Schmidt - New Lubrication Concepts for Environmentally Friedly Machines - BAM Coatings.
Forschungsbericht 277 Dr.-Ing. Roland Schmidt1 Dr. rer. nat. Günther Klingenberg1 Dr.-Ing. Mathias Woydt2 1 Physikalisch-Technische Bundesanstalt (PTB) Federal Institute for Materials Research and Testing (BAM) 2 New lubrication concepts for environmental friendly machines − Tribological, thermophysical and viscometric properties of lubricants interacting with triboactive materials − Research Report Nr. 277 Compiling the achievements of the project BMWA 14/02 Berlin and Braunschweig, Germany, 2006 Forschungsbericht 277 Berlin 2006 Forschungsbericht 277 Impressum Forschungsbericht 277: New lubrication concepts for environmental friendly machines − Tribological, thermophysical and viscometric properties of lubricants interacting with triboactive materials − 2006 Herausgeber: Bundesanstalt für Materialforschung und -prüfung (BAM) Unter den Eichen 87 12205 Berlin Telefon: +49 30 8104-0 Telefax: +49 30 8112029 E-Mail: [email protected] Internet: www.bam.de Copyright © 2006 by Bundesanstalt für Materialforschung und -prüfung (BAM) Verlag und Vertrieb: Wirtschaftsverlag NW Verlag für neue Wissenschaft GmbH 27568 Bremerhaven Telefon: +49 471 94544-0 Telefax: +49 471 94544-77 Umschlag: Lutz Mittenzwei Layout: BAM-Arbeitsgruppe Z.67 ISSN 0938-5533 ISBN 3-86509-528-3 2 Forschungsbericht 277 Summary The present research report was elaborated in close cooperation with Renault SAS, FUCHS Petrolub AG and Ingenieurgesellschaft Auto und Verkehr (IAV). The use of alternative oils for the lubrication of automobile engines has a potential of ecological and technical advantages. It requires the detailed knowledge of several thermophysical and viscometric properties in a large temperature range (mapping). Therefore, the following properties of up to twenty-eight different oils have been measured in the temperature range from 22 °C to 150 °C: density, heat capacity, thermal conductivity, viscosity at ambient pressure, viscosity under shear rates above 106 s-1, and the viscosity at elevated pressures (maximum 100 MPa). The last two have been measured with a substantially improved and a newly developed apparatus, respectively. The pressure-viscosity coefﬁcient has been measured on four hydrocarbon-based, factory-ﬁll oils, a parafﬁn oil and twenty-three alternative oils. Nine of the alternative oils are based partly or completely on esters, the other fourteen on polyglycols, two of them additionally on water. Based on the piston ring/cylinder liner simulation tests of BAM performed outside of engines and the SRV® tests both performed only under conditions of mixed/boundary lubrication, it is reasonable that thermally sprayed TiOx-based, Tin-2Cr2O2n-1 and (Ti,Mo)(C,N)+23NiMo piston ring coatings, so called “lubricious or triboactive oxides”, can substitute common materials and serve as a promising alternative to commercial piston ring coatings made of strategic Molybdenum and super-ﬁnishing intensive blends of WC/Cr3C2. Some couples qualiﬁed for “zero” wear. In combination with bionotox ester- and polyglycol-based lubricants the coefﬁcient of friction can be reduced fulﬁlling simultaneously stronger European exhaust emission regulations. Thermally sprayed Ti-based coatings with their high wear resistance can additionally be used on aluminium liners to increase the resistance of critical components against wear, adhesive wear and thermomechanical stresses. For given tribological test conditions all APS1 coatings on piston rings showed no friction reducing effect. The coefﬁcient of friction is more determined by the lubricants than by the materials or by an individual interaction between lubricants and a speciﬁc material or tribopairing. Lubricious oxides or triboactive materials and/or polar base oils may substitute the extreme pressure (EP) and anti-wear (AW) properties realized by the additives, thus enabling long drains and responding to “eco-tox” or “bio-no-tox” requirements as well as restrictions from the “chemical box”. Overall, the different polymer-free bionotox and low-ash prototype engine oils with reduced additive contents displayed isoperformance regarding the tribological behaviour against cast iron with high carbon content and triboreactive materials. Keywords Ester, polyglycol, PAG, PPG, factory ﬁll, hydrocarbon, engine oil, bio-oils, eco-lubricants, EAL, bio-no-tox oils, heat capacity, density, viscosity, pressure-viscosity, viscosity at high shear rate, thermal conductivity, mixed, boundary, lubrication, low sap, mid sap, wear, friction, triboactive materials, water-based oils, steam 1 abbeviation for “atmospheric plasma spraying” 3 Forschungsbericht 277 Zusammenfassung Der vorliegende Forschungsbericht entstand in enger Zusammenarbeit mit der Renault SAS, der FUCHS Petrolub AG und der Ingenieurgesellschaft Auto und Verkehr (IAV). Die Anwendungsfähigkeit alternativer Schmierstoffformulierungen in Verbrennungsmotoren hängt von der umfassenden Kenntnis des funktionalen Eigenschaftsproﬁles ab. Dazu ist die detaillierte Kenntnis thermophysikalischer und viskosimetrischer Größen in einem weiten Temperatur- und Druckbereich erforderlich. Daher wurden folgende Größen an bis zu 28 verschiedenen Ölen im Temperaturbereich von 22 °C bis 150 °C gemessen: Dichte, Wärmekapazität, Wärmeleitfähigkeit, Viskosität bei Atmosphärendruck, Viskosität bei Schergeschwindigkeiten bis 106 s-1 und die Viskosität bei erhöhten Drücken (maximal 100 MPa). Die beiden letzten Größen wurden mit einer grundlegend verbesserten bzw. mit einer neu entwickelten Apparatur gemessen. An vier Ölen auf Kohlenwasserstoffbasis, einem unadditivierten Parafﬁnöl und 23 alternativen Ölen wurde der Druckkoefﬁzient der Viskosität gemessen. Neun der alternativen Öle basierten teilweise oder vollständig auf Estern, die anderen 14 auf Polyglykolen, zwei davon zusätzlich auf Wasser. Die außermotorische Charakterisierung des tribologischen Verhaltens des Tribosystems „Kolbenring/Zylinderbahn“ unter Misch-/Grenzreibung beruhte auf zwei völlig verschiedenen Testphilosophien: dem BAM- sowie dem SRV®-Test. Im Rahmen des Projektes neuentwickelte, thermisch gespritzte, TiOx – und Tin-2Cr2O2n-1 – basierte und (Ti,Mo)(C,N)+23NiMo Kolbenringbeschichtungen, so genannte „schmierwirksame oder triboaktive Oxide“, offenbarten sich als vielversprechende Alternativen zu den kommerziellen Kolbenringbeschichtungen auf Basis von Molybdän und der endbearbeitungsintensiven Hartmetallbeschichtung aus WC/Cr3C2. Einige neuentwickelte Werkstoffpaarungen offerieren sogar „Null-Verschleiß“. In Verbindung mit den biologisch schnell abbaubaren und Bionotox-Schmiermitteln auf Ester- und Polyglykol–Basis können die Misch-/Grenzreibungszahlen nachhaltig reduziert werden und außerdem können die strengeren europäischen Abgasemissionsvorschriften eingehalten werden, da diese Formulierungen entweder aschearm oder aschefrei sind und/oder über „lean burn“-Eigenschaften verfügen. Thermisch gespritzte Beschichtungen auf Ti-Basis mit ihrer hohen Verschleißbeständigkeit können zusätzlich auf Aluminium–Zylinderbahnen aufgebracht werden, um den Verschleißwiderstand kritischer Komponenten auf das Niveau von hochgekohltem Grauguß zu bringen. Alle APS1-Beschichtungen auf Kolbenringen zeigten unter den verwendeten tribologischen Testbedingungen keinen die Reibungszahl verringernden Effekt. Unter Misch-/Grenzreibung bestimmen eher die Schmierstoffformulierungen die Reibungszahl, wobei in bestimmten Kombinationen durch individuelle Wechselwirkungen zwischen den Schmierstoffen und Werkstoffoberﬂächen niedrige Reibungszahlen gemessen wurden. Schmierwirksame Oxide oder triboaktive Materialien und/oder polare Basisöle können die Hochdruck(EP) – und Verschleißschutz(AW) – Eigenschaften der Additive substituieren. So sind verlängerte Ölwechselintervalle möglich, die Erfüllung der Zielforderungen „eco–tox“ oder „bio–no–tox“ sowie die jüngst sich aus der „chemical box“ ableitenden Restriktionen können funktional eingehalten werden. Trotz des abgesenkten Additivgehaltes zeigten die verschiedenen polymerfreien, biologisch schnell-abbaubaren Prototypenformulierungen mit reduzierten Aschegehalten und verringertem Additivkonzentrationen gegenüber hochgekohltem Grauguß und den triboaktiven Werkstoffen keine tribologischen Nachteile im Vergleich zu Erstbefüllungsölen auf Basis von Kohlenwasserstoffen. Schlüsselwörter Ester, Polyglykole, PAG, PPG, Erstbefüllung, Kohlenwasserstoff, Bioöl, Bio-no-tox-Öl, Wärmekapazität, Dichte, Viskosität, Druckviskosität, Viskosität unter hohen Scherraten, Wärmeleitfähigkeit, Misch/Grenzreibung, Schmierung, lowsap, midsap, Verschleiß, Reibung, triboaktive Werkstoffe, wasserbasiertes Öl, Dampf 1) Abkürzung für „atmospheric plasma spraying“ 4 Forschungsbericht 277 Contents 1 Introduction 7 1.1 General context for internal combustion engines 7 1.2 Steam technology 7 2 Tested Lubricants 8 3 Equipment used for the measurements of viscometric and thermophysical properties and tribological behavior 10 3.1 Viscosity at ambient pressure 10 3.2 Density 10 3.3 High-pressure viscosity 10 3.4 Heat capacity 11 3.5 Thermal conductivity 11 3.6 Tribological testing outside of engines 12 3.7 Tribological materials 13 3.7.1 Spray powder 13 3.7.2 Cylinder liner materials 14 3.7.3 Piston ring materials 14 3.7.4 Unlubricated sliding wear 17 4 Results of the measurements of viscometric and thermophysical properties 17 4.1 Density 17 4.2 Heat capacity 19 4.3 Thermal conductivity 21 4.4 Viscosity at ambient pressure 21 4.4.1 Viscosity of the oils in group 1 21 4.4.2 Viscosity of the oils in group 2 23 4.4.3 Viscosity of the oils in group 3 23 4.4.4 The function η(T) 25 4.5 High-Pressure-viscosity 25 4.5.1 Measurement program 25 4.5.2 Qualitative results 25 4.5.3 Data analysis and presentation 25 4.5.4 Shape of the function α(p) 27 4.5.5 The function α(T) 27 4.5.6 Results for α(T) 28 4.6 Film-forming behavior 28 4.6.1 Equations describing minimum ﬁlm thickness 29 4.6.2 Parameters 29 4.6.3 Inﬂuence of lubricant properties 30 5 Forschungsbericht 277 4.7 Relative ﬁlm thicknesses 30 5 Viscosity measurement at high shear rates up to 3,4 ⋅ 106 s-1 33 5.1 Description of the apparatus 33 5.2 Results 36 5.3 Estimation of the measurement uncertainty 37 6 Tribological behavior under continuous sliding (BAM-method) 37 6.1 TOTAL HC 5W-30 fresh oil and as engine aged with soot 37 6.2 FUCHS Titan GT1 37 6.3 TOTAL HCE midSAP 37 6.4 FUCHS HCE lowSAP 38 6.5 PPG 32-2 38 6.6 PAG 46-4 38 6.7 GGL20HCN 39 6.8 (Ti,Mo)(C,N)-23NiMo liner coating 39 6.9 Ti2-nCr2O2n-1 liner coating 40 6.10 TinO2n-1 ring coatings 40 6.11 Ester oil 41 6.12 Zero wear target 41 6.13 Summarizing friction and wear behavior in BAM test 42 7 Tribological behavior under linear, oscillating sliding (SRV®-method) 52 ® 7.1 Extreme pressure behavior in the SRV test 52 7.2 Friction and wear 52 7.3 Precision of SRV® test 53 8 Concluding summary 56 9 Literature/References 58 6 Forschungsbericht 277 1 Introduction More and more, the impact of engine oils on durability of particulate ﬁlters and catalysts has to be minimized or avoided, on fuel economy (FE) maximized, as well as their impact on terrestrial and aquatic environment. Replacing hydrocarbonbased oils with environmental friendly products is one of the ways to reduce adverse effects on the ecosystem caused by the use of lubricants. The competition between hydrocarbons and new alternative base oils is not yet technologically decided in favor for hydrocarbons or esters or polyglycols. 1.1 General context for internal combustion engines Original equipment manufacturers (OEMs) are more and more interested in passenger car engine oils (PCMO) with reduced metal-organic additives thus contributing to the vision of an environmentally friendly and sustainable car. This is necessary in order to reduce the ash build-up in the after-treatment system caused by engine oils and therefore improve its ﬁlter efﬁciency and lifetime. High fuel efﬁciency retention and long drain intervals are expected, as well, from the engine oils. Easy removal of bio-no-tox-ﬂuids and recycling supports a sustainable development. As displayed by the RENAULT demonstrator ELLYPSE  and the FORD Model U, additional requirements may be in the future demanded, like a. biodegradability and non-toxicity and/or b. a content of renewables. The criteria for attribution of the european environmental label “EUROMARGUERITE” require for hydraulic ﬂuids a content of >50 % of renewables. A smaller ﬁgure was proposed for engine oils . Besides, the fragmentation of standardized oil speciﬁcations between Europe, Asia and US persists, and the diversiﬁcation in original equipment manufacturer (OEM) speciﬁcations is spreading more and more since engine designs requiring speciﬁc oil formulations or using speciﬁc combustion processes have been released. Pure hydrocarbons it self can be US-FDA proof. The additive packages, which make hydrocarbons functional, determine the eco-tox and/or bio-no-tox and/or ash formation properties of hydrocarbon based formulations. It is obvious to look for the substitution of critical additives by others or new functional concepts, a. EP/AW properties by triboactice materials and coatings and/or b. Viscosity improvers by the high VI of base oils, like esters and polygycols and/or c. Polar base oil molecules for lubricity. One of the key questions is: How will the 2010+ engine oil look like? The two main tasks of engine lubricants are energy saving (friction, FE) and wear prevention. The ﬁrst task requires comparatively low viscosities at low temperatures and reduced coefﬁcients of friction under mixed/boundary lubrication. The second task − which is another key issue of this research report - is connected with the ability of the lubricant to form a liquid ﬁlm that separates the moving surfaces of the engine tribosystems from each other at high temperatures (in the case of IC engines at up to 150 °C). The higher the ﬁlm thickness, the lower is the risk of direct contact of surface asperities which might damage the surface. The most critical tribosystems in an engine are: a. the cam/follower (highest contact pressures, moderate sliding speed), b. the piston ring /cylinder (lower contact pressures, high sliding speed) and c. the crank shaft (highest sliding speeds, moderate pressure). Engine designers seeking for alternative engine oils need a methodology to compare the hydrodynamic ﬁlm forming behavior of base oils and formulations which are chemically completely different. The existing criteria in the oil speciﬁcations (ν40C, ν100C and HTHS) seem to be not descriptive enough. Heat capacity and thermal conductivity are other important issues for comparing alternative oils [3, 4], and are therefore discussed in this research report, too. The frictional and wear behavior of alternative lubricants when interacting with current state-of-the-art materials and new, triboactive materials need to be mapped. 1.2 Steam technology The thermodynamic and caloric properties of steam makes it attractive for heat conversion and propulsion systems. Combining a 19th century technology (steam Tmax.