J Synth Lube 2008

JOURNAL OF SYNTHETIC LUBRICATION
J. Synthetic Lubrication 2008; 25: 31–41
Published online in Wiley InterScience
(www.interscience.wiley.com) DOI: 10.1002/jsl.45

Improved additives for multiply alkylated
cyclopentane-based lubricants

Stephen C. Peterangelo1, Lois Gschwender2, Carl E. Snyder, Jr.3,*,†,
William R. Jones, Jr.4, QuynhGiao Nguyen5 and Mark J. Jansen6
1
Research, AK Steel Westchester, OH, USA
2
Materials Directorate, US Air Force Wright-Patterson Air Force Base, OH, USA
3
Editorial, Journal of Synthetic Lubricants; Lubricants, US Air Force Base, OH, USA
4
Sest, Inc. Middleburg Heights, OH, USA
5
Durability and Protective Coatings Branch/Structures and materials Division, NASA Glenn Research Center
Cleveland, OH, USA
6
Aurora Flight Sciences, MANASSA, VA, USA

ABSTRACT

Multiply alkylated cyclopentanes (MACs) are replacing heritage (mineral oil-based) spacecraft lubricants because
of their excellent performance and low volatility. While MACs have acquired an increasingly prominent role,
soluble additives with similarly low volatility are lacking. In this study, the performance of specially designed
candidate high-molecular-weight/low-volatility phosphate additives was compared with the performance of con-
ventional phosphate and lead naphthenate additives currently used in space. The candidate additives were
equivalent or superior to the currently used additives in both conventional (atmospheric) and vacuum wear tests.
Volatility studies revealed superior candidate additive performance compared with currently used additives. In
addition, surface chemical analysis of the wear surfaces provided a better understanding of the anti-wear protec-
tive ﬁlms formed by these additives. Copyright © 2008 John Wiley & Sons, Ltd.

KEY WORDS: additives; anti-wear; adverse environment tribology; aerospace tribology; space vehicles; synthetic
lubricants

BACKGROUND AND INTRODUCTION

The increasing cost and indispensability of satellite systems has necessitated an increase in their life
expectancy, from an average of 5–8 years to 15 years and beyond [1]. To accommodate these longer-
life missions, lubricants used in these craft must also advance. Recently, there has been a shift away
from mineral oil-based lubricants to synthetic multiply alkylated cyclopentane (MAC)-based lubri-
cants [1, 2]. Unfortunately, lubricant additives possessing low volatility have not been developed.

*Correspondence to: Carl E. Snyder, Jr, US Air Force Base, OH, USA.
†
E-mail: ed.snyder@wpafb.af.mil

Copyright © 2008 John Wiley & Sons, Ltd.

32 S. C. PETERANGELO ET AL.

In many cases, arylphosphates and lead naphthenate additives present in heritage mineral oils are being
used in these synthetic spacecraft lubricants.
In addition to wear and friction reduction in an ambient atmosphere setting, volatility and vacuum
wear characteristics must also be considered when evaluating lubricants for space [3, 4]. MACs are
superior to mineral oil-based lubricants in these categories [5, 6]. Their performance can be further
enhanced by the inclusion of less-volatile additives [3, 7]. These additives must have several essential
characteristics. Firstly, and most importantly, they must reduce wear and friction. They must also be
soluble in the base fluid (although in greases, this characteristic is not as important), and must be
chemically stable in both space and terrestrial environments [3].
In the past, experimental low-volatility additives were not commercially available. However, the
situation is changing because lubricant effectiveness (including additive volatility) is becoming an
increasingly important factor in spacecraft lifetime, and improved, lower-volatility additives are one
of the best ways of improving boundary lubrication in space [4, 8]. The benefits of additive research
are now becoming increasingly apparent and necessary.

EXPERIMENTAL

Materials
Base Fluid. The base fluid employed in this study was used as received from Nye Lubricants, Inc.,
Fourhaven, MA, USA as Pennzane X2000. This is an MAC with a nominal structure tris(2-octydo-
decyl) cyclopentane. Various additives were blended in percentages ranging from 0.25 to 3.00% by
weight.

