling J. Chem. Phys. 126(1) 014101. [34] Berendsen H J C, Postma J P M, van Gunsteren W F, Dinola A and Haak J R 1984 Molecular dynamics with coupling to an external bath J. Chemi. Phys. 81(8) 3684- 3690. [35] Daura X, Mark A E and van Gunsteren W F J 1998 Parametrization of Aliphatic CHn United Atoms of GRO- MOS96 Force Field J. Comput. Chem. 19(5) 535-547. [36] Oostenbrink C, Villa A, Mark A E and van Gunsteren W F 2004 A Biomolecular Force Field Based on the Free Enthalpy of Hydration and Solvation: The GROMOS Force-Field Parameter Sets 53A5 and 53A6 J. Comp. Chem. 25(13) 1656-1676. [37] Schuler L D, Daura X and van Gunsteren W F J 2001 An Improved GROMOS96 Force Field for Aliphatic Hydrocarbons in the Condensed Phase J. Comput. Chem. 22(11) 1205-1218. [38] Fritz S 2021 Considering semi-crystallinity in molecular simulations of mechanical polymer properties - using nanoindentation of polyethylene as an example CMMS21(1) 35-50. [39] Gromacs 2024.2 Manual (https: / / doi.org/ 10.5281/ zenodo.11148638). [40] Cheng L, Xia X, Scriven L E and Gerberich W W 2005 Spherical-tip indentation of viscoelastic material Mech. Mater. 37 213-226. Science and Research 37 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0010 lubricant systems, matching conventional lubricants such as polyalphaolefin (PAO-09), polyethylene glycol (PEG400), and an 86 % glycerol solution (G-86). Sheardependent viscosity was characterized and modeled using a Carreau-Yasuda approach. Friction properties were evaluated with a Mini Traction Machine (MTM2) tribometer, complemented by optical interferometry measurements to determine central and minimum lubricant film thickness under various conditions. Results and Discussion Rheological investigations revealed that original high molecular weight HEC derivatives (HCL and HCS) exhibited pronounced shear-thinning behavior, resulting in thin lubricant films. Enzymatic hydrolysis significantly reduced the molecular weight, eliminating shear-thinning and enhancing suitability for EHD lubrication. Resultant Science and Research 38 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0011 Introduction The development of environmentally friendly lubricants is a key concern in modern tribology. In view of the ecological challenges posed by fossil lubricants and increasing regulatory pressure, the use of bio-based, watersoluble systems is becoming an increasingly important focus of development ii,iii . Lubricants with a functional water content offer an attractive alternative to mineral oil-based products due to their environmental friendliness, high availability, and, last but not least, their nonflammable nature iv,v,vi . The combination of water with polymeric viscosity modifiers makes it possible to create hydrodynamic lubricating films even in highly stressed tribological contacts. The present study deals with the development and investigation of such systems based on hydroxyethyl cellulose (HEC) and glycerin, with a particular focus on their behavior under elastohydrodynamic (EHD) lubrication. Materials and Methods Commercially available hydroxyethyl celluloses, specifically adjusted to various molecular weights through enzymatic hydrolysis, were used in this study. The weight-average molecular mass was reduced to approximately 19 kg/ mol (see Figure 1). Additionally, glycerol (G) was used in weight proportions between 40 % and 70 % to adjust the viscosity of the Cellulose in Motion: Enzymatically Modified Biopolymers and Glycerol in Tribological Interaction Sandra Kiese, Daniela Leistl, Jan Ulrich Michaelis, Stefan Hofmann, Thomas Lohner* environmentally friendly, bio-based, water-soluble, hydroxyethyl cellulose, elastohydrodynamic Keywords * Dr. Sandra Kiese Dr.-Ing. Daniela Leistl M.Sc. Jan Ulrich Michaelis Fraunhofer Institute for Process Engineering and Packaging IVV Giggenhauser Straße 35, 85354 Freising, Germany M.Sc. Stefan Hofmann Dr.-Ing. Thomas Lohner Gear Research Center (FZG), Department of Mechanical Engineering School of Engineering and Design Technical University of Munich Boltzmannstraße 15, 85748 Garching, Germany Note: This article is an abridged and adapted version of the original article published under the title “Elastohydrodynamic lubrication of aqueous hydroxyethyl cellulose-glycerol lubricants” in the journal *Tribology International* i . Figure 1: Differential distribution as a function of molecular weight for HCL, HCS, and HCH, adapted from Michaelis et al. i . hydroxyethyl cellulose hydrolysates (HCH) showed stable viscosity across relevant shear rates, substantially improving film formation capabilities. Glycerol’s role is notable as it serves not only as a viscosity modifier but also increases the pressure-viscosity coefficient. Figure 2 shows the influence of different