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Design and Development of an Experimental Apparatus toStudy Jet Fuel Coking in Small Gas Turbine Fuel Nozzles
by
Jason Jian Liang
A thesis submitted in conformity with the requirementsfor the degree of Master of Applied Science
Graduate Department of Aerospace Science and EngineeringUniversity of Toronto
© Copyright 2013 by Jason Jian Liang
Abstract
Design and Development of an Experimental Apparatus to Study Jet Fuel Coking in
Small Gas Turbine Fuel Nozzles
Jason Jian Liang
Master of Applied Science
Graduate Department of Aerospace Science and Engineering
University of Toronto
2013
An experimental apparatus was designed and built to study the thermal autoxidative
carbon deposition, or coking, in the fuel injection nozzles of small gas turbine engines.
The apparatus is a simplified representation of an aircraft fuel system, consisting of a
preheating section and a test section, which is a passage that simulates the geometry,
temperatures, pressures and flow rates seen by the fuel injection nozzles. Preliminary
experiments were performed to verify the functionality of the apparatus. Pressure drop
across the test section was measured throughout the experiments to monitor deposit
buildup, and an effective reduction in test section diameter due to deposit blockage was
calculated. The preliminary experiments showed that the pressure drop increased more
significantly for higher test section temperatures, and that pressure drop measurement
is an effective method of monitoring and quantifying deposit buildup.
ii
Acknowledgements
Setting up a test rig comes with a steep learning curve, as I have come to realize. How-
ever, I am thankful that many people where there to help make it easier. First, many
thanks to Dr. Omer Gulder, for providing me with valuable guidance and advice, and for
entrusting me with major project decisions; to Pratt & Whitney Canada, for providing
funding and feedback for the project.
I want to give thanks to Ivo Fabris, who patiently taught me to operate the equipment
and answered my many questions; to Owen Wong, who put together the previous test
rig, without which my task would have been much more difficult.
I am thankful for Frank Yuen, whose expertise and assistance in troubleshooting, equip-
ment purchasing and safety were most valuable; for John Liu, who provided important
numerical simulation data and covered my lunch breaks during the long test runs; for
Leon Li and the Flow Control and Experimental Turbulence lab at UTIAS, for loaning
me testing equipment and helping me to set up data acquisition.
Lastly, to my parents, and Jessie, who kept me going through rain and shine, I am always
grateful.
iii
Contents
1 Introduction 1
1.1 Jet Fuel and Aircraft Thermal Management . . . . . . . . . . . . . . . . 1
1.2 Jet Fuel Thermal Stability . . . . . . . . . . . . . . . . . . . . . . . . . . 2
2 Project Definition 4
2.1 Fuel Injector Coking . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4
2.2 General Experimental Approach . . . . . . . . . . . . . . . . . . . . . . . 5
3 Background and Literature Review 10
3.1 Thermal Stability of Jet Fuels . . . . . . . . . . . . . . . . . . . . . . . . 10
3.2 Mechanics and Chemistry of Thermal Stability . . . . . . . . . . . . . . . 11
3.2.1 Autoxidation Mechanism . . . . . . . . . . . . . . . . . . . . . . . 11
3.2.2 Pyrolysis Mechanism . . . . . . . . . . . . . . . . . . . . . . . . . 12
3.3 Factors that Affect Thermal Stability . . . . . . . . . . . . . . . . . . . . 12
3.3.1 Dissolved Oxygen . . . . . . . . . . . . . . . . . . . . . . . . . . . 12
3.3.2 Temperature . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13
3.3.3 Flow Rate and Residence Time . . . . . . . . . . . . . . . . . . . 14
3.4 Past Experimental Studies and Apparatuses . . . . . . . . . . . . . . . . 14
3.4.1 Dynamic Flow Tests . . . . . . . . . . . . . . . . . . . . . . . . . 14
3.4.2 Fuel Injector Studies . . . . . . . . . . . . . . . . . . . . . . . . . 15
3.5 Analytical Techniques . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16
iv
3.5.1 Temperature Programmed Oxidation . . . . . . . . . . . . . . . . 16
3.5.2 Pressure Drop Measurement . . . . . . . . . . . . . . . . . . . . 17
3.5.3 Spectroscopic Techniques for Chemical Analysis . . . . . . . . . . 19
3.6 Jet Fuel Thermal Stability Research at UTIAS . . . . . . . . . . . . . . 20
4 Experimental Apparatus 22
4.1 Overview of the Experimental Apparatus . . . . . . . . . . . . . . . . . 23
4.2 Fuel Pump and Back Pressure Regulator . . . . . . . . . . . . . . . . . . 27