≈ 280 °C and <28 bar) with the advanced materials, design tools and manufacturing processes of the 21st century for steam with 600 °C and up to 100 bar could truly result in revolutionary new steam applications. A sound understanding and tribological data base is needed to ensure the success of these machines. The success of the development of advanced water-lubricated steam engine systems depends strongly on the identiﬁcation of triboactive materials and water-based crank shaft lubricants. This approach was nowadays pushed by IAV  GmbH (Ingenieurgesellschaft Auto und Verkehr GmbH, www.iav.de, ca. 50 patent applications for steam engines) with the development of a three cylinder reciprocating steam engine (Zero Emission Engine) using the Rankine cycle. This work is today continued by EGINION/AMOVIS for APUs (Auxiliary Power Units) in passenger cars and trucks as well as for cogeneration (SteamCell®), a venture capital ﬁnanced company. Also in Germany was recently marketed of linear, reciprocating steam engine “Lion” for cogeneration (www.otag.de). Spilling Energiesysteme  (see DE29906867U; EP1045128) markets since 2001 oil-free, reciprocating steam engines up to 2 MW for heat recovery using steam at 30 bar and 300 °C, but displayed a clear trend to use 450 °C and 60 bar. 7 Forschungsbericht 277 Yankee Scientiﬁc, Inc., (www.yankeescientiﬁc.com e.g. www. climate-energy.com) develops a steam scroll expander for cogeneration or energy supply. The miniaturization, simpliﬁcation and cost reduction of system components is achieved through the use of a two-phase working ﬂuid and an oil-free positive-displacement scroll expander. Beginning 2006, BMW AG unveiled the concept of a hydrid propulsion system for passenger cars combining inline “classic” IC engine with a reciprocating steam expander directly linked to the crank shaft using waste exhaust heat to generate steam. Steam technology is considered to be from the material science point of view more proven than fuel cells and much cheaper per kW. 2 Tested Lubricants In future, lubricants will more and more determine the functional performances and environmental properties of machineries and engines. Esters and polyglycols were identiﬁed as alternative base oils and blended to environmental friendly prototype engine oils meeting following properties: Group 1 a. low viscosity, b. low contributions to exhaust emissions (lean burning), c. high oxidative stability, d. high biodegradability and e. low toxicity (bio-no-tox) as well as f. low ash formation or ash-free and g. polymer-free. Three factory-ﬁll, hydrocarbon-based engine oils as high performance formulations Titan SL PCX 0W-30, Castrol SLX HC 0W-30 and TOTAL HC 5W-30 served as references with a HTHS of ca. 3.0 mPas (target for the prototype oils), also for the tribological properties under mixed/boundary lubrication. The results from the eco-toxicological tests (Erebio-ECproject) are published elsewhere [7, 8]. Polymer-free lubricants are advantageous for direct injecting engines in view of deposit formation on intake valves. The content of ash and metal must be limited because the exhaust treatment devices and in consequence the fuel economy might be inﬂuenced by it. Supplementarily, the requirement of reduced sulfur and phosphorus contents was taken into consideration. Bio-no-tox engine oils offer the chance of better fuel efﬁciency, i.e. lower fuel consumption. This is a relevant contribution to the actual European “Climate Change Policy”. Additionally, environmentally compatible engine oils based on synthetic esters can be formulated on renewable raw material, consequently offering further CO2 savings. Vegetable oils are one of the major source of these synthetic base ﬂuids. Group 1 Oils based on hydrocarbons and /or blends with esters HC 5W-30 HC 5W-30 + 3.7% soot Fuchs HCE 0W-20 Total 100E 100E 0W-20 Total HCE Titan SL PCX 0W-30 Castrol HC 0W-30 (SLX) Fuchs HCE-Low-SAP Total HCE-Mid-SAP 100E Aero Fuchs HCE-Low-SAP2 0W-20 Fuchs 100E-Low-SAP 10W-30 (HDDO) 8 Group 2 Oils for the Steam Rankine power cycle IAV-PAS 8 IAV 65-2 IAV 65-2 + water IAV 65-3 IAV 65-3 + water The results of the viscometric and thermophysical measurements will be presented in diagrams in chapter 4. Most diagrams show data for one of the three groups of oils. The names of the oils are listed in the following table. The commercially available Fuchs Titan GT1 0W-20 (1.2 wt.% ash) with a portion of 50 % ester is listed on the positive list of the German Market Introduction Programme (MIP) for „Biolubricants and Biofuels“, funded by the Ministry of Consumer Protection, Food and Agriculture (BMVEL). Also, fully ester-based, prototype oils of Fuchs Titan 100E SAE 0W-20, 100E HDDO and TOTAL 100E, were used. The formulations GT1, GTE/100E and HCE low SAP of FUCHS conform with the requirement of >50% of renewables. The GT1 and 100E/GTE are polymer-free. The TOTAL HCE midSAP (0.75 wt.-% of sulfated ash) is a blend of hydrocarbons with esters. The FUCHS GT1, HCE low SAP as well as the Total 100E and Total HCE (SAE 0W-30) comply with the bio-no-tox criteria in EC/1999/45. The Fuchs HCE lowSAPs, HCE 0W-20 and 100E 0W-20/10W-30 are zinc-free. The Fuchs HCE lowSAPs form only 0.5 wt.-% sulfated ash. Group 3 Polyglycols and others PAG 46-2 PAG 46-4 PPG 32-2 PPG 32-3 Triol-PO Triol-EO Diol-PO Paraffin 46 PG WS55 PAG 68 Forschungsbericht 277 The 100E aero is a pentaerythritester-based engine oil developed by SHELL in the eighties for adiabatic engines and displays an outstanding oxidation resistance forming 0.96 wt.-% sulfated ash. The Fuchs 100E-Low-SAP 10W-30 (HDDO) is a 100 % esterbased engine oil forming 0.8 wt.-% sulfated ash. The HC 5W-30 having 3.7 wt.-% soot was aged in a ﬁred 1.9 liter turbodiesel engine with 89 kW by Renault SAS. The aim of this oil sample was to investigate the inﬂuence of soot on friction and wear. The parafﬁnic oil (MERCK) in VG 46 was unadditivated. to DE 100 49 175 for reciprocating steam expanders using the Rankine cycle, which is considered as a competitor to the fuel cells, since the combustion process applied by IAV fulﬁls “zero-emission” or “lean-burn” criteria, except for CO2, H2O and N2. Due to steam blow-by, the crank case oil needs to be tolerant vis-à-vis water. The IAV 65 formulations contains a base oil composed of 50 wt.-% PEG 450 and 50 wt.-% triethyleneglycol. The very high viscosity index of 266 associated with a low viscosity at 40 °C underlines this robust concept, which is designed not to suffer under a high water take-up. Group 2 The PAG 68 is a polyethyleneglycol-based formulation from FUCHS with 20 wt.-% water. The boiling temperature was above 120 °C. The IAV-PAS 8 is a water-based polyethylene glycol (PEG 3350 g/mol) crank case oil with 50 wt.-% water according The polyethyleneglycols (CAS: 25322-58-3) comply with bio-no-tox and pharmaceutical requirements. The aim of all Table 1 Properties of engine and prototype oils Lubricants Noack Pour evapopoint ration in °C in % VI ν40 in mm²/s ν100 in mm²/s ν150 HTHS at 150 °C in in mm²/s mPa·s 159 55.15 9.57 4.197 Factory fill oils Total HC 5W-30 12.8 -42 3.0 Total HC 5W-30+ 3,7% soot ./. ./. 162 67.85 11.48 4.93 ./. Castrol SLX 0W-30 8.1 -57 168 57.0 10.2 4.42 3.0 Titan SL PCX 0W-30 9 - 45 162 53.19 9.44 4.14 2.95 Paraffin 46 b.o. ./. ./. 118 51.96 7.39 2.4 ./. TOTAL 100E 4.8 <-42 153 40.98 7.6 3.46 2.95 100 E 0W-20 5.5 -39 167 43.26 8.23 3.64 2.95 90 118.74 12.22 4.34 Ester oils 100E aero 100E LowSAP (HDDO) 7 <-48 151 56.22 9.44 4.03 3.0 TOTAL HCE ./. <-42 159 46.32 8.41 3.73 2.98 Fuchs HCE 0W-20 6 - 45 160 47.03 8.64 3.78 2.95 TOTAL HCE midSAP 6.6 <-48 165 57.8 10.4 4.53 2.99 FUCHS HCE LowSAP 5.2 -45 184 44 8.8 4.26 2.9 PAG 46-2 19.3 -31 203 47.4 9.94 4.81 4.3 PAG 46-3 19.5 -<27 207 46.7 9.95 ./. 4.5 PAG 46-4 (Base oil) 11.3 -33 146 52.2 8.56 3.61 3.6 Polyalkyleneglycols Diol-PO b.o. ./. ./. 147 39.8 7.31 3.39 ./. PPG 32-2 4.8 -45 149 34.3 6.7 3.2 2.78 PO-Triol b.o. ./. -29 41 119.1 8.95 3.04 ./. EO-Triol b.o. ./. ./. 89 104.1 11.1 4.0 ./. 265 39.5 10.4 ./. ./. 172 70.12 12.39 ./. ./. -27 61 26.11 4.76 2.11 ./. -55 ./. 9,21 2,57 1,38 Water-based formulations on polyethyleneglycols IAV PAS-8 ./. -30 PAG 68 IAV 65 PAG WS 55 ./. 9 Forschungsbericht 277 water-based or water diluable formulations was to achieve at 100 °C a kinematic viscosity comparable to engine oils at 150 °C, thus enabling the use of state-of-the-art crank shaft bearings in engines. Group 3 The polymer-free polyalkylene glycols (PAG 46-2/PAG 46-3 (Mw= 1 205/1 280 g/mol) and 46-4) were diols with EO: PO = 1:1 or 7:1 (EO= ethylene oxide, PO= propylene oxide). The polyalkylene glycols (PAG 46-2/46-3 and PAG 46-4) have different molecular weight distributions and EO:PO portions and were ﬁrst blended with a gear/hydraulic oil additive according to patent US 6,194,359 and then modiﬁed by BAM in view of oxidation resistance and tribological properties. The PAG 46-4 is a custom-made prototype PAG elaborated by DOW Europe SA with Mw= 664 g/mol. The PAGs are not obliged to be labeled with the symbol „N“, they are polymerfree, ash-free, and do not contain any zinc or calcium. The PPG 32-2 formulation uses a polypropylene glycol monobutyl ether base oil and a gear/hydraulic oil package which does not contain any polymers, is free of Zn, Ca and sulfur, and does not have to be labeled with the symbol “N” (Bio-no-tox). This PPG 32-2 has a comparatively low kinematic viscosity (only about 34 mm²/s at 40 °C) and at the same time a low NOACK-volatility of less than 5 %. The polymer-free PPG 32-2 contains 1 700 ppm sulfur and 200 ppm phosphorus respecting bio-no-tox criteria. Polypropylene glycol monobutyl ethers are classiﬁed as “slightly hazard“ to water (WGK 1) by the German Environmental Agency (www. umweltbundesamt.de) under the number #3530. Additionally, the oxidation resistance of the PAG 46-4 and PPG 32-2 was boosted by proprietary additive packages “Phepani”, “Phopani”, “Chopani” or “Papani”. The amount of phosphorus and sulfur is reduced to about 650-780 ppm [P] and to about 600-800 ppm [S]. The “triols” are trifunctional polyglycols, either based on 100 % ethylenoxide (EO-Triol) or 100 % propylene oxide (PO-Triol). All viscosity indices of the polyglycols were labeled in italic, as they don’t follow a linear relation between viscosity and temperature in a logarithmic plot. The Diol-PO is a polypropylene glycol without a butanol starter having a molar mass Mw of 490 g/mol and a surprisingly high α at 20 °C of 19.4 GPa-1 (compare with 19.2 GPa-1 of PPG 32-2 having Mw = 900 g/mol). The PAG WS 55 is a water-soluble, linear polymer with a very low viscosity having a molecular weight of ca. 250 g/mol. Some relevant properties of the different lubricants are summarized in Table 1. The eco-toxicological properties of most formulations presented in Table 1 are detailed in [7, 8]. 3 Equipment used for the measurements of viscometric and thermophysical properties and tribological behavior 3.1 Viscosity at ambient pressure For a part of the engine oils, the viscosity at ambient pressure has been measured at seven different temperatures using capillary viscometers of the Ubbelohde type. This type of measurement yields the kinematic viscosity. For other engine oils, the dynamic viscosity at ambient pressure has been measured in the same temperature range, using the rolling-ball viscometer being described in the next section. It is sufﬁcient to measure either the kinematic viscosity ν or the dynamic viscosity η. The conversion (η = ρν ) requires only the density ρ. 3.2 liquid. The latter must be known, as the buoyancy of the ball has an effect on its speed. The density has to be measured in a separate experiment. Density The density has been measured at ambient pressure using pycnometers in the temperature range from 20 °C to 150 °C. 3.3 High-pressure viscosity Figure 1 shows the rolling-ball viscometer that is used at PTB. Driven by the gravitational force, a hardened steel ball rolls downwards in a tube that is slightly (10°) inclined against the vertical. The diameter of the ball (15.721 mm) is only 214 μm smaller than the inner diameter of the tube (15.935 mm). The tube is ﬁlled with the liquid under test which has a lower density than the ball. The speed of the ball depends on the dynamic viscosity and on the density of the 10 Figure 1 Photo and schematic diagram of the rolling-ball viscometer used at PTB: 1: fall tube, 2: steel ball, 3: double-coil, 4: sample, 5: cylinder, 6: piston, 7: pressure vessel, 8: pressure-transmitting oil, 9: pressure connection to screw press, 10: oil, 11 thermostat ﬂow, 12: return of oil to thermostat, 13: insulation, 14: axis Forschungsbericht 277 The tube is located in a pressure vessel ﬁlled with oil. The pressure inside the vessel can be regulated by a screw press. A piston moving in a cylinder separates the liquid under test from the pressure-transmitting oil, guaranteeing the pressure equilibrium inside and outside the tube. The rolling-ball viscometer has been calibrated with a special oil provided by the Fuchs Petrolub AG, one of the project partners. The viscosity and density of this oil in the temperature range of interest have been measured with capillary viscometers and pycnometers, respectively. Thermostatisation is performed by a silicone oil which circulates around the pressure vessel in a channel which forms a spiral. The temperature of this oil is kept constant by a thermostat. A calculation of the uncertainty following the GUM  resulted in a relative uncertainty (k = 2) for the viscosity of 1 % to 1.5 %. The most important contribution to this uncertainty is caused by the uncertainty of the temperature measurement. The ball‘s position is detected inductively by coils which enclose the tube. The ball, which consists of magnetic steel, augments temporarily the inductivity of the coil because it acts as an iron core as long as it is in the coil. A combination of two coils forming a differential transformer allows to determine exactly the point of time at which the ball passes through the beginning or the end of a measurement section. In order to get a clear signal from the differential transformers, the pressure vessel and the tube are made of stainless, non-magnetic steel. On the one hand, the uncertainty of the temperature measurement is signiﬁcantly higher at high temperatures, compared to the ambient temperature. On the other hand, the viscositytemperature-coefﬁcient ß When a measurement is ﬁnished, the pressure vessel is turned round to bring the ball back into its original position. All parts of the apparatus containing liquid under pressure have to be turned round, too. For this purpose, the pressure vessel, the screw press, the manometers and the pressure valves are mounted on a common axis. The following quantities are recorded during a measurement: β (T ) = − 1 η η T (1) p = const. is signiﬁcantly smaller at high temperatures, resulting in a roughly constant contribution to the total measurement uncertainty. The uncertainty of the estimation of the oil compressibility contributes only slightly to the total uncertainty because the difference to the high density of the steel ball is of interest, and this difference is known with an uncertainty of 0.25 %, even if the density of the oil is only known with an uncertainty of 2 %. − pressure p by a digital manometer − temperature T by two platinum resistors which are located in the thermostat spiral channel − the runtime t of the ball A fourth quantity of importance is the density of the ﬂuid which is known at ambient pressure. A simple experiment has been assembled to measure the increase in density caused by the pressure. In this experiment, a screw press is ﬁlled with the tested oil. The number of rotations necessary to achieve the maximal pressure of 100 MPa is dependent on the compressibility at ambient temperature. The higher compressibility at higher temperatures is estimated on the basis of the measured value at ambient temperature. A feature of the measurements is the large range of viscosity. The maximal viscosity (at maximal pressure and ambient temperature) can be 200 times the minimal viscosity (at ambient pressure and maximal temperature). Exchanging the ball several times in order to adapt the size to the expected viscosity would be too time-consuming because the apparatus would have to be opened. Thus, the complete viscosity range is covered with just one ball, which leads to runtimes of up to 65 min. As the measurement values are recorded automatically by a computer, most of this time can be spent on other purposes. To prepare the experiment, a long thermostatisation time is required. This is due to the pressure vessel which is located between the tube and the thermostatisation oil. It has a large mass and a limited thermal conductivity (about 15 W/(m ⋅ K)). The thermostatisation process can be started in the early morning by means of a time switch so that a thermal equilibrium will have developed at the beginning of a working day. 3.4 Heat capacity The heat capacity measurements have been carried out using a power-compensated differential scanning calorimeter. This apparatus contains two crucibles. One of them is ﬁlled with the sample, the other one is empty. Both crucibles are heated with the same heating rate. The additional power that is necessary for the crucible which contains the sample is used to calculate the heat capacity of the sample. More details about the apparatus are given by Watson et al.  and by Höhne et al. . 