Additives. Hydrocarbon lubricants are a mature technology. Consequently, there is a wide variety of
commercially available additives available to enhance their performance [9]. The majority of these
additives, however, were developed for terrestrial applications, and the additives are therefore often
unsuitable for the unique demands of the space environment [10]. The additives explored in this study
consisted of two groups: two novel additives synthesized in-house (additives in formulations A and
B) and two commercial additives currently used in space lubricants (additives in formulations C and
D). A matrix consisting of five formulations was developed, and is shown in Table I.

Table I. Formulations developed for this study.

Test fluid Weight % Description Additive formula

X2000 Base — Synthetic hydrocarbon tris(2-octyldodecyl) N/A
cyclopentane
A 0.25 Base fluid + Tris(p-chlorophenoxyphenyl) [Cl—φ—O—φ—O]3P:O
phosphate
B 1 Base + chlorinated alkylated tris(phenoxyphenyl) Mixture
phosphate
C 3 Base + lead naphthenate Pb salt of naphthenic acids
D 1 Base + mixture of arylphosphates [R—φ—O]3P:O
(Syn-o-Ad 8478) R=H, t-butyl

A120 Nobel Chemicals, Chicago, IL, USA.

Copyright © 2008 John Wiley & Sons, Ltd. J. Synthetic Lubrication 2008; 25: 31–41
DOI: 10.1002/jsl

ADDITIVES FOR MULTIPLY ALKYLATED CYCLOPENTANE-BASED LUBRICANTS 33

Analysis
Vacuum Thermogravimetric Analysis (TGA). A small (∼10 mg) sample of the fluid or additive was
placed in an aluminium sample pan. The apparatus was swept with nitrogen for approximately 10
minutes, and subsequently evacuated to 33 Pa (0.25 torr). The temperature was programmed to increase
at 10 °C/min, from 25 to 600 °C. Values of greatest importance are the onset of volatilization (T0) and
the temperature at which half the sample has been volatilized (T1/2). The higher the T1/2 and T0, the
less volatile the sample. Vacuum simulates the space environment and ensures that evaporation occurs
before lubricant degradation. Thermal decomposition results in an uneven weight loss that is difficult
to repeat, whereas evaporation provides a smoother, repeatable weight loss. Differential scanning
calorimetry experiments confirmed that thermal degradation of the lubricant did not occur during the
vacuum TGA experiments [1].

Spectroscopy. X-ray photoelectron spectroscopy (XPS) was performed on the reciprocating tribom-
eter discs following the tests to obtain chemical information about wear surfaces. A spot size of 300 µm
was used. Measurements were taken inside and outside of the wear track. No sputtering was used.
Prior to insertion into the vacuum chamber, samples were cleaned sequentially for 10 minutes each
in reagent-grade hexanes, acetone and methanol in an ultrasonic bath. Of primary interest was the
presence of phosphate residue (for formulations A, B and D) or lead (for formulation C) in the wear
track (and the absence of those same elements outside of the wear track), indicating chemical reaction
and bonding of the additive with the metal.

Tribometry
Four-ball Wear. Four-ball wear was conducted per ASTM D4172 with the following deviations. A
fitted plexiglas cage was placed around the apparatus and purged with dry nitrogen for at least 15
minutes prior to and during the course of the test, and 440C stainless steel balls were used. The cal-
culated initial stress of the test was 3.41 GPa (495 kpsi). Friction was measured on a torsion bar 75 mm
from and orthogonal to the axis of rotation. Test conditions were: 392 N (40 kg) load; 12.7 mm (0.50 in.)
440C steel balls; 500 rpm; 1 hour; 75 °C; dry nitrogen [relative humidity (RH) < 8%].

Vacuum Four-ball. A four-ball tribometer, operating in the boundary lubrication regime, was used to
determine the steady wear rates in vacuum for each test fluid. Specimen configuration is the same as
the conventional four-ball apparatus, except that 9.5-mm (3/8-in.)-diameter 440C stainless steel bearing
balls (grade 10) were used. A complete description of this device appears in Masuko et al. [11].
Test conditions included: 200 N load (initial Hertzian stress of 3.5 GPa); vacuum (<6.7 × 10−4 Pa);
28.8 mm/s sliding speed; and 23 °C. A complete test was 4 hours in length, with interruptions every
hour for wear scar measurements. Steady-state wear rates were determined from the slope of the
straight line, which was obtained from a plot of wear volume (calculated from the wear scars) as
a function of sliding distance.

