Boundary Lubrication Mechanisms of Diamond

Transcription

NAGOYA UNIVERSITY
Boundary Lubrication Mechanisms of
Diamond-Like Carbon Coatings with Oil
Additives
by
Hacı Abdullah Taşdemir
A thesis submitted in partial fulfillment for the degree of
Doctor of Engineering
in the
DEPARTMENT OF MECHANICAL SCIENCE AND ENGINEERING
GRADUATE SCHOOL OF ENGINEERING
May 2014
“Run from what’s comfortable. Forget safety. Live where you fear to live. Destroy your
reputation. Be notorious... I have tried prudent planning long enough. From now on
I’ll be mad”
”I want to sing like the birds sing, not worrying about who hears or what they think”
- Mevlana Celaleddin Rumi-
This thesis is dedicated to
My lovely wife, Nur Hatice Kübra,
My parents, Ökkeş and Fatma Taşdemir,
My grandfather, Mehmet Balbaba,
And all of my family,
For their endless love, support and encouragement. . .
ii
Declaration of Authorship
I, Hacı Abdullah Taşdemir, declare that this thesis titled, ‘Boundary Lubrication Mechanisms of Diamond-Like Carbon Coatings with Oil Additives’ and the work presented
in it are my own. I confirm that:
This work was done wholly or mainly while in candidature for a research degree
at Nagoya University.
Where any part of this thesis has previously been submitted for a degree or any
other qualification at this University or any other institution, this has been clearly
stated.
Where I have consulted the published work of others, this is always clearly attributed.
Where I have quoted from the work of others, the source is always given. With
the exception of such quotations, this thesis is entirely my own work.
I have acknowledged all main sources of help.
Where the thesis is based on work done by myself jointly with others, I have made
clear exactly what was done by others and what I have contributed myself.
Signed: Hacı Abdullah Taşdemir
Date: May 2014
c
2014
Hacı Abdullah Taşdemir
iii
Dissertation Committee
The thesis of Haci Abdullah Tasdemir was reviewed and approved by the following:
Chair of Committee, Thesis Advisor:
Prof. Dr. Noritsugu umehara
Advanced Materials and Manufacturing Laboratory
Department of Mechanical Science and Engineering
Committee Member:
Prof. Dr. Kenji Fukuzawa
Micro-Nano Metrology Integrated Mechatronics Devices (Fukuzawa Lab.)
Department of Micro-Nano Systems Engineering
Committee Member:
Prof. Dr. Takashi Ishikawa
Department of Aerospace Engineering
Committee Member:
Assoc. Prof. Dr. Hiroyuki Kousaka
Reviewer:
Asst. Prof. Dr. Takayuki Tokoroyama
iv
NAGOYA UNIVERSITY
Abstract
DEPARTMENT OF MECHANICAL SCIENCE AND ENGINEERING
GRADUATE SCHOOL OF ENGINEERING
Doctor of Engineering
by Hacı Abdullah Taşdemir
Huge amount of money and energy have been lost in the world due to the friction
and wear in mechanical components. Diamond-like carbon coatings have desirable mechanical and tribological properties for many industrial applications like hard hardness,
chemical inertness, low friction and high wear resistance. These coatings can be prepared by various deposition techniques. Mechanical and chemical properties of DLC
coating strongly depend on the coating methods, hydrogen content, hybridization of
carbon and dopant elements. Besides, tribological properties of DLC coating are significantly affected by extrinsic factors or test conditions such as humidity, temperature,
surrounding environment and counter material. The excellent mechanical and tribological properties of DLC coatings make them promising candidate for engine components
to control friction and wear in passenger cars. However, most of the engine components
need to work under lubricated conditions and commercially available engine oils are formulated for ferrous surfaces. Therefore, interaction of DLC surfaces with the oils and
the lubricant additives is not yet fully understood. On the other hand, it is known that
engine components operate in a range of temperatures and it is an imported parameter
for tribological properties of DLC, lubricants and lubricant additives.
The aim of this study is to clarify the ultra-low friction and wear mechanism of ta-C DLC
under boundary lubricated conditions by testing in synthetic base oil poly alpha-olefin
(PAO4), PAO+GMO, PAO+ZnDTP and PAO+GMO+ZnDTP. Besides, the role of
temperature, additive concentrations and counter-material on the ultra-low friction and
wear of non-hydrogenated ta-C DLC coating will be analyzed. Additionally, in order
to better understanding of tribological properties and interactions between the DLC
surfaces and oil additives, a wide range of DLC coating will be tested and compared
with self-mated DLC/DLC contacts under same conditions in PAO and PAO+ZnDTP
oils.
vi
Tribological tests were performed in a pin-on-disc tribometer. Atomic Force Microscopy
(AFM), Field Emission Scanning Electron Microscopy (FESEM), Nano-indenter, Xray Photoelectron Spectroscopy (XPS), Raman spectroscopy and scanning white light
interferometry were used for characterization of ta-C DLC and worn surface analysis.
The results exhibit that ta-C give ultra-low friction in pure PAO for DLC/steel and
DLC/DLC tribopair. GMO additivated PAO provide smooth run-in period for transition
to ultra-low friction regime and also enhance the durability of coating. ZnDTP behave
differently depending on the presence of ferrous surfaces on the contact. It forms padlike wear protective tribofilm both on ta-C and steel surfaces for DLC/steel contact,
while it form thin white layer on ta-C surfaces for DLC/DLC contact.
The results show that ta-C DLC were totally worn out in DLC/steel contact tested
in base oil. The ta-C DLC exhibited totally different wear behavior in DLC/steel and
DLC/DLC contact depending on lubricant formulation. The wear performance of ta-C
DLC was found to have a clear dependence on combination of lubricant formulation,
concentration of the lubricant additives and counterbody material. Using GMO and
ZnDTP together does not show any synergistic correlation for steel/steel, DLC/steel and
DLC/DLC combinations. The results obtained at high temperature show the significant
and beneficial influence of oil additives on the wear performance of the coating. The
results revealed that ta-C coated pin was experienced less wear, up to one order of
magnitude, when rubbed against self-mated ta-C DLC and germanium disc compared
to steel counterpart in base oil. It is explained that observed high wear against steel
disc is due to thermally promoted tribo-chemical wear and rubbing against self-mated
ta-C DLC and germanium eliminated this phenomenon.
The effects of hydrogen, doping elements and surface morphology on reactivity of DLC
coatings have been studied in terms of ZnDTP tribofilm formation and tribological performance of DLC coatings under boundary lubrication conditions. Six types of DLC
coatings were tested: one non-hydrogenated amorphous carbon (a-C) coating, one nonhydrogenated tetrahedral amorphous carbon (ta-C) coating, two hydrogenated amorphous carbon (a-C:H) coatings, one silicon-doped hydrogenated amorphous carbon (SiDLC) coating, and one chromium-doped hydrogenated amorphous carbon (Cr-DLC)
coating. The results confirmed the ZnDTP derived pad-like or patchy tribofilm formation on the surfaces depending on kinds of DLC coatings. It is observed that hydrogen
content and doping elements increased the pad-like tribofilm formation ability of DLC
coatings. Doped DLC coatings exhibited better wear resistance than nondoped DLC
coatings. Addition of ZnDTP additives in to the base oil significantly improved the
wear resistance of hydrogenated DLC , silicon-doped hydrogenated DLC and chromiumdoped hydrogenated DLC. Hydrogen-free tetrahedral amorphous DLC coating provided
the lowest friction coefficient both in PAO (poly-alpha-olefin) and PAO +ZnDTP oils.
Acknowledgements
I would never have been able to finish my thesis without the guidance of my supervisor,
committee members, help from laboratory friends, and support from my family and wife.
First of all, I would like to express my sincere gratitude and respect to my supervisor,
Prof. Dr. Noritsugu Umehara, for all I have learned from him and for his excellent
guidance, continuous help, patience, motivation and support in all stages of this thesis.
I would also like to thank him for being an open person to ideas, and for encouraging
and helping me to shape my interest and ideas. Without his guidance and persistent
help this dissertation would not have been possible.
I would like to thank my committee members, Prof. Dr Kenji Fukuzawa, Prof. Dr.
Takashi Ishikawa and Assoc. Prof. Dr. Hiroyuki Kousaka for sharing their ideas on
the improvement of the thesis, and for their encouragement, insightful comments, and
suggestions.
The success of this thesis is attributed to the extensive support and assistance from
all members of Advanced Material and Manufacturing Laboratory. Especially, I would
like to express my grateful gratitude and sincere appreciation to Assoc. Prof. Dr.
Hiroyuki Kousaka, Asst. Prof. Dr. Takayuki Tokoroyama and Mr. Shinkoh Senda for
their kindness in examining the research work and providing technical suggestions for
improvement and encouragement during my time here. I am indebted to Mr. Yutaka
Mabuchi, Nissan Motor Co., Ltd., for his kind support and help in my work.
My special thanks are given to my labmates for their cheerful cooperation, encouragement, help and support and enjoyment during the course of my study, Mr. Azmmi, Mr.
Deng, Mr. Kawara, Mr. Nishimura, Mr. Inoue and all others who I can’t write their
name in here. I hope our friendship will remain same in coming years.
I would also like to offer my special thanks to my family, especially my mother and father
for always believing in me, for their continuous love and their supports in my decisions.
My special thanks are extended to my dearest sisters, Gülizar Bayraktar, Yasemin Sayar,
Ebru Balbaba and their family. I am particularly grateful for the support, love and
good times given by all my relatives and friend. They were always there cheering me
up, praying for my health and stood by me through the good times and bad. Without
whom I could not have made it here.
Above all, I owe it all to Almighty ALLAH (subhana wa taala), creator and sustainer of
the universe, for endowing me with wisdom, health, patience, and knowledge in exploring
things to pursue this dissertation.
vii
Contents
Declaration of Authorship
iii
Dissertation Committee
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Abstract
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Acknowledgements
vii
List of Figures
x
List of Tables
xiii
Symbols
xiv
1 Introduction
1.1 Tribology and industrial needs of DLC coatings .
1.2 Lubrication Theory . . . . . . . . . . . . . . . . . .
1.2.1 Hydrodynamic Lubrication . . . . . . . . .
1.2.2 Elastohydrodynamic Lubrication . . . . . .
1.2.3 Boundary Lubrication . . . . . . . . . . . .
1.3 Diamond-Like Carbon Coatings . . . . . . . . . . .
1.3.1 Oil Boundary Lubrication of DLC coatings
1.4 Purpose of This Study . . . . . . . . . . . . . . . .
1.5 Outline of Dissertation . . . . . . . . . . . . . . . .
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Ultra-low friction of ta-C DLC under boundary lubrication
2.1 Experimental details . . . . . . . . . . . . . . . . . . . . . . . .
2.1.1 Material characterization and lubricants . . . . . . . . .
2.1.2 Tribological experiments . . . . . . . . . . . . . . . . . .
2.1.3 Surface analysis . . . . . . . . . . . . . . . . . . . . . . .
2.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
2.2.1 Steel-Steel tribopair at 80◦ C . . . . . . . . . . . . . . .
2.2.2 DLC-Steel tribopair at 80◦ C . . . . . . . . . . . . . . .
2.2.3 DLC-DLC tribopair at 80◦ C . . . . . . . . . . . . . . .
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Contents
ix
2.2.4
2.3
2.4
Effect of oil temperature on the friction coefficinent for DLC/ steel
contact . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
2.2.5 Effect of additive concentration on the friction coefficinent for DLC/steel contact . . . . . . . . . . . . . . . . . . . . . . . . . . . .
Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3 Wear behaviour of ta-C DLC under boundary lubrication
3.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3.3 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3.3.1 Wear results at 80◦ C . . . . . . . . . . . . . . . . . . . .
3.3.2 Wear behavior depending on counter-body material and
tives at 80◦ C . . . . . . . . . . . . . . . . . . . . . . . .
3.3.2.1 DLC against steel contact . . . . . . . . . . . .
3.3.2.2 DLC against DLC contact . . . . . . . . . . .
3.3.3 Surface analysis on worn surfaces tested at 80◦ C . . . .
3.3.4 Effect of oil temperature on wear of ta-C . . . . . . . .
3.3.5 Effect of additive concentration . . . . . . . . . . . . . .
3.3.6 DLC v.s. Germanium disc . . . . . . . . . . . . . . . . .
3.4 Discussions on the wear mechanism of ta-C coating . . . . . . .
3.5 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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4 Boundary Lubrication of self-mated DLC/DLC contacts in synthetic
base oil and influence of ZnDTP tribofilm formation
4.1 Introduction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.2.1 Material characterization and Lubricants . . . . . . . . . . . . . .
4.2.2 Tribological experiments . . . . . . . . . . . . . . . . . . . . . . . .
4.2.3 Surface analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.3 Results and discussions . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.3.1 Coatings durability and ZnDTP derived tribofilm formation . . .
4.3.2 Wear results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.3.3 Friction results . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.3.4 Surface analysis with Raman and XPS spectroscopy . . . . . . . .
4.4 Conclusions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
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5 Conclusions and future outlook
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Bibliography
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oil addi. . . . .
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Publication List
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International Conferences
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List of Figures
1.1
1.2
1.3
1.4
1.5
1.6
1.7
Striberk Curve and Lubrication Regimes . . . . . . . . . . . . . . . . . .
Additives for better lubrication performance . . . . . . . . . . . . . . .
Caption without citation that appears in the List of Figures or Tables .
Structure od DLC coatings and Ternary phase diagram of amorphous
carbons . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
DLC coated automotive components . . . . . . . . . . . . . . . . . . . .
Structure of (a) GMO and (b) ZnDTP additives . . . . . . . . . . . . .
Outline of dissertation . . . . . . . . . . . . . . . . . . . . . . . . . . . .
2.1
2.2
2.3
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Raman Spectra of ta-C DLC coated pin and disk . . . . . . . . . . . . . . 20
Pin on disc type tribotester . . . . . . . . . . . . . . . . . . . . . . . . . . 21
Representative friction coefficients as a number of sliding cycles between
SUJ2 steel pin against SUJ2 steel disc in PAO, PAO+GMO, PAO+ZnDTP
and PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . . . 23
2.4 Fe-SEM images of steel pin and steel disc surfaces after rubbing 10700
cycles in steel/steel tribopair lubricated with a) PAO+ZnDTP and b)
PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24
2.5 Representative friction coefficients as a number of sliding cycles between
ta-DLC pin against SUJ2 steel disc in PAO, PAO+GMO, PAO+ZnDTP
and PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
2.6 Four individual tests results in PAO+GMO and PAO+ZnDTP for DLC/steel contact . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25
2.7 Fe-SEM images of DLC and Steel surfaces after rubbing 10700 cycles in
DLC/Steel tribopair lubricated with a) PAO, b) PAO+GMO, c) PAO+ZnDTP,
and d) PAO+GMO+ZnDTP. The arrows indicate sliding directions. . . . 26
2.8 Representative friction coefficients as a number of sliding cycles between
ta-DLC pin against ta-DLC disc in PAO, PAO+GMO, PAO+ZnDTP and
PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28
2.9 Comparison of steady-state friction of steel/steel, DLC/steel and DLC/DLC contacts for four different oil combinations after running 10700
cycles. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29
2.10 Fe-SEM images of DLC pin surfaces after rubbing 10700 cycles in DLC/DLC tribopair lubricated with a) PAO, b) PAO+GMO, c) PAO+ZnDTP,
and d) PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . 30
2.11 Generated wear grooves on DLC pin . . . . . . . . . . . . . . . . . . . . . 30
2.12 Fe-SEM and AFM images of DLC disc surfaces after rubbing 10700 cycles
in DLC/DLC tribopair lubricated with PAO+ZnDTP; a) washed with
acetone in ultrasonic bath b) rinsed with benzene and acetone not washed
in ultrasonic bath c) AFM topography d) AFM lateral force . . . . . . . 31
x
List of Figures
2.13 Variation of friction coefficients as a function of temperature using Group
A lubricants. The curves are for guidance only . . . . . . . . . . . . . .
2.14 Effect of additive concentration on the friction coefficient of ta-C DLC. .
2.15 Raman spectra of ta-C DLC pin rubbing in pure PAO for DLC/steel
tribopair a) before total wear occur b) after partially wear out of coating
from the topmost surface . . . . . . . . . . . . . . . . . . . . . . . . . .
2.16 Raman spectra of as-deposited ta-C DLC pin and after rubbing 10700
cycles in PAO, PAO+GMO, PAO+ZnDTP and PAO+GMO+ZnDTP for
DLC/DLC tribopair . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
2.17 Caption without citation that appears in the List of Figures or Tables .
2.18 XPS spectra for C1s state on the worn surface of ta-C DLC coating tested
in PAO+GMO compared with as deposited ta-C surface (a) and deconvolution of XPS C1s peak . . . . . . . . . . . . . . . . . . . . . . . . . .
3.1
3.2
3.3
3.4
3.5
3.6
3.7
3.8
3.9
3.10
3.11
3.12
3.13
3.14
3.15
3.16
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Measurement of the width of the wear scar on pin speciments using Zygo,
Newview. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46
Total wear rate of pin for steel/steel, DLC/steel and DLC/DLC contacts
tested at 80◦ C with different lubricants after 405 m sliding distance .
(steel/steel contacts are given for comparison) . . . . . . . . . . . . . . . . 47
Wear scar on ta-C DLC pins at DLC/Steel contact after sliding in different
lubricants tested at 80◦ C . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48
Wear scar on ta-C DLC pins at DLC/DLC contact after sliding in different
lubricants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48
Steady state friction coefficients for steel/steel, DLC/steel and DLC/DLC
X : ta-C pin exhibited ultra-low friccontacts as a function of lubricants [X
tion of 0.025 for DLC/steel in PAO before total wear and then friction
coefficient jumped to 0.09 after total wear of coating] . . . . . . . . . . . . 49
Wear volume loss of ta-C DLC pin vs. sliding distance for DLC/steel
contact at 80◦ C with different lubricants . . . . . . . . . . . . . . . . . . 50
Wear scar on ta-C DLC pin depending on sliding distance at DLC/steel
contact after sliding in PAO . . . . . . . . . . . . . . . . . . . . . . . . . . 51
Deep scratch lines paralel to sliding direction on steel disc (a) optical
microscopy and (b) AFM images . . . . . . . . . . . . . . . . . . . . . . . 51
Wear volume loss of ta-C DLC pin vs. sliding distance for DLC/DLC
contact at different lubricants . . . . . . . . . . . . . . . . . . . . . . . . 52
Deep abrasive scratches on ta-C pin generated in all lubricants at DLC/DLC contacts (a) optical microscopy and (b) FESEM images . . . . . 53
FESEM images of the DLC disc surface tested with PAO . . . . . . . . . 53
Raman spectra with excitation wavelength for ta-C DLC before and after
rubbing in PAO . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55
FeSEM micrograph (a), AFM topography (b) and AFM lateral force (c)
images of steel disc tested in PAO+GMO+ZnDTP . . . . . . . . . . . . . 56
ZnDTP derived tribofilm formation on ta-C DLC surfaces tested in (a)PAO+ZnDTP
and (b) PAO+GMO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . 56
XPS spectra of P 2p (with Zn 3s), S 2p, O1s and Zn 2p peaks obtained
from (a) PAO+ZnDTP and (b) PAO+GMO+ZnDTP lubricated ta-C
DLC surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57
Wear rate of ta-C coated pin for DLC/steel contact as a function of temperature. The curves are for guidance only . . . . . . . . . . . . . . . . . 58
List of Figures
3.17 Microscopic images of the wear scar on the ta-C coated pin when tested
in pure PAO at (a) 50 ◦ C and (b) 80 ◦ C . . . . . . . . . . . . . . . . . .