3.5 Thermal conductivity The conductivity of eight oils has been measured using a plate apparatus. In this experiment, a known ﬂow of thermal energy ˙. Q is driven through a gap between two parallel plates. The gap is ﬁlled with the sample. The temperature difference ΔT that is necessary for this heat ﬂux is measured. From these data and some geometrical information (area of the circular plates Ac, width of the gap d), the thermal conductivity λ can be calculated using λ= Q ⋅ d Ac ⋅ ΔT (2) For more details about the apparatus, see Hammerschmidt . 11 Forschungsbericht 277 3.6 Tribological testing outside of engines The tribological properties of a new triboreactive coating interacting with prototype engine oils based on esters and polyglycols were tribologically characterized outside of engines only under mixed/boundary lubrication using the SRV®  and BAM  test method in order to rank by two distinct different test methods. SRV® tests are complementary to those the BAM-tests and were additionally performed according to a new ASTM Dyyyy-xx draft method  as cross-check to the BAM test method. For the comparison of the results achieved with both tests, it has to be noted that the load in the SRV® test is six times higher than in the BAM test and both differ in the oil temperature (See Figure 64 and Figure 65). Piston ring/cylinder liner simulation tests were performed under mixed lubrication conditions in different lubricants at 170 °C and 0.3 m/s, whereby a thermal sprayed piston ring segment was pressed with 50 N against the rotating Figure 2 BAM test rig and piston ring / cylinder liner test conﬁgurations for mixed/boundary lubrication conditions under unidirectional sliding FN Body 1 Body 2 Figure 3 SRV® test rig with piston ring-on-disk (or cylinder liner) conﬁguration 12 Δx Forschungsbericht 277 Novel and non-commercial “triboactive” or “triboreactive” materials were selected from Magnéli-type phases, like TinO2n-1 and Tin-2Cr2O2n-1 (see FR 2 793 812), as well as substrates, like (Ti,Mo)(C,N)+23NiMo-binder (see DE 195 30 517), which forms by triboxidation these, namely γ-Ti3O5, Ti5O9, Ti9O17 and Mo0.975Ti0.025O2 as well as double oxides like NiTiO3 and β-NiMoO4. cylinder liner segment (or ﬂat disk) up to a sliding distance of 24 000 m. An fresh oil amount of 0.3 to 0.4 l was used for each test. The test rig and the piston ring/cylinder liner conﬁguration are shown in  using liner segments and in Figure 2. The SRV® sample conﬁguration with piston ring segments and wear scars  used here is shown in Figure 3. The resistance of different lubricants against scufﬁng was determined with 100Cr6/100Cr6 (AISI 52100) pairings according to modiﬁed ASTM D5706-05 . Experimental TinO2n-1, Tin-2Cr2O2n-1 and (Ti,Mo)(C,N)-23NiMo as novel and non-commercial, triboactive powders applied by thermal spraying for automotive applications as cylinder liner and/or ring coatings were developed, produced and deposited for the ﬁrst time on piston rings and cylinder liner samples. The grey cast iron SRV-disks (GGL20HCN) corresponding to Renault GL1 have been produced and lapped to C.L.A. (Ra) = 0.343 μm, Rz = 2.483 μm, RpK = 0.451 μm and RvK = 0.596 μm. 3.7 With 550-880 HV0,2, the triboactive coatings follow not a metallurgically “hard” concept (See Table 2). Tribological materials 3.7.1 Wear protection represent another concern while using “midSAP” or even “lowSAP” oils or oils without or low contents of extreme pressure (EP) and/or anti-wear (AW) additives associated with bio-no-tox-properties according to directive EC/1999/45. Spray powder For the deposition of these coatings different Ti-based and substoichiometric powders were developed and purchased from H.C. Starck GmbH and FhG-IKTS (both Germany). Lubricious oxides (LO) and triboactive materials appeared recently in scientiﬁc literature  and display estimated functional properties by different approaches. There exists within the scientiﬁc community no ofﬁcial consensus about their meaning. 20 The term of “lubricious oxides” was created 1989 by Michael N. Gardos [18, 19] for TiO2-x as well as thematized by  and aim low wear with may be associated low dry coefﬁcient of friction. The correct term for TiO2-x is Magnéli-phases of titania, TinO2n-1 with 4≤n≤9, whereas TiO2-x, with x≤0.01, describe “Wadsley”-defects. 10 The term “triboactive materials” appeared in Europe end of the nineties describing more a beneﬁcial reaction between the surface and the lubricant or the ambient, thus indicating a more overall functional approach. Oxides, hydroxides or hydrates cover this understanding. 0 % 100 TinO2n-1 90 (Ti,Mo)(C,N)+NiMo 80 Tin-2Cr2O2n-1 60 70 50 40 30 20 1.0 10.0 Particle diameter [μm] 100.0 10 0 Figure 4 Particle size distribution of three different triboactive spray powders (agglomerated and sintered) Table 2 Porosity and hardness of plasma sprayed Ti-based coatings on piston rings and on cylinder liners Coating Porosity by volume MKP81A® Vickers hardness 566 144 HV0,2 2% 657 88 HV0,2 APS Tin-2Cr2O2n-1 on piston ring (TARABUSI) 5% 530 55 HV0,5 APS (Ti,Mo)(C,N)+23NiMo (TM23-1) on piston ring (TARABUSI) 15 % 699 99 HV0,5 APS (Ti,Mo)(C,N)+23NiMo (TM23-2) on piston ring (TARABUSI) 10 % 650 59 HV0,5 APS TinO2n-1 (n = 4 …6; TiO1,60 to TiO1,80) on cylinder liner (FhG-IWS) 2% 846 54 HV0,2 VPS Amperit 782.1 (TiO1,95) on cylinder liner (FhG-IWS) 2% 785 39 HV0,2 APS Tin-2Cr2O2n-1 on cylinder liner (FhG-IWS) 2% 851 36 HV0,2 HVOF (Ti,Mo)(C,N)+23NiMo on cylinder liner (FhG-IWS) 10 % 816 36 HV0,2 APS TinO2n-1 (n = 4 …6; TiO1,60 to TiO1,80) on piston ring (TARABUSI) 13 Forschungsbericht 277 Before thermal spraying the spray powders were thoroughly characterized. Spray powder composition, particle morphology and resulting phase composition of coatings as well as particle size distribution are compiled in Figure 4. 3.7.2 Cylinder liner materials The deposition of substoichiometric TiO1.93, TinO2n-1 (TiO1.60 to TiO1.80) and of (Ti,Mo)(C,N) cylinder liner coatings with different thermal spray processes (APS: Atmospheric plasma spraying, VPS: Vacuum plasma spraying, HVOF: High-Velocity-OxyFuel) was described in more detail elsewhere [21, 22, 23]. The Tin-2Cr2O2n-1- and (Ti,Mo)(C,N)-23NiMo-powders were sprayed on GG20HCN disks by FhG-IWS using MF-P-1000 plasma spray equipment with F6 spray gun. A TiO1.95-x coating was deposited with a vacuum plasma spray (VPS) process using a commercial, fused and crushed TiO1.95 powder (Amperit 782.1 (22.5-45 μm; H.C. Starck GmbH, Goslar, Germany) directly on GGL20HCN without a bond layer as economic alternative to TinO2n-1. The advantage of vacuum plasma spraying (VPS) is that in this process no re-oxidation of the TiOx powders can occur. Analysis with XRD on the APS sprayed TinO2n-1 coatings exhibit besides Rutile and Anatase peaks of Magnéli phases, mainly of Ti4O7 (Compare with ca. 66 wt.-% Ti5O9, ca. 17 wt.-% Ti6O11 and ca. 17 wt.-% Ti4O7 in the spray powder). The surfaces of Tin-2Cr2O2n-1-coatings were lapped to RPK ca. 0.61 μm and in some cases smoothly ﬁnished to RPK < 0.05 μm as well as those of (Ti,Mo)(C,N)-23NiMo-coatings to RPK < 0.03 μm. In the sprayed Tin-2Cr2O2n-1-coatings were analyzed by means of ESMA 26.2 at.-% Titanium, 9.80 at.-% Chromium and 64.0 at.-% Oxygen. Those newly developed triboactive Ti-based coatings were tribologically compared with uncoated grey cast irons (GGL20HCN with 3.7 wt.-% C) and GG26Cr  used as cylinder liner materials. It has to be noted, that dry running brake disks in passenger cars apply these cast iron grades with a high carbon contents. The triboactive liner coatings were deposited without intermediate layer/bond coating on grey cast iron GGL20HCN supplied from Schwäbische Hüttenwerke (SHW) used as reference material. The structure of GGL20HCN is charac- terized by a perlitic matrix, a small amount of ferrite (<5 %), Fe3P and small MnS inclusions as well as homogeneous distributed lamellar graphite. In comparison to common grey cast iron materials, GGL20HCN contains a relative high carbon content (3.66 wt.-% C) besides 2.0 % Si, 0.236 % Cr, 0.564 % Mn, 0.5 % Ni, 0.39 % Mo, 0.057 % S, 0.045 % P, 0.206 % Cu (all in wt.-%). 3.7.3 Piston ring materials The ring diameters ranged from 79 mm to 89 mm. On piston rings (∅ = 80 mm x 2.5 mm) triboactive TinO2n-1, Tin-2Cr2O2n-1 and (Ti,Mo)(C,N)+23NiMo coatings were deposited by TARABUSI on a 94(NiCr)6Al bond layer using the APS process with a plasma gun SG 100 (Miller Thermal Inc.) in an Ar/He mixture as plasma forming gases. Piston ring substrate material is “AT 126”, a grey cast iron with spheroidal graphite and high carbon content. These newly developed coatings were compared with two Mo coatings (PL72; APS-MKP81A® composed of Mo-NiCrBSi). The molybdenum in the VL® (Mo, ∅ = 122 mm) coating was deposited by means of ﬂame spraying. The commercially available Mo-based piston rings differ mainly in the content of oxygen and in the amount of hard phases. Porosity, hardness and roughness values before tribotesting are given in Table 2 and Table 3. The roughness of TinO2n-1 and MKP81A® coated piston rings are similar, but the hardness of TinO2n-1 is much greater (Table 4). Figure 5 and Figure 6 present the cross sections of triboactive APS sprayed coatings with NiCrAl bond coatings on piston rings. Ring substrate material is a grey cast iron with modular graphite (AT126). APS Mo (PL72) has a dense and lamellar structure and is used as reference coating and sprayed and complies with MKP81A®. To improve the bonding a NiCrAl intermediate layer was used. Concerning structure, phase composition and tribological behaviour APS Mo (PL72) and additionally investigated APS Mo (MKP81A®) (Figure 7) coatings are very similar. The PCF251® as well as PCF278® were supplied by DANA Corp. (Perfect Cycle Division). The PCF 278® consists of 59 wt.-% of Mo and 35 wt.-% CrNi as well as the PCF 251® of 80 wt.-% Mo and 18 wt.-% NiCr as well as the MKP81A® of 67-77 wt.-% Mo and 19-31 wt.-% NiCr. Table 3 Roughness of uncoated grey cast iron and coated cylinder liner (or disk) Specimen Machining Ra in μm Rz in μm Rpk in μm Rvk in μm GGL20HCN (BAM disks) Lapped 1.07 5.9 0.53 1.45 VPS TiO1,95-x (Amperit 782.1) Lapped 0.18 1.37 0.21 0.32 APS TinO2n-1 (n = 4…6) Lapped 0.135 0.863 0.164 0.200 APS- Tin-2Cr2O2n-1 Lapped or polished 0.57 0.152 4.49 1.44 0.61 0.054 1.25 0.571 HVOF (Ti,Mo)(C,N)+23NiMo Lapped or polished 0.38 0.05 2.73 0.39 0.38 0.03 0.70 0.17 14 Forschungsbericht 277 Table 4 Roughness of different piston rings Metallurgy of running surfaces Rz Ra / C.L.A. Rpk Rvk HV0,2 Supplier APS 67-77 wt.-% Mo + Ni (MKP 81A) APS 67-77 wt.-% Mo + Ni (PL72) APS 80 wt.-% Mo +Ni (PCF-251) Flame Mo (VL) APS 59 wt.-% Mo +Cr+Ni (PCF-278) APS TinO2n-1 (n = 4 – 6) APS (Ti,Mo)(C,N) + 23NiMo (TM23-1) APS (Ti,Mo)(C,N) + 23NiMo (TM23-2) APS Tin-2Cr2O2n-1 HVOF WC/Cr3C2 (MkJet 502) CKS-36 (Cr+ 2-6 Vol.-% Al2O3) GDC-50 (Cr+0.5-2.0 Vol.% diamond) Nitrided GGG with 3.7-4.2 wt.-% C (KV1) 1.9 0.5 0.2 2.2 570 Goetze 0.98 0.149 0.105 0.559 530 TARABUSI 0.455 0.090 0.069 0.137 1.34 0.902 0.2 0.190 0.22 0.311 1.35 0.210 >900 Goetze DANA 2.3 15.1 0.5 3.1 0.2 0.8 1.6 6.00 830 700 TARABUSI TARABUSI 1.76 0.36 0.18 1.24 660 TARABUSI 0.81 0.574 0.14 0.111 0.14 0.08 0.52 0.305 550 1 200 TARABUSI Goetze 0.142 0.027 0.014 0.065 660 Goetze 0.204 0.040 0.026 0.108 830 Goetze 0.176 1.207 0.030 0.250 0.018 0.346 0.071 0.878 310 Goetze Goetze Scale: 500 μm APS Mo (PL 72) coating DANA Scale: 20 μm a) Coating thickness: 269 ± 6.5 μm APS TinO2n-1 coating b) Coating thickness: 204 ± 9.6 μm Figure 5 Optical micrographs of cross sections of APS coatings (TARABUSI) on grey cast iron piston rings: a) Mo and b) TinO2n-1 15 Forschungsbericht 277 50 μm Figure 6 Cross sections of HVOF-MKJet502® coated piston ring Figure 7 Optical micrograph of cross section of commercial Mo (MKP81A®) coated piston ring Figure 8 APS-(Ti,Mo(C,N)-23NiMo piston ring coating deposited by TARABUSI (TM23-2) Figure 9 APS Tin-2Cr2O2n-1 piston ring coating deposited by TARABUSI 16 Forschungsbericht 277 Besides a good bonding APS TinO2n-1 is characterized by a dense, typical lamellar morphology. Dark and bright grey areas visible in cross sections of APS TinO2n-1 coating can be attributed to different oxygen contents. The hardness of the triboactive piston ring coating is slightly lower than for the triboactive cylinder liner coatings. The HVOF-sprayed WC/Cr3C2-based hard metal ring coating (MKJet502®) with NiMo-binder has a dense structure. The tribological surfaces were superpolished. Due to the at least two hard phases, the hardness values of MKJet502® vary between 879 and 1 330 HV0.3. GDC-50®, a plated chromium with up to 2.0 % by volume of ﬁne diamond particles and CKS-36®, also plated chromium with up to 6 % by volume of ﬁne alumina particles, were tribological characterised, too. KV1 is a cast iron with globular graphite (3.5-4.0 wt.-% C). The metallurgical characterizations of the piston rings supplied by Federal Mogul Goetze can be found in reference . Despite the functional references of molybdenum-based piston rings, some concerns grow more and more. Common sense precautions are necessary in repeated exposures of human beings especially in dusts and fumes of molybdenum and trioxide products as they occur during thermal spraying. Another aspect are the stock exchange prices for Ferromolybdenum and Molyoxide as they increased from ca. 3 US-$/lb in ﬁrst quarter 1999 to ca. 7 US-$/lb in third quarter 2003. By 30th September 2005, the average price for Molyoxide reached 34 US-$/lb and by 17. March 2006 25 US-$/lb or 62 US-$/kg for Ferromolybdenum by 26th May 2006. There exist in consequence drivers for substituting molybdenum. Coated Tin-2Cr2O2n-1-rings were machined to RPK ca.0.14 μm remaining a functional thickness of ca. 187 μm. 3.7.4 Unlubricated sliding wear The tribological behavior under unlubricated sliding conditions of APS-Tin-2Cr2O2n-1 machined to RpK ca. 0.61 μm was characterized up to 800 °C and 7.5 m/s using stationary specimen in 99.7 % alumina. The results were elaborated in a high-temperature tribometer described elsewhere in reference . In the work presented in references [27, 28], it is clearly visible, that the APS-Tin-2Cr2O2n-1, even with a quite elevated roughness of RpK ca. 0.61 μm, runs best with low wear rates above 400 °C and 1 m/s associated with wear rates of the alumina below 10-7 mm³/Nm. At 800 °C and 7.5 m/s under unlubricated sliding conditions, the lowest wear rate was kv = 1.62 10-6 mm³/Nm and of the alumina toroid kv = 9.9 10-8 mm³/Nm associated with P ⋅ Vvalues of 60.8 MPam/s and a coefﬁcient of friction of 0.27. Under dry friction, the self-mated couples of monolithic (sinterHIPpped) (Ti,Mo)(C,N)-15NiMo [TM10] displayed in a temperature range up to 800 °C and sliding speeds up to 5 m/s total wear rates lower than 10-6 mm³/Nm, which lie on a level comparable to those known for mixed/boundary lubrication. Compared to ceramic-ceramic composites, the tribosystem composed of self-mated (Ti,Mo)(C,N)+15NiMo couples achieved a further wear reduction at 22 °C and up to 800 °C. At 800 °C, the wear rate of the stationary (pin, shell) specimen decreased from 3.82 ⋅ 10-7 mm3/Nm at 0.12 m/s down to 9.5 ⋅ 10-8 mm3/Nm at 3.68 m/s, whereas the wear rate of the rotating (disc, shaft) specimen was only detectable at 3.68 m/s with a wear rate of 3.5 ⋅ 10-7 mm3/Nm. The lowest wear rate for TM10 was achieved at 6.17 m/s and 800 °C with 2.9 ⋅ 10-7 mm³/Nm for the stationary specimen and 5.7 ⋅ 10-7 mm³/Nm for the rotating specimen or as total wear rate of 8.6 10-7 mm³/Nm with an associated PV-value of 42 MPam/s. Magnéli-phases of titania, TinO2n-1, are also part of the family of triboactive/-reactive materials with a prone dry running ability [17, 29]. Dry running or “oil-off” is a frequent operation mode for the piston ring/cylinder liner system requiring sliding couples free of adhesive wear. 4 Results of the measurements of viscometric and thermophysical properties For a set of 28 different lubricants, the following properties have been measured in the temperature range from 20 °C to 150 °C: − viscosity at ambient pressure − density at ambient pressure − dynamic viscosity at pressures up to 100 MPa = 1 kbar − thermal conductivity (10 oils) in the same temperature range − dynamic viscosity at 150 °C and at a shear rate of up to 3 .106 s-1 (6 oils) The upper temperature limit of 150 °C is considered to be the maximum temperature in the oil sump. For reasons discussed in chapter 4.6, functions α(T) and η(T) have been derived. 