Reciprocating Tribometer. The reciprocating tribometer conditions were as follows: 100 °C; 9 mm
stroke; 6 Hz; 250 N load; 5 hours test duration; 6.35 mm (0.25 in.) 440C steel ball on a 440C steel disc;
N2 purge; RH 0.02% or less. The discs were polished to a surface roughness (Ra) of between 0.025
and 0.051 µm, with no significant differences in wear area being noted at the varying levels of

DOI: 10.1002/jsl


roughness utilized in this test. The calculated initial stress of this test was 3.95 GPa. Test time was 5
hours in order to obtain meaningful differences in data for the base fluid vs the formulated fluids.
Three tests were performed on each sample, with the average and standard deviation reported.

Vacuum Spiral Orbit Tribometer (SOT). The SOT was developed by Pepper et al. [12]. It simulates
conditions observed in angular contact bearings. It is a retainer-less steel thrust bearing with a single
440C bearing ball (12.7 mm), loaded between two 50.8-mm-diameter 440C flat plates. The lower plate
is stationary while the upper plate is rotated at ∼200 rpm. The ball is lubricated with ∼50 µg of lubri-
cant. The lubricant is consumed during the rolling/pivoting process, which eventually results in test
failure, characterized by a friction coefficient of >0.28. Normalized lubricant lifetime was calculated
by dividing the number of ball orbits by the weight of the lubricant on the ball. Test conditions included
a Hertzian stress of 1.5 GPa, vacuum (<1.3 × 10−6 Pa) and room temperature (∼23 °C). A complete
description appears in Jones et al. [13].

RESULTS AND DISCUSSION

Tribometry
Four-ball Wear (Dry Nitrogen). Reduction of friction and wear in the boundary lubrication regime
are the two primary reasons for using additives [3, 9]. The results obtained by the authors indicate
that these two properties are rather closely linked. As can be seen in Figure 1, a plot of the average
coefficient of friction vs wear scar area produced for the standard four-ball test data shows a linear

0.50

0.45

0.40

0.35

Base Fluid
Coefficient of Friction

0.30
R2 = 0.9363
0.25

0.20

Additive B Additive D
0.15

0.10

0.05
Additive C
Additive A
0.00
0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0
Wear Scar Area (mm2)

Figure 1. Four-ball friction vs wear (500 rpm, 75 °C, N2 purge, 392 N load, 1 hour, 440C steel).

DOI: 10.1002/jsl


relationship, as shown by the computer plotted trend line (R2 = 0.9363). With this in mind, compari-
sons between friction and wear correlation can be used to screen lubricants (Figure 2). The unformu-
lated MAC base fluid showed a relatively high coefficient of friction (µ) throughout the test, with an
average of 0.299 ± 0.106 standard deviation. In-house additives in formulations A and B showed
marked friction reduction over the entire course of the test, with an average µ of 0.084 ± 0.004 and
0.099 ± 0.019, respectively, indicating superior boundary lubrication. Frictional characteristics of the
additive in C (lead naphthenate) were actually superior to the additives in A and B (µ = 0.070 ± 0.004).
The additive in C was an anomaly, in terms of correlation of four-ball friction and wear, however,
because its wear was higher than the wear generated by A and B, while showing lower frictional
characteristics (C gave a 67% wear reduction from the base fluid vs 80% for A and 77% for B — see
Figure 3). The lack of correlation between the wear scar and frictional forces experienced by the addi-
tive in C as compared with the additives in A, B and D is likely related to the different chemistries of
the additives (lead vs phosphorus). The commercial phosphate additive (in formulation D) showed
some improvement over the base fluid, but the friction was high from 8 to 21 minutes before it reduced
to a lower level for the remainder of the test. The average coefficient of friction was lower than the
base, at 0.177 ± 0.098, but still higher than formulations A and B. The standard deviation was almost
an order of magnitude higher as a result of the period of high friction at the beginning of the test. This
is probably due to the chemical reaction of the additive with the metal surface, requiring more activa-
tion energy to form a lubricious film than with the other additives. There was no correlation between
friction and wear for any formulation in the reciprocating tribometer test.