3.18 Effect of additive concentration on the wear rate of ta-C DLC. . . . . .
3.19 Friction coefficients of ta-C DLC when rubbed against Germanium as a
function of temperature in PAO oil. . . . . . . . . . . . . . . . . . . . .
3.20 Wear rate comparison of ta-C DLC when rubbed against Germanium and
Steel discs as a function of temperature in PAO oil. The curves are for
guidance only . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
3.21 Hypothesis on the wear mechanism of ta-C DLC in pure PAO lubrication:
(a) Abrasive wear by generated wear particles, (b) Tribo-chemical wear by
the degraded oil molecules, and (c) Tribo-chemical wear by graphitization
and following carbon diffusion . . . . . . . . . . . . . . . . . . . . . . . .
4.1
4.2
xii
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Schematic of tribotester and pin-on-disc configuration . . . . . . . . . . .
Worn surfaces of DLC coated pins after rubbing in PAO and PAO+ZnDTP
oils: (a) a-C , (b) a-C:H 1 and (c) a-C:H 2 (arrows show sliding direction)
4.3 FESEM images of ZnDTP derived tribofilm formation on various DLC
surfaces . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4.4 AFM topography and lateral force images of ZnDTP derived tribofilm
formation on various DLC surfaces . . . . . . . . . . . . . . . . . . . . . .
4.5 Wear rate of the all DLC coated pins for self-mated DLC/DLC contacts .
4.6 Friction coefficient curves with sliding cycles for self-mated DLC/DLC
contacts (a) in PAO and (b) in PAO+ZnDTP . . . . . . . . . . . . . . . .
4.7 Steady state friction coefficients of self-mated DLC/DLC contacts for
PAO and PAO+ZnDTP oils.pdf . . . . . . . . . . . . . . . . . . . . . . .
4.8 Raman spectra of (a) a-C:H 2 coating and (b) ta-C coating scanned before
and after sliding in base oil. . . . . . . . . . . . . . . . . . . . . . . . . . .
4.9 XPS P 2p (with Zn 3s), S 2p and Zn 2p peaks recorded in the tribofilm
formed on tested DLC coatings . . . . . . . . . . . . . . . . . . . . . . . .
4.10 Detailed XPS spectra of oxygen 1s of tested DLC coatings after sliding
in PAO+ZnDTP . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
71
74
75
76
78
79
80
82
84
85
List of Tables
2.1
2.2
2.3
2.4
4.1
4.2
4.3
Properties of ta-C DLC coated disc and pin . . . . . . . . . . . . . . . . .
Boundary lubricant components and additive composition. . . . . . . . . .
EDS measurements on boxed area of DLC pin surfaces in Fig. 2.7 lubricated with ZnDTP and GMO+ZnDTP containing oil for DLC/Steel
tribopairs . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
EDS measurement on the worn surface of DLC disc lubricated with ZnDTP
containing oil for DLC/DLC tribopairs . . . . . . . . . . . . . . . . . . .
19
20
27
31
Important characteristic properties of DLC coatings . . . . . . . . . . . . 69
Lubricant components and additive composition. . . . . . . . . . . . . . . 70
Binding energies and concentrations of XPS peaks recorded in the tribofilm formed on tested DLC surfaces. . . . . . . . . . . . . . . . . . . . 86
xiii
Symbols
Λ
Dimensionless film parameter
hmin
Minimum film thickness
(m)
hm
fillm thickness where dp/dx = 0
(m)
Rq,a
Root-mean-square (rms) surface roughness of solid a
(m)
Rq,b
Root-mean-square (rms) surface roughness of solid b
(m)
ρ
Density
(kg/m3 )
ρ0
Density when p = 0
(kg/m3 )
h
Film thickness
(m)
p
Pressure
(N/m2 )
η
Absolute viscosity
(N.s/m2 )
η0
Absolute viscosity at p=0 and at constant temperature
(N.s/m2 )
Cartesian coordinate in direction of motion
(m)
y
Cartesian coordinate
(m)
um
mean surface velocity in x direction
(ua + ub )/2, m/s)
vm
mean surface velocity in y direction
(va + vb )/2, m/s)
t
time
s
R0
Effective radii of contact
(m)
E0
Effective elastic modulus
(Pa)
U
Dimensionless speed
α
Pressure-viscosity coefficient
W
Dimensionless normal load
k
ellipticity parameter
p
Pressure
(N/m2 )
V
Wear volume loss
(m3 )
k
Wear rate
(m3 /N.m)
x
xiv
(m2 /N )
Symbols
xv
F
Normal load
(N)
s
Sliding distance
(m)
sp1 , sp2 orsp3
Hybridisation of carbon atoms
HL
Hydrodaynamic Lubrication
EHL
Elastohydrodaynamic Lubrication
T EHL
Thermo-elastohydrodaynamic Lubrication
BL
Boundary Lubrication
FM
Friction Modifier
AW
Anti-wear additive
DLC
Diamond-Like Carbon
CV D
Chemical vapor deposition
PV D
Physical vapor deposition
M oDT C
Molybdenum dithiocarbamate, a kind of FM
GM O
Glycerol mono-oleate, a kind of organic FM
ZnDT P/ZDDP
Zinc dialkydithiophospahte, commonly used AW addtive
Chapter 1
Introduction
1.1
Tribology and industrial needs of DLC coatings
Tribology is defined as the science, technology and engineering of interacting surfaces
in relative motion which includes the fields of friction, lubrication and wear. The study
of tribology plays an important role in many aspects of industry and daily life since the
successful application of many mechanical, electrical, electromechanical and biological
systems depends on the appropriate tribology knowledge. Some examples of tribological
applications include bearings, gears, clutches, artificial hip implants and prostheses, micro/nanoelectromechanical systems (MEMS/NEMS), lubricant formulations, hydrophobic coatings and wind turbines [1]. The application of tribology is a rapidly growing
field and has attracted numerous scientist from physics, chemistry, material science and
mechanical engineering.
The general purpose of tribological study is to control friction and wear to desired
levels. Friction is a force that resists the motion whenever two solid surfaces touch and
slide against each other. It is the major source of energy dissipation. One-third of the
world’s energy resources is used to overcome friction in mechanical applications [2, 3].
Wear is described as removal or loss of materials which occurs at the interface between
interacting bodies. It is a principle cause of material wastage and loss of mechanical
performance [4]. Hence, controlling the friction and wear contributes the big saving of
energy, materials and maintenance cost. It should be noted that friction and wear are
not only material properties but also responses of tribo-systems. They are sensitive to
various system parameters such as temperature, counter-material, environmental gases
and contact geometry. Potential actions to control friction and wear in interaction
surfaces include liquid/gas lubrication, application of surface coatings, modification of
surface topography and altering surface structure.
1
Introduction
2
The automotive and oil industries are two main driving force behind tribological research. These two industries are facing tough international competition, government
regulations and rapid technological developments. The modern cars should be powerful,
fuel efficient and comfortable as well as environmentally friendly. Reducing the friction
and wear in mechanical components of passenger cars with environmentally friendly
materials could produce tremendous cost saving and reduce the release of hazardous
chemicals. In recent years, many automotive companies are extensively seeking out new
environmentally friendly materials and lubricants for reduced carbon emission, increased
fuel efficiency and improved durability of the powertrain components in their passenger
car [5, 6]. Accordingly, Diamond-Like Carbon (DLC) coatings are becoming attractive
protective films as they offer high hardness, ultra-low friction and good wear resistance
under dry or lubricated contacts [7]. With the proper design and applications, they can
exhibits ultra-low friction and low wear rate with the biodegradable base oil and organic
friction modifier which provide environmentally friendly lubrication by eliminating use
of harmful lubricant additives in engine oil [8].
Excellent mechanical and tribological properties of DLC coatings under dry, lubricant
free conditions have been known for several decades [9–11]. In the last 10 years, DLC
coatings have been applied to several components of engines and power trains in automobiles which work under boundary lubricated conditions [12, 13]. On the other hand,
commercially available fully formulated oils are designed for ferrous surfaces and contain
numerous additives to enhance performance of base oil. Therefore, although these coatings have exhibited excellent properties, more research in their tribological performance
and their tribo-chemical interactions with conventional lubricant and lubricant additives
are still needed for successful operation of these coatings under lubricated conditions.
This study is focused on the boundary lubrication of DLC coatings and analysis of
DLC surfaces. In this first chapter, I will attempt to cover fundamental descriptions for
this dissertation. I will give brief descriptions of lubrication theory and DLC coatings
as well as literature survey on oil lubrication of DLC coatings. Following these brief
descriptions, the purpose of this dissertation and outline will be explained.
Introduction
1.2
3
Lubrication Theory
Lubrication is essential for the modern machinery to increase the efficiency and longterm performance of moving parts operating under different conditions of speed, pressure
and temperature. The main functions of lubrication are friction reduction, wear protection, cooling, corrosion inhibition and contaminants removal. Lubricant materials can
be classified in two group; solid lubricants and fluid lubricants. Solid lubricants have
inherent self-lubricating capability such as graphite, molybdenum disulfide, PTFE etc.
and they can be applied to contacting surfaces directly where liquid lubrication is not
possible. Fluid lubricants can be any type of fluid like oils, greases, gases and water.
Fluid lubricants can be dragged into the loaded contact due to the rotation and pressure generation between the surfaces in contact. In classic lubrication theory, depending
on the lubricant viscosity, velocity and normal load, three types of fluid film lubrication
can be defined: hydrodynamic lubrication (HL), elastohydrodynamic lubrication (EHL),
boundary lubrication (BL).
Figure 1.1: Striberk Curve and Lubrication Regimes
The variation of friction coefficient and transition between lubrication regimes can be
graphically illustrated by use of Striberck curve (Fig. 1.1) [14]. The horizontal axis show
a dimensionless parameter (Λ) that combines minimum film thickness and roughness of
surfaces (eq. 1.1). The vertical axis on the right side is the film thickness, while the
vertical axis on left side is the friction coefficient. The thickness of the fluid film between
two solid surfaces determines the lubrication regime.
Introduction
4
Λ= q
hmin
(1.1)
2 + R2
Rq,a
q,b
The generalized Reynolds equation that governs the pressure distribution in thin lubricant film is the fundamental equation in fluid film lubrication theory (eq. 1.2). It
can be employed in the analysis of any fluid-film bearings design. It is a second order
partial differential equation which was derived from Navier–Stokes equations by Osborn
Reynolds. This equation don’t have analytical solution. However, with the acceptable
assumptions, it can be reduced to any forms which can be solved either analytically or
numerically.
∂
∂x
1.2.1
ρh3 ∂p
12η ∂x
+
∂
∂y
ρh3 ∂p
12η ∂y
= um
∂(ρh)
∂x
+ vm
∂(ρh)
∂y
+
∂(ρh)
∂t
(1.2)
Hydrodynamic Lubrication
HL is the ideal state of lubrication. As the relative speed of two surfaces and viscosity
of lubricant increase, or the bearing load decreases, a hydrodynamic pressure is build
up and the load bearing surfaces are completely separated by a relatively thick lubricant film. Since two surfaces do not contact directly, the bearing load is supported
by pressure generated in the lubricant films and shearing occurs within the lubricant.
Accordingly, lubricant properties and geometry of contact determine the tribological
characteristics. In the HL regime, lubricant properties such as viscosity and density
don’t change throughout the contact zone and can be accepted as a constant. Also, if
the motion is pure sliding , the vm term become zero. Thus, the Reynolds equation can
be reduced in form
∂
∂x
3 ∂p
h
∂x
∂
+
∂y
3 ∂p
h
∂y
= 12um η0
∂h
∂x
(1.3)
this form of Reynolds equation also do not have analytical solution, but it can be easily
solved numerically. If the sliding motion takes place in x direction, the flow of lubricant
in the y direction is known as side-leakage. If the bearing is long enough, side-leakage
can be neglected. Now, the Reynolds equation can be further reduced the form
Introduction
5
dp
h − hm
= 12um η0
dx
h3
(1.4)
This is integrated form of the two-dimensional Reynolds equation. This equation has
analytical solutions and can be applied the simplest bearing design problems. HL theory
is well developed and almost all journal and thrust bearings design can be made based
on this theory. Detailed analysis of HL theory and different application of Reynolds
equation can be found elsewhere [14, 15]. As can be seen in Fig. 1.1, the thinner the
fluid film thickness, the lower the friction coefficient in HL regime. Low viscosity oils
are preferable for the low friction performance in HL. Lastly, since there is no actual
surface contact, wear hardly takes place in this lubrication regime.
1.2.2
Elastohydrodynamic Lubrication
This type of lubrication regime can take place in many critical, heavily loaded contacts
such as ball or rolling element bearings, gears, cam and followers. It is also governing
lubrication regime for soft bearing systems such as seals and synovial joints [16]. In EHL
regime, the load is sufficiently high enough that the contact zone elastically deform.
Also, liquid lubricant in contact zone is subjected to high pressure that the viscosity of
lubricant can increase by several orders of magnitude due to pressure effect. With the
further increase of bearing load, some contact between asperities of both surface takes
place. This type of EHL lubrication is called partial or mixed lubrication (ML).
The main variables in EHL regimes are pressure distribution and film thickness. For
the calculation of this variables, the Reynolds equation must be solved simultaneously
with the equation of elasticity, the pressure- viscosity equation and the pressure-density
equations [17]. Surface deformation in EHL contact can be estimated by Hertz’s equations of elastic deformation, Hertzian contacts. The Roelands equation is generally used
for the pressure- viscosity equation
η = exp (lnη0 + 9, 67) −1 + (1 + 5.1x10−9 p)z
(1.5)
where z is a material parameter. The Dowson and Higginson formula can be used for
pressure-density relation
Introduction
6
ρ = ρ0 1 +
0.6x10−3 p
1 + 1.7x10−3 p
(1.6)
Simultaneous solutions of all this equations are very complex and it is only possible with
numerical methods. To solve the EHL equations several different numerical methods
have been developed in the literature [18]. Temperature is also important parameter
in EHL regimes. Therefore, the energy equation of the lubricant film should be included to the solution of Reynold equation. This type of EHL regime is called Thermoelastohydrodynamic lubrication (TEHD). Reliable TEHD models are also developed in
the literature [19].
The EHL theory is also well developed as like HL theory. Today, engineers are using
empirical formulas for the calculation of pressure distribution and minimum film thickness in EHL contacts which are derived from the numerical models by curve fitting. For
example, one commonly used, reliable empirical formula for minimum film thickness was
proposed by Dowson and Hamrock
hmin
= 3.63
R0
U η0
E 0 R0
0.68
αE
0 0.49
W
E 0 R02
−0.073
(1 − e−0.68k )
(1.7)
where R0 is reduced radius of curvature, U is entraining surface velocity, W is normal
load, E0 is reduced Young’s modulus, η0 is dynamic viscosity, α is the pressure-viscosity
coefficient.
The EHL theory has reached a high level of complexity and is well consistent with
experimental study. It is capable of predicting the fluid film pressure supporting the load,
surface heating, lubricant film thickness, rolling contact fatigue and wear of components.
Research progress in EHL theory is focused on partial lubrication, starved lubrication
and modeling of actual surface roughness [20, 21].
1.2.3
Boundary Lubrication
Boundary lubrication occurs whenever HL and EHL lubrication fails due to the high
bearing loads, low operating speed and low viscosity of lubricant. Under BL conditions,
most of lubricating film is squeezed out between the solid surfaces, and solid-to-solid
contact occurs. Since the fluid film formation is not possible, the Reynolds equation and
related EHL theory is not valid anymore. In contrast to HL and EHL regimes, the bearing load is primarily supported by contacting surfaces and the shearing occurs within
Introduction
7
contact interface. Thus, with the severity of operating conditions in BL regimes, high
friction and wear become inevitable which cause seizure and failure of contact surfaces.
Engineers apply special surface modification techniques such as lubricant additives, surface texturing, use of solid lubricant coatings for better controlling friction and wear
under boundary lubricated rolling, sliding or rotating contact conditions.
Lubricants are blended with approximately 10% of additives package in order to enhance
their lubricating capabilities under BL and ML regimes [22]. Certain additives are used
to impart special properties to the oil such as antioxidants, viscosity index improvers,
detergents, corrosion inhibitors, friction modifier (FM), demulsifiers/emulsifiers, antiwear (AW) and extreme pressure additives (Fig. 1.2). Antioxidants are used in engine
lubricants to prevent oxidation and oil degradation. Antiwear and extreme pressure
additives are added to lubricating oil to prevent welding of moving parts, high wear
rate, and seizure under high pressure conditions. As temperature increases, viscosity
decreases, and vice versa. Viscosity index improvers allow the lubricants to attain different viscosity depending on temperature to achieve optimum film thickness. Friction
modifiers are added to lubricants to meet requirements for reduced friction coefficient,
smooth transition from static to dynamic condition at start-up as well as reduced noise,
frictional heat and start-up torque.
Figure 1.2: Additives for better lubrication performance
Introduction
8
Lubricant and lubricant additives interact with contacting surfaces and form boundary
lubrication films by several mechanisms [23]. For example, such friction modifiers have
a polar end which attaches to the surfaces and a non-polar end which points out into oil
solution (Fig.1.3a). This type of additives physically or chemically adsorb to surfaces
which prevents direct contact and allows motion without high friction and wear. Viscosity improvers form globular or continuous thick layers by the reaction of oil components
in the presence of rubbed surfaces which gives a hydrodynamic effect (Fig.1.3b). Thin
reacted layers are formed by chemical reaction of additives with clean metal surfaces
that provide low shear strength, reduced adhesion and ploughing (Fig.1.3c). These previous three films cannot be expected to provide effective lubrication under high contact
loads and high temperature conditions. On the other hands, thick reacted inorganic
layers are formed only at high temperature and loads (Fig.1.3d). These films tend to
be low-shear-strength friction films or semi-plastic deposit anti-wear films depending on
their individual structure.
Figure 1.3: Boundary lubricating films on contacting surfaces by applying lubricant
and additives: (a) a layer of molecules adsorption, (b) High viscosity layer formation,
(c) Thin reacted layer and smoothing, (d) Thick reacted inorganic layer [23]
Formulation of well balanced and optimized additive systems is a real challenge to minimize friction whilst controlling wear at the same time. Temperature, additive concentration, nature of contacting surfaces, time and operating environment have huge influence
on the boundary film formation and the effectiveness of each particular additive. Also
there is a possibility that some additives may interfere each other negatively.
Introduction
1.3
9
Diamond-Like Carbon Coatings
Carbon is one of the most naturally abundant element in the Earth. It is able to form
chemicals bonds with number of different elements that can be found in many inorganic
and organic materials. It is also able to form single, double and triple bonds between
its atoms creating several allotropes such as diamond, graphite, graphene, fullerenes,
buckyballs, carbon nanotubes and amorphous carbons [24]. Graphite has the most
stable structure with sp2 hybridization within the carbon allotropes. Its other alloprotes
are in metastable state and can transform quickly to graphitic structure under certain
conditions. Carbon-based materials have unique properties depending on the allotropic
form like high harness (diamond), softness and lubricity (graphite) and high thermal
conductivity (graphene and carbon nanotube) [25].