4.1 In addition, three other properties have been measured for some of the oils: The results of the density measurements are presented in Figure 10, Figure 11 and Figure 12. These data are needed for several calculations, namely: − speciﬁc heat capacity (20 oils) in the temperature range from 20 °C to 150 °C Density − volumetric heat capacity cp⋅ρ 17 Forschungsbericht 277 Group 1: Density HC 5W-30 0.95 HC 5W-30 + 3.7% Soot Common curve for 100E 0W-20 and Fuchs HCE-Low-SAP2 0W-20 Curve for 3 oils Density in g/cm³ 0.90 Total 100E Total HCE 0.85 Titan SL PCX 0W30 HC 0W-30 100E Aero 0.80 Fuchs 100E-LowSAP 10W-30 Common curve for Fuchs HCE 0W-20, Total HCE-Mid-SAP and Fuchs HCE-Low-SAP Curve for 2 oils 0.75 20 40 60 80 100 120 140 Temperature in °C Figure 10 Density of the oils in group 1 Group 2: Density 1.14 1.12 Density in g/cm³ 1.10 IAV-PAS 8 IAV 65-2 1.08 IAV 65-2 + water IAV 65-3 IAV 65-3 + water 1.06 1.04 1.02 1.00 20 40 60 80 100 120 140 Temperature in °C Figure 11 Density of the oils in group 2 Group 3: Density 1.20 1.15 1.10 PAG 46-2 Density in g/cm³ 1.05 PAG 46-4 1.00 Triol-PO Triol-EO 0.95 Diol-PO 0.90 Paraffin 46 Common curve for PPG 32-2, PPG 32-3 and PG WS55 0.85 PAG 68 0.80 Curve for 3 oils 0.75 20 40 60 80 100 Temperature in °C Figure 12 Density of the oils in group 3 18 120 140 Forschungsbericht 277 − conversion between dynamic and kinematic viscosity, and − buoyancy of the ball in the rolling-ball viscometer Comparison of the diagrams shows that the oils in group 1 have a lower density than those in group 2 and 3, with the exception of Parafﬁn 46 from group 3. 4.2 Heat capacity In most diesel engines, the piston crown is cooled by an oil jet to the piston bowl. Assuming a constant volume ﬂow of the oil pump, the volumetric heat capacity (the product of the speciﬁc heat capacity cp and the density ρ) has to be determined therefore to obtain the ability of oils to carry out heat. Engine designers are concerned about possible losses in the cooling efﬁciency of the piston by alternative ﬂuids. Another aspect related to the heat capacity is the heating-up of the lubricating ﬁlm during shearing, which results in an individual loss of viscosity and ﬁlm-forming ability. Also, the ﬁlm itself is deﬁned by the volume of the bearing gap. Figure 13 and Figure 14 show the speciﬁc heat capacity of 9 oils from group 1, 3 oils from group 2, and 7 oils from group 3 as a function of the temperature. The oils of group 1 have a very similar heat capacity, with the exception of Total 100E. The lubricants IAV-PAS 8, IAV 65-2 + water and PAG 68 contain 20 wt.-% water, which results in a high heat capacity. The product of speciﬁc heat capacity cp and density ρ is plotted in Figure 15 and Figure 16. The ranking of the oils is different, compared to the situation of in Figure 13 and Figure 14. The diagrams show that the volumetric heat capacities of the hydrocarbon-based oils HC 5W-30 and Titan SL PCX 5W-30 are the lowest that occurred in the experiments. Group 1: Specific heat capacity 2.6 The heat capacities of Fuchs HCE 0W-20, 100E 0W-20, Titan SL PCX 0W-30, Total HCE-Mid-SAP, and Fuchs 100E-Low-Sap 10W-30 are in the interval formed by the highest and the lowest heat capacity of these four oils. 2.5 cp in J/(g K) 2.4 2.3 HC 5W-30 Total 100E Total HCE 2.2 2.1 Fuchs HCE-LowSAP 2.0 Fuchs HCE-LowSAP2 0W-20 1.9 20 40 60 80 100 120 140 Temperature in °C Figure 13 Results of the heat capacity measurements having been carried out with a power-compensated differential scanning calorimeter (10 oils of group 1). Group 2 and 3: Specific heat capacity 3.6 IAV-PAS 8 3.4 IAV 65/2 3.2 IAV 65/2 + water PAG 46-2 cp in J/(g K) 3.0 PAG 46-4 2.8 PPG 32-2 2.6 Triol-PO 2.4 Triol-EO 2.2 Diol-PO 2.0 PAG 68 1.8 20 40 60 80 100 120 140 Temperature in °C Figure 14 Results of heat capacity measurements for 10 oils of group 2 and group 3 19 Forschungsbericht 277 Group 1: Volumetric heat capacity 2.00 HC 5W-30 Fuchs HCE 0W20 1.95 Total 100E (cp*ρ) in J / (cm³ K) 1.90 100E 0W-20 1.85 Total HCE Titan SL PCX 0W-30 1.80 HCE-Mid-SAP 1.75 Fuchs HCE-LowSAP 1.70 Fuchs HCE-LowSAP2 0W-20 Fuchs 100E-LowSAP 10W-30 1.65 20 40 60 80 100 120 140 Temperature in °C Figure 15 Volumetric heat capacity for 10 oils of group 1 Group 2 and 3: Volumetric heat capacity 3.8 3.6 3.4 IAV-PAS 8 (cp*ρ) in J/(cm³ K) 3.2 IAV 65/2 3.0 IAV 65/2 + water PAG 46-2 2.8 PAG 46-4 2.6 PPG 32-2 Triol-PO 2.4 Triol-EO 2.2 Diol-PO 2.0 PAG 68 1.8 20 40 60 80 100 120 140 Temperature in °C Figure 16 Volumetric heat capacity for 10 oils of the groups 2 and 3 Thermal conductivity of 8 oils 0.16 Fuchs HCE 0W-20 and 100E 0W-20 λ in W / (m K) 0.155 Titan SL PCS 0W30 0.15 HC 5W-30 0.145 Total 100E Total HCE 0.14 PAG 46-2 0.135 Triol-PO 0.13 30 50 70 90 Temperature in °C Figure 17 Thermal conductivity of 8 lubricants 20 110 130 150 Forschungsbericht 277 Thermal conductivity of IAV-PAS 8 and PAG 68 0.36 0.34 λ in W / (m K) 0.32 0.3 IAV-PAS 8 0.28 PAG 68 0.26 0.24 0.22 0.2 20 30 40 50 60 70 80 90 100 Temperature in °C Figure 18 Thermal conductivity of IAV-PAS 8 (group 2) and PAG 68 (group 3) The oils Total 100E, PAG 46-2 and PPG 32-2 compensate their low speciﬁc heat capacity with a comparatively high density, resulting in volumetric heat capacities that surpass that of the factory-ﬁll oils signiﬁcantly. The GT1, GTE and HCE oils, too, present a volumetric heat capacity that is slightly higher than that of the hydrocarbon-based factory-ﬁll oils. Consequently, at the thermal management of the pistons, there should be no problems related to the tested esters and polyglycols. As oil ﬁlms are sheared adiabatically, the loss in viscosity by heating is minimized by an increased oil heat capacity. Consequently, the thermal management of the pistons should be improved by the tested esters and hydrocarbons. 4.3 Thermal conductivity The thermal conductivity affects the heat transfer at the liquidsolid interface of the oil ﬁlm. The results of the measurements are given in Figure 17 for eight oils and in Figure 18 for two additional lubricants. These samples could be examined only at temperatures up to 80 °C or 100 °C because they approached their boiling point. The thermal conductivity of these two lubricants is by far higher than that of the other samples. As shown in Figure 17, the thermal conductivities of these seven oils did not show any signiﬁcant differences. It is therefore considered not to determine the ﬁnal oil selection between hydrocarbons, esters and polyglycols. The relatively high conductivity of PAG 46-2 will enhance the transfer of heat generated during ﬁlm shearing to the rubbing surfaces. Thus, the loss in viscosity due to the higher temperature will be minimal in comparison. are needed. In most cases, 2.9 to 3.0 mPas are required, in some others even more than 3.5 mPas. On the other hand, low viscosities at low temperatures are advantageous for minimizing the fuel consumption in city driving and for short drives. To respect these requirements, a low change in viscosity with temperature is required, which is equivalent to a high viscosity index (VI). A polymer-free formulation with a viscosity index of VI > 170 would represent a challenge to hydrocarbons in ISO VG 32-46. 4.4.1 Viscosity of the oils in group 1 Figure 19 presents the kinematic viscosities of the hydrocarbon- or ester-based oils of group 1 as a function of the temperature. The measurements have been carried out at ambient pressure using capillary viscometers of the Ubbelohde type or the rolling-ball viscometer. Hence, the diagram shows the viscosity at very low shear rates. The oils show a very similar kinematic viscosity in the temperature range of interest, with two exceptions: − HC5W-30+3.7 % soot shows a signiﬁcantly higher viscosity in the whole temperature range. Only the rolling-ball viscometer could be used for this measurement because the soot would not pass through a capillary. − 100E Aero has a viscosity which is signiﬁcantly higher at 40 °C, but similar to the other oils at 150 °C. The viscosity is more temperature-dependent than that of the other oils. The viscosity itself represents a key property for safe and durable operation, especially for the crankshaft bearings. Compared to fresh HC5W-30, the 3.7 wt.-% soot increased the viscosity and also the pressure-viscosity coefﬁcient. From the hydrodynamic point of view, soot presented no disadvantages, but under mixed/boundary conditions the wear increased signiﬁcantly (see chapter 6.12 and Figure 46), which affects the top dead region and/or the cams. On the one hand, the OEMs require comparatively high values of the high-temperature high-shear (HTHS) viscosity at 150 °C. Depending on the operation cycle, at least 2.6 mPas The average viscosity νavg of the remaining 11 oils of group 1 has been calculated for all of the seven measurement temperatures. In Figure 20, the relative deviation to this 4.4 Viscosity at ambient pressure 21 Forschungsbericht 277 Group 1: Kinematic viscosity 1000 100E Aero 100 ν in mm²/s HC 5W-30 + 3.7% soot Max of 11 others 10 Min of 11 others 1 20 40 60 80 100 120 140 Temperature in °C Figure 19 Kinematic viscosity of the oils in group 1 in a logarithmic scale Group 1: Deviation to the average kinematic viscosity 1.25 HC 5W-30 1.2 Fuchs HCE 0W-20 1.15 Total 100E 100E 0W-20 ν / νavg 1.1 Total HCE 1.05 Titan SL PCX 0W-30 1 Total HCE-Mid-SAP 0.95 Fuchs HCE-Low-SAP 0.9 HC 0W-30 0.85 Fuchs HCE-LowSAP2 0W-20 0.8 20 40 60 80 100 120 140 Fuchs 100E-LowSAP 10W-30 Temperature in °C Figure 20 Deviation of the kinematic viscosity to the average value. For 100E Aero and HC 5W-30 + 3.7% soot, refer to Figure 23. Group 1: Deviation to the average dynamic viscosity 1.30 HC 5W-30 Fuchs HCE 0W-20 1.20 Total 100E 100E 0W-20 Total HCE 1.00 Titan SL PCX 0W30 Total HCE-MidSAP Fuchs HCE-LowSAP HC 0W-30 η / ηavg 1.10 0.90 Fuchs HCE-LowSAP2 0W-20 Fuchs 100E-LowSAP 10W-30 0.80 20 40 60 80 100 120 140 Temperature in °C Figure 21 Deviation of the dynamic viscosity to the average value for 11 oils of group 1. For 100E Aero and HC 5W-30 + 3.7 % soot, refer to Figure 24. 22 Forschungsbericht 277 average viscosity is plotted as a function of the temperature. Oils with δ(ν/νavg)/δT > 0 show a comparatively low change of viscosity with temperature, which is equivalent to a high viscosity index and advantageous for the use as a lubricant in an engine. A big positive deviation of ν/νavg is found for Fuchs HCE-Low-SAP2 0W-20 and Fuchs HCE-Low-SAP. The relation of the dynamic viscosity η to the average dynamic viscosity of the 11 prementioned ﬂuids ηavg is plotted in Figure 21. The lubricants with a comparatively high density show higher values η/ηavg, compared to ν/νavg. These ﬂuids are mainly the ester-based lubricants Fuchs 100E-Low-SAP 10W-30 and Total 100E. 4.4.2 Viscosity of the oils in group 2 Figure 22 shows measurement data for the ﬁve lubricants which have to be tolerant to water. Three of the samples contain water and could not be examined at 130 °C and 150 °C. IAV 65-2 and IAV 65-3 both show a loss of viscosity if they are mixed with water. The relative loss of viscosity is roughly the same in the temperature range from 20 °C to 100 °C. Figure 23 shows the relation ν/νavg for the ﬁve oils of group 2 as well as two oils of group 1. For νavg, the same values as in Figure 20 have been used. The high viscosity index of the water-based lubricant IAV-PAS 8 (VI= 265!) is shown by the positive derivation of the corresponding curve, as well as the low viscosity index of 100E Aero by the negative derivation. The lubricants IAV 65-2, IAV 65-2 + water, IAV 65-3, IAV 65-3 + water and HC 5W-30 + 3.7% soot have a viscosity index similar to the average value of group 1. The corresponding diagram for the dynamic viscosity (Figure 24) shows higher values η/ηavg for all lubricants with the exception of HC 5W-30 + 3.7% soot, which is due to the relatively high density of these ﬂuids. 4.4.3 Viscosity of the oils in group 3 In group 3 (see Figure 25), the viscosity at 22 °C is between 17.1 mm²/s (PG WS 55) and 443 mm²/s (Triol-PO). Thus, the range is by far bigger than in the groups 1 and 2. Group 2: Kinematic viscosity ν in mm²/s 100 IAV-PAS 8 IAV 65-2 10 IAV 65-2 + water IAV 65-3 IAV 65-3 + water 1 20 40 60 80 100 120 140 Temperature in °C Figure 22 Kinematic viscosity of the lubricants for the Rankine power cycle Group 1 and 2: Deviation to the average kinematic viscosity 2.5 HC 5W-30 + 3.7% soot 2 100E Aero IAV-PAS 8 ν / νavg 1.5 IAV 65-2 1 IAV 65-2 + water 0.5 IAV 65-3 IAV 65-3 + water 0 20 40 60 80 100 120 140 Temperature in °C Figure 23 Deviation of the kinematic viscosity to the average value for the oils of group 2 as well as two oils of group 1. 23 Forschungsbericht 277 Group 1 and 2: Deviation to the average dynamic viscosity 2.8 2.4 HC 5W-30 + 3.7% soot 100E Aero 2.0 η / ηavg IAV-PAS 8 1.6 IAV 65-2 IAV 65-2 + water 1.2 IAV 65-3 0.8 IAV 65-3 + water 0.4 20 40 60 80 100 120 140 Temperature in °C Figure 24 Deviation to the average dynamic viscosity for the oils of group 2, for 100E Aero, and HC 5W-30 + 3.7% soot (Group 1) Group 3: Kinematic viscosity 1000 PAG 46-2 PAG 46-4 PPG 32-2 / 32-3 100 ν in mm²/s Triol-PO Triol-EO Diol-PO 10 Paraffin 46 PG WS 55 PAG 68 1 20 40 60 80 100 120 140 Temperature in °C Figure 25 Kinematic viscosity of the lubricants in group 3 Group 3: Deviation to the average kinematic viscosity 1.8 1.6 1.4 PAG 46-2 ν / νavg 1.2 PAG 46-4 1 PPG 32-2 / 32-3 Triol-PO 0.8 Triol-EO 0.6 Diol-PO 0.4 Paraffin 46 0.2 PG WS 55 PAG 68 0 20 40 60 80 100 Temperature in °C Figure 26 Deviation of the kinematic viscosity to the average value νavg(T) 24 120 140 Forschungsbericht 277 Group 3: Deviation to the average dynamic viscosity 2.0 1.8 1.6 PAG 46-2 1.4 PAG 46-4 PPG 32-2 η / ηavg 1.2 PPG 32-3 1.0 Triol-PO Triol-EO 0.8 Diol-PO 0.6 Paraffin 46 PG WS 55 0.4 PAG 68 0.2 0.0 20 40 60 80 100 120 140 Temperature in °C Figure 27 Deviation to the average dynamic viscosity to the average value ηavg(T) for the oils of group 3. Figure 26 shows that PAG 46-2, PAG 46-4, PPG 32-2, PPG 32-3 and Parafﬁn 46 have a kinematic viscosity ν(T) which is similar to the reference average of group1, while the other ﬁve oils show a deviation of more than 40 % to this average in at least a part of the temperature range. (IAV-PAS 8, IAV 65-2 + water, IAV 65-3 + water) could not be examined at 150 °C, because their boiling point is below 150 °C. As most oils in group 3 have a higher density than those of group 1, Figure 27 shows for most oils higher values for η/ηavg, compared to the previously mentioned values for ν/νavg. An exception is Parafﬁn 46, which has a density similar to the oils determining the average. For all lubricants, the relative increase in viscosity with pressure is the highest at the minimum measurement temperature and the lowest at maximum temperature. 4.4.4 The function η(T) For the calculation of η(T), which is needed to calculate the ﬁlm thickness (see 4.6), the Vogel equation has been used: η (T ) = 1 mPas ⋅ exp c T −a T −b (3) 4.5.2 At 22 °C, the viscosity of many lubricants increases by a factor of typically 6 if the pressure increases by 100 MPa. At 150 °C, this factor is only about 3. These qualitative results are in accordance with measurements of other authors for other oils, for example those of Schmidt  and Blume . Note that some of the measurement results presented in this publication are fairly different from this behavior, namely those for lubricants containing water. 4.5.3 An equivalent form of this equation is given in Bauer . The coefﬁcients a, b, and c have been calculated based on the measurement results at seven different temperatures in the range from 22 °C to 150 °C. 4.5 High-Pressure-viscosity The viscosity at high pressures up to 100 MPa and temperatures up to 150 °C has been measured using the rolling-ball viscometer that has already been described. The calculation of the pressure-viscosity-coefﬁcient α is based on the measurement results. 