HEC types with varying molecular weights and glycerin content on the coefficient of friction in aqueous formulations as a function of the mean velocity v m . As expected, friction decreases with increasing speed in all formulations - albeit to very different degrees. Compared to the reference PEG400, all HEC-glycerol mixtures show improved friction performance at medium to high speeds. Figure 3 shows the influence of different glycerol concentrations in combination with hydrolyzed hydroxyethyl cellulose (HCH) on the coefficient of friction. The results show that all glycerin-HCH systems have significantly lower friction coefficients than the reference oils PAO-09 and PEG400 and the non-hydrolyzed samples. Compared to PAO-09, friction was reduced by more than 90 %. While PEG400 and PAO-09 show only moderate friction reductions with increasing speed, the glycerin-containing systems - especially those with glycerin contents above 50 wt% - exhibit a clearly pronounced speed dependence and a transition from mixed friction to fluid friction at speeds above approximately 1000 mm/ s. In combination with HCH, a friction coefficient of less than µ = 0.01 was achieved in formulations with a glycerin content of 50 % by weight - a range that is considered liquid superlubrication according to the current definition. Even with higher glycerin contents, favorable friction behavior is achieved, although a slight increase in friction coefficients was observed from approx. 70 % glycerin by weight, which is attributed to an increasing pressure viscosity coefficient. In addition, the central lubricating film thickness was measured in EHD contact (Figure 4). This showed that the hydrolyzed HCH types in combination with glycerin were able to form significantly higher lubricating film thicknesses than the non-hydrolyzed starting materials, whose lubricating film thicknesses were below 50 nm. The formation of a lubricating film correlates strongly with the effective viscosity in the inlet area of the EHD Science and Research 39 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0011 Figure 2: Friction coefficients as a function of mean velocity for HEC-glycerol formulations, adapted from Michaelis et al. i . Figure 3: Friction coefficients as a function of mean velocity for HCH-glycerol formulations, adapted from Michaelis et al. i . Figure 4: Central lubricant film thicknesses at varying mean velocities for different formulations, adapted from Michaelis et al. i . i Michaelis, J. U., Kiese, S., Hofmann, S., Lohner, T., & Eisner, P. (2025). Elastohydrodynamic lubrication of aqueous hydroxyethyl cellulose-glycerol lubricants. Tribology International, 110563. Volltext unter: https: / / www. sciencedirect.com/ science/ article/ pii/ S0301679X25000581 ii Shah, R., Woydt, M., & Zhang, S. (2021). The economic and environmental significance of sustainable lubricants. Lubricants, 9(2), 21. iii Pichler, J., Maria Eder, R., Besser, C., Pisarova, L., Dörr, N., Marchetti-Deschmann, M., & Frauscher, M. (2023). A comprehensive review of sustainable approaches for synthetic lubricant components. Green Chemistry Letters and Reviews, 16(1), 2185547. iv Schmidt, R., Klingenberg, G., & Woydt, M. (2006). Thermophysical and viscosimetric properties of environmentally acceptable lubricants. Industrial Lubrication and Tribology, 58(4), 210-224. v Shetty, P., Mu, L., & Shi, Y. (2020). Polyelectrolyte cellu lose gel with PEG/ water: Toward fully green lubricating grease. Carbohydrate polymers, 230, 115670. vi Rahman, M. H., Warneke, H., Webbert, H., Rodriguez, J., Austin, E., Tokunaga, K., ... & Menezes, P. L. (2021). Water-based lubricants: Development, properties, and performances. Lubricants, 9(8), 73. Science and Research 40 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 DOI 10.24053/ TuS-2025-0011 contact, with the glycerine concentration playing a particularly dominant role here. In the systems examined, the range for a good balance between low friction and sufficient lubricating film thickness was around 50 - 60 % glycerine in combination with 3 - 5 % HCH. Summary and Outlook The presented findings demonstrate that lubricants with functional water content based on hydroxyethyl cellulose and glycerol hold significant potential for technical applications, contributing substantially to bio-based tribological systems development and their use in numerous fields of application in which environmental aspects are playing an increasingly central role. The targeted adjustment of the polymer structure via enzymatic hydrolysis enables significant friction reduction, improved lubricating film formation, and favorable rheological behavior. At the same time, glycerin allows flexible control of viscosity and supports lubricating film formation