Diamond-like carbon (DLC) coatings are metastable form of amorphous carbon that
can be deposited by various advanced chemical vapor deposition (CVD) and physical
vapor deposition (PVD) techniques [10, 26, 27]. These coatings consist of mixture of
sp2 and sp3 hybridized carbon and can be alloyed with certain elements such as hydrogen, nitrogen, silicon, titanium, boron, fluorine to improve their properties [28]. DLC
coatings can be accepted as solid lubricant and have been used widely in many areas
as surface coating due to their promising mechanical and tribological properties such as
high hardness, chemical inertness, low coefficient of friction and high wear resistance [7].
Figure 1.4 shows the structure of DLC coatings and a thernary phase diagram of the
sp2 , sp3 and hydrogen contents of various DLC coatings.
The various types of DLC coatings have shown unique tribological properties depending
on coating methods, structural and chemical nature, dopant elements, humidity, temperature, working conditions, surrounding environment and substrate materials [11, 29, 30].
Hydrogen-free tetrahedral amorphous carbon (ta-C) and amorphous carbon (a-C) DLC
coatings exhibit low friction coefficient in humid environments, but wear rate of a-C DLC
is higher than ta-C DLC [31, 32]. Hydrogenated (a-C:H) DLC coatings show super-low
friction in dry and inert environments, but the presence of dopants in hydrogenated DLC
coatings greatly affect the tribological performance [33–35]. Low friction performance
of DLC coatings have been attributed either the surface passivation by -H, -OH, water
vapor and oxygen or the transformation of top most surfaces to graphitic structure [11].
Detailed overview of deposition methods, characterization, mechanical and tribological
properties for DLC coatings can be found in the books edited by C. Donnet and A.
Erdemir [36].
Introduction
10
Figure 1.4: Structure od DLC coatings and Ternary phase diagram of amorphous
carbons
1.3.1
Oil Boundary Lubrication of DLC coatings
Several components of engines and power trains in automobiles such as valve train, piston ring assembly and transmission clutch operate under severe boundary lubrication
condition which result in high friction and wear losses. Almost 30% of total energy
generated in an engine is lost in these components due to high friction coefficient. Automotive industries apply special surface technologies on these components to control the
friction and wear for improved fuel efficiency, durability and environmental concerns. In
recent years, application of hard DLC coatings on the surfaces of these components are
becoming one of the potential solutions (Fig. 1.5). However, the use of DLC coating
Introduction
11
in these tribo-components generate compatibility concern with existing lubricant and
lubricant additives as they originally developed for uncoated metallic surfaces.
Figure 1.5: DLC coated automotive components
Systematic studies on the friction and wear behavior of DLC coatings under lubricated
conditions were intensified in the last decade [37–43]. Early studies on the boundary
lubrication of DLC coatings were scarce and contradictory on the view points of tribological performance and tribofilm formation. Nevile et al. in 2007 and Kalin et al.
in 2008 published comprehensive review papers on this area [8, 39]. Vengudusamy et
al. studied a large number of different DLC coatings lubricated with API III base oil
[44]. In their observation, the ta-C type DLC coatings gave lower boundary friction than
any other type. They found that wear resistance has a clear dependence on DLC type.
Masripan et al. investigated the effect of hardness on the DLC coating in additive-free
mineral base oil [45]. They observed that the hardest DLC showed the lowest friction
and wear.
Studies on the boundary lubrication of DLC coatings have been shown that hydrogenated form of DLC coatings reach ultra-low friction values in the lubricant containing
friction modifier Molybdenum Dithiocarbamates (MoDTC) due to the formation of selflubricating MoS2 sheets [37, 46]. However, it has also recognized that the wear rates of
hydrogenated DLC coatings in MoDTC containing oils were higher than pure base oil
lubrication [47–49]. It has shown that MoDTC and Zinc Dialkyldithiophosphate (ZDDP/ZnDTP) additivated oils further improve the friction and wear performance of DLC
coatings under boundary lubricated conditions [50, 51]. ZDDP is a widely used antiwear additive in engine oils (Fig. 1.6) which forms wear protective tribofilms on ferrous
Introduction
12
surfaces through tribochemical reactions [52]. Even though some studies have reported
that ZDDP don’t have anti-wear performance on DLC surfaces and no tribochemical
reaction occurs between ZDDP additives and DLC surfaces [38, 53, 54], there are numerous studies which report the formation of weak ZDDP tribofilm on DLC coatings
[55–58].
On the other hand, many studies have confirmed that ultra-low friction and even superlow friction can also be achievable with hydrogen-free DLC or ta-C coatings under boundary lubricated contacts when lubricated with ester containing lubricants [13, 59, 60].
Surface-sensitive analysis using X-ray photoelectron spectroscopy (XPS) [59, 61, 62]
and time-of-flight secondary ion mass spectrometer (ToF-SIMS) [63] have confirmed a
thin absorbed OH layer on the surface of ta-C coating. Based on these findings, it is
suggested that mechanically activated surface dangling bonds of ta-C surfaces are terminated by alcohol function groups of ester which resulted in ultra-low friction. Mabuchi et
al. reported that dangling bonds in ta-C surfaces are key factor for the ultra-low friction
mechanism of ta-C coating, when lubricated in glycerol-mono-oleate (GMO) containing
oil [64]. Recently, it was observed that the wear rate of ta-C coating under the lubricated
condition tented to decrease with increasing surface hardness [65] and it was shown that
addition of GMO greatly enhanced the wear resistance of ta-C coating [66].
Figure 1.6: Structure of (a) GMO and (b) ZnDTP additives
Introduction
1.4
13
Purpose of This Study
Classifying the lubrication of tribological components, since HL and EHL contacts are
completely separated by a continuous lubricant film, boundary-lubricated contacts are
most crucial that predefine the performance, durability and effectiveness of many engineering systems due to the direct solid surfaces contact which leads to increased friction, energy loss, high wear and material damage. Several components of engines and
power trains in automobiles operate under severe boundary lubricated conditions such
as piston rings and cams/followers as well as engine bearings during start up, stopping,
shock-loads and direction changes. Most of the fuel energy generated in an engine is
dissipated in those components to overcome high friction coefficient. On the other hand,
automotive industries are one of the major contributors of atmospheric pollution. Governments regulations force automakers to reduce greenhouse gas (CO2 ) emissions and
fuel consumption.
Accordingly, for better friction and wear performance, the application of DLC coatings
on engine components with biodegradable oil is becoming one of the very effective ways
to meet both the increasingly tighter emission control standards and higher fuel efficiency
requirements of engines by automotive industries. However, commonly used engine oils
are very complex systems with different additives that used for different tasks. Also,
there are several different types of DLC coating available depending on the sp2 / sp3
ratio, hydrogen content and presence or absence of doping agents. Therefore, due to the
variety of DLC coatings and lubricants, a generalized boundary lubrication mechanism
for all DLC coatings has not been proposed yet. Thus, the oil lubrication mechanisms
of DLC coatings and interaction between lubricant additives with various DLC surfaces
are still not clear enough.
As previously noted above, hydrogen-free ta-C DLC able to provide ultra-low friction
under boundary lubricated conditions when lubricated with GMO type FM containing
oil. On the other hand, it is obvious that GMO alone will not be as effective in wear
reduction as it is effective in friction reduction. The engine oil must also contain an AW
additive which should work synergistically with FM and ta-C DLC. No one has been
systematically studied the boundary lubrication mechanism of ta-C DLC for DLC/steel
and DLC/DLC contacts when lubricated with a base oil and base oil containing both
FM and AW additive. So the first objective of this study is to propose a friction and
wear mechanism for hydrogen-free ta-C DLC when lubricated with base oil and base
oil containing FM alone, AW additive alone and FM+AW. In addition, the effects of
operating conditions and additive concentrations on the performance of ta-C DLC will
also be studied.
Introduction
14
DLC coatings have been known to be chemically inert. However, recent studies have
shown that various types of DLC coatings may interact with certain lubricant additives.
Despite having significant research progress, results are still controversial in terms of interaction between additives and DLC surfaces. Therefore, in order to better understand
the boundary lubrication mechanism of ta-C DLC and interactions between the DLC
surfaces and AW additive, a wide range of DLC coatings will be tested with self-mated
DLC/DLC contacts under same conditions and be compared with ta-C DLC coating.
To summary, the main objectives of this study are as follows:
1. To clarify the ultra-low friction and wear mechanism of ta-C DLC under boundary
lubricated conditions in base oil and base oil blended with a FM, an anti-wear (AW)
additive and mixture of FM+AW additives.
2. To study the effects of oil temperature, additive concentrations and countermaterial on the boundary lubrication performance of ta-C DLC.
3. To investigate the reactivity of various DLC coatings with an commonly used AW
additive, and to identify governing parameters such as hydrogen, doping elements
and surface morphologies in terms of interaction between additives and DLC surfaces.
Introduction
1.5
15
Outline of Dissertation
This dissertation presents the latest research in the field of oil boundary lubrication of
DLC coatings. This first chapter begins with a description of tribology and industrial
needs of DLC coatings. The chapter presents a short introduction of lubrication theory
with regard to the three lubrication regimes of Stribeck curve; hydrodynamic lubrication,
elastohydrodynamic lubrication and boundary lubrication. The chapter also reviews
briefly the DLC coatings and oil boundary lubrication of DLC coated surfaces. The
organization of dissertation is presented graphically in Figure 1.7.
Chapter 2 presents the ultra-low friction behaviour of non-hydrogenated ta-C DLC coating under boundary lubrication by testing in environmentally friendly base oil poly
alpha-olefin (PAO4) and PAO4 containing GMO type FM, ZnDTP type AW addtive
and mixture of GMO+ZnDTP. Friction properties, role of temperature and functions of
additives were analysed after tribological tests. To clarify the ultra-low friction mechanism of ta-C DLC and effectiveness of lubricant additives, various surface sensitive
spectroscopic and microscopic analysis were applied on the tested DLC surfaces. A
mechanism of friction for each condition is also discussed in details.
Chapter 3 discuss the wear mechanism of same ta-C DLC in same testing conditions as
described in chapter 2, along with the effects of running distance, counterbody material
and oil temperature. With the light of current research work and understanding, the
wear mechanism of ta-C under oil boundary lubricated condition have tried to clarify
with the suggested hypothesis.
Chapter 4 compares the friction and wear behaviour of six types of DLC, a-C, ta-C, two
kind of a-C:H, Si-DLC, Cr-DLC lubricated with same poly alpha-olefin (PAO4) base oil
in order to clarify more clearly the oil boundary lubrication mechanism of DLC coatings.
Additionally, this chapter will investigate the reactivity and ZnDTP tribofilm formation
on various DLC surfaces by FE-SEM, AFM and XPS analysis. The effects of hydrogen,
doping elements and surface morphologies will be studied in terms of ZnDTP tribofilm
formation and tribological performance of DLC coatings under boundary lubricated
conditions.
Finally, all findings of this dissertation will be summarized in chapter 5.
Introduction
16
Chapter 1: Introduction
This chapter gives a historical overview on tribology, lubrication theory and DLC coatings, literature survey on oil lubrication of DLC coatings, purpose of this dissertation and outline
Chapter 2:
Ultra-low friction of ta-C DLC
under boundary lubrication
• This chapter presents the ultralow friction mechanism of nonhydrogenated ta-C DLC coating
under boundary lubrication by testing in base oil poly alpha-olefin
(PAO4) and PAO4 containing GMO,
ZnDTP and GMO+ZnDTP and
analyze the effect of temperature, additive concentration and load bearing
capacity.
Chapter 4:
Boundary Lubrication of selfmated DLC/DLC contacts in
synthetic base oil and influence
of ZnDTP tribofilm formation
• This chapter discusses the tribological behavior of six different types of
DLC coatings in synthetic base oil
under boundary lubricated condition.
• Investigation of ZnDTP tribofilm
formation on various DLC surfaces.
Chapter 3:
Wear behaviour of ta-C DLC
• This chapter focused on the wear
behavior of ta-C DLC coating in
additive containing lubricants against
steel and self-mated ta-C DLC under
boundary lubricated conditions to
analyze the effects of additives, temperature, applied load and counter
materials on the wear of ta-C DLC.
• The effects of hydrogen, doping elements and surface morphologies on
the ZnDTP tribofilm formation on
DLC surfaces.
Chapter 5: Summary, Conclusions and Future outlook
Figure 1.7: Outline of dissertation
Chapter 2
Ultra-low friction of ta-C DLC
Advanced mechanical systems and modern engines need to work in extreme environments
such as high contact pressure, high or low temperature, high vacuum, high or low speeds
etc. On the other hand, fuel efficiency and government’s restrictions on the emission of
harmful elements generate serious concerns in the automotive industry [2]. Extensive
research is being devoted to overcome these challenges by controlling the friction and
wear between moving parts. In recent years use of diamond-like carbon (DLC) coatings
on the engine components which works under boundary lubrication conditions with
biodegradable oils is becoming a potential solution to achieve ultra-low friction and low
wear rates [8] .
Use of oils is a common application to reduce the friction and the wear in the contacting
surfaces. Typically, commercially available fully formulated oils are designed for ferrous
surfaces and contain numerous additives to enhance performance of base oils. Hence,
interaction of DLC surfaces with oils and additives is not yet fully understood. On
the other hand, complex shapes of mechanical parts make it difficult to coat surfaces
in good quality, adhesion and uniformity. Therefore, comprehensive studies on the
performance of DLC surfaces under oil boundary lubricated conditions with DLC/steel
and DLC/DLC mating contacts are needed.
DLC coatings have been used more than two decades as solid lubricants in industry
due to their exceptional tribological and mechanical properties such as low friction, high
wear resistance and high hardness. There are variety of DLC coatings and some of them
are extremely hard as like diamond. Friction and wear properties of DLC coatings show
significant fluctuation depending on the deposition method, chemical and structural
nature, substrate material, contact pressure, working temperature and test environment
17
18
[7]. Basically, hydrogenated DLC coatings give ultra-low friction in inert or vacuum
environments whereas hydrogen-free DLC coatings provide ultra-low friction and wear
in the presence of oxygen, hydrogen or water molecules[10, 31].
Tribological performance of DLC coatings under oil boundary lubrication have been
studied by many researchers [38, 42, 44, 67] . It has been showed that various DLC
coatings exhibit ultra-low friction and excellent wear resistance when used with proper
base oil and lubricant additives. De Barros et al. showed that hydrogenated DLC provide ultra-low friction under boundary lubrication when lubricated by poly alpha-olefin
(PAO) containing MoDTC and ZDDP+MoDTC [37]. It is concluded that formation
of low friction MoS2 sheets on DLC surface is the mechanism of ultra-low friction and
presence of ZDDP enhance the formation of MoS2 sheets. It is commonly known that
ZDDP forms a tribofilm on steel surface to protect against wear [52]. However, role of
ZDDP (or ZnDTP) and formation of ZDDP related surface protective tribofilm on DLC
is still not clear [55, 56] .
Studies showed that hydrogen free ta-C DLC lubricated with PAO base oil containing organic friction modifier glycerol mono-oleate (GMO) also give ultra-low friction
for DLC/steel and DLC/DLC contacts [59, 61, 63] . It is stated that termination of
friction activated dangling bonds of ta-C DLC surface through tribochemical reaction
by lubricant alcohol function groups and resulting low energy interaction between OHterminated surfaces sliding on each other is the origin of ultra-low friction.
In the present chapter, we aimed to investigate ultra-low friction mechanism of nonhydrogenated ta-C coating under boundary lubrication by testing in base oil poly alphaolefin (PAO4) and PAO4 containing GMO, ZnDTP and GMO+ZnDTP. Friction properties and functions of additives were analyzed after tribological tests. The effectiveness
of coating one of the contact surfaces or both contacting surfaces was compared by
testing steel/steel, DLC/Steel and DLC/DLC sliding pair at 80◦ C . In order to better
understand the friction mechanism of ta-C coating in lubricated condition, in this study
we present the effect of oil temperature and additive concentrations on the lubrication
performance of ta-C coating when rubbed against steel counterpart.
2.1
2.1.1
19
Experimental details
Material characterization and lubricants
Non-hydrogenated ta-C DLC coatings were deposited by FCVA (filtered cathodic vacuum arc) method on polished high carbon chrome steel (SUJ2) pins and high carbon steel
(S55C) discs with the thickness of 0.7 and 1 µm, respectively. A thin metal interlayer was
used to increase the adhesion between the coating and the substrate. Coatings were supplied by Nippon ITF Inc., Japan. Average hardness of substrates was 65 HRC for SUJ2
and 60 HRC for S55C. Surface roughness, hardness and Young modulus were measured
by atomic force microscopy (SEIKO, Nanopics 1000) and Nano indenter (NANOPICS
1000 Elionix ENT-1100a), respectively. The properties of ta-C coated pin and disc are
listed in Table 2.1. Raman spectroscopy (NRS-1000 Laser, Jasco Inc.., Japan) measurements with 532 nm Ne laser radiation were carried out to characterize coating. Nearly
symmetrical G band was observed at 1560 cm−1 as it is seen Fig. 2.1 which state very
high sp3 value [10]. Elastic recoil detection analysis (ERDA) was also performed on
DLC coating for the verification of hydrogen absence in the coating structure.
Table 2.1: Properties of ta-C DLC coated disc and pin
Properties
Disc
Pin
Dimension (mm)
Substrate
Coating Method
Thickness (µm)
Surface roughness, Ra (nm)
Hardness (GPa)
Young Modulus (GPa)
Hydrogen content (at.%)
22.5 X 4
S55C
FCVA
1.0
15 ±5
75 ±5
900 ±50
<1
5X5
SUJ2
FCVA
0.7
25 ±5
75 ±5
900 ±50
<1
The base oil used in this study was synthetic polyalpha-olefin (PAO4) having viscosity
of 19 mm2 /s and pressure-viscosity coefficient of 17.08 GPa−1 at 40 ◦ C. Glycerol monooleate (GMO) was added to the base oil as an organic friction modifier. Anti-wear
additive zinc dithiophosphate (ZnDTP) was also used to enhance the wear performance
of tribosystem. The ZnDTP was a secondary type which is generally used for commercial
engine oils. Incorporating the additives did not result in significant change of base oil
viscosity. The lubricants used are defined in Table 2.2.
20
Figure 2.1: Raman Spectra of ta-C DLC coated pin and disk
Table 2.2: Boundary lubricant components and additive composition.
Lubricants
Base oil
(PAO)
wt%
GMO
wt%
ZnDTP
wt%
100
99
99.82
98.82
1
1
0.08
0.08
99.5
99.86
98.86
0.5
0.5
0.04
0.04
Group A
PAO
PAO+GMO
PAO+ZnDTP
PAO+GMO+ZnDTP
Group B
PAO+GMO
PAO+ZnDTP
PAO+GMO+ZnDTP
2.1.2
21
Tribological experiments
Boundary lubrication friction tests were carried out using standard pin-on-disc type
unidirectional tribotester (Fig. 2.2). The DLC-coated or non-coated flat ended circular
cylinder pin, measuring 5 mm in diameter and 5 mm in length, was loaded and rubbed
against DLC-coated or non-coated steel disc, measuring 22.5 mm in diameter and 4 mm
in thickness, under pure sliding condition. For the DLC/steel contact, DLC coated pin
was rubbed on steel disc. The pin was located 6 mm in diameter from the center of the
disc and fixed to prevent it from rotating. A load of 5 N was applied (corresponding
to a maximum initial Hertzian contact pressure of 150 MPa) with 0.1 m/s entrainment
speed. Both pin and disc were totally immersed in lubricant solution. The tests were
performed at 25, 50, 80 and 110◦ C with 0.1 m/s average linear speed. The average speed
was calculated in accordance with the middle of the line contact. The oil temperature
was controlled by thermocouple which is fitted to just below the holder. Test duration
was 1 hour and total sliding distance was 405 m. Before and after the friction tests, all
samples were rinsed in benzene and washed with acetone in an ultrasonic bath (unless
otherwise stated) to remove contaminants and oil species. The tests were repeated at
least four times for reproducibility.