4.5.1 Measurement program The measurements are performed for most oils at three temperatures. These are 22 °C (or 40 °C), 80 °C (or 100 °C) and 150 °C. The viscosity has been measured for every temperature at ambient pressure, at 10 MPa, at 100 MPa and at least at two additional pressures between 10 MPa and 100 MPa (e.g. at 40 MPa and 70 MPa). Three lubricants Qualitative results Data analysis and presentation The pressure-viscosity-coefﬁcient α for a deﬁned temperature T is a coefﬁcient in the following equation: η ( p ) = η 0 ⋅ exp(αp ) (4) where η0(T) is the dynamic viscosity at ambient pressure and at the temperature T and η is the dynamic viscosity at the same temperature and at the pressure p. This equation can only be used as a rough estimation if a constant value for α is used. Therefore, the α-value for a deﬁned temperature is given in dependence on the pressure. For high pressures, the calculation results in lower α-values. This means that the viscosity increases less than exponentially if the pressure is increased. The pressure-viscosity-coefﬁcient α is calculated using the formula α ( p, T ) = 1 η ( p, T ) ln η 0 (T ) p (5) 25 Forschungsbericht 277 Group 1: Pressure-viscosity-coefficient at 22°C 23 HC 5W-30 22 α in 1/GPa HC 5W-30 + 3.7% soot 21 Fuchs HCE 0W-20 20 Total 100E and 100E 0W-20 Total HCE 19 Titan SL PCX 0W-30 Total HCE-Mid-SAP 18 Fuchs HCE-Low-SAP 17 HC 0W-30 16 0 20 40 60 80 100 Pressure in MPa Figure 28 Pressure-viscosity-coefﬁcient α(p) for ten oils of group 1 at 22 °C Group 1: Pressure-viscosity-coefficient at 80°C 18 HC 5W-30 HC 5W-30 +3.7% soot 17 Fuchs HCE 0W-20 α in 1/GPa 16 Total 100E 100E 0W-20 15 Total HCE Titan SL PCX 0W30 14 Total HCE-Mid-SAP Fuchs HCE-LowSAP 13 HC 0W-30 12 0 20 40 60 80 100 Pressure in MPa Figure 29 Pressure-viscosity-coefﬁcient α(p) for ten oils of group 1 at 80 °C Group 1: Pressure-viscosity coefficient at 150°C 16 HC 5W-30 15 HC 5W-30 + 3.7% soot Fuchs HCE 0W-20 14 α in 1 / GPa Total 100E 100E 0W-20 13 Total HCE Titan SL PCX 0W-30 12 Total HCE-Mid-SAP Fuchs HCE-Low-SAP 11 HC 0W-30 10 0 20 40 60 Pressure in MPa 80 Figure 30 Pressure-viscosity-coefﬁcient α(p) for ten oils of group 1 at 150 °C 26 100 Forschungsbericht 277 Shape of the function α(p) Function α(T) for two lubricants First, for three measurement temperatures (22 °C, 80 °C and 150 °C), the values for α(p) for ten oils of group 1 are presented (see Figure 28, Figure 29, and Figure 30). This is done to show the qualitative shape of the curves α(p), which is typical for all oils which have been examined in this measurement program. As discussed later, the function α(T) for p = pmax = 100 MPa is by far more important for tribological calculations. Therefore, the function α(p) is not presented for the remaining oils of group 1, which have been measured at other temperatures, and for the oils of group 2 and group 3. 10 17 17.0 16 9.0 15 α in GPa -1 100E 0W-20 (right axis) 13.1 The diagrams show that the hydrocarbon-based oils HC 5W-30, HC 5W-30 + 3.7 % soot, HC 0W-30 and Titan SL PCX 0W-30 have high pressure-viscosity coefﬁcients. But among the 18 oils which are not represented in the diagram, ﬁve have values of the same range. 14 8 13 12 10.8 7 IAV 65/2 (left axis) 10 6.5 6 9 8 5.6 Additionally, it can be seen that the pressure-viscosity-coefﬁcient of HC 5W-30 is slightly increased by soot in the oil, especially at high temperatures. 11 -1 9 α in GPa 4.5.4 5 20 40 60 80 100 120 140 160 7 180 Temperature in °C 4.5.5 The function α(T) Figure 31 Measured α-values and possible functions α(T) for two of the tested lubricants. Between 80 °C and 150 °C, the average of the parabolic and the linear ﬁt was used as a function α(T). At temperatures below 80 °C, the parabolic ﬁt is used. The calculation of ﬁlm thicknesses requires a function α(T) for p = pmax = constant. The pressure-viscosity coefﬁcient α has been measured at only three temperatures (22 °C, 80 °C, and 150 °C or 40 °C, 100 °C and 150 °C). For all lubricants, α decreases, if the temperature increases. Based on these measurement results, it is possible to calculate for every lubricant three coefﬁcients A, B, and C to achieve α (T ) = AT 2 + BT + C If the minimum is only slightly above 150 °C, the value of δα/δT has the correct sign, but it is unrealisticly small. These problems have been avoided by using the average of the parabolic ﬁt and a linear ﬁt in the temperature range of 80 °C<T<150 °C (see Figure 31). Between 22 °C and 80 °C, the parabolic ﬁt didn‘t cause any problems. (6) The pressure-viscosity coefﬁcient of IAV-PAS 8, IAV 65-2 + water, IAV 65-3 + water and PAG 68 has been measured only at temperatures up to 100 °C because the boiling point of these water-containing lubricants is below 130 °C. If measurements have been carried out only at two temperatures, a linear ﬁt was used for calculating α(T). where T is the temperature in °C. An analysis of the resulting parabolas α(T) shows that for all ﬂuids, the parameter A is positive. Thus, the parabola has a minimum which is found at temperatures between 140 °C (IAV 65-2) and 176 °C (100E 0W-20). A minimum of the parabola at T<150 °C results in δα/δT>0 at T ≈150 °C, which is physically wrong. Group 1: Pressure coefficient α as a function of T at p=100 MPa 22 HC 5W-30 HC 5W-30 + 3.7% soot 20 Fuchs HCE 0W-20 and Total HCE α in 1/GPa 18 Total 100E and 100E 0W-20 Titan SL PCX 0W30 and HC 0W-30 16 Total HCE-Mid-SAP Fuchs HCE-LowSAP 14 100E Aero 12 Fuchs HCE-LowSAP2 0W-20 Fuchs 100E-LowSAP 10W-30 10 20 40 60 80 100 120 140 Temperature in °C Figure 32 α(T) for p=100 MPa and group 1 27 Forschungsbericht 277 Group 2: Pressure coefficient α as a function of T at p=100 MPa 10 9 8 IAV-PAS 8 α in 1/GPa 7 IAV 65-2 6 IAV 65-2 + water IAV 65-3 5 IAV 65-3 + water 4 3 0 20 40 60 80 100 120 140 160 Temperature in °C Figure 33 α(T) for p=100 MPa and group 2 Group 3: Pressure coefficient α as a function of T at p=100 MPa α in 1/GPa 22 20 PAG 46-2 18 PAG 46-4 16 PPG 32-2 and PPG 32-3 14 Triol-PO and Diol-PO 12 Triol-EO 10 Paraffin 46 8 PG WS55 6 PAG 68 4 20 40 60 80 100 120 140 Temperature in °C Figure 34 α(T) for p=100 MPa and group 3 4.5.6 Results for α(T) The functions α(T) are presented in Figure 32, Figure 33, and Figure 34. The oils of group 1, based on hydrocarbons and/or esters, have all very high and very similar pressure-viscosity coefﬁcients. Therefore, additional points had to be drawn into Figure 32 so that the curves can be distinguished. The water-tolerant lubricants of group 2 have by far lower α-values, especially if they are really mixed with water. IAV-PAS 8 consists to nearly one half of water and has the lowest pressure-viscosity-coefﬁcient that occurred in the test program. In group 3, a wide variety of α-values can be seen. Parafﬁn 46 (which is a hydrocarbon (aliphatic), too) has, after 100E Aero, the second highest pressure-viscosity coefﬁcient of all oils tested. The polypropylene glycols show a behavior similar to most oils of group 1, while the polyalkylene glycols have quite low pressure-viscosity coefﬁcients. This is especially true for PAG 68, a water-containing oil for hydraulics, which has α-values only slightly higher than those of IAV-PAS 8. 28 4.6 Film-forming behavior The engine oil speciﬁcations of standardization bodies and OEMs refer to the kinematic viscosities in mm²/s and the high-temperature-high shear viscosity (HTHS in mPas). The latter can be determined using the ASTM D4741 (rotating tapered-plug viscometer) and ASTM D4683 (tapered bearing simulator-viscometer). The kinematic viscosity and the HTHS are considered by OEMs to be key properties for safe and durable operation, especially for the crankshaft bearings, as two main tasks for engine lubricants are energy saving (friction) and wear prevention. The hydrodynamic design of engine components is actually based only onto the kinematic viscosity and HTHS. This might be no problem, if only hydrocarbon-based formulations are considered, because there exists a practical background for correlation. In order to differentiate hydrodynamic ﬁlm forming behavior of alternative oils (hydrocarbons, esters and polyglycols), the dynamic viscosity taking into account the differences in density and the pressure-viscosity-coefﬁcients “α” have to be used, because the equations for ﬁlm forming behavior of ﬂuids contain these two parameters. Forschungsbericht 277 The pressure-viscosity coefﬁcient is up to now not mentioned in engine oil speciﬁcations, but it has a strong inﬂuence on the ﬁlm thickness and in consequence also on the frictional losses associated to the ﬁlm shearing. It was recently shown that the fuel efﬁciency [33, 34, 35] of an engine is correlated with the pressure-viscosity coefﬁcient. Further correlations exist with the high-temperature-high-shear (HTHS) viscosity (a second viscometric property of the lubricant) and with the coefﬁcient of friction under mixed/boundary lubrication. 4.6.1 ( A detailed discussion of EHD contacts yields different equations for the ﬁlm thickness, depending on the geometrical situation: The contact region might be − a line, if a cylinder contacts a planar surface or if two parallel cylinders contact each other − a rectangle if the height of at least one of the cylinders is small − an ellipse if an elliptic body contacts a plane, a sphere or another elliptic body − a point, if a sphere contacts a planar surface or another sphere. The point contact is a special case of the elliptic contact. The individual increase of contact area as function of load or oil ﬁlm pressure conducts to different exponents in the equations (13) to (15). ) (8) where E is the elasticity modulus and μ is the Poisson ratio. 1.) The materials parameter G: G = αE ' Jacobson  and Jones  outlined a numerical method that allows the calculation of the ﬂuid dynamics in a Hertzian contact, taking into account the shape and elasticity of the solids as well as their speed and the load on them. Further, three properties of the lubricant are used in the calculation: the viscosity, the compressibility and the increase of its viscosity due to the pressure. To shorten the notation, two abbreviations are introduced: The equivalent radius of curvature r for a line contact of two parallel cylinders A and B or a point contact of two spheres A and B which have the radii of curvature rA and rB is deﬁned by the following equation: (7) Replacing cylinder or sphere B by a plane (rB→∝) yields r = rA. (9) where α is the pressure-viscosity coefﬁcient of the lubricant. 2.) The speed parameter Ue: Ue = η 0u (10) rE ' where η0 is the dynamic viscosity at ambient pressure and operating temperature, while u is the entraining velocity: u= 1 (u A + u B ) 2 (11) 3.) The load parameter We‘ (for a line contact) or We (for a point or elliptical contact) We '= Parameters 1 1 1 = + r rA rB ) ( 1 − B2 1 1 1 − A2 = + E' 2 EA EB The equations for the ﬁlm thickness depend on three nondimensional groups of parameters: Equations describing minimum ﬁlm thickness Tribological problems are considered to be elastohydrodynamic if the deformation of the solid bodies in the contact region is not negligible compared to the ﬁlm thickness. Two of the three critical tribosystems have full EHD lubrication: the cam/follower system and the crank shaft. The piston ring/cylinder system has a mixed lubrication which is not the object of this paper, but at surface asperities, the contact can be elastohydrodynamic, too. 4.6.2 In a similar way, the equivalent elasticity modulus E‘ is calculated using F' E' r and We = F E' r 2 (12) where F‘ is the force per unit length, while F is the total force on the contact. Equations A repeated simulation of the minimum ﬁlm thickness with varied dimensionless parameters G, Ue and We‘ results in data for the minimum ﬁlm thickness which can be correlated by the following equations: Line contact [37, 38]: U 0.70 G 0.54 hmin = 2,65 e 0.13 r We ' (13) Rectangular contact : U 0.71G 0.57 hmin = 3,07 e 0.11 r We ' (14) 29 Forschungsbericht 277 Elliptical contact : 4.7 0.68 0.49 e 0.073 e U G hmin = P* r W (15) P* is the following parameter: Relative ﬁlm thicknesses The comparison is based on the fact that HC 5W-30 which is widely used as factory ﬁll oil in passenger cars of a European car manufacturer obviously limits the wear of the engine elements in a satisfactory way at temperatures of up to 150 °C. Therefore, HC 5W-30 at 150 °C is used as reference oil. The reference ﬁlm height href in an engine tribosystem with a constant CTr is href = CTrη ref α ref j r P * = 3.68 ⋅ 1 − exp − 0.67 s r 2 3 (16) where r is the radius parallel to the ﬂuid entrainment, while rs is the radius normal to this direction. For a point contact, rs = r yields P*= 1.797, while rs= 20⋅r yields P*= 3.65. 4.6.3 Inﬂuence of lubricant properties The pressure-viscosity coefﬁcient is up to now not mentioned in engine oil speciﬁcations, but it has a strong inﬂuence on the ﬁlm thickness and in consequence also on the frictional losses associated to the ﬁlm shearing. Two of the previously mentioned parameters are inﬂuenced by ﬂuid properties: G is proportional to the pressure-viscosity-coefﬁcient α of the lubricant U is proportional to the viscosity at ambient pressure η0. The other elements in the deﬁnitions of G, Ue and We (or We‘) depend on the geometry, the solid material properties, the speed of the moving surfaces and the load on them. For a comparison of different lubricants which shall be used in the same engine, they can be treated as constants. This yields to three equations of the form hmin = CTrη jα k (17) where CTr is a constant of the tribosystem. As shown in equations (13) to (15), different exponents j and k are used for different contact situations: Table 5 Exponents j and k for different contact situations Contact situation Line Rectangle Ellipse or point Exponent j 0.70 0.71 0.68 Exponent k 0.54 0.57 0.49 The exponents are quite similar in the equations that have been derived for different contact situations. The discussion shows that a comparison of several oils requires data for the dynamic viscosity and for the pressure-viscosity-coefﬁcient at high temperatures (e.g. 150 °C). Both η and α are temperature-dependent and have a negative temperature coefﬁcient. Thus, the ﬁlm thickness hmin decreases if the temperature increases. Functions η(T) and α(T) which are based on measurement results have been discussed in 4.4.4 and 4.5.5, respectively. k (18) where αref and ηref are α and η of HC 5W-30 at 150 °C. The relative ﬁlm thickness h* for any lubricant at any temperature T is deﬁned by h (T ) CTrη (T ) α (T ) η (T ) h (T ) = min = = j k href η ref CTrη ref α ref j k * j α (T ) α ref k (19) and can be calculated without knowledge of the constant CTr, which depends on the properties of the tribosystem. The contact situation must be known to calculate h*, because there are different exponents j and k for different situations. For a line contact, j = 0.70 and k = 0.54: hmin ∝ η 0.70α 0.54 (20) For all tested ﬂuids, the relative ﬁlm thicknesses have been calculated in dependence of the temperature. In Figures 35 to 37, the results are plotted for low temperatures of maximal 100 °C. At 22 °C, most ﬂuids have relative ﬁlm thicknesses between 10 and 16. The higher values for 100E Aero, HC 5W-30 + 3.7% soot, EO-Triol, and PO-Triol are due to a high viscosity at low temperatures, while the low values of ﬁlm thickness for all oils of group 2 as well as for PAG 68 are mainly caused by their low pressure-viscosity coefﬁcient α. At any oil temperature, only a relative ﬁlm thickness of h* ≈1 is needed. This supports the strategy to use engine oils with a high intrinsic viscosity index in order to reduce the fuel consumption especially during cold engine operation. The relative ﬁlm thicknesses at high temperatures are plotted in Figure 38, Figure 39, and Figure 40. Three lubricants show relative ﬁlm thicknesses h*(150 °C)>1: − 100E Aero, mainly because of its high pressure-viscosity coefﬁcient, − HC 5W-30 + 3.7 wt.