even under high pressure and shear loads. News 41 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 As you could already read in the previous issue of “Tribologie und Schmierungstechnik”, the 66 th German Tribology Conference will take place for the first time in Wernigerode in the Harz Mountains. After the abstract submission is closed, the program committee has selected 56 scientific presentations on cutting-edge tribological topics, in particular on this year's main topic “Cost Reduction by Tribology”. Together with plenary lectures and separately organized sessions on special topics, you can look forward to more than 100 lectures in 6 parallel sessions. Among them are lectures from the research field "Tribology", which is funded by the Federal Ministry for Economic Affairs and Energy. In addition, the final colloquium of the DFG priority programme “Liquid-free lubrication systems with high mechanical loads” will take place as part of the conference. For the first time, two further sessions will also be offered, each with a special focus. One of these sessions will focus specifically on the interaction of additives with metallic surfaces, highlighting the acting mechanisms, formation of reaction layers, experimental work and the corresponding theories and models. An important topic, which will be particularly highlighted this year, is standardization and regulation. Lectures on standardization in tribology in general, but also in special areas such as carbon layers and hard ceramic layers will be given. Furthermore, the definition of wear parts and tribological reliability design are dealt with. You are still welcome to register for the TriboSlam, the unique science slam of tribology. Since 2022, the German Tribology Conference has proven every year how much fun tribology is and how exciting research can be. In just 360 seconds, researchers transform complex research into exciting stories that will inspire you too. As every year, an award ceremony will take place during the plenary session. An outstanding personality will be awarded by the Georg Vogelpohl Medal of Honour. Furthermore, young scientists and engineers who have achieved above-average performance in their bachelor's, master's and doctoral theses will receive GfT Sponsorship Awards. During the closing event, the GfT awards the prize “Tribology is everywhere”, donated by the company Werner Stehr Tribologie. This year’s winner will vividly present his research subject in a plenary lecture. A poster session and a trade exhibition, for which still places are available, complete the program. The GfT is looking forward to welcoming you to the new conference location. A registration form and further information can be found in conference program, which will the soon be published, and on the website www.gft-ev.de/ en/ german-tribology-conference-2025. Gesellschaft für Tribologie 66 th German Tribology Conference - September 29 to October 1, 2025 It is probably inevitable - on a 66 th birthday, we reflexively think of the classic song by Udo Jürgens: “Life begins at sixty-six, the best is yet to come”. In this respect, we see it as a sign of vitality that, after many years in Göttingen, the annual conference of the Society for Tribology GfT is venturing to move to a new location for its 66 th edition - to Wernigerode, the “colourful town on the Harz Mountains”, as the writer Hermann Löns once called the place. Is the change a good sign? Don’t we love reliability, predictability, simplicity and clarity? Instead, the Greek philosopher Heraclitus is said to have stated 2,500 years ago: “Nothing is more constant than change.” And let’s take a look at the current world situation - isn’t it characterised by volatility, uncertainty, complexity and ambiguity? Abbreviated boldly as VUCA, these terms have now found their way into trendy management seminars. Everyone should put their own house in order - what does this mean for tribology? Between constant change and the desire for stability, there is perhaps a key concept: Sustainability, or more precisely: thinking about ecological, economic and social sustainability together as a long-term anchor of stability. While CO 2 reduction in vehicles has long been a central driver of tribological development, other focal points will come into focus in the future: energy efficiency is increasingly in conflict with resource scarcity, political framework conditions and environmental regulations. Issues such as environmental compatibility, durability and recyclability must be given greater weight. Market distortions caused by unstable framework conditions require more robust products and faster new developments. The reduction in available solutions necessitates new approaches to materials and low-wear systems. Wear