0.5

0.4
Coefficient of Friction

0.3
Base

Base
Additive D
A
0.2 Additive B
B
C
D

0.1

Additive A
Additive C

0.0
0
2
4
6
8
10
12
14
16
18
20
22
24
26
28
30
32
34
36
38
40
42
44
46
48
50
52
54
56
58

Time (min)

Figure 2. Friction of the four-ball test (75 °C, 40 kg, 440C steel, 1 hour, N2).

DOI: 10.1002/jsl


3.0

2.5
2.31

2.0

1.64
Wear Scar

Reciprocating Tribometer (mm^2)
1.5
1.38 Four Ball (diameter in mm)

1.0

0.75

0.59 0.61 0.59
0.52
0.46
0.5
0.33

0.0
MAC A B C D
Fluid

Figure 3. Four-ball and reciprocating tribometer wear results.

Reciprocating Tribometer. Similar wear-reducing characteristics were noted in both the four-ball and
reciprocating tribometer tests (Figure 3). In both tests, the greatest wear reduction, in comparison with
the base fluid, was obtained with the additive in the specially formulated B fluid. In the reciprocating
tribometer test, the wear area of the formulation B was nearly 76% less than the base fluid. This was
compared with the 55% reduction for the lead naphthenate (formulation C) and the 57% for the com-
mercial phosphate (formulation D). In all tests, the untreated base fluid showed the highest wear as
expected.

Vacuum Four-ball. Vacuum four-ball wear rates for all test fluids appear in Figure 4. The unformulated
base stock yielded the highest wear rate. All additive formulations reduced the wear rate, with formu-
lation B yielding the best results. Except for the fact that all additives reduced wear in both series of
tests, there was no correlation between the standard four-ball tests and the vacuum four-ball wear
results. This is not surprising since the contact conditions are quite different (500 vs 100 rpm and 75
vs 23 °C). The standard four-ball conditions cause much higher contact temperatures, resulting in more
rapid or different additive/surface reactions.

SOT Results. Relative lifetimes for all test fluids appear in Figure 5. Formulations B, C and D showed
no improvements in lifetime compared with the base fluid. Formulation A yielded a lifetime twice that
of the base fluid. This correlates well with the standard four-ball test in that the lowest wear scar
was obtained with formulation A. However, this may be simply fortuitous in that the SOT measures

DOI: 10.1002/jsl


2.00E-10

1.80E-10

1.60E-10
Mean Wear Rate (mm3 / mm) 1.40E-10

1.20E-10

1.00E-10

8.00E-11

6.00E-11

4.00E-11

2.00E-11

0.00E+00
Base A B C D
Formulation

Figure 4. Mean wear rate of the vacuum four-ball test (3.5 GPa, vacuum <6.7 × 10−4 Pa, room temperature).

10000

9000

8000

7000
Ball Orbits /mg Lubricant

6000

5000

4000

3000

2000

1000

0
Base A B C D
Formulation

Figure 5. Relative lifetime results from the spiral orbit rolling contact tribometer
(1.5 GPa, vacuum <1 × 10−6 torr, room temperature).

lubricant consumption and not wear. (It should be noted that most space bearings that fail or whose
performance deteriorates are due to loss of lubricant and not due to wear.)

Surface Analysis
XPS data supports the hypothesis of chemical (especially phosphate) bonding taking place (Table II),
as can be seen by the phosphate residue observed exclusively in the wear track of the additive-rich
specimens. The atomic percentages of phosphate were observed for the various formulations. It was
found that the additive in B left a higher amount of phosphorous in the wear track (8.3% of surface

DOI: 10.1002/jsl


Table II. XPS analysis of the surface of the reciprocating tribometer test specimen (atomic percentage of
reactive elements on the surface).

Formulation Element Inside wear track (%) Outside wear track (%)

Base — 0 0
A P 2.82 0
B P 8.28 0
C Pb 5.74 3.91
D P 1.54 0

400
378

351 357
350
328

300

250
T 1/2(°C)

204
200

150

100

50

0
Base A B C D
Fluid

Figure 6. Volatility; vacuum TGA (33 Pa, 10 °C/min to 600 °C).

atoms) than the additive in A (2.8%) or D (1.5%). No phosphorous was found outside the wear track
on any of these specimens. Lead was found in a concentration of 5.7% inside the wear track and 3.9
% outside the wear track with formulation D. This lends credence to the hypothesis of different chem-
istries of the additives leading to different relationships between friction and wear.

