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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 [1]
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 [2].
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 [5] 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 [6] (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 [9] 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. [10]
and by Höhne et al. [11].
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
[12].
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®
[13] and BAM [14] 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 [13] 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 [14] using liner segments and in
Figure 2.
The SRV® sample configuration with piston ring segments and
wear scars [15] 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 [16].
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 [17] 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 [20]
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 [24] 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
[25].
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 [26].
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 [31] and Blume [32].
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 [30]. 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 [36] and Jones [37] 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 [36]:
U 0.71G 0.57
hmin
= 3,07 e 0.11
r
We '
(14)
29
Forschungsbericht 277
Elliptical contact [37]:
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 [39]. 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 [40], 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 [41]
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 [41] 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
[41] 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 [40]:
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 [40]:
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 [41]:
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 [9] 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 [14]. 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
[1]
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)
[10]
E.S. Watson, J.J. O‘Neill, and N. Brenner
A Differential Scanning Calorimeter for Quantitative
Differential Thermal Analysis
Analytical Chemistry 36 (1963), pp. 1233-1238
[11]
G. Höhne, W. Hemminger, and H.-J. Flammersheim
Differential Scanning Calorimetry
2nd edition, Springer-Verlag 2003
[12]
U. Hammerschmidt
Thermal Conductivity of a Wide Range of Alternative
Refrigerants. Measured with an Improved Guarded
Hot-Plate Apparatus
Int. J. Thermophys.16 (1995), pp. 1203-1211
[2]
Environmental Protection Agency (EPA)
“Oil pollution prevention and responses, non-transportation-related facilities, final rule”, of 30. June 2000,
40 CFR part 112,
[3]
R. Schmidt, G. Klingenberg, and M. Woydt
Thermophysical and viscometric properties of environmentally acceptable lubricants
Industrial Lubrication and Tribology, 4/2006, Vol. 58,
p. 210-224
[13]
ASTM D xxxx.yy draft „Tribological Characterization
of Piston Ring and Cylinder Liner Materials and Lubricants using the translatory oscillation apparatus
(SRV®)“, supported by ASTM D02 “L” and “B” subcommittees
[4]
R. Schmidt, M. Woydt
Viskosimetrische und thermophysikalische Eigenschaften umweltverträglicher Motorschmierstoffe
Tribologie und Schmierungstechnik Vol. 53 (2006),
No 2, pp. 16-20
[14]
M. Woydt and N. Kelling
Testing the tribological properties of lubricants and
materials for the system “piston ring / cylinder liner”
outside of engines
Industrial Lubrication and Tribology 55 (2003), No.5,
pp. 213-222
[5]
G. Buschmann, H. Clemens, M. Hoetger, and
B. Mayr
The Steam Engine – Status of Development and
Market Potential
Motortechnische Zeitschrift Vol. 62 (2001) No. 5, pp.
2-10
[15]
M. Woydt and J. Ebrecht
SRV-Testing of the Tribosystem Piston Ring and
Cylinder Liner Outside of Engines
Proc. KSTLE – 41st Autumn Conference, 25. Nov.
2005, Seoul, pp. 158-168, publisher: Korea Military
Academy.
[16]
ASTM D5706-05 “Standard Test Method for Tribological Characterization of Piston Ring and Cylinder Liner
Materials and Lubricants using SRV® Test Machine”
(2005)
[17]
M. Woydt
Review on Lubricious Oxides and Their Practical
Importance
In: Handbook of Surface Modifications and Processing: Physical & Chemical Tribological Methodologies,
edited by: G.E. Totten; Marcel Dekker, New York,
ISBN 0-9247-4872-7 (2004)
[18]
M.N. Gardos
The effect of anion vacancies on the tribological properties of rutile (TiO2-x)
Tribology Transactions 32 (1989), pp. 30-31
[19]
M.N. Gardos
The Effect of Magnéli Phases on the Tribological
Properties of Polycrystalline Rutile
Proc. 6th Int. Congress on Tribology (1993), Vol. 3,
pp 201-206
[6]
[7]
[8]
[9]
58
Spillingwerk GmbH
Werftstrasse 5, D-20457 Hamburg
Tel. +49/(0)40-789175-0, Fax +49/(0)40-7892836,
Internet www.spilling.de
M. Woydt, N. Kelling, M. Hartelt , A.Igartua,
O. Areitioaurtena, C. Seyfert, and R. Luther
Tribological performance of bio-no-tox engine oils and
new triboactive materials for piston ring/cylinder liner
systems
Proc. 15th Int. Coll. Tribology, TAE Esslingen, 17.-19.