-% soot, an oil with a similar α-value but a higher viscosity, compared to HC 5W-30 and − PAG 46-2, an oil with a low pressure-viscosity coefﬁcient and a high dynamic viscosity at 150 °C (4.45 mPas, HTHS = 4.3 mPas). For the other oils, h*(150 °C) < 1. For these oils, a critical temperature T* is deﬁned by h*(T*) = 1. The use of these lubricants with the same tribologic safety that is provided by HC 5W-30 30 Forschungsbericht 277 Group 1: Relative film thickness 40 100E Aero 35 Relative film thickness 30 HC 5W-30 + 3.7% soot 25 HC 5W-30 20 Max of 10 others 15 10 Min of 10 others 5 0 20 30 40 50 60 70 Temperature in ˚C 80 90 100 Figure 35 Relative ﬁlm thickness of the oils in group 1 at low temperatures Group 2: Relative film thickness 10 9 Relative film thickness 8 7 IAV-PAS 8 IAV 65-2 6 IAV 65-2 + water 5 IAV 65-3 IAV 65-3 + water 4 3 2 1 0 20 30 40 50 60 70 80 90 100 Temperature in °C Figure 36 Relative ﬁlm thickness of the oils in group 2 at low temperatures Group 3: Relative film thickness 45 40 35 Relative film thickness Triol-PO 30 Triol-EO 25 Paraffin 46 20 PG WS55 15 Max of 6 others 10 Min of 6 others 5 0 20 30 40 50 60 70 80 90 100 Temperature in °C Figure 37 Relative ﬁlm thickness of the oils in group 3 at low temperatures 31 Forschungsbericht 277 Group 1: Film thickness at high temperatures HC 5W-30 1.30 HC 5W-30 + 3.7% soot 1.25 Common curve for HC 5W-30 and Fuchs 100E-Low-SAP 10W-30 1.20 Fuchs HCE 0W-20 Relative film thickness Total 100E 1.15 100E 0W-20 1.10 Total HCE 1.05 Titan SL PCX 0W30 1.00 Total HCE-MidSAP 0.95 Fuchs HCE-LowSAP HC 0W-30 0.90 0.85 130 100E Aero 132 134 136 138 140 142 144 146 148 150 Temperature in °C Fuchs HCE-LowSAP2 0W-20 Figure 38 Relative ﬁlm thickness of the oils in group 1 at high temperatures Group 2: Film thickness at high temperatures 2.0 1.8 Relative film thickness 1.6 1.4 IAV-PAS 8 IAV 65-2 1.2 IAV 65-2 + water 1.0 IAV 65-3 IAV 65-3 + water 0.8 0.6 0.4 70 75 80 85 90 95 100 105 110 Temperature in °C Figure 39 Relative ﬁlm thickness of the oils in group 2 at high temperatures Group 3: Film thickness at high temperatures 1.20 1.15 Relative film thickness 1.10 PAG 46-2 1.05 PAG 46-4 1.00 PPG 32-2 0.95 PPG 32-3 0.90 Triol-PO 0.85 Triol-EO 0.80 Paraffin 46 0.75 130 132 134 136 138 140 142 Temperature in °C 144 Figure 40 Relative ﬁlm thickness of the oils in group 3 at high temperatures 32 146 148 150 Forschungsbericht 277 requires a limitation of oil sump temperature to T*, which can be achieved by an enlargement of the cooling system, high oil ﬁll volume or by a reduction of the maximal power in case of a high engine temperature. But to avoid misunderstandings, three things should be pointed out: is used at a low temperature. Thus, lubricants with h*<1 can contribute to decrease the fuel consumption of the engine if they show relatively low ﬁlm thicknesses in the whole temperature range and if the lowest h*-values which occur in operation are not dangerous for the engine. − The oil Titan SL PCX 0W-30 with T*=148 °C is actually used without problems in a lot of car engines, just like the reference oil HC 5W-30. This proves that a relative ﬁlm height h* slightly below 1 is not dangerous for the engines which are actually in use. − For engines which will be manufactured in future, it is possible to reduce the required minimum ﬁlm height of the lubricant by lowering the surface roughness as well as the clearance of the bearings. In this way, the potential to reduce the fuel consumption can be used. − Low ﬁlm thicknesses are equivalent to a lower power loss by friction, which is mainly important if the engine Viscosity measurement at high shear rates up to 3,4 ⋅ 106 s-1 5 The determination of the HTHS viscosity using a capillary viscometer is standardized in ASTM D4624 for capillaries made of glass or stainless steel. The experience shows that the rotational method (D4741, D4683) has a better repeatability and reproducibility, compared to the capillary method. However, it is by far easier to extend the measuring range to high shear rates using the capillary method. The heating of the oil passing through the PTB capillary viscometer under very high shear rates is limited to 2-3 K. This is one of the main causes for the comparatively small uncertainty. The instruments described in ASTM D4624, ASTM D4741 and ASTM D4683 allow viscosity measurements at shear rates up to 1 ⋅ 106 s-1 provided the instruments are calibrated using special reference oils. The viscosity of these reference oils is certiﬁed at high shear rates of up to about 1 ⋅ 106 s-1. It is important that there is no need to calibrate the PTB capillary viscometer using a reference oil certiﬁed at high shear rates. 5.1 Description of the apparatus The high-shear viscometer HVA 6, manufactured by Anton Paar (www.anton-paar.com), Austria, was used at PTB to determine the Newtonian range of numerous viscosity reference liquids . The working principle is as follows: The liquid under test is pressurized by nitrogen from a cylinder by means of a pressure regulator and a specially formed membrane made of viton for the separation of nitrogen and liquid. When the ball valve is opened, the liquid is driven through the capillary and a parabolic distribution of velocity develops. Behind the capillary, the volume ﬂow of the liquid is measured using a precision glass tube of wide diameter in conjunction with two optical sensors and an interface for the separation of the liquid under test and the liquid used in the precision glass tube. According to the standard , the shear stress at the wall of the capillary is = R ⋅ Δp 2⋅L (21) , with R being the inner radius of the capillary, Δp the differential pressure along the capillary, and L the length of the capillary. In the case of a Newtonian liquid, the shear rate at the wall is D= 4⋅Q ⋅ R3 (22) . Q is the volume ﬂow rate. The dynamic viscosity is η= D (23) . The instrument is designed for a maximum shear rate of 1⋅106 s-1, which can be achieved with a capillary of 100 mm in length, 0.15 mm in diameter and a differential pressure of 12 MPa for a typical engine oil. This instrument had to be modiﬁed in order to meet the following requirements: measuring temperature: up to 150 °C measuring shear rate: up to 3⋅106 s-1 From equation 22 it can be seen that at a constant ﬂow rate smaller diameters result in higher values of D. An increase of pressure to keep Q constant is not possible due to safety reasons. Therefore, a shorter capillary with 25 mm in length and 0.1 mm in diameter is used. The small bore is advantageous to reduce the ﬂow rate (during one measurement series, a total volume of only about 300 cm3 of test ﬂuid is available) as well as the Reynolds number  Re = 2⋅Q ⋅ ρ ⋅η ⋅ R (24) . ρ is the density of the liquid. The radius of the capillary is determined from measurements with a viscometric reference liquid at low shear rates: R=4 8 ⋅η ⋅ Q ⋅ L ⋅ Δp (25) The maximum value Re =1256 occurred in the measurement series on 100 E 0W-20. This is in accordance with  or D 4624, where Re < 2000 is stated. The resulting volume ﬂow rate at pressures of the order of 0.1 MPa is too small to detect the ﬂow rate using the optical system. It is not designed for a slowly moving meniscus. For that reason, and 33 Forschungsbericht 277 due to problems with decoupling the ﬂow rate measuring device from that part of the apparatus, that has to be kept at a temperature of 150 °C, the mass ﬂow is measured instead of the volume ﬂow. A stainless steel tube was installed at the outlet of the capillary with 425 mm in length and an inner diameter of 1.5 mm, allowing the ﬂuid to ﬂow into a glass cup situated on a 200g-balance. The mass ﬂow measurement is carried out as follows: Until stable temperatures are reached, the ﬂow is directed into an additional cup. The reading of the balance is taken. With the start of the measurement, the ﬂow is directed into the cup on the balance. At the end, the ﬂow is switched to the additional cup and the second reading of the balance is taken. The time is taken by means of an electronic stopwatch. The buoyancy correction is used. Using the density, the volume ﬂow is obtained from the mass ﬂow. It is advantageous to have a considerably larger measuring range and no problems with thermostatisation. The small pressure loss in the stainless steel tubing is indeed very small, but it is taken into account in the analysis. Three calibrated 100 Ω platinum resistance thermometers are installed to measure the temperature of the vessel containing the liquid and in the volumes at the inlet and the outlet of the capillary. At 150 °C, the ambient temperature seriously inﬂuenced the thermometer readings. The thermal insulation of just the thermometer bearings was not sufﬁcient. For that reason, an air thermostat with a fan, an electric heater and a pid-controller was installed on the top plate of the HVAinstrument. The vessel containing the liquid is thermostated using a circulating thermostat, and the capillary is − in contrast to the original design − directly thermostated using the air thermostat. Smart piezo-resistive pressure transducers with the ranges 1.6 MPa and 16 MPa and a relative measurement uncertainty of 0.05 % of the maximum value are installed outside the air thermostat. A needle valve is installed in the tubing to the pressure transducer in order to allow a changing of the transducers during one measurement series. This tubing is connected to the vessel containing the liquid at the top. Although the oil under test was heated up to 150 °C before the measurement in order to degas it, gas bubbles sometimes occur and can be allowed to escape through the needle valve (with the pressure transducers removed) before starting the experiment. The hydrostatic head ρ ⋅ g ⋅ h, with g being the acceleration of free fall and h the difference in the height between the pressure transducer and the capillary, is taken into account and is important at small pressures. At a temperature of 150 °C and elevated pressures, considerable diffusion of nitrogen occurred through the Viton membrane which was intended for the gas-liquid separation. A two-phase mixture came out of the capillary which did not allow any reproducible measurements. Changing the material of the membrane did not turn out to be successful. The problem was solved by replacing the membrane by a piston-cylinder assembly made of stainless steel. The gap between the piston and the cylinder is only about 0.1 mm and an O-ring seal (Viton) was used. During the experiment (which took several hours), no more gas bubbles were observed leaving the capillary. Figure 41 is a schematic diagram of the modiﬁed HVA 6 viscometer. The most important correction at high shear rates is the Bagley-correction. It is determined according to the standard  with the aid of a reference liquid on hydrocarbon basis. Three capillaries, 10 mm, 25 mm and 40 mm in length, all with the same inner diameter of 0.1 mm, were used to measure the volume ﬂow rate as a function of the driving differential pressure along the capillary. The result is approximated by the ﬁt ρ = a ⋅ Q + b ⋅ Q2 for each capillary with coefﬁcients a, b and the corresponding standard deviations sa, sb. Thus it is possible to calculate the pressure p and its standard deviation sp at additional values of Q within the measuring range. The result is shown in Figure 42 with the differential pressure p as a function of the length L of the capillaries at constant volume ﬂow rate (parameter). The slope of the curve is calculated from the difference of the pressures between 10 mm and 40 mm. This is identical with the arithmetic mean value of all three slopes at equidistant points: p − p1 1 p 3 − p 2 p 2 − p1 p 3 − p1 + + = 3 3 15 mm 15 mm 30 mm 30 mm (26) From each of the three pressures p1, p2, and p3 at 10, 25, and 40 mm one intersection with the ordinate is calculated. The arithmetic mean value of these three results is the Bagley-correction. The corresponding standard deviation to the degree of freedom 1 is sp1. The Bagley-correction as a function of the ﬂow rate, as shown in Figure 43, is p B = 4.8616 ⋅1010 Q + 1.1057 ⋅1019 Q 2 , (27) with Q in m³/s and pB in Pa. This equation is obtained from measurements with the reference liquid. The maximum difference in viscosity at high shear rates between the ﬂuids and the reference liquid is only 33 %. Therefore it is assumed that equation 27 is also valid for the ﬂuids under investigation. To achieve smallest uncertainties it is necessary to determine the Bagley-correction for each individual liquid. Figure 41 Schematic diagram of the modiﬁed HVA 6 viscometer: 1: pressure vessel, 2: cylinder for gas-liquid separation, 3: piston with O-ring seal, 4: liquid under test, 5: driving nitrogen, 6: ball valve, 7: capillary, 8: pressure transducer, 9: balance with cup, 10: 100 Ω platinum resistance thermometers, 11: thermostated oil, 12: heat-insulated box of the air-thermostat. 34 From the measurements with the ﬂuids under investigation equation 21 is used with Δp = p − pB and the apparent shear rate is calculated : D ap = 4⋅Q = A⋅ + B ⋅ ⋅ R3 2 (28) Forschungsbericht 277 Bagley Correction 14000 12000 Pressure in kPa 10000 8000 6000 4000 2000 0 0 5 10 15 20 25 30 35 40 45 Length of capillary in mm Figure 42 Pressure as a function of the length of the capillary for different ﬂow rates: : 6.0⋅10-8 m³/s, Δ: 1,0⋅10-7 m³/s, x : 1,4⋅10-7 m³/s, ◊ : 2.0⋅10-8 m³/s, ∗ : 1,8⋅10-7 m³/s, 0 : 2,2⋅10-7 m³/s, + : 2,6⋅10-7 m³/s Bagley correction as a function of the flow rate 800 700 y = 1.105731E-02x2 + 4.861572E-02x Bagley correction in kPa 600 500 400 300 200 100 0 0 50 100 150 200 Figure 43 Parabola of Bagley and ﬁtted with the coefﬁcients A and B. From the apparent shear rate, the true shear rate is calculated : Dtrue = D ap 4 ⋅ 3+ 1 1 + ⋅ A⋅ + ⋅ B ⋅ 4 2 250 300 Volume flow rate in 10-9 m3/s D ap 2 ⋅ dD ap d = pressure pm (arithmetic mean pressure in the capillary) has to be corrected according to : η (T , p amb ) = η (T , p m ) ⋅ 1 − p m ⋅ (α − 3 ⋅ D ap + 4 (29) In the case of the liquids under test, the difference between the apparent and the true shear rate is moderate. The maximum difference occurs in the liquid Titan SL PCX 0W-30. The relative difference between the ﬂow curves is only 3 %. Equation 23 with D = Dtrue yields the dynamic viscosity. The dynamic viscosity HTHS150 °C is related to the ambient pressure pamb. For that reason, the viscosity measured at the ) (30) , with α being the pressure coefﬁcient of viscosity and Κ the compressibility of the liquid under test. Values for α are available from measurements with the high pressure rolling-ball viscometer. From the compression of the liquid at 100 MPa at room temperature Τroom from the volume V to V-ΔV, only a rough estimate of the compressibility is possible according to the empirical formula (150 C) = 2.5 ⋅ ΔV 100 MPa ⋅ V (31) , which is based on investigations with oils used for heat exchange. 