remains a key factor limiting service life - new lubricants, digital twins and AI-based condition monitoring are becoming increasingly important. The latter, in its current form, has been known for many years to reduce downtimes and repair costs. New methods for AI-supported early damage detection will further safeguard the reliable operation of energy systems. In the past, development times were heavily dominated by the scope and duration of experimental test series; in future, comprehensive simulation tools and efficient design methods will be essential to drastically reduce testing costs and remain competitive. This requires close co-operation between research and industry. The Research Association for Drive Technology (FVA) and the “Research Field on Tribology” are examples of long-term supporters. Let’s move on to the GfT’s own activities - there is some news to report from the Working Groups (WG). Under the overall direction of Prof. Bartel and in cooperation with the FVA, the Training and Further Education Working Group has continued the new seminar series. The seminar “Plain bearings - basics, calculation and use” took place for the second time in May 2025. New to the programme is the seminar “Fundamentals of Tribology - Friction, Wear & Lubrication”, which was held online in May 2025. There will be a continuation, also online, in September 2025 with “Fundamentals of Tribology - Materials, Damage & Methods”. Further events, e.g. on gearbox lubrication and rolling bearing lubrication, are in preparation. There is a new addition to the Regional Working Groups: the Hanover WG is being expanded to become the Hanover/ Northwest Regional WG under the leadership of Prof Max Marian. In addition to Leibniz Universität Hannover, Emden/ Leer University of Applied Sciences and Arts, Clausthal University of Technology and Paderborn University will also be participating. The Munich, Rhine-Neckar and Berlin-Brandenburg Working Groups are now offering their regular lecture events in a hybrid format, which will also enable supraregional participation - significantly increasing their reach. The Climate Protection & Sustainability Working Group, for which Dr Mathias Woydt is responsible, recently published the position paper “The concepts of ‘wear and tear’ and ‘wearing parts’ in the field of tension with ‘obsolescence’”. This activity should be seen in the context of emerging standardisation approaches. To draw your attention to this and already successfully running activities in the field of tribological standardization, a separate session on this topic will be offered for the first time at the German Tribology Conference. The Public Relations Working Group, headed by Dr Mirjam Bäse, is committed to presenting the tasks and benefits of the GfT even more prominently in the modern media world - informative and easily accessible on the News 42 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 Ladies and Gentlemen, Gesellschaft für Tribologie News 43 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 one hand, attractive for the younger generation on the other. To this end, we have launched a new website that endeavors more than ever to provide our target groups with comprehensive and up-to-date information. In addition, the GfT’s presence on social media will be further expanded to reach young tribologists in particular “on their channels”. The Young Tribologists Working Group has a new leadership with Klara Feile, Christian Orgeldinger and Dr. Benedict Rothammer. This WG will organize the 8 th Young Tribological Researcher Symposium on July 21 and 22, 2025. This year’s host is the Rhineland-Palatinate Technical University Kaiserslautern-Landau (RPTU). As always, this year’s GfT Annual Conference aims to offer you plenty of inspiration for new tribological approaches and solutions. In addition to already established special sessions, e.g., on the research field of tribology, it includes many exciting contributions, poster, and technical exhibitions and, of course, stimulating discussions “in between” in the circle of the GfT family! In addition to the strictly scientific contributions, you can also expect tribological curiosities, e.g., a plenary lecture on the tribology of wine tasting. And for friends of the brew, there will be an opportunity to get to know the TU Clausthal, which is celebrating its 250 th birthday this year, from a different angle on the evening before the GfT conference (28.09.): At the Institute of Tribology, we can put the results from the TU Clausthal research brewery to a practical test. Finally - after the conference is always before the conference - here is a reminder of the 3 rd International Tribology Conference nextlub, which will once again take place in Leipzig from January 