Volatility
Volatility studies are very important when considering lubricants for space applications. Ideally, addi-
tives should have volatility comparable with the base ﬂuid volatility. If the additive evaporates faster
than the base ﬂuid, its usefulness will diminish. Additives need to be present in a minimum concentra-
tion to be effective. For example, lead naphthenate must be used at a 3% or higher concentration.
High-additive concentrations are undesirable because additives may be corrosive or evaporate onto
critical optical surfaces. For these reasons, the additive in D is usually used at 1%. As can be seen
from Figure 6, additives in A, B and C had a higher T1/2 (lower volatility) than the additive in D and

DOI: 10.1002/jsl


Figure 7. Vacuum TGA of additive C vs base fluid.

the base fluid. This is significant because it implies that the additive in D may evaporate more quickly,
leaving behind a fluid with compromised lubricating ability. The additive in C showed a high T1/2 but
left a significant residue after the TGA was complete (20% of original weight), and further heating
did not volatilize this residue. In addition, it demonstrated multi-step evaporation (Figure 7), because
the lead naphthenate is in a mineral oil carrier (∼25%) that evaporates first. All additives, except the
additive in C, showed a volatilization pattern similar to that shown by the base fluid in Figure 4.
Although the additive in A imparted marked friction reduction and had excellent volatility, over
several months’ time it precipitated out of solution, even at 0.25 wt% concentration. Moderate success
was obtained by the addition of a high-molecular-weight ester to the solution, but the ester had to be
used in such high concentrations (30–40%) that its use became impractical. The additive in A would
therefore be more suited to grease formulations rather than in liquid lubricants.

SUMMARY

As longer life and better performance are increasingly required of lubricants, new soluble additives
must be developed to fully exploit their capability. The two experimental additives explored in
this paper showed improvement in areas tested over the commercially used additives for MAC
base fluids.

DOI: 10.1002/jsl


The additives in A and B showed noticeable improvement over the commercially available lubri-
cants in volatility and reduction of friction and wear in a dry nitrogen atmosphere. These findings were
supported by XPS data showing that phosphate chemically reacted to the test discs used in the recip-
rocating tribometer. In vacuum tribology tests, formulation A yielded a lower wear rate than the base
fluid and an increased lifetime in a bearing simulator compared with all other fluids. B exhibited the
lowest wear rate of all fluids, but no improvement in lubricant lifetime. Although C exhibited friction-
reducing properties, it was found that it did not reduce wear. C also demonstrated low volatility, but
again, upon closer inspection of the data, there was a significant residue left behind, and the evapora-
tion occurred in a manner indicative of a non-uniform molecular weight composition.

CONCLUSIONS

• High-molecular-weight/low-volatility phosphate additives offer a marked improvement over the
currently used commercial additives in volatility and wear reduction considerations.
• Specifically matching lubricant and additive is a profitable endeavour when optimal lubricant effec-
tiveness is a priority.
• A variety of characteristics must be considered when evaluating additives, as some show improve-
ments in certain areas while having flaws in others.

ACKNOWLEDGEMENTS

The authors thank Dr K. C. Eapen, Dr Loomis Chen and Ms Grace Chen for the synthesis of the in-house addi-
tives used in this study.

REFERENCES

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Lubrication Engineering 2000; 56(12):25–31.
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Lubrication Engineering 1991; 47(7):586–591.
3. Fleischauer PD, Hilton MR. Assessment of the tribological requirements of advanced spacecraft mechanisms. Materials
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4. Hilton MR, Fleischauer PD. Lubricants for high-vacuum applications. In Friction, Lubrication, and Wear Technology, Dlau
P (ed.), ASM Materials: Materials Park, OH, 1992; 150–161.
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8. Jones WR Jr., Jansen MJ. Space tribology. NASA TM 2000-209924, March 2000.
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11. Masuko M, Jones WR Jr., Jansen R, Ebihara B, Pepper SV, Helmick LR. A vacuum four-ball tribometer to evaluate liquid
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TP-3629, 1996.
13. Jones WR Jr., Pepper SV, Jansen MJ, Nguyen Q, Kingsbury E, Loewenthal SH, Predmore R. A new apparatus to evaluate
lubricants for space applications — the spiral orbit tribometer (SOT). NASA TM 2000-209935, 2000.

DOI: 10.1002/jsl

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