35 Forschungsbericht 277 During the measurement it is not possible to keep the temperature exactly at 150 °C. At high shear rates, the temperature increases at the capillary outlet due to friction. After starting the ﬂow through the capillary, it takes about 20s until a stable temperature is reached and the ﬂow measurement can be started. In the analysis, the arithmetic mean of the temperatures, measured by the platinum resistance thermometers, is used. The temperature differences to 150 °C are corrected using the temperature-viscosity coefﬁcient β obtained from a Vogel equation, which is set up for each liquid under investigation by measurements at ambient pressure: K 150qC, p amb K T , p amb >1 E T 150qC @ 5.2 Results The result of the corrected HVA 6 measurements is shown in Figure 44. The polyalkylene glycol PAG 46-4 shows an almost perfect Newtonion behavior with no indications for a shear thinning. The polymer-free Fuchs HCE 0W-20 and 100E 0W-20 are quite similar. The shear thinning of the polymer-containing FUCHS HCE-Low-SAP and Titan SL PCX 0W-30 is visible. All four liquids show a very weak pseudoplastical behavior. In rheology, Figure 44 is normally shown with a double logarithmic axis. As a result, 5 nearly horizontal curves are obtained. If only the D-axis is logarithmic, Figure 44 changes to Figure 45. (32) Viscosity as a function of the shear rate at 150°C 36 PAG 46-4 32 Titan SL PCX 0W-30 -4 η in 10 Pa s Fuchs HCE-Low-SAP 28 100E 0W-20 Fuchs HCE 0W-20 24 0 500 1000 1500 2000 3 D in 10 s 2500 3000 3500 4000 -1 Figure 44 Viscosity curves depending on the shear rate for ﬁve different liquids Viscosity as a function of the shear rate 36 Titan SL PCX 0W-30 32 -4 η in 10 Pa s Fuchs HCE 0W-20 100E 0W20 28 Fuchs HCELow-SAP PAG 46-4 24 Polynomisc 10 100 1000 D in 103 1/s Figure 45 Viscosity curves depending on the shear rate for ﬁve different liquids 36 10000 Forschungsbericht 277 5.3 Estimation of the measurement uncertainty The uncertainty calculations according to  are performed with the aid of the GUM Workbench. The working equations 25, 21, 28, 29 and 23 are entered, and after that the input quantities including an appropriate measure of uncertainty. One contribution to the Bagley-correction is the already deﬁned sp1, which is correlated with the uniformity of the inner diameter of the three capillaries of different length. In addition, the standard deviation sp2 belonging to sa ⋅ Q + sb ⋅ Q2 is calculated. This contribution is a measure for the quality of the ﬁt p = a ⋅ Q + b ⋅ Q2. For each capillary, at least 11 measurements are carried out to determine the coefﬁcients a and b. Hence, the degree of freedom 9 is used for sp2. At high shear rates the Bagley-correction yields the leading term in the uncertainty budget. Another important contribution is due to coefﬁcient B in equation 28. The corresponding standard deviation sB is calculated from the measurements of the liquid under investigation. For example, in the case of the liquid Titan SL PCX 0W-30 at a volume ﬂow rate of 340 mm³/s (corresponding shear rate: 3.4⋅106 s-1) in the uncertainty budget, the index of sp1 is 37.5 %, that of sp2 24.9 %, and that of sB 16.1 %. The contribution for the determination of the capillary radius yields 4.2 % (small volume ﬂow rate of 5.56 mm³/s) and the measurement of low pressure 4.3 %. The remaining three inﬂuence 6 The uncertainty of viscosity as a function of the shear rate η = η(D) and not as a function of the ﬂow rate η = η(Q) is of interest. For that reason the relative standard uncertainty (denoted by ‘) along the viscosity curve is s ' 2 (η (D )) = s '2 (η (Q )) + D dη ⋅ ⋅ s ' 2 (D ) η dD + s '2 (fit ) , (33) provided the input quantities are not correlated. The relative standard deviation s‘(ﬁt) is calculated from the differences between the viscosity curve and the measuring points. The GUM Workbench yields the result 2.7 % for the expanded relative measurement uncertainty of the shear rate (same example as before). Thus, the expanded relative measurement uncertainty at D = 3.45⋅106 s-1 is 5.1 %. At D = 2.90⋅106 s-1 this value is reduced to 4.2 % and at D = 0.31⋅106 s-1 to 3.3 %. These results and uncertainties for the other tested liquids are shown in Figure 44 as error bars. Tribological behavior under continuous sliding (BAM-method) The following ﬁgures in this chapter contain not only the wear rates under mixed/boundary lubrication for each triboelement, but also to each couple the coefﬁcient of friction. They were achieved by using the BAM-test procedure . The wear rate (or wear coefﬁcient) is deﬁned by the wear volume divided by sliding distance and load. 6.1 quantities with indices higher than 0.7 % are: the volumetric ﬂow rate at the high shear rate (4.5 %), temperature measurement including control (4.3 %), and the coefﬁcient A (2.7 %). As a result, the expanded relative measurement uncertainty for the viscosity 4.9 % is obtained. TOTAL HC 5W-30 fresh oil and as engine aged with soot Depending from the combustion process in an internal combustion engine, the generation of soot in the oil concerns wear, as the primary particles of 20-50 nm agglomerate to some micrometers and became larger than the oil ﬁlm thicknesses. The Figure 46 bench mark nine metallurgically different piston rings under mixed/boundary lubrication. In any cases, the 3.7 wt.-% soot increased the ring wear by one to two orders of magnitude. The only exception was the “GDC50®”, a hard chrome coating with diamond particles. The GGL20HCN liner wear was accelerated by up to one order of magnitude. Based on the viscometric results, it can be concluded, that soot has an adverse effect on component life only under the regime of mixed/boundary lubrication. The abrasive action of soot can be limited, but not eliminated, by using hard metal or cermet (see also Figure 58) as well as GDC50® ring coatings. 6.2 FUCHS Titan GT1 The ester-based TITAN GT1 (HCE) displayed no indications for accelerated wear under mixed/boundary lubrication of the piston ring coatings and cylinder liner materials. The wear rates and coefﬁcients of friction of cast iron (GGL20HCN) in GT1 are in the same range as those in HC 5W-30. MKJet 502® presented the lowest ring wear, but associated also the highest liner wear of triboactive liner coatings (See Figure 47). MKJet 502® and TM23-1 are in GT1 responsible for high/excessive liner wear. The “zero” liner wear of quite rough ﬁnished (RPK ≈ 0.36 μm) TM23-liner was associated with intensive TinO2n-1- and MKP81A®-ring wear. 6.3 TOTAL HCE midSAP The wear rates of the reference couple MKP81A®/GG20HCN in HCE midSAP were similar to those measured in HC 5W-30. A friction reducing action of HCE lowSAP was not observed. “Zero” liner wear in HCE mid SAP was observed when using smoothly machined APS-Tin-2Cr2O2n-1 and HVOF-(Ti,Mo)(C,N)23NiMo liner coatings (See Figure 48). The wear rates of APS-Tin-2Cr2O2n-1 coated rings are in HCE midSAP in the same order of magnitude than those of MKP81A®, even the wear rates of GG20HCN- and TinO2n-1-liner tend to be slightly increased by the APS-Tin-2Cr2O2n-1 ring coating. 37 Forschungsbericht 277 Figure 46 Volumetric wear coefﬁcients of coated piston rings and uncoated grey cast iron using HC 5W-30 fresh oil and diesel engine aged HC 5W-30 (1.9 liter TD, 89 kW) under mixed lubrication conditions 6.4 FUCHS HCE lowSAP The “hard/hard” couples MKJet 502®/TM23 were not signiﬁcantly more wear resistant than MKP81A®/GG20HCN using the HCE LowSAP formulation. The lowest liner and system wear was observed in HCE LowSAP with the smooth ﬁnished (Ti,Mo(C,N)-23NiMo liner coating (See Figure 49). The HCE lowSAP formulation displayed a trend to increase under mixed/boundary lubrication the coefﬁcients of friction. 6.5 PPG 32-2 As with the PAG46-4 (see chapter 6.6), the coefﬁcients of friction were signiﬁcantly lowered also in PPG32-2. The GGL20HCN liner wear seemed to be slightly increased by PPG32-2. “Zero” wear and low friction can be achieved by using triboactive coatings and PPG32-2. 38 6.6 PAG 46-4 As the PAG46-4+2.65 Phopani contained no friction modiﬁers, the PAG-base oil molecules tend to halven down to ca. 0.04 the coefﬁcients of friction compared to most other formulations of this test programme (See Figure 51). Smoothly ﬁnished (Ti,Mo)(C,N)-23NiMo offer in PAG46-4 “zero” liner wear associated with reduced coefﬁcients of friction. The lubrication of GGL20HCN with PAG46-4 did not affect the wear rates, but reduced the coefﬁcients of friction for different ring metallurgies. In PAG46-4, the APS- TinO2n-1liner was as wear resistant as the GGL20HCN. It was surprising to observe, that in PAG46-4 the “nitrided” ring presented one of the highest wear resistances of the rings, which has an economic impact. Forschungsbericht 277 Figure 47 Volumetric wear coefﬁcients of coated piston rings and of thermal sprayed triboactive cylinder liner coatings and uncoated grey cast iron using TITAN GT1 and factory ﬁll oils under mixed lubrication conditions 6.7 GGL20HCN the coefﬁcient of friction even though no friction modiﬁers were added, as well as in the PAG 46-1. Additionally, the wear resistance of piston rings in low additivated PAGs46 are comparable or even higher than in the hydrocarbon and ester based lubricants. Figure 52 summarizes the wear rates of the different piston rings (Mo (MKP81A® and VL®), TinO2n-1, TM23-1 and MKJet502®) sliding against an uncoated grey cast iron cylinder material with high carbon content (HCN) in different lubricants. For these combinations the wear rates of the piston rings with two Mo-based coatings are quite similar indicating a high reproducibility of the test method and coating process. As trend the wear rates of the GGL20HCN cylinder liner are “independent” of the used ring material and lubricant or “robust” against the selected ring materials or lubricant with exception in Titan GT1 sliding against MKP81A. The MKJet 502® seems to be sensitive to the lubricant used and does not always presents signiﬁcant tribological beneﬁts. On the other hand, the lowest friction was observed for the couples MKJet502®/GG20HCN and MKP81A®/GG20HCN in PAGs 46-3/-4. The PCX and the two PAGs displayed the lowest coefﬁcients of friction. The TM23-1 ring was ﬁnished to RPK ≈ 0.8 μm and the second TM23-2 grade to RPK ≈ 0.18 μm. Overall, the tribological data in Figure 53 show clearly, that rings coated with (Ti,Mo)(C,N)23NiMo present signiﬁcantly reduced ring and liner wear rates, when they are smoothly machined. The wear rates of APS coated piston rings are similar to that of the reference Mo-based coating with a trend of higher wear rates for APS-TinO2n-1 (See chapter 6.10). The coefﬁcient of friction can be halvened with Supersyn SL PCX and by a factor 2 – 3 with PAGs 46-3/-4. The PAGs 46-3/-4 reduces The wear resistance of triboactive (Ti,Mo)(C,N)+23NiMo (TM23) piston ring coatings is similar or even higher than for widespread used Mo-based coatings against grey cast iron in different lubricants. In Zinc-free PCX and PAG46-4 slightly reduced coefﬁcients of friction were observed. The ﬁgure of 4.51 10-8 mm³/Nm in Figure 55 represents the average of 230 tests performed with GGL20HCN against different ring metallurgies and in several oils. This may act as a base line for validation of other liner metallurgies in the BAM test. 6.8 (Ti,Mo)(C,N)-23NiMo liner coating 39 Forschungsbericht 277 Figure 48 Volumetric wear coefﬁcients of coated MKP81A® and APS-Tin-2Cr2O2n-1 piston rings and of thermal sprayed triboactive cylinder liner coatings and uncoated grey cast iron using Total midSAP under mixed lubrication conditions Mating TM23 ring coatings with liners coated with TiO1.93, TinO2n-1 or Ti2-nCr2O2n-1 resulted in all formulations in excessive liner wear of the laters. 6.9 Ti2-nCr2O2n-1 liner coating The APS-Tin-2Cr2O2n-1 liner wear rates (RpK ≈ 0.61 μm) in Figure 54 displayed either no signiﬁcant advantages in respect to GGL20HCN regarding wear resistance within the precision of the BAM test procedure between the fully-formulated hydrocarbon- and ester-based oils nor positive effects regarding a ﬂuid-surface interaction. Liners coated with APS-Tin-2Cr2O2n-1 presented in Figure 54 and in Figure 55 a signiﬁcant wear reduction by nearly two orders of magnitude in “HC” 5W-30 using all of the three types of Mo-based rings, when the liner surfaces were smoothly ﬁnished to RPK ≈ 0.05-0.1 μm. The same wear reduction was found in PCX and HC 5W-30 only mating with the PCF 251 and PCF 278 rings. Mating with the MKP81A® piston rings, the APS-Tin-2Cr2O2n-1 liner coating with RpK ≈ 0.61 μm is as wear resistant as the GG20HCN. Such a wear resistance was also determined for TinO2n-1 liner coatings. This qualiﬁes APS-Tin-2Cr2O2n-1 liner as a candidate liner coating for aluminium substrates. 40 The low APS-Tin-2Cr2O2n-1 liner wear in “HC” 5W-30 was conﬁrmed by means of repeated tests. The low coefﬁcients of friction for “PCX” are in general not untypical for this Mo-free formulation. On the other hand, the friction modiﬁer used in “PCX” not always interacts with all “ceramics” and “hard metals” resulting in low friction. The most signiﬁcant reduction in “system” wear was demonstrated by mating the APS-Tin-2Cr2O2n-1 coated piston rings with smooth machined HVOF-(Ti,Mo)(C,N) liner coating (See Figure 58), which was conﬁrmed in BAM- and SRV-type tests. The APS-Tin-2Cr2O2n-1 coated piston rings offer a potential for substituing molybdenum and hard metal based rings. In the case of PCX-oil, the friction modiﬁer of PCX not acted beneﬁcial in couples sliding against Ti2-nCr2O2n-1-liners. 6.10 TinO2n-1 ring coatings Figure 56 presents effects of liner materials for APS-TinO2n-1 coated piston ring wear in factory ﬁll HC 5W-30 as well as in the ester containing Titan GT1. Against GGH20HCN, the wear of the TinO2n-1 coated piston ring is increased as well as slightly the wear rate of the cast iron using GT1, but tribocouples with TinO2n-1 coated piston rings sliding against Forschungsbericht 277 Figure 49 Volumetric wear coefﬁcients of coated piston rings and of thermal sprayed triboactive cylinder liner (Ti,Mo)(C,N) coatings and uncoated grey cast iron using FUCHS lowSAP under mixed lubrication conditions TiOx liner coatings exhibit the same high wear resistance as commercial Mo/GGL20HCN couples. The wear coefﬁcient of TinO2n-1 coated piston rings is similar to that of Mo-based ring coating MKP81A® in HC 5W-30 and the wear of grey cast iron is not affected. TM23 liner coatings illuminated the highest wear resistance close to “zero” wear independent of liner roughness. The smoothing of TM23 liner roughness from Rpk ≈ 0.53 μm to Rpk ≈ 0.03 μm reduced the wear rates of TinO2n-1-rings by more than one order of magnitude. Under unidirectional sliding a slightly higher system wear and under oscillating sliding conditions (SRV®) a reduced system wear was measured for APS-TinO2n-1/GG20HCN compared to Mo/GGL20HCN. Overall, both ring coatings wear in the same order of magnitude. It is reasonable that thermally sprayed TiOx-based coatings can substitute common materials and serve as a promising alternative to commercial piston rings coated of strategic molybdenum. 6.11 Ester oil The wear rates of MKP81A® coated rings and those of the smoothly machined (Ti,Mo)(C,N)-23NiMo liner coatings were in GTE (100E) also close to the “zero” wear limit. In order to meet a “zero” wear target, MKJet502® is not consequently necessary to use this ester formulation. Also in GTE (100E), the MKJet502® increased the wear rates of coated TinO2n-1 and TiO1.93 liners. It has to be noted, that the APS-Tin-2Cr2O2n-1 liners presented on “abrasive” action against the hard metalbased MKJet 502®. A speciﬁc friction reducing action under mixed/boundary lubrication against many couples was not observed with this ester-oil formulation. 6.12 Zero wear target Compared to the MKP81A®, the APS-Tin-2Cr2O2n-1 ring coating (see Figure 54) sliding on GG20HCN displayed no advantages in wear resistance, except when lubed by the PCX-oil. A synergistic, signiﬁcant reduction of the „system wear“ (summ of ring and liner) can be achieved by mating the APSTin-2Cr2O2n-1 coated piston rings with smooth ﬁnished HVOF(Ti,Mo)(C,N)-23NiMo liner samples as the APS-Tin-2Cr2O2n-1 ring wear is reduced by one order of magnitude and those of HVOF-(Ti,Mo)(C,N)-23NiMo liner samples up to two orders of magnitude down to the resolution limit of the BAM-test using 24 km (see Figure 58). 41 Forschungsbericht 277 Figure 50 Volumetric wear coefﬁcients of coated piston rings and of thermal sprayed triboactive cylinder liner coatings and uncoated grey cast iron using PPG32-2+2.65 Phopani and factory ﬁll oils under mixed lubrication conditions If the HC 5W-30 contains 3.7 wt.-% of soot (Diesel engine aged HC 5W-30), the wear rates of (Ti,Mo)(C,N) coated liner and APS-Tin-2Cr2O2n-1 coated piston ring increase under mixed/boundary lubrication by one order of magnitude close to the wear rates of MKP81A®/GGL20HCN in fresh HC 5W-30 without soot. The friction reducing effect of soot is probably related to the lapping movement (slip-rolling) of the soot particle in the tribocontact. By using APS-Tin-2Cr2O2n-1/(Ti,Mo(C,N)-23NiMo, the zero wear target can be achieved in fully additivated hydrocabonbased or with alternative formulations with reduced contents of additives. 6.13 Summarizing friction and wear behavior in BAM test Figure 59 and Figure 60 summarize in two plots the coefﬁcients of friction under mixed/boundary lubrication versus wear rate of different triboactive and state-of-the-art ring and liner coatings in ﬁve oils using the BAM test procedure. 