20-21, 2027 as a joint format of Uniti, FVA and GfT. Preparations are already underway! We are now looking forward to seeing you again at our GfT annual conference, for the first time in beautiful Wernigerode am Harz! Rolf Luther, Chairman of the Executive Board of GfT Dr. Thomas Gradt, Managing Director of GfT News 44 Tribologie + Schmierungstechnik · volume 72 · issue 2/ 2025 BOOK RECOMMENDATION expert verlag - Ein Unternehmen der Narr Francke Attempto Verlag GmbH + Co. KG Dischingerweg 5 \ 72070 Tübingen \ Germany \ Tel. +49 (0)7071 97 97 0 \ info@narr.de \ www.narr.de This monograph takes a new look at tribology with its basic concepts of friction and wear using the example of lubricating greases. The consideration of the phenomenon of occurring instabilities and the introduction of the entropy concept into lubricating grease tribology provide a new perspective on known phenomena. The second part of this book presents a wide range of experimental possibilities for investigating lubricating greases. Contents Introduction to Instability and Postmodern Tribology - On the Phenomenon of Self - Organization - Postmodern Grease Tribology - Lubricating Grease - Rheological behavior of Lubricating greases - A Selected Traditional Wear Model - The Extension of the Wear Concept Erik Kuhn On the Tribology of Lubricating Greases An energetic approach to post-modern tribology Tribologie - Schmierung, Reibung, Verschleiß 1st edition 2025, approx. 210 p. €[D] 118,00 ISBN 978-3-381-14171-5 eISBN 978-3-381-14172-2 Checklist Author information Corresponding author: F Mailing address F Telephone and fax number F eMail All authors: F Academic titles F Full name F ORCID (optional) F Research instititute / company F Location and zip code Length F Approximately: 3,500 words Data F Word and pdf documents (both with images + captions) F Additionally, please send images as tif or jpg / 300 dpi / Please send vector data as eps Manuscript F Short and concise title F Keywords: 6-8 terms F Abstract (100 words) F Numbered pictures/ diagrams/ tables (please refer to the numbers in your text) F List of works cited After the typesetting is completed, you will receive the proofs, which you are requested to review and then give your approval to start the printing process. Changes to the manuscript are no longer possible at this stage. Please also consider The editors and the publisher assume that the authors are authorized to publish all data used, that the provided texts and all visual material (images/ pictures/ illustrations) do not violate any (copy)rights of a third party, and that, where necessary, source references are provided for visual material. In cases of doubt, please obtain a printing permission from the copyright holder. Editors and publisher cannot assume liability for potential copyright infringements. Open Access Free access to knowledge is important to us. That is why you also have the opportunity to make your contribution immediately available digitally to all interested parties. This not only benefits you with an increased reach, but also researchers worldwide. In order to guarantee the high quality and substantial indexing, we are unfortunately unable to offer this service free of charge. You can obtain the full open access service for a one-off article processing charge of € 1,850.00 (plus VAT). Editor in chief Dr. Manfred Jungk eMail: jungk@verlag.expert Publisher expert verlag Ein Unternehmen der Narr Francke Attempto Verlag GmbH + Co. KG Dischingerweg 5 D-72070 Tübingen Tel.: +49 (0)7071 97 556 0 eMail: info@verlag.expert www.expertverlag.de Editor Patrick Sorg eMail: sorg@verlag.expert Tel.: +49 (0)7071 97 556 57 Tribologie und Schmierungstechnik Tribology—Lubrication Friction Wear An Official Journal of Gesellschaft für Tribologie | An Official Journal of Österreichische Tribologische Gesellschaft | An Official Journal of Swiss Tribology We’re looking forward to your contribution! ISSN 0724-3472 Science and Research www.expertverlag.de Christof Koplin, Bernadette Schlüter, Raimund Jaeger The sensitive dependence of sealing elastomers on lubricants Alexander Roegnitz, Elias Merz, Hubert Mantz, Carsten Siemers, Andreas Haeger Tribological Investigation of the Novel Titanium Alloy TNTZ-O for Dental Implant Applications: Subsequent Results of a Comparative Study Felix Farrenkopf, Thomas Lohner, Karsten Stahl, Kirsten Bobzin, Christian Kalscheuer, Max Möbius, Marta Miranda Marti Dry Lubrication of Spur Gears Coated with CrAlN+Mo: W: S Susanne Fritz Insight into large deformation contacts of soft polymers with molecular dynamics simulations Sandra Kiese, Daniela Leistl, Jan Ulrich Michaelis, Stefan Hofmann, Thomas Lohner Cellulose in Motion: Enzymatically Modified Biopolymers and Glycerol in Tribological Interaction