January 2006, ISBN 3-924813-62-0
M. Woydt and A. Igartua
Ashfree and bionotox engine oils
Virtual Tribology Institute (VTI) Internet site, 2005,
www.vti-europe.org, ISBN 83-7204-449-X
“Guide to the expression of uncertainty in measurement”, International Organization for Standardization
(Geneva, Switzerland) 1995
Forschungsbericht 277
[20]
M. Woydt, J. Kadoori, H. Hausner, and K.-H. Habig
Development of engineering ceramics according to
tribological considerations (bilingual)
Cfi/Ber. DKG, Vol. 67 (1990), No. 4, pp.123-130
(Journal of the German Ceramic Society)
[30]
H. Bauer (1986)
Flüssigkeiten mit NEWTONschem Fließverhalten
In: Fließverhalten von Stoffen und Stoffgemischen,
Hüthig & Wepf Verlag, Basel, Heidelberg, New York,
Kulicke, W.-M., ed., pp 100-146, ISBN 3-85739-115-4
[21]
O. Storz, H. Gasthuber, and M. Woydt
Tribological properties of thermal sprayed Magnélitype coatings with different stoichiometries (TinO2n-1)
Surface and Coatings Technology 140 (2001), pp.
76-81
[31]
A. Schmidt
Untersuchung des Viskosität-Druck-Temperatur-Verhaltens der Bioöle B, K, N und S im Hochdruckbereich
Forschungsvereinigung Antriebstechnik, Vorhaben Nr.
265, (1998), Frankfurt a. M.
[22]
M. Woydt, N. Kelling, and M. Buchmann
Thermisch gespritzte, keramische Zylinderlaufbahnen
unter Misch-/Grenzreibung
DGM-Tagung „Reibung und Verschleiß“,
09.-11.03.2004, Fürth, in: Materialwissenschaft und
Werkstofftechnik, ISSN: 0933-5137 (print), 1521-4052
(online), Vol. 35 (2004), No.10, pp. 824-829
[32]
J. Blume
Druck-und Temperatureinfluß auf Viskosität und
Kompressibilität von flüssigen Schmierstoffen Ph.D.
Thesis, RWTH Aachen University (1987)
[33]
J. Sorab, S. Korcek, C.B. McCollum, and K. W.
Schriewer
Sequence VIB engine test for evaluation of fuel
efficiency of engine oils- Part II: Stage selection and
time factor determination
SAE Technical paper 982624
[34]
J. Igarashi
The mineral oil industry in Japan
Proc. 13th Int. Coll. Tribology, 15.-17. Jan. 2002,
Esslingen, ISBN 3-924813-48-5, Vol. I, pp. 13-17
[35]
R.I. Taylor, R.T. Dixon, F.D. Wayne, and S. Gunsel
Lubricants&Energy Efficiency: Life-Cycle Analysis
Leeds-Lyon Symposium on Tribology, September
2004
[36]
B.O. Jacobson
Rheology and elastohydrodynamic lubrication
Elsevier Science Publishers, Amsterdam, The Netherlands, ISBN 0-444-88146-8 (1991), pp. 263-288
[37]
D. A. Jones
Elastohydrodynamic Lubrication Theory
In Engine Tribology, Elsevier Science Publishers,
Amsterdam, The Netherlands, Taylor, C.M., ed, ISBN
0-444-89755-0 (1993), pp. 15-50
[38]
D. Dowson
Elastohydrodynamics, Proc. Inst. Mech. Eng., London,
Vol. 182 (3A) (1968), pp 151-167
[39]
N. Böse and H. Broeke
Reducing Measurement Uncertainty of a High Pressure Capillary Viscometer
Rheology 95 (1995), pp.190-195
[40]
DIN 53014 Part 1: Viscometry, capillary viscometers
with circular and rectangular cross section for determination of flow curves; principles, concepts, definitions
(1994)
[41]
DIN 53014 Part 2: Viscometry, capillary viscometers
with circular and rectangular cross section for determination of flow curves; systematic deviations, sources
and corrections (1994)
[23]
[24]
[25]
T. Naumann, L.-M. Berger, M. Ingwerth, and
P. Vuoristo
Titanium suboxide coatings prepared by VPS spraying
Thermal Spray 2003: Advancing the Science & Applying the Technology, C. Moreau and B. Marple (eds.),
ASM International, Materials Park, Ohio, USA (2003),
pp. 1441-1445
„Hochwertig legierter Grauguß“ (“High value alloyed
grey cast iron”), Supply Specification DBL 4401 of
Mercedes-Benz, December 2002, 4 pages
Kolbenringhandbuch (Piston ring Handbook) , April
2003, Federal Mogul Burscheid GmbH, D-51399
Burscheid (Germany)
[26]
M. Woydt and K.-H. Habig
High temperature tribology of ceramics
Tribology International, Vol. 22 (1989), No. 2, pp. 75-88
[27]
M. Woydt, N. Köhler, and N. Kelling
Magnéli-type phases for dry high-speed sliding applications up to 800 °C
World Tribology Conference, WTC2005-63236,
12.-16. September 2005, Washington D.C., USA,
In German in: Tribology&Schmierungstechnik,
Jg. 52 (2005), Heft 4, pp. 5-12 as well as in English
in: Proc. 15th Int. Coll. Tribology, TAE Esslingen,
17.-19. January 2006, ISBN 3-924813-62-0
[28]
M. Woydt
Tribomaterials for axial and radial foil bearings,
DE 10 2004 046 320.4, filed 17.09.2004
[29]
M. Woydt
Tribological characteristics of polycrystalline titaniumdioxides with planar defects
Tribology Letters, Vol. 8 (2000), No. 2-3, Special issue
„Lubricious Oxides“, pp. 117-130
59
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