42 The two polyglycols without and the FUCHS PCX containing an organic friction modiﬁer displayed as a trend the lowest coefﬁcients of friction under the regime of mixed/boundary lubrication. Depending from the portion of mixed/boundary lubrication, they will contribute to improve Fuel Economy. The PPG32-2, PAG46-4 and the PCX offer, lubing the appropriate materials, a potential for “zero liner wear”, even they are polymer-, Zn- and Mo-free and respect bionotox criteria (Except bionotox for PCX!) and follow different strategies to reduce Zinc, phosphorus, sulphur and low ash. Such lubrication concepts avoiding a different number of additives (VI, EP, AW) enable a retention of tribological properties over drain. For given test conditions all APS coatings on piston rings showed no friction reducing effect. The coefﬁcient of friction is more determined by the lubricants than by the materials or by an individual interaction between lubricants and a speciﬁc material or tribopairing. Forschungsbericht 277 Figure 51 Volumetric wear coefﬁcients of coated piston rings and of thermal sprayed triboactive cylinder liner coatings and uncoated grey cast iron using PAG46-4+2.65 Phopani and factory ﬁll oils under mixed lubrication conditions 43 Forschungsbericht 277 Figure 52 Volumetric wear coefﬁcients of coated piston rings and uncoated grey cast iron GGL20HCN in different oils under mixed lubrication conditions 44 Forschungsbericht 277 Figure 53 Volumetric wear coefﬁcients of two ﬁnishing grades of APS-(Ti,Mo)(C,N) coated piston rings sliding under mixed lubrication conditions against cast irons and different coatings in different oils 45 Forschungsbericht 277 Figure 54 Volumetric wear coefﬁcients of TinO2n-1, MKP81A® and MKJet502® coated piston rings sliding under mixed lubrication conditions against cast irons and APS-Tin-2Cr2O2n-1 in different oils 46 Forschungsbericht 277 Figure 55 Volumetric wear coefﬁcients of Mo-based and MKJet502® piston rings sliding under mixed lubrication conditions against cast iron and APS-Tin-2Cr2O2n-1 in two factory ﬁll hydrocarbon-based lubricants. 47 Forschungsbericht 277 Figure 56 Volumetric wear coefﬁcients of APS TinO2n-1 coated piston rings and of different triboactive cylinder liner coatings in comparison to uncoated GGL20HCN using factory ﬁll HC 5W-30 and an ester-containing Titan GT1 under conditions of mixed lubrication. 48 Forschungsbericht 277 Figure 57 Volumetric wear coefﬁcients of coated molybdenum-based and MKJet502® coated piston rings sliding against of thermal sprayed triboactive cylinder liner coatings and uncoated grey cast iron using TITAN GTE (100E) under mixed lubrication conditions 49 Forschungsbericht 277 Figure 58 Volumetric wear coefﬁcients of MKP81A® and APS-Tin-2Cr2O2n-1 coated piston rings sliding under mixed lubrication conditions against cast iron GGL20HCN and HVOF-(Ti,Mo)(C,N) in different oils 50 Forschungsbericht 277 Figure 59 Summarizing plot of “coefﬁcient of friction at test end” versus “Wear rate for ring” of sets of different tribo-couples in PAG 46-4+2.6 Phopani, PPG32-2+2.6 Phopani, SAE 5W-30 (HC), PCX 0W-30 and GT1 using the BAM test (FN = 50 N ; v = 0.3 m/s ; T = 170 °C ; s = 24 km) Figure 60 Summarizing plot of “coefﬁcient of friction at test end” versus “Wear rate for liner” of a set of different tribo-couples in PAG 46-4+2.6 Phopani, PPG32-2+2.6 Phopani, SAE 5W-30 (HC), PCX 0W-30 and GT1 using the BAM test (FN = 50 N ; v = 0.3 m/s ; T = 170 °C ; s = 24 km) 51 Forschungsbericht 277 7 Tribological behavior under linear, oscillating sliding (SRV®-method) 7.1 Extreme pressure behavior in the SRV® test The resistance against seizure of an iron based alloy (ball bearing 100Cr6H = AISI 52100) was determined for different lubricants with the SRV® test rig2 according to ASTM D5706-05 under conditions of mixed lubrication and quoted as Hertzian contact pressure (for the last O.K. pressure before failure, see Figure 61). At 135 °C, the factory-ﬁll oils ranged acceptable from 3000 MPa to 3500 MPa. The unadditivated polyglycols (base oil= b.o.) PAG46-4 b.o. and PPG32-2 b.o. itself achieved highest values of ~ 3700 MPa, which were lowered by antioxidants, and followed or on the same level by the ester-based formulations FUCHS 100E and TOTAL HCE. The formulations having low content of EP-additives or containing no “classic” EP-additives displayed no disadvantages exceeding the maximum design limit of today of 2000 MPa. 7.2 Friction and wear ® The SRV procedure applies a higher load of 300 N associated with a lower oil temperature of 135 °C than the BAM test procedure using 50 N and 0.3 m/s at 170 °C oil temperature. The different deposition techniques applied for the liner samples using HVOF in BAM test versus APS in SRV® test of the same spray powder play a key role and SRV ® , n – Schwingung, Reibung, Verschleiß (German); oscilation, friction, wear (English translation). Optimol Instruments GmbH, Westendstr. 125, D-80339 Munich, Germany. See ASTM D5706, D5707, D6425 and D7217. 2 have to be taken into account as additional factors the lower oil temperature of 135 °C and higher roughness of the APS(Ti,Mo)(C,N)-23NiMo liner samples. The cast iron liner wear didn´t increase using the APSTin-2Cr2O2n-1 piston ring coating compared to MKP81A®. Also the APS-Tin-2Cr2O2n-1 piston ring coating wear remain on the same level as the MKP81A® coating ring wear. Thus, also the SRV® tests conﬁrm the potential for substituting molybdenum-based rings by APS-Tin-2Cr2O2n-1 and also the signiﬁcant reduction of liner wear when APS-Tin-2Cr2O2n-1 coated rings are mated with APS-(Ti,Mo)(C,N)-23NiMo. The wear rates of the liner samples coated with APS-(Ti,Mo)(C,N)-23NiMo lie in the range of kV = 0.7 to 0.9 10-9 mm³/Nm (see Figure 62). Figure 63 compiles friction and wear results determined with SRV® test rig under linear, oscillating sliding motion for mixed lubricated conditions for APS-Mo (MKP81A®) and APS-TinO2n-1 coated piston rings running against GGL20HCN cylinder liner material (lamellar cast iron with high carbon content). The coefﬁcient of friction is overall not affected by the used Mo and TinO2n-1 piston ring coatings with a small scatter of about 1 % between both coatings. Nevertheless both piston ring coatings exhibit the same dependency for the coefﬁcient of friction in different lubricants. The lowest coefﬁcient of friction was measured with polyalkylene glycoles, which corresponds to the results with unidirectional sliding motion (BAM test, see Figure 56). Compared with the factory ﬁll hydrocarbon based HC 5W-30 the ester containing GT1 and GTE can reduce the coefﬁcient of friction under oscillating and unidirectional sliding motion by about 0.01-0.03. PPG32-2 and modiﬁed PPG32-2 have in SRV tests a higher coefﬁcient of friction than in the BAM test. The FUCHS Supersyn SL PCX presents Figure 61 Resistance against seizure for different lubricants according to ASTM D5706-05 using 100/Cr6H/100Cr6H (AISI 52100) at 135 °C 52 Forschungsbericht 277 Figure 62 SRV® test results for MKP81A®- and APS-Tin-2Cr2O2n-1-coated piston rings sliding on GGL20HCN and APS-(Ti,Mo)(C,N)-23NiMo disks in different lubricants (Top: coefﬁcients of friction; bottom: wear rates; FN = 300 N, Δx = 2 mm, ν = 50 Hz, s = 1440 m) under oscillation sliding a higher coefﬁcient of friction as it was not found under unidirectional sliding according to the BAM test method. The wear data under linear oscillation display, that the molybdenum coating can be substituted by the TinO2n-1 coating, since the wear rates of the TinO2n-1 piston ring coating are comparable or lower than those of the Mo-coating. The ranking was conﬁrmed by BAM tests. Furthermore, the TinO2n-1 piston ring coating promotes a beneﬁcial wear reducing action when lubed with alternative bio-no-tox oils as Titan GTE, PAGs 46 and PPG32-2 with 2.6 Phopani. 7.3 Precision of SRV® test The Figure 64 and Figure 65 show the inﬂuence of two different test conditions (BAM and SRV® test) on friction and wear of molybdenum-based MKP81A® against cast iron GGL20HCN in the SRV® test with the associated standard deviation from ﬁve consecutive tests. Both, load and temperature were changed. For these comparisons in a meaningful two hours SRV® test, the stroke and frequency were identical. The standard deviation bars indicate a high repeatability for these SRV® tests which is superior to those known from engine tests. 53 Forschungsbericht 277 Figure 63 SRV® test results for MKP81A® coated and TinO2n-1 coated piston rings against GGL20HCN in different lubricants (Top: coefﬁcient of friction, bottom: wear rates] 54 Forschungsbericht 277 Figure 64 Repeatability and inﬂuence of test conditions on friction using SRV® for piston ring cylinder liner evaluation Figure 65 Repeatability and inﬂuence of test conditions on wear using SRV® for piston ring cylinder liner evaluation 55 Forschungsbericht 277 8 Concluding summary The present results from the test program revealed that in engine oil speciﬁcations the dynamic viscosity, especially measured under higher shear rates than 106 s-1, the heat capacity and the pressure-viscosity-coefﬁcients have to be introduced, especially when alternative oils of different chemistries have to be ranked. With these data, the oil ﬁlm thickness of an individual formulation can be calculated. In order to differentiate viscometric properties of alternative oils, the dynamic viscosity taking into account the differences in density has also to be used. All viscometric lubricant properties should be determined at least at 150 °C. Some polymer-free ester-type and polyglycol-based engine oils presented thermo-physical and viscometric properties conforming with hydrocarbon-based factory-ﬁll oils or exceeding them. The cooling ability (volumetric heat capacity cp⋅ρ) of polyglycols was the highest. The alternative ester- and polyglycol-based formulations offer additional beneﬁts when criteria related to biono-tox, low NOACK evaporation, high VI, low/no ash content and reduced additive concentrations have to be respected. Important properties such as price or polymer compatibility were not considered here, but they are the subject of other, parallel, validations. A favorable economical forecast can be seen in light of the sum of functional properties displayed by the alternative formulations. Based on the piston ring/cylinder liner simulation tests performed outside of engines by means of the BAM and the SRV® tests, both performed under conditions of only mixed/ boundary lubrication, it is reasonable to conclude, that a. thermally sprayed TiOx-based coatings (Tin-2Cr2O2n-1, TiO1,93, TinO2n-1) can substitute common materials and serve as a promising alternative to commercial piston ring coatings using strategic molybdenum b. “Zero wear” was displayed by mating the APSTin-2Cr2O2n-1 coated piston rings with smooth machined HVOF-(Ti,Mo)(C,N) liner coatings or molybdenum-based rings against smooth machined APS-Tin-2Cr2O2n-1 and (Ti,Mo(C,N) coated liners c. The coefﬁcient of friction is more determined by the lubricants than by the materials or by an individual interaction between lubricants and a speciﬁc material or tribopairing. For given tribological test conditions all APS coatings on piston rings showed no friction reducing effect. The different bionotox and low ash prototype engine oils with reduced additive contents displayed isoperformance regarding the tribological behavior when lubing common and triboreactive materials. They presented no visible weakness in wear resistance, coefﬁcient of friction and extreme pressure properties, but a distinct potential to reduce the coefﬁcient of friction and to reduce the system wear. The outcomes and data generated in the frame of this project, showing that ecological compatibility and technical performance can be reached simultaneously, supported the German Environmental Agency (www.umweltbundesamt.de) to draw the draft document with the criteria for the attribution of an ecolabel for engine oils. 56 Acknowledgements: The authors are grateful to the German Ministry of Economics and Labour (www.BMWA.bund.de) funding the project BMWA14/02 “New lubrication concepts for environmentally friendly machines” related to thermophysical and viscometric properties of alternative lubricants interacting tribologically with triboreactive materials. Some of the protoype oils were supplied within the framework of the parallel EC-funded project GROWTH Contract N° G3RD-CT-2002-00796-EREBIO, “-Emission reduction from engines and transmissions substituting harmful additives in biolubricants by triboreactive materials” focussing on triboreactive materials and bio-no-tox-properties. The supply of oils is gratefully acknowledged by the authors. The alternative bio-no-tox-lubricants were supplied by the project partners Fuchs Petrolub AG (Mannheim, Germany) and via Renault SAS from Total SA (Paris, France). Factory-ﬁll oils used at the project partner Renault SAS (Paris, France) were supplied by Total SA (HC 5W-30 fresh oil, aged in a turbodiesel engine having 3.7 wt.-% soot) and Fuchs Petrolub AG (Titan SL PCX 0W-30). Prototype oils were supplied by the same companies (Total: 100E, HCE and HCE-Low-SAP; Fuchs: Fuchs HCE 0W-20 and GTE (100E 0W-20), HCE low SAPs, 100E HDDO). All pre-blended polyglycols were modiﬁed by the Federal Institute for Materials Research and Testing (BAM), Berlin. The authors are grateful for the active support received from industry and OEM members, namely to Dr. Michael Berg and Dr. Hubert Schultheiß of IAV GmbH as well as to Tom Linnemann, Gérard Desplanches, Bernard Criqui and Nathalie Davias of Renault SAS and to Rolf Luther of FUCHS Petrolub AG. The experimental APS-TinO2n-1, APS-Tin-2Cr2O2n-1 and APS(Ti,Mo)(C,N)+23NiMo coated piston rings were provided by Jesu Landa/Dr. Iñaki Illaramendi from Grupo CIE Automotive (Tarabusi), Barrio Urquizu 58, E-48140 Igorre, Spain. Within BAM: Mr. Norbert Kelling and Manfred Hartelt are gratefully acknowledged for performing the tribological tests and proﬁlometry. The assistance of our colleagues Ms. Sigrid Binkowski, Ms. Silvia Benemann and Ms. Birgit Strauß is gratefully acknowledged in carefully performing metallography, recording optical and SEM micrographs and Ms. Dagmar Nicolaides for XRD analysis and the measurement of particle size distribution. Within PTB: The authors wish to thank Mr. Karl-Heinz Metzing for the construction and manufacturing of the rolling-ball viscometer as well as of important parts of the high-shear viscometer. He also did a lot of measurements with this two instruments. Thanks also go to Ms. Nicole Wloczek for carrying out the kinematic viscosity measurements using Ubbelohde viscometers. We are also grateful to Mr. Jörg Matthis, who performed the heat conductivity measurements according to Dr. Ulf Hammerschmidt. We have to thank Dr. Stefan Sarge for the speciﬁc heat measurements carried out by Mr. Peter Bartling, Dr. Dirk Boghun and Dr. Michael Müller–Wiegand as well. Dr. Harro Bauer is acknowledged for numerous discussions. Forschungsbericht 277 List of variables Symbol a, b c Dimension K - A m² s-1 N-1 Ac B m² m4 s-2 N-2 cp kJ kg-1 K-1 CTr d D Dap Dtrue E f F G h* hmin href m s-1 s-1 s-1 MPa N mm mm kV L Mw p pB pm pamb r mm³/(Nm) mm g mol-1 MPa MPa MPa MPa mm rA mm Definition Coefficients of the Vogel equation unless otherwise defined in the text Linear coefficient of the apparent flow curve Area of circular plates Quadratic coefficient of the apparent flow curve Specific heat capacity at constant pressure Constant, tribosystem Gap Shear rate Apparent shear rate True shear rate Elasticity modulus Coefficient of friction Force on the contact Material parameter Relative film thickness Minimum film thickness Film thickness of reference oil at 150°C Wear rate Capillary length Molar mass Pressure Bagley correction Pressure, arithmetic mean Ambient pressure Equivalent radius of curvature Radius of curvature of cylinder/sphere A Symbol rB Dimension mm R Re s sz mm m Dimension of quantity z m³ s-1 W s’ Q Q T Ue VI We D °C GPa-1 E K-1 K Kavg mPa s mPa s N O Q GPa-1 W m-1 K-1 mm²s-1 mm²s-1 μ — kg m-3 N m-2 Qavg U W Definition Radius of curvature of cylinder/sphere B Capillary radius Reynolds number Sliding distance Standard deviation Relative standard deviation Volume flow rate Flow of thermal energy Temperature Speed parameter Viscosity index Load parameter Pressure coefficient of the dynamic viscosity Temperature coefficient of the dynamic viscosity Dynamic viscosity Arithmetic mean of the dynamic viscosity within a group of oils Compressibility Thermal conductivity Kinematic viscosity Arithmetic mean of the kinematic viscosity within a group of oils Poisson ratio Density Shear tension 57 Forschungsbericht 277 9  Literature/References RENAULT SAS Dossier Zukunftssichere Entwicklung – “ELLYPSE”, Radikal konstruiert-R&D –Wege der Innovation-, Das Magazin für Forschung und Entwicklung, Nr. 26, Oktober 2002, Publisher: Renault SAS, Direction de la Communication, rue du Vieux-Pont-de-Sèvres, F-92109 Boulogne-Billancourt (France), ISSN: 1289009X or in press kit for the „Mondial de l´Automobile, 2002, Paris (www.planeterenault.com go to Protos go to Ellypse)  E.S. 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