Figure 2.2: Pin on disc type tribotester
22
Tests were performed on steel/steel, DLC/steel and DLC/DLC tribopair. Non-coated
SUJ2 cylindrical steel pins with the surface roughness of Ra 25 ±5 nm and non-coated
SUJ2 steel disc which polished in several steps to surface roughness Ra 3 ±1 nm were
used as steel pairs for steel/steel and DLC/steel contacts. The initial tests were conducted with Group A lubricants at 80◦ C. In order to obtain oil temperature effect, tests
were performed with DLC/steel couple at 25, 50, 80 and 110◦ C in Group A lubricants.
The tests were repeated at 80 and 110◦ C with Group B lubricants which contain half
the additives concentration of Group A lubricants to analyze the effect of additive concentration.
The minimum film thickness (hmin ) and dimensionless lambda ratio (Λ) were calculated
using equations 2.1 and 2.2, respectively, by Hamrock and Dowson [14].
hmin = 3.63RU 0.68 G0.49 W −0.073 (1 − e−0.68k )
Λ= q
hmin
(2.1)
(2.2)
2 + R2
Rq,a
q,b
where R is radius of pin, U is dimensionless speed parameter, G is dimensionless materials
parameter, W is dimensionless load parameter, Rq,a is the surface roughness of pin and
Rq,b is the surface roughness of disc. The calculated lambda ratio was 1.2 for 20◦ C and
less than unity for 50, 80 and 110 ◦ C which means that operating lubrication regime
was mixed lubrication for 20 ◦ C and boundary lubrication for 50, 80 and 110 ◦ C.
2.1.3
Surface analysis
Rubbed surfaces of pins and discs were studied using optical microscopy, field emission scanning electron microscopy (JEOL, JSM-7000FK), Energy-dispersive X-ray spectroscopy (EDS), X-Ray photoelectron spectroscopy (PHI Quantera II, ULVAC-PHI,
Inc.) and AFM (Nanopics 1000, SEIKO instruments and SPA400, SII Nanotechnology
Inc.). XPS measurements on rubbed ta-C surfaces were conducted using a monochromatized AlKα source in an area of 500 µm with 45◦ take-out angle. CasaXPS software
was used for analysis of all the XPS spectra. Raman spectroscopy were also used to
analyze structural transformation of DLC coating.
2.2
23
Results
Friction tests were performed for steel/steel, DLC/Steel and DLC/DLC tribopair in
non-additivated and additivated PAO oil to understand effects of material combination
and role of additives.
2.2.1
Steel-Steel tribopair at 80◦ C
Fig. 2.3 shows representative friction coefficients of steel/steel pairs lubricated with
four different lubricant solutions. When lubricated with pure PAO, friction coefficient
of steel/steel pair was reduced to 0.09 value from the initial value of 0.12 after rubbing
10700 cycles. Although GMO is a friction modifier, use of GMO with PAO caused
higher friction coefficient than pure PAO for steel/steel contact. ZnDTP additivated
PAO showed friction coefficient ranged between 0.1 and 0.11. Friction tests of steel/steel
contact performed with GMO+ZnDTP containing PAO also exhibited quite high friction
coefficient than non-additivated PAO base oil tests.
Friction Coefficient
0.14
0.12
0.1
0.08
0.06
0.04
PAO
PAO+GMO
PAO+ZnDTP
PAO+GMO+ZnDTP
Sliding Speed : 0.1 m/s
Applied load : 5 N
Max. Pressure : 150 MPa
Temperature
: 80 °C
Steel pin vs. Steel Disc
0.02
0
0
2000
4000
6000
8000
10000
Number of Sliding cycles N, cycle
Figure 2.3: Representative friction coefficients as a number of sliding cycles between SUJ2 steel pin against SUJ2 steel disc in PAO, PAO+GMO, PAO+ZnDTP and
PAO+GMO+ZnDTP
An interesting result of steel/steel tribopair was observed in the friction tests performed
with PAO+GMO+ZnDTP. As it is known, ZnDTP containing oils form surface protective pad-like glassy phosphate tribofilm on steel surfaces [52] and this type of tribofilm
24
formation was also observed in our tests with PAO+ZnDTP (Fig. 2.4a). However,
ZnDTP related pad-like tribofilm formation was not observed on steel surfaces when
lubricated with PAO+GMO+ZnDTP as it is seen in Fig. 2.4b.
Figure 2.4: Fe-SEM images of steel pin and steel disc surfaces after rubbing 10700 cycles in steel/steel tribopair lubricated with a) PAO+ZnDTP and b)
PAO+GMO+ZnDTP
2.2.2
DLC-Steel tribopair at 80◦ C
Representative friction coefficients for DLC/steel tribopair in pure PAO and PAO containing GMO, ZnDTP and GMO+ZnDTP are shown in Fig. 2.5. Four individual tests
results in PAO+GMO and PAO+ZnDTP for DLC/steel contact are also presented in
Fig. 2.6 as examples of repeatability of the test results. Introduction of DLC coated
pins on DLC/steel contact immediately reduced the high initial friction coefficient values
from 0.11 to approximately 0.08 comparing to steel/steel contacts for all oil combinations. After running in period, tests in PAO and PAO additivated with GMO and
GMO+ZnDTP were reached ultra-low friction values for DLC/steel contact. On the
other hand, tests in base oil and additivated oils exhibited different initial periods.
25
0.14
0.12
0.1
PAO
PAO+GMO
PAO+ZnDTP
PAO+GMO+ZnDTP
Applied Load : 5 N
Temperature : 80 °C
DLC pin vs. steel disk
0.08
0.06
0.04
0.02
0
0
2000
4000
6000
8000
10000
Number of sliding cycles N, cycle
Figure 2.5: Representative friction coefficients as a number of sliding cycles between ta-DLC pin against SUJ2 steel disc in PAO, PAO+GMO, PAO+ZnDTP and
PAO+GMO+ZnDTP
0.12
PAO+GMO
PAO+ZnDTP
0.1
0.08
0.06
0.04
0.02
0
0
2000
4000
6000
8000
10000
Number of Sliding cycles N, cycle
Figure 2.6: Four individual tests results in PAO+GMO and PAO+ZnDTP for DLC/steel contact
Figure 2.7: Fe-SEM images of DLC and Steel surfaces after rubbing 10700 cycles in
DLC/Steel tribopair lubricated with a) PAO, b) PAO+GMO, c) PAO+ZnDTP, and d)
PAO+GMO+ZnDTP. The arrows indicate sliding directions.
26
27
In pure PAO oil at DLC/Steel contact, the coefficient of friction was reduced from the
initial value of around 0.08 to 0.025 after initial period. However, the DLC coating
was not stable when lubricated with pure PAO. After 10700 friction cycles, only in
PAO condition, the DLC coating totally wore out (Fig. 2.7a). After 5000 cycles, it
was assumed that the DLC coating partially wore out from the topmost surface, then
the friction coefficient showed vibration due to the steel/steel contact at some local
contacting points. Finally at the 8000 cycles, friction coefficient became high because
of the total wear of the DLC and removal of carbon species from the contact area.
Addition of GMO additives increased the stability of DLC coating and protected it from
total wear, see Fig. 2.7b. Compared to non-additivated PAO oil, PAO+GMO changed
the initial period and provided smooth transition to ultra-low friction values reducing
the friction from 0.08 to 0.025 for DLC/steel contact.
In the case of PAO+ZnDTP with DLC/steel combination, the coefficient of friction was
0.075 in initial cycles and slowly increased to 0.085 levels after rubbing 10700 cycles.
On the other side, it should be noted that this correspond to a 20% reduction from
the levels of steel/steel contact. Surface protective pad-like tribofilm formations were
observed both on DLC pin and steel disc surfaces (Fig. 2.7c). The lubrication with
GMO+ZnDTP additivated PAO oil exhibited similar friction property with PAO+GMO
oil for DLC/steel contact. ZnDTP related pad-like tribofilm were not formed neither on
steel disc nor on DLC pin surfaces (Fig. 2.7d). EDS measurement on DLC pin surfaces
lubricated with PAO+ZnDTP and PAO+GMO+ZnDTP were performed and elemental
compositions are listed in Table 2.3. Although pad-like tribofilm not formed on DLC
pin and steel disc surfaces in PAO+GMO+ZnDTP solution, small amount of ZnDTP
elements were detected on rubbed surfaces by EDS measurement.
Table 2.3: EDS measurements on boxed area of DLC pin surfaces in Fig. 2.7 lubricated with ZnDTP and GMO+ZnDTP containing oil for DLC/Steel tribopairs
Element
Concentration (%)
ZnDTP (X)
Carbon
Oxygen
Phosphorus
Sulphur
Iron
Zinc
83.35
4.54
1.70
1.16
0.81
8.43
GMO+ZnDTP (Y)
87.63
10.38
0.53
0.20
0.94
0.33
2.2.3
28
DLC-DLC tribopair at 80◦ C
Friction curves versus number of sliding cycles for DLC/DLC tribopair lubricated with
different oil solutions are shown in Fig.2.8. Initial friction coefficients were similar to
DLC/Steel tribopair which was lower than steel/steel combination. Friction reduction
to the ultra-low range was observed in all oil solutions. Average values of the steadystate friction coefficients with standard error are also compared in Fig. 2.9 for different
material combinations under pure PAO and PAO containing additives. The steady
state friction coefficients for each data point were calculated taking the average friction
coefficient of last 100 cycles of each test.
0.16
0.14
0.12
PAO
PAO+GMO
PAO+ZnDTP
PAO+GMO+ZnDTP
Applied Load : 5 N
Temperature : 80 °C
DLC Pin vs. DLC disc
0.1
0.08
0.06
0.04
0.02
0
0
2000
4000
6000
8000
10000
Number of sliding cycles N, cycle
Figure 2.8: Representative friction coefficients as a number of sliding cycles between ta-DLC pin against ta-DLC disc in PAO, PAO+GMO, PAO+ZnDTP and
PAO+GMO+ZnDTP
Tests in non-additivated PAO for DLC/DLC tribopair also exhibited ultra-low friction
but following different initial period than tests for DLC/steel. At DLC/DLC contact
in PAO, friction coefficient slightly increased from inital 0.09 value to 0.1 level and
then reduced around 0.025 after rubbing 10700 cycles. Comparing to DLC/DLC and
DLC/Steel tests in PAO, using DLC coated disc prevented the total wear of DLC pin.
In the case of PAO+GMO oil, the friction behavior for DLC/DLC combination was
similar to DLC/steel combination which reduced from 0.085 to 0.025 with smooth and
29
Figure 2.9: Comparison of steady-state friction of steel/steel, DLC/steel and DLC/DLC contacts for four different oil combinations after running 10700 cycles.
stable running-in period. Deep groove like wear scars were observed on DLC coated pin
surfaces for DLC/DLC contact after rubbing in all oil combinations as seen in Fig. 2.10.
In some cases, series of those wear grooves were formed on pin surfaces (Fig. 2.11). It is
believed that generations of these wear grooves on DLC pins at DLC/DLC contacts could
cause surface irregularities in the contact which could also be responsible for some local
friction spikes as it is seen in Fig. 2.8. After smoothening of these surface irregularities
by following running cycles, the friction coefficients were returned to the normal values.
These friction spikes were genuine tribo-couple behavior and not a measurement artifact
since they were observed in all four repeats of each test for DLC/DLC contacts.
ZnDTP is an anti-wear additive, but surprisingly the lowest friction was obtained in
PAO+ZnDTP oil at the DLC/DLC combination which showed very different initial
period than other oil solutions in any material combinations (Fig. 2.8). The high initial
coefficient of friction reduced sharply from 0.09 to approximately 0.04 after around 1000
cycles and finally reduced to less than 0.02 after rubbing 10700 cycles. The pad-like
tribofilm formation that was observed on DLC pin surfaces for DLC/steel tribopair
rubbed in PAO+ZnDTP (Fig. 2.7c) didn’t form on DLC pin surfaces for DLC/DLC
tribopair (Fig. 2.10c). ZnDTP related pad-like tribofilm was not also found on DLC
disc (Fig. 2.12). However, thin ZnDTP derived white layer was imaged with Fe-SEM on
the DLC disc, if it is just rinsed in benzene and acetone after rubbing in PAO+ZnDTP
solution (Fig. 2.12b). Table 2.4 presents the elemental composition of ZnDTP derived
white layer on DLC surfaces.
Figure 2.10: Fe-SEM images of DLC pin surfaces after rubbing 10700 cycles in DLC/DLC tribopair lubricated with a) PAO, b) PAO+GMO, c) PAO+ZnDTP, and d)
PAO+GMO+ZnDTP
Figure 2.11: Generated wear grooves on DLC pin
30
Figure 2.12: Fe-SEM and AFM images of DLC disc surfaces after rubbing 10700
cycles in DLC/DLC tribopair lubricated with PAO+ZnDTP; a) washed with acetone
in ultrasonic bath b) rinsed with benzene and acetone not washed in ultrasonic bath c)
AFM topography d) AFM lateral force
Table 2.4: EDS measurement on the worn surface of DLC disc lubricated with ZnDTP
containing oil for DLC/DLC tribopairs
Element
Carbon
Oxygen
Phosphorus
Sulphur
Zinc
Concentration (%)
83.55
9.29
1.01
2.61
3.54
31
2.2.4
32
Effect of oil temperature on the friction coefficinent for DLC/
steel contact
The average values of the steady-state coefficient of friction, over the last 5 min of
test period, as a function of oil temperature is given in Fig. 2.13 for DLC/steel contact
lubricated with Group A lubricants. The ta-C coating exhibited low coefficient of friction
about 0.05-0.06 in all lubricants at 25 ◦ C. When tested in pure PAO at 50 ◦ C, the ta-C
coating gave ultra-low boundary friction value of 0.018. As it is mentioned above, the
durability of ta-C coating was poor at 80 ◦ C and above in PAO. Therefore, although
the friction initially decreased reaching the lowest levels of 0.03-0.04 before the total
wear, it jumped to 0.08 levels due to the total wear of ta-C coating and resulting direct
steel/steel contacts at the end of the tests at 80 ◦ C and above in PAO.
Figure 2.13: Variation of friction coefficients as a function of temperature using Group
A lubricants. The curves are for guidance only
The PAO+GMO oil showed ultra-low friction coefficient with values around 0.025 at
50 and 80 ◦ C. The lowest friction value in PAO+GMO observed at 110 ◦ C, which was
around 0.016. For PAO+ZnDTP the friction consistently showed the relatively high
friction at all temperature range; with a stable value slightly below 0.09. In the case
of PAO+GMO+ZnDTP oil, friction coefficient reduced to around 0.05 at 50 ◦ C which
was higher than GMO alone and lower than ZnDTP alone. At higher temperature,
GMO+ZnDTP oil exhibited similar friction performance with GMO alone.
2.2.5
33
Effect of additive concentration on the friction coefficinent for
DLC/steel contact
The effect of additive concentration on the friction coefficients was compared in Fig. 2.14.
In all tests, a very small change in friction was seen when the additive concentration
was reduced by half. In the case of PAO+GMO, low concentration of GMO additive
resulted in lower friction coefficient. This effect was more pronounced at 80 ◦ C tests. It
is clear that reducing the concentration of ZnDTP additive in PAO+ZnDTP lubricant
did not change the friction performance. In the PAO+GMO+ZnDTP lubricants, low
concentration of additives provided a significantly lower coefficient of friction at 80 ◦ C
than high concentration of additives. However, this effect was not observed when the
temperature increased to 110 ◦ C.
Figure 2.14: Effect of additive concentration on the friction coefficient of ta-C DLC.
2.3
34
Discussion
Friction behavior of non-hydrogenated ta-C DLC in a synthetic base oil and effect of
additives on the tribological properties of boundary-lubricated ta-C DLC were investigated in this study. With the use of DLC in the contact, the coefficient of friction
reduced substantially for DLC/steel and DLC/DLC combinations. The results showed
a clear effect of oil temperature and additive concentrations on the friction coefficient
of ta-C coating. Firstly, the ta-C coating was able to provide ultra-low friction value of
0.018 with decent durability at 50 ◦ C in pure PAO. This is the lowest boundary friction
coefficient of DLC coating in base oil that recorded in literature and one of the major
finding in this work. However, when the oil temperature was increased over 50 ◦ C, the
ta-C coating exhibited limited lifetime against steel disc in pure PAO. Additionally, as
it will be examined in chapter 4, when tested and compared under same test conditions
with PAO oil, hydrogen-free ta-C DLC exhibited the lowest friction coefficient than any
other types of DLC coatings.
The low friction of diamond and ta-C DLC without lubrication in gaseous environments
has been explained by two hypothesis: (1) the transformation of sp3 and sp1 carbon
bonds on the topmost surface to more stable graphitic sp2 layers due to the heavy
pressure, high shear deformation and frictional heating, graphitic structure is capable
of reducing friction or have high anti-friction properties, or (2) passivation of dangling
σ-bonds generated during sliding by species from the surrounding environment, which
reduce or even eliminate the adhesive interactions across the sliding interfaces and result
in low friction [11, 68, 69].
In order to understand the ultra-low friction mechanism in pure PAO, all worn surfaces
tested at 80 ◦ C were analyzed using Raman spectroscopy to determine any structural
transformations. Since DLC coatings were totally worn out after rubbing 10700 cycles
for DLC/steel contact, friction test was stopped after 3000 cycles and 5000 cycles (after
wear out of coating from the topmost surface) to measure Raman spectra. Fig. 2.15
presents the Raman spectra of DLC pin surfaces before and after total wear. There
were no significant structural transformation as it seen Fig. 2.15a. The only structural
transformation observed on the partially wear out dark areas with very low intensity (Fig.
2.15b). Raman spectra measurements were also performed on DLC pin and DLC disc
after rubbing in lubricant solutions for DLC/DLC combinations. In Fig. 2.16, Raman
spectra of the as-deposited ta-C DLC pin compared to those obtained after sliding in
pure PAO and PAO containing additives. No significant structural transformation was
also observed neither on ta-C pin nor ta-C disc when rubbed in lubricant solutions for
DLC/DLC tribopair. On the other hand, it is important to note that graphitization
35
may occur on nanometer scale of top-most DLC surfaces and Raman spectroscopy may
not probe this modification.
Figure 2.15: Raman spectra of ta-C DLC pin rubbing in pure PAO for DLC/steel
tribopair a) before total wear occur b) after partially wear out of coating from the
topmost surface
In the literature, ta-C was tested under boundary-lubricated condition with base oil by
few research groups [44, 59, 70]. Vengudusamy et al. examined 12 different DLC coating
with an API Group III base oil [44]. It was found that ta-C give lower boundary friction
than other type of DLCs in base oil. It was concluded that hydrogen content and high
sp3 /sp2 ratio are key factor for low friction of ta-C in base oil rather than structural
transformation. Our colleagues Masripan and Ohara studied one ta-C and two a-C:H
DLC in an additive-free mineral oil [45, 71]. They found that ta-C DLC show the lowest
36
Figure 2.16: Raman spectra of as-deposited ta-C DLC pin and after rubbing 10700
cycles in PAO, PAO+GMO, PAO+ZnDTP and PAO+GMO+ZnDTP for DLC/DLC
tribopair
friction coefficient. Based on the their surface analysis with Raman spectroscopy and
spectroscopic ellipsometry, it was concluded that thickness of structurally transformed
graphitic layer can be important factor for the low friction performance of DLC coatings.
Based on literature survey, our results and Raman spectroscopy analysis, the ultra-low
friction of ta-C DLC under oil boundary lubrication can be explained by the passivation
of nascent ta-C surface and oil interaction after the removal of surface contaminants
during initial period. However, this explanation is insufficient since the PAO oil is
non-polar which has high degradation stability at elevated temperature and cannot
interact with ta-C DLC surfaces. The second and strongest explanation for the ultralow friction of ta-C in PAO oil is the structural transformation of topmost surface to
graphite like structure. But, it should be noted that higher sp3 hybridization, hydrogenfree nature of coating, deposition methods and thickness of transformed layer are also
important factors for the ultra-low friction performance of DLC coating under boundary
lubrication.
37
It is suggested that dangling bonds created at the top surfaces of DLC due to friction
and wear easily react with environmental species such as O, water molecules and H
in the dry sliding and resulting terminated surfaces reduce friction significantly [31].
Similar mechanism is reported for boundary-lubricated ta-C with GMO containing PAO
[59, 61, 63]. Some authors reported that dangling bonds of carbon atoms on the taC surfaces were terminated by hydroxyl -OH group of GMO additive and then the
formation of a hydrogen network leads to ultralow friction [59, 61]. Besides this model,
some other authors have proposed that the GMO molecules were assembled and bonded
with C atoms on the hydrogen-free DLC surfaces through the agent of O atoms of
hydroxyl -OH to form a monomolecular layer tribofilm (Fig.2.17) [62, 63].
Figure 2.17: Schematic presentation of GMO tribofilm formation on ta-C coating.
Termination of dangling bond of C atoms by GMO additive and its decomposed products [59, 62].
In our analysis, ultra-low friction behavior of same ta-C in GMO containing oil is very
different than the behavior observed in pure PAO. The initial period is same and stable
both in DLC/steel and DLC/DLC combinations lubricated with PAO+GMO but different than pure PAO cases. Also, GMO additivated PAO increased the durability of DLC
substantially and prevented the ta-C coated pins from total wear for DLC/steel pair.
These indicate that GMO interact with the ta-C and sustain the low friction behavior
of ta-C DLC in boundary-lubricated condition. Figure 2.18a shows the XPS spectra of
C1s on the worn surface of ta-C DLC after testing in PAO+GMO oil. The C1s spectra
was fitted with three peaks (2.18b ) around 284.9, 286.5 and 289.0 eV, respectively. The
38
main peak around 284.9 eV corresponds to C-C or C-H bonds, and the peaks around
286.5 and 289.0 eV correspond to C-O and C=O bonds, respectively. It was noted that
intensity of C-O and C=O peaks were slightly higher in the inside of wear track than in
the outside of wear track and as deposited ta-C surfaces. The presence of the C-O and
C=O bonds are mainly attributed to the hydroxylation of ta-C surfaces during sliding
in GMO containing lubricants [59, 61, 62].
The friction coefficient in 1 wt% GMO containing oil at 50 and 80 ◦ C is slightly higher
than in pure PAO at 50 ◦ C (Fig. 2.13). This friction behavior can be associated with
the efficient passivation of the ta-C surface by GMO additive. However, when tested in 1
wt% GMO containing oil at 110 ◦ C, the friction coefficient is reduced to similar level with
pure PAO at 50 ◦ C. This behavior can be explained by reduced adsorption/desorption
rate resulting the inefficiency of the GMO additive and its passivation mechanism at 110
◦ C.
Relatively higher coefficient of friction in PAO+ZnDTP oil for DLC/steel combination compared to other three oil solution is attributed to formation of pad-like ZnDTP
related tribofilm both on steel and ta-C pin surfaces (Fig. 2.7c). The formation of
pad-like ZnDTP related tribofilm prevents direct solid-to-solid contact and also supress
the graphitization of ta-C DLC surfaces. Friction properties of ZnDTP films on steel
surfaces is discussed in detail by H. Spikes [52]. In the literature, it is still not clear
whether ZnDTP in oil form a tribofilm on DLC surface to protect against wear as like it
does on steel surface or not. While some authors have reported that ZnDTP didn’t react
with DLC surface and no ZnDTP related tribofilm form on DLC coating [38, 53], some
have reported that ZnDTP related tribofilm did form on DLC surfaces but patchy like
not pad-like structure and it can be easily removed from surface by washing in ultrasonic
cleaning [55, 56]. It was found in our tests that ZnDTP related pad-like wear protective tribofilm form on ta-C pin surface for DLC/steel combination. Even after washing
15 minutes in acetone with ultrasonic cleaning, the tribofilm was still on the surface
which was evidenced by Fe-SEM and EDS observation as mentioned above. However,
the formation of pad-like tribofilm on ta-C pin may not be the result of the tribochemical reaction between ta-C and ZnDTP. Pad-like tribofilm formed on ta-C pin when the
counterpart was steel (Fig. 2.7c) and not formed on ta-C pin when the counterpart was
DLC disc (Fig. 2.10c). Hence, it is thought that ferrous molecules may be transferred
to ta-C surfaces at the initial cycles of friction test and pad-like tribofilm may form on
those transferred ferrous molecules.
In the case of DLC/DLC tribopair in PAO+ZnDTP oil, no pad-like tribofilm was observed neither on ta-C pin nor on ta-C disc surfaces in contrast to DLC/steel tribopair.
Figure 2.18: XPS spectra for C1s state on the worn surface of ta-C DLC coating
tested in PAO+GMO compared with as deposited ta-C surface (a) and deconvolution
of XPS C1s peak
39
40
On the other side, lowest friction was recorded in PAO+ ZnDTP for DLC/DLC combination in all our tests. Analysis indicated that very thin ZnDTP-derived white layer
form on DLC disc surfaces and it can be readily cleaned from surface when it was washed
in ultrasonic bath as it is shown in Fig 2.12. These support also our theory that transferred ferrous molecules can be the reason of pad-like tribofilm formation on ta-C pin
for DLC/steel combination. Basically, ZnDTP behave differently depending on material
combination; it form pad-like wear protective tribofilm on steel surface which is the
reason of high friction for DLC/steel tribopair, while it form thin white layer on ta-C
surfaces with DLC/DLC tribopair. Therefore, it is derived from results that ZnDTP
don not react with ta-C surfaces as like it react with ferrous surfaces.
A general observation in this work was that sliding in PAO+GMO+ZnDTP have different friction mechanism from those sliding in PAO+ZnDTP for all material combinations.
Under PAO+GMO+ZnDTP lubricated condition initial period is slightly different but
friction values after rubbing 10700 cycles reduce to the ultra-low values both for DLC/steel and DLC/DLC tribopair. No pad-like ZnDTP-derived tribofilm on steel surfaces
and no ZnDTP-derived white layer formation on ta-C surfaces were found. It gives similar friction mechanism with PAO+GMO for DLC/steel and DLC/DLC combinations.
In the study by Miclozic et al., it is stated that ZnDTP-derived tribofilm thickness is
reduced when ZnDTP generated steel surfaces rubbed in GMO containing oil solution
[72]. In our tests, use of GMO and ZnDTP together prohibits the formation of ZnDTPderived tribofilm and changes the friction behavior as a result. This may be due to
thermal or tribochemical reactions between GMO and ZnDTP molecules during sliding.
It is suggested that the GMO additive suppresses the ZnDTP tribofilm formation by
reacting with ZnDTP in oil or adsorbing and blocking the solid surfaces.
2.4
41
Conclusions
Friction performance of ta-C in a synthetic base oil, with and without additives, have
been studied. Basically, we observed two distinct ultra-low friction mechanism of ta-C
DLC in boundary lubricated conditions; graphitization of topmost surface in pure PAO
and surface passivation by -OH, -H molecules in PAO+GMO oil. Following conclusions
can be drawn from this chapter.
• In PAO base oil, ta-C shows ultra-low friction for DLC/steel and DLC/DLC combination. Duration of ultra-low friction behavior was limited because of the total
wear of coating, but use of self-mated DLC/DLC tribopair prevents the ta-C pin
from total wear. The strongest explanation for the ultra-low friction of ta-C in
PAO oil is the structural transformation of very thin topmost surface to graphite
like structure.
• Addition of GMO further improved the boundary-lubricating effect of ta-C surfaces by providing smooth transition to the ultra-low friction and increased the
durability of ta-C pin for DLC/steel by preventing total wear of ta-C pin. The
mechanism of low friction in presence of GMO additive can be attributed to surface
passivation of ta-C surfaces by GMO molecules. Effectiveness of GMO on friction
is signiffcantly affected by temperature and additive concentration.
• In the presence of ZnDTP additive, tribological properties are totally different
depending on material combination. Relatively high friction are attributed to
the formation of pad-like tribofilm both on ta-C and steel surfaces for DLC/steel
combination. On the other hand, formation of ZnDTP-derived thin white layer on
ta-C DLC generates the lowest friction for DLC/DLC contact.
• GMO and ZnDTP do not have synergistic correlation. No additional friction
reduction or no ZnDTP-derived wear protective tribofilm formations are recorded
under PAO+GMO+ZnDTP lubricated conditions for DLC/steel and DLC/DLC
combinations.
• Further surface sensitive analysis is needed in order to clarify ultra-low friction
mechanism of ta-C in oil boundary-lubricated condition with and without additives.
42
Chapter 3
Wear behaviour of ta-C DLC
3.1
Introduction
Increasing human population, demands for higher energy consumption, destruction of
natural resources, pollution and global warming have been generated serious concern on
sustainability of the Earth. Governments, scientists and engineers are now focusing on
the environmentally friendly materials and designs on new technologies to reduce the
harmful impact of human activities on environment [73, 74]. Automotive industries are
one of the major sources of energy consumption, greenhouse gases emissions and environmental pollution [2, 3]. Controlling the friction and wear in mechanical components
of passenger cars could save huge amount of energy and reduce the release of hazardous
chemicals [5].
Diamond-like carbon films are eco-friendly hard coatings with the superior mechanical
and tribological properties which can be one of the solution to control friction and wear
in many applications [6, 10]. Previous studies have showed that DLC coatings can
provide ultra-low friction and high wear resistance [7]. In recent years, these coatings
have been applied to the mechanical components in cars which work under oil-lubricated
conditions [8, 75]. Since then, studies have been focusing on the interaction of various
DLC coatings with the different kinds of lubricants and lubricant additives to enhance
tribological properties of the mechanical systems [39–41, 43, 51, 59].
Generally, DLC coatings provide low friction and high wear resistance in oil-lubricated
conditions and it has been reported that tribofilm formation through the chemical or
physical reaction between DLC surfaces and oil additives is feasible depending on kinds
43
Wear behaviour of ta-C DLC under boundary lubrication
44
of DLC and structure of oil additives [13, 48, 58, 76, 77]. Furthermore, it has showed
in the literature that the friction coefficient of non-hydrogenated tetrahedral amorphous
diamond-like carbon (ta-C DLC) coatings are much lower than in any other kinds of
DLC under oil boundary lubrication, reaching the super-low values [44, 59].
Studies so far mostly have focused on the friction properties of ta-C DLC under boundary
lubricated conditions. Low friction mechanism of ta-C DLC in oil environments has
been attributed either to the structural transformation of upper most surfaces or to the
passivation of dangling bonds through oil additives [44, 60, 62, 64]. So far, very little
emphasis has been put on the wear properties of ta-C DLC in lubricated conditions [65].
In the present study, we focus on the wear behavior of ta-C DLC coating in additive
containing lubricants against steel and self-mated ta-C DLC under boundary lubricated
conditions to analyze the effects of additives and counter materials on the wear of ta-C
DLC. Synthetic Poly-alphaolefin (PAO) was used as base lubricants. The organic friction
modifier Glycerol mono-oleate (GMO) and commercial ZnDTP anti-wear additives were
added into the base oil to measure the influence of these two additives on the wear of
ta-C DLC.
3.2
45
Experimental
All materials, lubricants and tribological tests are same in section 2.2. Additionally,
Raman spectroscopy measurement (Jasco NRS-1000 and Renishaw) was performed with
two different systems on DLC coating for sp3 hybridisation. Additional wear tests were
conducted at 80◦ C with DLC/steel and DLC/DLC contacts for 75, 150, 225, 300 and 405
meters to evaluate the wear characteristic of ta-C DLC coating depending on running
distance. The wear rates of steel/steel contacts were also calculated for comparison.
For the clarification of wear mode and failure mechanism of the ta-C coating in PAO,
the ta-C coating was also rubbed against single-crystalline, pure germanium disc which
is hard metaloid with a diamond-like crystalline structure. Pure germanium is a semiconductor material which has less reactivity than ferrous surfaces with carbon materials.
The Ge-C phase diagram indicate that the solid solubility of carbon in germanium is
extremely low [78]. The surface roughness of the tested germanium disc was 6 ±2
nm. Surface hardness of the germanium disc was 11 ±2 GPa which measured by nanoindentation method with a Berkovich indenter (Elionix ENT-1100a).
Wear tracks of pins and discs were studied using optical microscopy, field emission
scanning electron microscopy (JEOL, JSM-7000FK), non-contact, three-dimensional,
scanning white light interferometry (Zygo, Newview), X-Ray photoelectron spectroscopy
( PHI Quantera II, ULVAC-PHI, Inc.) and AFM (Nanopics 1000, SEIKO instruments
and SPA400, SII Nanotechnology Inc.). Since the accurate wear volume measurements
on discs were not possible, wear calculation was performed only on pin specimens.
The wear rates were calculated using Archard wear equation (Eq. (3.1)).
V = kF s
(3.1)
where k is the dimensional wear rate (m3 /Nm), F the normal load (N), s the sliding
distance (m), V the wear volume loss (m3 ). The wear volume loss of pin specimens was
calculated by measuring the width of the wear scar. Optical microscope and scanning
white light interferometry were used to measure the width of the wear scar. Figure
3.1 shows the schematic measurements of the wear width using scanning white light
interferometry.
Figure 3.1: Measurement of the width of the wear scar on pin speciments using Zygo,
Newview.
46
3.3
3.3.1
47
Results
Wear results at 80◦ C
Friction and wear are described as function of a tribo-system [79]. They are not constant
property of materials and can differ depending on operating conditions, working environment (dry contact, lubrication, additives, temperature), material properties and counter
surfaces. Figure 3.2 displays the effects of counter surface and lubricant additives on
wear of hydrogen free ta-C DLC when tested in both base oil and additive containing
oils at 80◦ C for DLC/steel and DLC/DLC tribopair. Steel/steel contact results were
also given as a reference and comparison.
Figure 3.2: Total wear rate of pin for steel/steel, DLC/steel and DLC/DLC contacts
tested at 80◦ C with different lubricants after 405 m sliding distance . (steel/steel
contacts are given for comparison)
Hydrogen free ta-C DLC coatings exhibited severe wear when tested against steel in
pure PAO and coating was total wear out after 405 m sliding distance. GMO and
GMO+ZnDTP additivated PAO provided similar performance and greatly increased
the wear resistance of ta-C DLC pin against steel reducing the wear rate more than one
order of magnitude as compared to pure PAO. The lowest wear rate of ta-C DLC pin was
48
provided by PAO+ZnDTP oil solution for DLC/steel tribopair. The wear of ta-C DLC
coated pin was always lower than uncoated steel pin when tested against steel disc in
additive containing oils. However, wear rate of ta-C coated pin was higher than uncoated
pin when rubbed against steel in pure PAO. Figure 3.3 shows the optical microscopy
images of ta-C DLC worn surfaces tested against steel disc in different lubricants. As it
is seen in figure 3.3, ta-C coating on pin was worn out in PAO exposing the bright color
of substrate surface, while with the use of additives, total wear of coating was eliminated
for DLC/steel contact. Raman and XPS analysis after the tribotests also verified that
total wear of coating was occurred only in PAO for DLC/steel contact. The wear scars
obtained using GMO and GMO+ZnDTP additivated base oil exhibit polishing wear. No
significant wear was observed on ta-C pin when lubricated PAO+ZnDTP for DLC/Steel
contact (Figure 3.3).
Figure 3.3: Wear scar on ta-C DLC pins at DLC/Steel contact after sliding in different
lubricants tested at 80◦ C
In the case of DLC/DLC tribopair, the highest wear was provided by pure PAO and
PAO+ZnDTP oil solutions. The lowest wear rate was observed by PAO+GMO and
PAO+GMO+ZnDTP, respectively. As it is seen in Figure 3.2, with a DLC counterpart,
ta-C DLC pins exhibited higher wear rate than corresponding DLC/steel tribosystem in
additive containing lubricants. Conversely, wear rate of ta-C pin at DLC/DLC contact
was lower than DLC/steel contact in pure PAO. Use of DLC counterpart elaminated
the total wear of ta-C pin in pure PAO as compared to steel counterpart. Figure 3.4
shows the wear scar on ta-C DLC pins at DLC/DLC contact after sliding in different
lubricants. Polishing wear along with local dark abrasive scratches were observed in all
ta-C pin surface at DLC/DLC contact by using optical microscopy (Figure 3.4).
Figure 3.4: Wear scar on ta-C DLC pins at DLC/DLC contact after sliding in different
lubricants
49
Figure 3.5 presents the steady state friction coefficients as a function of different lubricants for steel/steel, DLC/steel and DLC/DLC contacts. The coefficients of friction for
steel/steel contacts were almost the same level of about 0.1 ±0.015 for different lubricant
solutions. In the case of DLC/steel combination, friction coefficient reached ultra-low
level of 0.025 in PAO but increased around 0.1 level after total wear of ta-C pin which
was the same level as that of steel/steel contact. GMO and GMO+ZnDTP additivated
lubricants gave ultra-low friction, but only ZnDTP containing lubricant exhibited relatively high friction coefficient at DLC/steel contacts. Ultra-low friction was measured
for the DLC/DLC contacts in all lubricant solutions.
Figure 3.5: Steady state friction coefficients for steel/steel, DLC/steel and DLC/DLC
X : ta-C pin exhibited ultra-low friction of 0.025
contacts as a function of lubricants [X
for DLC/steel in PAO before total wear and then friction coefficient jumped to 0.09
after total wear of coating]
3.3.2
50
Wear behavior depending on counter-body material and oil additives at 80◦ C
Wear volume measurement was calculated after the 75, 150, 300 and 405 m sliding
distance with new specimens for the specification of wear mechanisms and wear slope of
ta-C DLC depending on sliding distance.
3.3.2.1
DLC against steel contact
Figure 3.6: Wear volume loss of ta-C DLC pin vs. sliding distance for DLC/steel
contact at 80◦ C with different lubricants
The wear volume loss of ta-C DLC pins as a function of the sliding distance for DLC/steel
contact in four different lubricants are shown in Figure 3.6. The wear volume loss
increases linearly with increasing the sliding distance in PAO for DLC against steel
contact. The optical microscopy images of the wear scar on ta-C pin (Figure 3.7) reveal
that pure polishing wear occurred on sliding surfaces when lubricated with PAO. The
coating was worn out progressively and becomes thinner and started to break down
after running around 300 m which resulted exposure of bright substrate surface. With
51
the addition of GMO and GMO+ZnDTP additives, wear volume loss of ta-C decreased
significantly for DLC/steel contact and appeared to reach steady state condition with the
increasing sliding distance which state a constant wear rate (Figure 3.6). PAO+ZnDTP
lubricant showed exceptional wear protection for ta-C pin for DLC/steel contact and
generated wear volume loss was very low as it is seen in (Figure 3.6). Since ta-C DLC
are extremely hard as compared to steel, it generated rough abrasive wear scratch lines
parallel to the sliding distance on steel disc when lubricated with any lubricant solutions
used in this study (Figure 3.8).
Figure 3.7: Wear scar on ta-C DLC pin depending on sliding distance at DLC/steel
contact after sliding in PAO
Figure 3.8: Deep scratch lines paralel to sliding direction on steel disc (a) optical
microscopy and (b) AFM images
3.3.2.2
52
DLC against DLC contact
Figure 3.9 presents the wear volume loss of ta-C DLC pin for self-mated DLC coating as
a function of sliding distance when rubbed in four lubricant solutions. The wear volume
loss of ta-C pin depending on increasing sliding distance for DLC/DLC contacts scales
linearly with decreasing slope in all lubricants. The slope of PAO and PAO+ZnDTP
lubricated cases was similar and slightly higher than GMO and GMO+ZnDTP containing oil solutions. As can be seen in Figure 3.9, GMO additive was more effective than
ZnDTP for reduction of wear at DLC/DLC contacts. One of the difference comparing
the DLC/steel and DLC/DLC contact is that generated deep abrasive scratches on ta-C
pins at DLC/DLC contacts (Figure 3.10). These deep scratches were observed on ta-C
DLC pin in all lubricant solutions as it is seen dark spots in Figure 3.4. Since the counter
ta-C DLC surfaces were as hard as ta-C pin, wear scar could not be observed well with
optical microscopy. Figure 3.11 shows FESEM images of DLC disc worn surfaces with
the original scratches from the sample preparation after tested in pure PAO.
Figure 3.9: Wear volume loss of ta-C DLC pin vs. sliding distance for DLC/DLC
contact at different lubricants
Figure 3.10: Deep abrasive scratches on ta-C pin generated in all lubricants at DLC/DLC contacts (a) optical microscopy and (b) FESEM images
Figure 3.11: FESEM images of the DLC disc surface tested with PAO
53
3.3.3
54
Surface analysis on worn surfaces tested at 80◦ C
DLC coatings are a metastable form of amorphous carbon and tribo-induced graphitization of DLC films can be occurred in tribological tests [65, 80, 81]. Raman spectroscopy
measurements were taken on DLC surfaces before and after rubbing in base oil to identify graphitization. Graphitization can be interpreted from an increase in intensity ratio
(ID /IG ) indicated by the maximum disordered D-peak (ID ) intensity and the maximum
graphite G-peak (IG ) intensity in the Raman spectra. Figure 3.12 compares the Raman
spectra of ta-C before and after tests taken with micro-Raman (Jasco NRS-1000) and
ultra-violet (UV) Raman (Renishaw) with different wavelength. The change on the Raman spectra of rubbed ta-C samples was very slight. It is difficult to decide whether this
change may results from graphitization of top most surfaces or removal of pre-adsorbed
oxygen layer. However, the high wear rate that occurred in the pure PAO lubricated
tests can suggest the formation of graphite layer on the top most surface since it is softer
and more prone to wear.
ZnDTP produced wear protective pad-like tribofilm on steel surfaces during sliding tests
as like it is reported in literature [52]. However, FESEM and AFM analysis showed that
ZnDTP derived pad-like tribofilm didn’t form on steel surface when lubricated with
PAO+GMO+ZnDTP oil (Figure 3.13). It is believed that GMO may suppress the
ZnDTP and thus prohibit the formation of ZnDTP tribofilm. When friction and wear
results are evaluated broadly, it is apparent that PAO+GMO+ZnDTP provide similar
results with PAO+GMO. This also suggests that presence of GMO surpass the ZnDTP
in PAO+GMO+ZnDTP lubricant.
Formation of ZnDTP derived tribofilm on ta-C surfaces were examined by FESEM, AFM
and XPS. Not pad-like but white distributed ZnDTP tribofilm observed on ta-C surfaces
tested in PAO+ZnDTP (Figure 3.14a), but this white distributed tribofilm was not
found on ta-C when tested in PAO+GMO+ZnDTP (Figure 3.14b). XPS measurements
on rubbed ta-C surfaces were conducted using a monochromatized AlKα source in an
area of 500 µm with 45◦ take-out angle. CasaXPS software was used for analysis of
all the XPS spectra. XPS spectra were recorded on worn ta-C surfaces lubricated with
PAO+ZnDTP and PAO+GMO+ZnDTP. Figure 3.15 shows the XPS peaks of P 2p, S
2p, O 1s and Zn 2p obtained from ta-C disc surfaces. Results suggest that ZnDTP was
decomposed during sliding and ZnDTP derived Zn-, S- and P- containing species were
found in tribofilm for PAO+ZnDTP lubricated condition [57, 82]. However, when tested
in PAO+GMO+ZnDTP, it was found that peak intensity of Zn and P was much lower
than PAO+ZnDTP and S peak was absent on the ta-C surfaces.
Figure 3.12: Raman spectra with excitation wavelength for ta-C DLC before and
after rubbing in PAO
55
Figure 3.13: FeSEM micrograph (a), AFM topography (b) and AFM lateral force (c)
images of steel disc tested in PAO+GMO+ZnDTP
Figure 3.14: ZnDTP derived tribofilm formation on ta-C DLC surfaces tested in
(a)PAO+ZnDTP and (b) PAO+GMO+ZnDTP
56
Figure 3.15: XPS spectra of P 2p (with Zn 3s), S 2p, O1s and Zn 2p peaks obtained from (a) PAO+ZnDTP and (b) PAO+GMO+ZnDTP
lubricated ta-C DLC surfaces
57
3.3.4
58
Effect of oil temperature on wear of ta-C
In Figure 3.16, the wear rate of the ta-C coated pin as a function of oil temperature is
shown for DLC/steel contact lubricated with Group A lubricants. Generally, increased
oil temperature led to higher wear rate in all lubricants but PAO+ZnDTP. When tested
in pure PAO oil ta-C coated pin gave very low wear rate at 20 and 50 ◦ C, but further
increase in oil temperature to 80 and 110 ◦ C led to severe wear of the ta-C coating,
resulting in total wear of the ta-C coated pin against steel sample. As can be seen in
Fig. 3.17, the coating is totally worn out and thus the substrate material was exposed
with a brighter wear scar when tested in PAO at 80 ◦ C and above temperature. From
the wear analysis, it was noted that total wear of the ta-C coating in PAO observed
around 8000 cycles at 80 ◦ C and around 6000 cycles at 110 ◦ C.
Figure 3.16: Wear rate of ta-C coated pin for DLC/steel contact as a function of
temperature. The curves are for guidance only
The additive GMO significantly improved the wear performance and greatly enhanced
the durability of ta-C coating by eliminating the total wear at higher temperature. The
wear observed in GMO containing oil was more than 10 times lower than pure PAO oil at
80 and 110 ◦ C. The lowest pin wear occurred in ZnDTP containing oil in all temperature
range. It is interesting to note that GMO+ZnDTP showed similar wear performance
with GMO alone.
59
Figure 3.17: Microscopic images of the wear scar on the ta-C coated pin when tested
in pure PAO at (a) 50 ◦ C and (b) 80 ◦ C
3.3.5
Effect of additive concentration
The change of wear rate as a function of the additive concentration is presented in Fig.
3.18. In contrast to the friction behavior, reduction of GMO additive had significant
effect on the wear performance of ta-C coating. It is clear from the results that the wear
rate of ta-C coating was increased substantially when the use of GMO additive reduced
by half. In the PAO+ZnDTP lubricants, the amount of wear was very similar for low and
high concentration of ZnDTP additive. GMO+ZnDTP additivated lubricants exhibited
similar wear behaviour with GMO alone additivated lubricants in terms of the effect of
additive concentration.
Figure 3.18: Effect of additive concentration on the wear rate of ta-C DLC.
3.3.6
60
DLC v.s. Germanium disc
DLC vs. germanium disc tests were performed only in pure PAO lubricant. Figure 3.19
compares the average values of the steady-state friction coefficients for DLC/germanium
and DLC/Steel contacts as a function of the oil temperature. The DLC/germanium
system in general gave very low coefficient of friction. At room temperature (25 ◦ C), the
coefficient of friction was 0.035, lower than what was observed for DLC/steel contact.
When tested at 50, 80 and 110 ◦ C, the average friction of the DLC/germanium contacts
was noticeably lower than room temperature, reaching around 0.016 levels.
Figure 3.19: Friction coefficients of ta-C DLC when rubbed against Germanium as a
function of temperature in PAO oil.
The wear rates of the ta-C coated pins are compared in Fig. 3.20 for DLC/steel and DLC/germanium contacts when tested in PAO oil. The results showed that rubbing against
germanium disc significantly reduced the wear of ta-C coated pins at higher temperature. At 25 and 50 ◦ C tests, the wear rate was similar for both contacts. However, the
wear rates obtained in DLC/steel contacts at higher temperature was approximately 10
times higher than DLC/germanium contacts.
Figure 3.20: Wear rate comparison of ta-C DLC when rubbed against Germanium
and Steel discs as a function of temperature in PAO oil. The curves are for guidance
only
61
3.4
62
Discussions on the wear mechanism of ta-C coating
The tribological results reveal that ta-C coated pin undergoes severe wear and totally
wear out against steel counterpart in pure PAO lubricated case at high temperature
range. The results also showed a clear effect of oil temperature and additive concentrations on the wear behavior of ta-C coated pins. Wear behavior analysis of ta-C pin in
PAO (Figure 3.7) showed that wear take place as smooth polishing wear for DLC/steel
contact. On the other hand, addition of lubricant additives or use of DLC counterpart
eliminates the total wear of ta-C pin under boundary lubricated condition. Moreover,
the results are exhibited that rubbing against germanium disc caused an order of magnitude lower wear rate than rubbing against the SUJ2 steel disc at 80 and 110 ◦ C (Fig.
3.20). Among all other results, the excessive wear of harder ta-C DLC against softer
steel disc in PAO is one of major finding of this work.
For the explanation of excessive wear of ta-C DLC against steel disc in PAO and effect
of counter materials, we proposed three hypothesis. One of the hypothesis can be abrasive wear of ta-C DLC by generated, abrasive steel particles (Fig. 3.21a). According to
the this hypothesis, hard ta-C DLC coated pin scratches and generates wear particles
from the steel disc. Those wear particles are dragged into the contact area at every
running cycles and accelerate the mechanical (abrasive) wear of ta-C coated pin. However, same phenomenon can occur, when ta-C DLC rubbed against germanium disc and
also generated steel particles cannot be hard enough to scratch the ta-C DLC. Second
hypothesis is degradation of oil and attack of degraded oil molecules to ta-C coating
(Fig. 3.21b). For this hypothesis, oil molecules degrade at the contact center or near
the contact center due to the high flash temperature. Those degraded oil molecules react
with carbon molecules on top of the ta-C coating and cause tribo-chemical wear. This
hypothesis is also not strong. Because, same oil degradation can occur, when ta-C DLC
is tested against self-mated ta-DLC and germanium disc. If second hypothesis had been
valid, we would have observed same excessive wear of ta-C DLC, regardless of counter
materials. Besides, it is known that synthetic PAO oils withstands high temperatures
with minimum decompositions.
The both above hypothesis are not sufficiently convincing. The strongest explanation of
the excessive wear of hard ta-C DLC against softer steel disc in PAO is tribo-chemical
wear by diffusion or dissolving carbon into the steel disc (Fig. 3.21c). Carbon atoms can
easily diffuse/dissolve into Fe, Ni, Co and their alloys [83]. Experimental and molecular dynamics (MD) studies in literature indicate that wear rate of diamond extremely
high when rubbed against steel and carbon has strong affinity for ferrous materials to
form covalent bond [84–87]. These studies also states that graphitization can occur
on diamond surface and thermo-chemical interaction of iron with carbon atoms causes
63
Figure 3.21: Hypothesis on the wear mechanism of ta-C DLC in pure PAO lubrication: (a) Abrasive wear by generated wear particles, (b) Tribo-chemical wear by the
degraded oil molecules, and (c) Tribo-chemical wear by graphitization and following
carbon diffusion
removal of carbon from the diamond surface. Hence, we hypothesized that, with the
increasing temperature and additional frictional heating, mobility of carbon atoms are
increased and top most surfaces of ta-C coating is transformed to graphitic structure.
Subsequently, thermally activated carbon atoms on the topmost surfaces of ta-C coating
diffuse into the steel surface and thermo-chemical interaction occur between iron and
carbon atoms which cause higher wear rate. The much lower wear rate of ta-C coating
sliding against self-mated ta-C DLC and germanium than those against steel are attributed the carbon rich nature of self-mated ta-C disc and the extremely low solubility
of carbon in germanium, respectively [78]. From these results, it was suggested that the
64
wear of ta-C coating in DLC/DLC and DLC/Germanium contacts caused by polishing
wear. On the other hand, the wear of ta-C coating in DLC/steel contact caused by
polishing mechanical wear associated with thermally activated tribo-chemical wear at
higher temperature. Therefore, it is suggested that counterbody material is an important environmental parameter for the design of ta-C coated systems. Using counterbody
materials of stainless steels, hard alloys, oxide coatings, refractory ceramics and diamond
can eliminate the carbon diffusion/dissolving and resulting tribo-chemical wear [11, 83].
The poor durability of the ta-C coating in PAO at higher temperature was eliminated
with the addition of GMO. The effectiveness of GMO for friction reduction of ta-C
coating has been reported by several authors [59, 62, 63]. MD studies by Morita et al.
indicate that H- and OH-terminated surfaces repulsively interact with Fe surfaces and
thus weaken the covalent interaction between ferrous surface and diamond surface [88].
Based on our test results and XPS analysis, we believe that dangling bonds of ta-C
DLC surfaces were passivated by molecules of GMO through the agent of O atoms of
hydroxyl -OH and a monomolecular layer form on the ta-C surface as it is proposed
by Ye et al. [62]. Passivation of dangling bonds can also suppress the graphitization
[89]. In the present study, enhanced wear resistance was also achieved along with the
friction reduction in the presence of GMO. It is thought that formation of OH-terminated
tribofilm prevents direct contact of the carbon atoms with the ferrous surface which
eliminates the thermo-chemical interaction between iron and carbon atoms and reduces
the wear rate. Therefore, since the tribo-chemical wear is eliminated by passivation of
ta-C surfaces, the wear mechanism of ta-C pin in GMO and GMO+ZnDTP additivated
lubricants for DLC/steel contact was polishing mechanical wear.
In the case of PAO+ZnDTP for DLC/steel contact, the ta-C coated pin was found to
have excellent wear performance with very low wear rate and relatively high friction
coefficient, around 0.09, in all temperature range. Surface analysis showed the ZnDTP
tribofilm formation both on steel disc and ta-C coated pin. ZnDTP tribofilm formation on the steel was due to the chemical reaction between ferrous surface and ZnDTP
molecules [52]. However, as has been noted previously, ZnDTP tribofilm formation on
ta-C coated pin was due to the transfer of ferrous molecules form steel disc to ta-C
pin instead of chemical reaction between ZnDTP molecules and ta-C coating. These
ZnDTP tribofilm formed on both steel and ta-C surfaces prevented the direct ta-C and
steel contact which prevent carbon diffusion into the steel surface. Consequently, sliding occurs between ta-C pin and ZnDTP tribofilm which resulted in exceptional wear
protection. Lowered mechanical wear was the wear mechanism for DLC/steel contact
in PAO+ZnDTP lubricant.
65
In DLC/DLC contacts, although C also has strong affinity for C to form covalent bond,
use of ta-C DLC coated disc eliminate the total wear of ta-C pin in PAO as compared
to DLC/steel contact. This is due to the graphitization of both surfaces which inhibit the covalent bonding formation between C atoms of counter surfaces. GMO and
GMO+ZnDTP additivated lubricants in DLC/DLC contact passivate the ta-C surfaces
and generate repulsive interaction with each other owing to antibonding interactions as
it was discussed earlier which provide low wear rate than pure PAO lubricated case in
DLC/DLC contact. White distributed ZnDTP-derived tribofilm was formed on ta-C
surfaces in PAO+ZnDTP lubricant but it was not wear protective. Generally, polishing
mechanical wear was the main wear mechanism of ta-C pin in DLC/DLC contact along
with minor local abrasive wear.
3.5
66
Conclusions
Wear behavior of ta-C DLC in non-additivated and additivated oils for DLC/steel and
DLC/DLC contacts have been studied. The results show that friction and wear performance of ta-C DLC coating strongly depend on the combination of lubricant formulation
and counterpart material. Use of additives in DLC/steel contact greatly increased the
wear resistance and durability of coating. Passivation of either ta-C surface or ferrous
surface through the lubricant additives is a crucial factor for reduction of ta-C DLC pin
wear. Anti-wear additive ZnDTP is more effective for DLC/steel contact in term of wear
protection. On the other hand, friction modifier GMO provides better wear protection
than ZnDTP in DLC/DLC contact. Based on above considerations and detailed surface
analysis, it is concluded that ZnDTP anti-wear additive don’t react directly with the
ta-C DLC surfaces. Additionally, the following conclusions can be drawn from results
of the effect of oil temperature and additive concentrations on the tribological behavior
of hydrogen-free ta-C DLC coated pin rubbed against steel disc.
• The ta-C coated pin can provide ultra-low friction in pure PAO, but the wear
resistance of ta-C coated pin is very poor against steel disc at 80 ◦ C and above
temperature which can be attributed to the thermally activated tribo-chemical
interaction between carbon and ferrous atoms.
• The use of additives enhance the wear resistance of ta-C coated pin at high temperature by elaminating the tribo-chemical wear due to the termination of surfaces.
• Effectiveness of GMO on friction and wear is significantly affected by temperature
and additive concentration. GMO is found to more effective at 50 and 80 ◦ C. Reducing the concentration of GMO by half is caused almost one order of magnitude
higher wear rate at 80 and 110 ◦ C.
• The ZnDTP additive is very effective for the reduction of wear even with reduced
concentration in all temperature range, but at the same time it causes relatively
high friction for DLC/steel contact.
Chapter 4
Boundary Lubrication of
self-mated DLC/DLC contacts in
synthetic base oil and influence of
ZnDTP tribofilm formation
4.1
Introduction
Diamond-like carbon (DLC) coatings are amorphous, meta-stable forms of carbon that
have promising properties like high hardness, chemical inertness, high electrical resistivity, high wear resistance, and low friction coefficient [9, 10]. Properties of DLC coatings
strongly depend on the deposition method, deposition conditions, chemical composition and operating conditions [7]. Hydrogen-free DLC coatings are mostly deposited
by physical vapor deposition methods and provide better tribological properties in a
humid environment [31, 90, 91]. On the other hand, hydrogenated DLC coatings are deposited by chemical vapor deposition methods and provide better tribological properties
in vacuum and dry conditions [33, 92].
In recent years, DLC coatings have attracted much attention as protective hard coatings
on automotive components due to their superior mechanical, chemical and tribological
properties [8, 57, 59]. However, owing to the chemical inertness of DLC surfaces, there
are still controversial questions concerning tribofilm formation on DLC surfaces and tribochemistry between lubricants, lubricant additives and DLC surfaces. Therefore, much
effort need to be put into understanding of tribofilm formation on DLC surfaces and their
effects on the tribological performance of DLC coatings under lubricated conditions.
67
Boundary Lubrication of self-mated DLC/DLC contacts in synthetic base oil and
influence of ZnDTP tribofilm formation
68
The anti-wear additive Zinc Dialkyl-Dithiophosphates (ZnDTP) is extensively used in
engine oils to control wear, corrosion and oxidation [52, 93]. ZnDTP forms pad-like
tribofilm on ferrous surfaces through the tribochemical reaction in actual sliding contact
[94, 95]. These films impart excellent anti-wear properties. However, since ZnDTP
additive is designed for ferrous surfaces, it is still not clear whether this additive will
react with DLC surfaces and form similar pad-like tribofilm as like it forms on ferrous
surfaces.
A number of previous studies have investigated the ZnDTP tribofilm formation on DLC
surfaces. Some of these studies reported no ZnDTP tribofilm formation on DLC coatings
by noticing the chemical inertness of DLC surfaces [67, 96, 97], while some researcher
reported ZnDTP tribofilm formation on DLC surfaces but not in pad-like structure
[55, 82]. There are also other studies which report actual pad-like ZnDTP tribofilm
formation on DLC coatings [56, 57]. These earlier results are conflicting for the formation of tribofilm. Accordingly, this study investigated the ZnDTP tribofilm formation
on various DLC surfaces sliding against self-mated counterpart surfaces in ZnDTP containing lubricant. The effects of hydrogen, doping elements and surface morphologies
have been studied in terms of ZnDTP tribofilm formation and tribological performance
of DLC coatings under boundary lubricated conditions.
4.2
4.2.1
69
Experimental
Material characterization and Lubricants
In this study, six types of DLC coatings were tested: one non-hydrogenated amorphous
carbon (a-C) coating, one non-hydrogenated tetrahedral amorphous carbon (ta-C) coating, two hydrogenated amorphous carbon (a-C:H) coatings, one silicon-doped hydrogenated amorphous carbon (Si-DLC) coating, and one chromium-doped hydrogenated
amorphous carbon (Cr-DLC) coating. All DLC coatings were deposited on discs and
pins of high carbon chrome bearing steel (JIS SUJ2), which has an average hardness
of 62 HRC. For all coatings, a thin metal interlayer was used to increase the adhesion
between the coating and steel. The a-C coating was produced by magnetron sputtering
method method, while the ta-C coating was deposited by filtered cathodic vacuum arc
(FCVA) deposition. Plasma-enhanced chemical vapor deposition (PECVD) was used
for the deposition of a-C:H, Si-DLC, and Cr-DLC with the same interlayer. Coatings
were supplied by Nippon ITF Inc., Japan.
The hardness and Young’s modulus of all DLC coatings were determined by nano indentation with a Berkovich indenter (Elionix, ENT-1100a), while the surface roughness
was evaluated by atomic force microscopy (AFM; Nanopics 1000, SEIKO instruments,
Japan). Hydrogen content in the coatings was measured using elastic recoil detection
anaysis. Table 4.1 shows the important characteristic properties of the DLC coatings
used in this study.
Table 4.1: Important characteristic properties of DLC coatings
No Material
Hardness
(GPa)
1
2
3
4
5
6
15
75
26
18
24
20
a-C
ta-C
a-C:H 1
a-C:H 2
Si-DLC
Cr-DLC
±3
±5
±3
±3
±3
±2
Young’s
Modulus
(GPa)
192 ±10
900 ±50
203 ±15
122 ±10
189 ±12
248 ±12
Roughness Hydrogen, Thickness Deposition
Ra (nm) (at.%)
(µm)
Method
5 ±3
17 ±5
12 ±4
11 ±3
14 ±4
4 ±2
<1
<1
19
25
27
22
1
1
1-2
1-2
1-2
1-2
PVD
FCVA
PECVD
PECVD
PECVD
PECVD
The base oil used in this study was a synthetic poly-alphaolefin (PAO) having 19 mm2 /s
viscosity and 17.08 GPa−1 pressure-viscosity coefficient at 40◦ C. Commercial antiwear ZnDTP additive was used at a concentration of 0.08 wt% to investigate additive
reaction with DLC coatings and the effect of a tribofilm on tribological performance.
The lubricant components and additive composition are shown in Table 4.2.
70
Table 4.2: Lubricant components and additive composition.
Lubricants
Base oil (PAO)
wt%
100
99.82
PAO
PAO+ZnDTP
4.2.2
ZnDTP
wt%
0.08
Tribological experiments
Tribological tests were carried out using a pin-on-disc type unidirectional tribotester.
The same type of DLC coated pin and disc was used in each test. A DLC-coated
flat-ended circular cylindrical pin, measuring 5 mm in diameter and 5 mm in length,
was loaded and rubbed against the self-mated DLC-coated disc, measuring 22.5 mm in
diameter and 4 mm in thickness, under pure sliding conditions as shown in Fig. 4.1.
The lower, flat disc was mounted on a steel holder fixed to a rotary turnable, while the
upper, cylindrical pin sample was located 6 mm eccentrically from the center of the disc.
The pin was fixed to prevent it from rotating and ensure a pure sliding condition. A
load of 5 N was applied, which resulted in a maximum initial Hertzian contact pressure
of 150 MPa. Both the pin and disc were totally immersed in the lubricant solution
and the temperature was kept at 80◦ C during sliding. The entrainment speed and test
duration were 0.1 m/s and 1 h , respectively. Prior to all tribological testing, samples
were cleaned with acetone in an ultrasonic bath to remove contaminants. Tribological
tests were repeated three times for verification and reproducibility of results. Figure 4.1
displays the schematic of the tribotester and pin-on-disc configuration.
The theoretical minimum film thickness (hmin ) and dimensionless lambda (Λ) ratio
were calculated using equations (4.1) and (4.2), respectively, by Hamrock and Dowson
to ensure boundary lubricated regime [14].
hmin = 3.63RU 0.68 G0.49 W −0.073 (1 − e−0.68k )
Λ= q
hmin
(4.1)
(4.2)
2 + R2
Rq,a
q,b
where R is radius of pin, U is dimensionless speed parameter, G is dimensionless materials parameter, W is dimensionless load parameter, Rq,a is the surface roughness of
71
Figure 4.1: Schematic of tribotester and pin-on-disc configuration
pin and Rq,b is the surface roughness of disc. The calculated lambda ratio was less than
unity which means that operating lubrication regime was boundary lubrication.
4.2.3
Surface analysis
Wear tracks of pins and discs were studied using optical microscopy, field emission
scanning electron microscopy (FESEM; JEOL, JSM-7000FK) and AFM (SPA400, SII
Nanotechnology Inc.). Since an accurate wear volume measurement on discs was not
possible, a wear calculation was performed only on pin specimens. Wear rates were
calculated using the Archard wear equation as defined (Eq. (3.1)). Wear volume loss
of pin specimens was calculated by measuring the width of the wear scar. Optical
microscopy and FESEM were used to measure the width of wear scar.
Chemical analysis of tribofilms was performed by X-ray photoelectron spectroscopy
(XPS; PHI Quantera II, ULVAC-PHI, Inc.) using a monochromatized AlKα source
in an area of 500 µm with a 45◦ take-out angle, and a power of 25 W. The pass energy of the analyzer was 140 eV. CasaXPS software was used for analysis of all the
XPS spectra. Curve-fitting was performed with Gaussian/Lorentzian functions after
Shirley background subtraction. The value of position, and full-width at half-maximum
(FWHM) were constrained to obtain the most appropriate chemical meaning. Sample
72
charging was corrected by shifting all peaks, referring all binding energies to carbon
at 284.8 eV. Oxygen spectra of all DLC coatings were fitted with three peaks by constraining the FWHM of synthetic peak at 2.0: one main non-bringing oxygen (NBO)
peak, one bringing oxygen (BO) peak and one oxide peak. The binding energy of the
NBO peak was constrained to be 531.8 eV. The BO peak was fixed to be 1.6±0.1 eV
above the NBO peak, and the oxide peak 1.8±0.1 eV below the NBO peak. The spectra
of phosphorus, sulfur and zinc were fitted with two or three peaks by constraining the
main peaks at 133.6±0.1, 162.2±0.1 and 1022.5±0.1 eV, respectively. A handbook of
XPS has been used to find chemical species at respective binding energies [98]. Raman
spectroscopy (NRS-1000 Laser, Jasco Inc., Japan) with a 532 nm Ne laser was carried
out for characterization of coatings.
4.3
4.3.1
73
Results and discussions
Coatings durability and ZnDTP derived tribofilm formation
Analyses of the rubbed DLC surfaces revealed that all DLC coated discs showed no
significant, measurable wear tracks with the exception of few scratches oriented in the
sliding direction. No evidence of fracture or delamination was also found on any of DLC
coated discs. On the other hand, a-C coated pin showed catastrophic delamination
both in PAO and PAO+ZnDTP oils as it’s seen in Fig. 4.2a. The a-C:H 1 coated
pins exhibited severe peeling-off on the worn surfaces (Fig. 4.2b). Similar peeling-off
phenomena was also observed on the worn surfaces of a-C:H 2 coated pins with lower
amount (Fig. 4.2c). It is important to note that the peeling-off or delamination of these
coatings is due to an adhesion failure between the coating and substrate material of the
pin. Optical microscopy images in Fig. 4.2 reveal that ZnDTP additive reduced the
amount of worn-surface damage and peeling-off on the a-C:H 1 and a-C:H 2 coated pins.
Si-DLC, Cr-DLC and ta-C DLC coated pins showed no delamination and peeling off on
the worn surfaces (not shown in here). Improved coating durability and wear resistance
were achieved with Si or Cr doping into a-C:H coating which further verified by the wear
analysis as discussed in the following subsection.
Surfaces of rubbed DLC discs were imaged with FESEM and AFM after the tribological
tests in PAO+ZnDTP oil to know whether characteristic pad-like ZnDTP tribofilm
form on DLC surfaces. FESEM images (Fig. 4.3) show clear formation of white padlike tribofilm on all DLC coatings with the exception of ta-C coating. On a-C coating,
formed pad-like structures were very tiny (Fig. 4.3a) and this was further verified by
AFM topography and lateral force images (Fig. 4.4a). Tribofilm formed on ta-C disc
was not in the shape of pad-like structure but white distributed patchy ZnDTP tribofilm
formation was observed on ta-C discs instead, as can be seen in FESEM (Fig. 4.3b) and
AFM (Fig. 4.4b) images.
Fig. 4.3c and Fig. 4.3d show formation of pad-like tribofilm on two different hydrogenated DLC coatings which have different amount of hydrogen content. Pads structure
on both a-C:H coating were not oriented in the sliding direction and shape of pads were
slightly larger than those formed on a-C coating. Area density of pad-like structure
formation on a-C:H 2 (Fig. 4.3d) coating was substantially higher than a-C:H 1 (Fig.
4.3c) coating which contain less amount of hydrogen. Formation of pad-like tribofilm on
both a-C:H coatings was also supported by AFM topography and lateral force images
(Fig. 4.4c and Fig. 4.4d).
74
Figure 4.2: Worn surfaces of DLC coated pins after rubbing in PAO and
PAO+ZnDTP oils: (a) a-C , (b) a-C:H 1 and (c) a-C:H 2 (arrows show sliding direction)
Shape of pad-like surface morphologies on doped DLC coatings were different than nondoped DLC coatings due to the alloying elements. It appears that ZnDTP pads are able
to form on doped-DLC coatings. FESEM images indicate that ZnDTP pads may be
localized around the Si or Cr sites in the top-most surfaces of the films. Si-DLC coating
showed close-packed tribofilm formation on the surface (Fig. 4.3e). Pad-like structure
on Si-DLC were in irregular shape which can form larger pads structure surrounded with
smaller ones. ZnDTP also produced close-packed pad-like tribofilm on the surface of CrDLC (Fig. 4.3f) but shape of the pads were slightly different compared to Si-DLC and
non-doped DLC coatings. Pads structure on Cr-DLC resemble more like bubble foam
patterns. Fig. 4.4e and Fig. 4.4f show AFM images of pad-like tribofilm on Si-DLC and
Cr-DLC, respectively.
Figure 4.3: FESEM images of ZnDTP derived tribofilm formation on various DLC
surfaces
75
Figure 4.4: AFM topography and lateral force images of ZnDTP derived tribofilm
formation on various DLC surfaces
76
77
Delamination of a-C coating on pin specimen could be results of poor adhesion on the
curved surface or poor-load bearing capacity of coating. Since a-C coated pin delaminated during the rubbing test, some ferrous molecules may transfer and spread on the
counter a-C disc. Formation of tribofilm on a-C coating could be induced by those transfered ferrous molecules. Formation of pad-like ZnDTP tribofilm on non-hydrogenated
a-C coating was also reported by Haque et al., when a-C coated disc rubbed against cast
iron pin [57].
Tribofilm formation on hydrogenated DLC coatings are extensively researched by many
authors. It is widely reported that lubricant compounds like MoDTC (molybdenum
dithio-carbamates) and ZnDTP are able to form tribofilm on hydrogenated DLC coatings
[37, 48, 50]. Literature results are so far in agreement with the ZnDTP derived patchy
tribofilm formation on hydrogenated DLC coatings [56, 99]. Apart from these reports,
our results suggest that pad-like tribofilm can also form on hydrogenated DLC coating.
As it can be seen in FESEM and AFM images, formation rate of pad-like structure on
a-C:H 2 are higher than a-C:H 1 which contains less amount of hydrogen. This result
represent that hydrogen may have influence on the formation of pad-like tribofilm. On
the other hand, it seems that pad-like structures on a-C:H coatings only form on asperity
peaks in the contact area which may indicate that generation of pad-like structure were
induced by solid-solid rubbing.
Analysis of tribofilm formation on DLC surfaces revealed a clear difference in the tribofilm structure of doped and non-doped coatings. Features of the pads structure on
the doped-DLC coatings would suggest that metals within DLC coatings increase their
surface activity and their ability to form pad-like structures through the tribochemical
reaction. The enhanced ability of metal-doped DLC to form ZnDTP derived tribofilm
is also reported in the literature [54, 56, 67, 77]
4.3.2
Wear results
The wear rate of all DLC coated pins for self-mated DLC/ DLC contacts as a function
of the lubricants is given in Fig.4.5. Since a-C coated pins showed total delamination in
both lubricants, wear rate calculation was not performed for that coating. The results
show that doped-DLC coatings provided improved wear resistance than non-doped DLC.
It is thought that better adhesion of the coating is the reason for the large difference
in wear of non-doped and doped DLC in pure PAO. In pure PAO oil, the lowest wear
was provided by Cr-DLC tribopair followed by Si-DLC, ta-C DLC, a-C:H 2 and a-C:H 1
tribopair. In PAO+ZnDTP oil, the lowest wear rate was provided by Cr-DLC tribopair
followed by Si-DLC, a-C:H 2, ta-C DLC and a-C:H 1 tribopair. As can be seen in Fig.
78
4.5, the highest overall wear rate was given by a-C:H 1 in PAO and the lowest wear was
given by Cr-DLC in PAO+ZnDTP.
Figure 4.5: Wear rate of the all DLC coated pins for self-mated DLC/DLC contacts
As it is evidenced above that pad-like tribofilm formed on all DLC coated discs except
ta-C coating, and it is clear from the wear results that wear significantly decreased
when ZnDTP additive forms pad-like structure on DLC surfaces. The beneficial effect
of ZnDTP-derived pad-like tribofilm was more obvious in the case of a-C:H 2, Si-DLC
and Cr-DLC. Compared to pure PAO oil, reduction in wear was more than 60% in aC:H 2, Si-DLC and Cr-DLC when lubricated with ZnDTP containing oil. ta-C coating
showed same amount of wear in both PAO and PAO+ZnDTP oils. This suggest that
patchy like tribofilm on ta-C disc don’t have anti-wear performance.
4.3.3
Friction results
Figure 4.6a and 4.6b show the friction coefficient curves of DLC coatings as a function
of sliding cycles for PAO and PAO+ZnDTP oils, respectively. a-C couple gave a gradual
reduction of friction in PAO oil that finally reached to around 0.08 values. a-C couple also
provided similar friction coefficient of 0.08 in ZnDTP containing oil as pure PAO. The
friction coefficient of ta-C coating were high from beginning of the experiment (0.09)
for PAO oil lubrication, and then gradually reduced to 0.025 level. The high initial
Figure 4.6: Friction coefficient curves with sliding cycles for self-mated DLC/DLC
contacts (a) in PAO and (b) in PAO+ZnDTP
79
80
friction coefficient of ta-C under ZnDTP additivated oil reduced sharply from 0.09 to
approximately 0.04 after around 1000 cycles and finally reduced to less than 0.02 after
rubbing 10700 cycles. This sharp reduction in PAO+ZnDTP was probably because of
patchy tribofilm formation. The a-C:H 1 coatings exhibited immediate friction reduction
in the early stage of sliding regardless of oils and then increased slowly to a stable value
of 0.055 in PAO and 0.065 in PAO+ZnDTP. For a-C:H 2 coating, the friction coefficient
decreased rapidly to 0.55 value under PAO and to 0.065 value under PAO+ZnDTP then
remained at that values until end of the tests.
Figure 4.7: Steady state friction coefficients of self-mated DLC/DLC contacts for
PAO and PAO+ZnDTP oils.pdf
Doped DLC coatings exhibited totally different friction behavior than undoped hydrogenated DLC coatings. When Si-DLC was lubricated with base oil PAO, the friction
coefficient was around 0.11 at the initial cycles and then gradually reduced to 0.1 at
the end of the test. On the other hand, the friction coefficient of Si-DLC showed small
reduction first , and then increased quickly to a stationary value of around 0.12 when
ZnDTP additive was added to base oil. For Cr-DLC, gradual friction reduction was
observed in PAO oil, reaching the level observed for a-C:H 1 and a-C:H 2 in pure PAO.
However, friction coefficient of Cr-DLC in PAO+ZnDTP started to rise from the beginning of the test, reaching the value of about 0.09 at the end of the test. The resulting
higher friction coefficient in doped DLC coatings when lubricated with PAO+ZnDTP
81
was probably quick interaction with ZnDTP and generation of tribofilm on the DLC
surfaces.
The average friction coefficients of self-mated DLC/DLC contacts for last 500 cycles are
given in Fig. 4.7. In all coatings except ta-C coating, ZnDTP containing oil showed
higher friction coefficient than pure PAO. This can be correlated with the formation of
pad-like tribofilm on DLC surfaces during rubbing. Formation of patchy like tribofilm
on ta-C in PAO+ZnDTP provided lower friction coefficient than in PAO.
In the case of pure PAO lubrication, ta-C coating showed the lowest friction followed by
Cr-DLC, a-C:H 1, a-C:H 2 and a-C coatings (see Fig. 4.7). The highest average friction
value in PAO oil was obtained with Si-DLC. In ZnDTP containing oil, ta-C coating also
provided the lowest friction followed by a-C:H 2, a-C:H 1, Cr-DLC and a-C coatings.
The highest friction in ZnDTP containing oil was also given by Si-DLC.
4.3.4
Surface analysis with Raman and XPS spectroscopy
Present results have shown that friction and wear properties of various DLC coatings
against self-mated counter material under oil-lubricated conditions vary largely depending on hydrogen content, sp3 hybridization, doping element, and the formation of a
tribofilm. It was generally observed that hydrogen-free ta-C coatings provided the lowest friction and Cr-DLC gave the lowest wear rate both in pure PAO and PAO+ZnDTP
oils.
Low friction behavior of DLC coatings in oil lubricated condition can be attributed to
graphitization of DLC surfaces or the inherent nature of DLC itself. Graphitization of
DLC coating involves transformation of sp3 carbon bonding to graphitic sp2 bonding
at the rubbing surfaces. It is commonly accepted that graphitization of DLC coatings
can be responsible for the low friction performance in dry sliding conditions. Raman
spectroscopy is a powerful, non-destructive method for identification of friction-induced
graphitization of DLC coatings. Degree of graphitization can be assessed by the increase
of ID /IG ratio after sliding, where ID and IG are maximum intensity of D-peak and
G-peak, respectively. Raman analysis of rubbed DLC surfaces were revealed that a
small degree of graphitization occurring on all DLC coatings. It was also observed that
increase of ID /IG ratio was significantly higher in a-C:H 2 coating than the other type
of coatings (Fig. 4.8). On the other hand, it is important to note that graphitization
may occur on nanometer scale of top-most DLC surfaces and Raman spectroscopy may
not probe this modification. Thus, when the friction and Raman analysis results are
interpreted together, it is thought that attribution of low friction performance to the
graphitization can be insufficient. It is believed that the inherent nature of DLC like
Figure 4.8: Raman spectra of (a) a-C:H 2 coating and (b) ta-C coating scanned before
and after sliding in base oil.
82
83
high sp3 hybridization and low hydrogen content is more important than graphitization
for low friction performance of DLC coatings under oil lubricated conditions [64] .
Surface analysis with FESEM and AFM showed that ZnDTP was able to form pad-like
or patchy tribofilms on the DLC surfaces. XPS measurements on rubbed DLC surfaces
were conducted for chemical analysis of tribofilm. XPS results exhibited that Zn, P,
and S were found in the wear scars of all DLC coatings, indicating the decomposition
of ZnDTP additive. No iron peak were detected in any of the DLC coatings. Detailed
spectra of phosphorus P 2p with Zn 3s, sulphur S 2p and Zinc Zn 2p are given in Fig.
4.9.
The oxygen 1s spectra of all DLC coatings and binding energies of O 1s, P 2p, S 2p
and Zn 2p were given in Fig. 4.10 and Table 4.3, respectively. The binding energies
of NBO and BO peaks are assigned to oxygen bonded to phosphorus (P=O, P-O-Zn,
P-O-P) [82, 94, 95]. Presence of Zn 2p peak also suggests the formation of ZnO, ZnS
and Zinc phosphate glass [57, 100]. The S 2p and P 2p peaks represent the formation
of sulfide and phosphate compounds [50, 95]. XPS results confirm the formation of
tribofilm on DLC surfaces. In the present study, the formation of a ZnDTP tribofilm on
DLC coatings and its effect on tribological performance was demonstrated. However, it
is difficult to determine the exact structure and composition of tribofilm. More detailed
analysis is needed to identify the structure of the tribofilm and the interaction between
ZnDTP and carbonaceous surfaces.
Figure 4.9: XPS P 2p (with Zn 3s), S 2p and Zn 2p peaks recorded in the tribofilm
formed on tested DLC coatings
84
Figure 4.10: Detailed XPS spectra of oxygen 1s of tested DLC coatings after sliding
in PAO+ZnDTP
85
Material
a-C
ta-C
a-C:H 1
a-C:H 2
Si-DLC
Cr-DLC
O 1s
530.1
531.8
533.4
530.0
531.8
533.3
529.9
531.8
533.3
530.1
531.8
533.4
530.1
531.8
533.3
530.1
531.8
533.3
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
eV
(%5.4)
(%62.3)
(%33.3)
(%6.4)
(%66.0)
(%27.6)
(%4.1)
(%67.4)
(%28.5)
(%2.9)
(%71.5)
(%25.6)
(%4.9)
(%73.9)
(%21.2)
(%8.3)
(%69.6)
(%22.1)
Binding energies of XPS peaks
P 2p
S 2p
133.5 eV (%78.2)
162.2 eV (%62.5)
134.7 eV (%21.8)
163.6 eV (%37.5)
Zn 2p
1022.5 eV (%92.2)
133.5 eV (%62.2)
134.6 eV (%37.8)
162.3 eV (%55.7)
163.8 eV (%44.3)
1022.5 eV (%83.1)
1024.5 eV (%16.9
133.6 eV (%71.1)
134.6 eV (%28.9)
162.2 eV (%58.4)
163.8 eV (%41.6)
1022.5 eV (%85.0)
133.6 eV (%78.1)
134.8 eV (%21.9)
162.3 eV (%60.6)
163.7 eV (%39.4)
1022.4 eV (%90.2)
133.5 eV (%78.0)
134.7 eV (%22.0)
162.3 eV (%54.9)
163.7 eV (%45.1)
1022.5eV (%85.4)
133.6 eV (%62.1)
134.9 eV (%37.9)
162.2 eV (%60.0)
163.8 eV (%40.0)
1022.4 eV (%85.8)
Table 4.3: Binding energies and concentrations of XPS peaks recorded in the tribofilm formed on tested DLC surfaces.
86
4.4
87
Conclusions
The boundary lubrication properties of six different DLC coatings at self-mated DLC/DLC contacts were evaluated under pure PAO and ZnDTP additivated PAO lubricants. The formation of ZnDTP derived pad-like or patchy tribofilm on various DLC
coatings was evidenced by means of FESEM, AFM and XPS analysis. In the case of pure
PAO lubrication, ta-C DLC exhibited the lowest friction coefficient. Graphitization may
contribute to this low friction behavior of ta-C, but high sp3 and hydrogen-free nature
of the coating are also important factor for the low friction behavior of the ta-C coating.
The friction and wear behaviors of hydrogenated coatings in base oil lubrication clearly
affected by the hydrogen content and doping elements. Higher hydrogen content resulted
in higher wear resistance. Si-DLC and Cr-DLC experienced enhanced wear resistance,
but Si doping into DLC coating resulted in higher friction coefficient.
The addition of ZnDTP greatly influenced the tribological performance of DLC coatings.
Wear of Si-DLC, Cr-DLC and a-C:H coatings reduced more than 50% with the addition
of ZnDTP due to the pad-like tribofilm formation. However, these pad-like tribofilm were
found to promote higher friction coefficients. The formation of patchy tribofilm on ta-C
didn’t provide anti-wear performance but gave the lowest friction coefficient. Hence, it is
suggested that pad-like tribofilm formation is necessary for the improved wear resistance
of DLC coatings. Finally, it was found that hydrogenated and doped DLC coatings were
more reactive with ZnDTP additive in favor of pad-like tribofilm formation. Doping
elements further increased the reactivity of DLC coatings for tribofilm formation.
88
Chapter 5
Conclusions and future outlook
This thesis deals with the experimental investigation of ultra-low friction and wear mechanisms of DLC coatings under oil boundary lubricated conditions. The ultra-low friction
and wear performance of hydrogen-free ta-C DLC lubricated with additivated and nonadditivated PAO oils are presented in chapter 2 and chapter 3, respectively, along with a
discussion of the effects of oil temperature, additive concentration and counter-materials.
In order to obtain better understanding of tribological properties and interactions between the DLC surfaces and oil additives, a wide range of DLC coatings have been
tested and compared with self-mated DLC/DLC contacts under same test conditions in
the same test rig lubricated with PAO oil and the same oil with ZnDTP type anti-wear
additive (chapter 4).
From the experimental results and surface analysis, the main findings of this work can
be summarized as follows:
(1) One of the important finding of this study is that the ta-C DLC coating is able
to provide ultra-low friction coefficient under pure PAO boundary lubrication. A
suggested mechanism for ultra-low friction behavior of ta-C DLC under oil boundary lubrication is the graphitisation of the top most surfaces. We hypothesized
that, with the increasing temperature and additional frictional heating, mobility
of carbon atoms are increased and top most surfaces of ta-C coating is transformed
to graphitic structure which provide ultra-low friction. It is generally believed that
graphite is capable of reducing friction due to layered structure. As noted in the
text, DLC coatings are amorphous, meta-stable form of carbon and they can convert from diamond-like sp3 hybridization to softer graphite-like sp2 hybridization
under the high pressure and frictional heat. Raman spectra analysis indicated
that a slight structural changes of ta-C DLC coatings occurred during the friction
89
90
tests. However, comparing to low friction behavior and Raman spectra analysis
of various DLC coatings in chapter 4, it is noted that graphitisation is not only
the crucial factor for low friction. It is believed that higher sp3 hybridization,
hydrogen content, thickness of transformed structure, and deposition methods of
DLC coatings are also important factors for low friction behavior of DLC coatings
under oil boundary lubricated conditions.
(2) The wear performance of ta-C DLC under pure PAO lubrication strongly depend
on the counterbody material. Tribological results reveal that even though ta-C
DLC coated pin can provide ultra-low friction, the wear resistance of ta-C DLC
coated pin is very poor against steel disc when tested at 80 ◦ C and above temperature in pure PAO. The results are exhibited that rubbing against self-mated ta-C
and germanium disc caused an order of magnitude lower wear rate than rubbing
against the SUJ2 steel disc at 80 and 110 ◦ C. Also, use of additives in DLC/steel
contact greatly increased the wear resistance and durability of ta-C DLC coated
pins. Therefore, it was thought that accelerated wear of the ta-C coating against
steel at higher temperature in PAO can be explained by the tribo-chemical wear.
It is hypothesized that, with the increasing temperature and additional frictional
heating, mobility of carbon atoms are increased and top most surfaces of ta-C
coating is transformed to graphitic structure. Subsequently, thermally activated
carbon atoms on the topmost surfaces of ta-C coating diffuse into the steel surface
and thermo-chemical interaction occur between iron and carbon atoms which cause
higher wear rate. Due to the the extremely low solubility of carbon in germanium
and carbon rich nature of the self-mated ta-C disc, rubbing against germanium and
self-mated ta-C disc eliminate the tribo-chemical wear of ta-C coated pin in PAO.
Besides, passivation of either ta-C surface or ferrous surface through the lubricant
additives prevent the direct contact in DLC/Steel contact and also eliminate the
tribo-chemical wear.
(3) The ta-C DLC also showed ultra-low friction and enhanced wear resistance when
PAO base oil blended with GMO additive. The findings clearly suggest that GMO
molecules physically interact with the dangling σ-bonds of surface carbon atoms of
ta-C DLC to form a thin physisorbed tribofilm. The passivation of these σ-bonds
by hydroxyl -OH group of GMO additive results in weak adhesive and chemical
interaction of the ta-C DLC with the counterbody materials. Also, H- and OHterminated surfaces repulsively interact with Fe surfaces and thus weaken the covalent interaction between the ferrous surface and the ta-C DLC surface. Therefore,
formation of thin physisorbed tribofilm and sliding between OH-terminated layer
and counterbody materials is the ultra-low friction and enhanced wear mechanism
of ta-C DLC in GMO blended PAO oil. It should be noted that effectiveness of
91
GMO on friction and wear is significantly affected by temperature and additive
concentration.
(4) ZnDTP anti-wear additive react directly on the ferrous surfaces. It is found in this
work that ZnDTP related pad-like wear protective tribofilm forms both on steel
disc and ta-C coated pin in DLC/steel combination which resulted in excellent wear
resistance and relatively higher coefficient. On the other hand, no ZnDTP related
pad-like tribofilm was observed neither on ta-C pin nor on ta-C disc surfaces for
DLC/DLC combination with PAO+ZnDTP oil in contrast to DLC/steel tribopair.
Based on these considerations and the detailed surface analysis in the text, it is
concluded that ZnDTP anti-wear additive don’t react directly on the ta-C DLC
surfaces. ZnDTP pad-like tribofilm formation on ta-C pin is explained by the
transferred ferrous molecules when the counterpart was steel. It is thought that
ferrous molecules are transferred to ta-C surfaces at the initial cycles of friction
test and pad-like tribofilm form on those transferred ferrous molecules.
(5) One of the interesting finding of this work that GMO and ZnDTP additives do not
work synergistically. ZnDTP related pad-like tribofilm formation was not observed
even on steel surfaces when lubricated with PAO+GMO+ZnDTP. In all tests and
surface analysis, GMO+ZnDTP containing oil behaved very similar to only GMO
containing oil. It is suggested that the GMO additive suppresses the ZnDTP
tribofilm formation by reacting with ZnDTP in oil or adsorbing and blocking the
solid surfaces.
(6) Overall, the use of additives enhance the wear resistance of ta-C coated pin at
high temperature by eliminating the tribo-chemical wear due to the termination
of surfaces and preventing direct solid-to-solid contact.
(7) Tribological performance of DLC coatings in oil boundary lubricated condition
greatly affected by the hydrogen content and dopant elements. Higher hydrogen
content resulted in higher wear resistance. It is demonstrated that doped DLC
coatings have better wear resistance than non-doped DLC coating which are deposited by same PECVD method. This improvement are attributed to better adhesion and mechanical properties undoubtedly. However, in term of friction, such
generalization cannot be done, since different dopant element resulted in totally
different friction coefficient.
(8) The results from this work clearly show that the ta-C DLC coating was the most
successful DLC coating with regard to the friction coefficient under pure PAO
boundary lubrication. When tested under same test conditions, hydrogen-free
ta-C DLC exhibited the lowest friction coefficient than any other types of DLC
coatings.
Appendices
92
(9) The formation of ZnDTP related tribofilms on all DLC surfaces is clearly evident
and has been demonstrated by means of Fe-SEM, AFM and XPS. It is found that
hydrogenated and doped DLC coatings are more reactive with ZnDTP additive
in favor of pad-like tribofilm formation. Doping elements further increase the
reactivity of DLC coatings for tribofilm formation. Wear of Si-DLC, Cr-DLC
and a-C:H coatings reduced more than 50% with the addition of ZnDTP due to
the pad-like tribofilm formation. However, these pad-like tribofilm are found to
promote higher friction coefficients.
(10) The formation of ZnDTP related patchy tribofilm on ta-C don’t provide antiwear performance but give the lowest friction coefficient. Comparing the ZnDTP
pad-like and patchy tribofilm formation, it is suggested that pad-like tribofilm
formation is necessary for the improved wear resistance of DLC coatings.
The finding in this research indicate that hydrogen-free ta-C DLC coating is the most
successful DLC coating than any other types of DLC coatings with regard to the friction coefficient under boundary lubricated conditions. However, we found that counterbody material is an important environmental parameter for the design of ta-C DLC
coated systems. The wear resistance of ta-C DLC depends very much on the carbon
diffusion/dissolving property of counterbody materials. Therefore, hydrogen-free ta-C
DLC coupling with counterbody materials of stainless steels, hard alloys, oxide coatings,
refractory ceramics or diamond can provide both ultra-low fiction and high wear resistance. Future direction of this research will be examination of the effect of counterbody
materials on the ultra-low friction and wear performance of DLC coatings under boundary lubrication condition. Besides, characterization and identification on the formation
mechanism of pad-like ZnDTP tribofilm on DLC surfaces will be continued.
In summary, the deposition parameter, chemical and structural nature, lubricants, lubricant additives and counterbody material is always a challenging issues for each type
of DLC coatings for optimum tribological performances under oil boundary lubricated
condition. It has been showed that various DLC coatings exhibit ultra-low friction
and excellent wear resistance when used with proper base oil and lubricant additives.
The findings in this work and previous research by different authors have shown that
boundary lubrication performance of DLC coatings can be optimized by controlling the
chemical and structural nature of DLC coatings and synergies between additives for
specific DLC coatings.
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Publication List
[1] Haci Abdullah Tasdemir, M. Wakayama, T. Tokoroyama, H. Kousaka, N. Umehara, Y. Mabuchi and T. Higuchi, “Ultra-low friction of tetrahedral amorphous
diamond-like Carbon (ta-C DLC) under oil boundary lubrication in Poly alphaolefin (PAO) with additives”, Tribology International 65 (2013), pp. 286-294
[2] Haci Abdullah Tasdemir, M. Wakayama, T. Tokoroyama, H. Kousaka, N.
Umehara, Y. Mabuchi and T. Higuchi, ”Wear behaviour of tetrahedral amorphous diamond-like carbon (ta-C DLC) in additive containing lubricants”, Wear
307 (2013), pp. 1-9
[3] Haci Abdullah Tasdemir, M. Wakayama, T. Tokoroyama, H. Kousaka, N. Umehara, Y. Mabuchi, and T. Higuchi, “The effect of oil Temperature and additive concentration on the wear of non-hydrogenated DLC coating”, Tribology International
77 (2014), pp. 65-71
[4] Haci Abdullah Tasdemir, T. Tokoroyama, H. Kousaka, N. Umehara, Y. Mabuchi,
“Influence of ZnDTP tribofilm formation on the tribological performance of selfmated DLC/DLC contacts in boundary lubrication”, Thin Solid Films (2014) in
press, http://dx.doi.org/10.1016/j.tsf.2014.05.004
[5] Haci Abdullah Tasdemir, Takayuki Tokoroyama, Hiroyuki Kousaka, Noritsugu
Umehara, Yutaka Mabuchi, “Friction and wear performance of boundary-lubricated
DLC/DLC contacts in synthetic base oil”, Procedia Engineering 68 ( 2013 ) 518524.
102
International Conferences
Umehara, Yutaka Mabuchi and Tsuyoshi Higuchi, “Ultra-low friction mechanism
of Diamond-Like Carbon (DLC) under oil boundary lubrication”, 39th Leeds-Lyon
Symposium on Tribology, 4-7 September 2012, Leeds, UK
Umehara, Yutaka Mabuchi and Tsuyoshi Higuchi, “Use of Diamond –Like Carbon (DLC) In Engine Components For Ultra-Low Friction and Environmentally
Friendly Lubrication”, 5th International Symposium on Advanced Plasma Science
and its Applications for Nitrides and Nanomaterials, pp. 91, ISPlasma2013, January 28-February 2013, Nagoya, Aichi, JAPAN
[3] Haci Abdullah Tasdemir, Masaharu Wakayama, Takayuki Tokoroyama, Hiroyuki Kousaka, Noritsugu Umehara, Yutaka Mabuchi and Tsuyoshi Higuchi,
“Tribological Properties of Tetrahedral Amorphous Carbon Coating under Boundary Lubrication: Effect of Load”, Proceedings of The 5th International Conference
on Manufacturing, Machine Design and Tribology (ICMDT 2013), pp. 264-266,
May 22-25, 2013, Busan, KOREA
[4] Haci Abdullah Tasdemir, Masaharu Wakayama, Takayuki Tokoroyama, Hiroyuki Kousaka, Noritsugu Umehara, Yutaka Mabuchi and Tsuyoshi Higuchi,
“Role of Temperature on the Ultra-low Friction and Wear of Diamond-like Carbon
under Oil Boundary Lubrication”, Proceedings of 5th World Tribology Congress
(WTC 2013), September 8-13, Torino, Italy, ISBN 978-88-908185-09
Umehara, Yutaka Mabuchi, “Friction and wear performance of boundary-lubricated
DLC/DLC contacts in synthetic base oil”. Malaysian International Tribology Conference (MITC2013), November 18-20, Sabah, Malaysia
103
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Boundary Lubrication Mechanisms of Diamond

Transcription

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