4.2.1 Syringe Pumps . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
4.2.2 Back Pressure Regulator . . . . . . . . . . . . . . . . . . . . . . . 27
4.3 Fuel Preheater . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27
4.3.1 Oil Bath for Preheating . . . . . . . . . . . . . . . . . . . . . . . 29
4.3.2 Silicone Bath Fluid . . . . . . . . . . . . . . . . . . . . . . . . . . 30
4.4 Test Section . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 31
4.4.1 Test Section Design Drivers . . . . . . . . . . . . . . . . . . . . . 31
4.4.2 Test Section Designs . . . . . . . . . . . . . . . . . . . . . . . . . 32
4.5 Test Section Heater . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35
4.5.1 Brass Heating Block . . . . . . . . . . . . . . . . . . . . . . . . . 36
4.5.2 Band Heater and Temperature Controller . . . . . . . . . . . . . 37
4.6 Pressure Drop Measurements . . . . . . . . . . . . . . . . . . . . . . . . 37
4.7 Data Acquisition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42
4.8 Other Components . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 42
5 Experimental Methodology and Procedures 44
5.1 Overview . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44
5.2 Numerical Simulation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44
5.3 Description of Experimental Procedures . . . . . . . . . . . . . . . . . . 46
5.3.1 Preparation and Set Up . . . . . . . . . . . . . . . . . . . . . . . 46
v
5.3.2 Run Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47
5.3.3 Heat-Up . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47
5.3.4 Steady State Operation . . . . . . . . . . . . . . . . . . . . . . . . 49
5.3.5 Shutdown and Purging . . . . . . . . . . . . . . . . . . . . . . . . 51
5.3.6 Data Collection . . . . . . . . . . . . . . . . . . . . . . . . . . . . 52
6 Results and Discussion 54
6.1 Fuel Batch and Jet Fuel Thermal Oxidation Tester (JFTOT) Results . . 54
6.2 Apparatus Verification Experiment Conditions . . . . . . . . . . . . . . . 55
6.3 Pressure Drop Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57
6.3.1 Test Section Temperature Measurement and Profiles . . . . . . . 62
6.4 Pressure Drop Dependence on Temperature . . . . . . . . . . . . . . . . 62
6.4.1 Viscosity as a Function of Temperature . . . . . . . . . . . . . . . 62
6.4.2 Handbook Viscosity Data . . . . . . . . . . . . . . . . . . . . . . 64
6.4.3 Semi-Empirical Approximation . . . . . . . . . . . . . . . . . . . 64
6.4.4 Effect on Pressure Drop Measurements . . . . . . . . . . . . . . . 65
6.5 Sources of Experimental Error . . . . . . . . . . . . . . . . . . . . . . . . 68
7 Conclusion 71
7.1 Recommendations and Future Work . . . . . . . . . . . . . . . . . . . . . 72
Bibliography 74
vi
List of Tables
2.1 Commercial aircraft fuel environment . . . . . . . . . . . . . . . . . . . . 7
3.1 Overview of paraffin autoxidation mechanism . . . . . . . . . . . . . . . 11
4.1 Commercial stainless steel tubing sizes . . . . . . . . . . . . . . . . . . . 33
5.1 Revised Pratt & Whitney Canada test matrix for fuel temperature tests . 45
5.2 Revised Pratt & Whitney Canada test matrix for wetted wall temperature
tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45
5.3 Oil bath temperature settings for various test section inlet temperatures . 49
6.1 Measured steady state Tin and Tout . . . . . . . . . . . . . . . . . . . . . 63
6.2 Comparison of pressure drop measurement against literature . . . . . . . 69
vii
List of Figures
3.1 Temperature regimes of jet fuel deposit formation . . . . . . . . . . . . . 13
3.2 Schematic of UTIAS single tube flow test apparatus . . . . . . . . . . . . 21
4.1 Photograph of the experimental apparatus . . . . . . . . . . . . . . . . . 24
4.2 Experimental apparatus overview . . . . . . . . . . . . . . . . . . . . . . 25
4.3 Insulation for non-heated components . . . . . . . . . . . . . . . . . . . . 25
4.4 Nitrogen purging schematic diagram . . . . . . . . . . . . . . . . . . . . 26
4.5 Memmert ONE 45 oil bath . . . . . . . . . . . . . . . . . . . . . . . . . . 29
4.6 EDM test section example (0.009 in.) . . . . . . . . . . . . . . . . . . . . 33
4.7 Brass block with thermocouple probes . . . . . . . . . . . . . . . . . . . 38
4.8 Thermocouple contact mechanism in brass heater block . . . . . . . . . . 39
4.9 Test section heater wiring diagram . . . . . . . . . . . . . . . . . . . . . 40
4.10 Placement of pressure taps for pressure drop measurement . . . . . . . . 41
4.11 Overpressure protection valve . . . . . . . . . . . . . . . . . . . . . . . . 41
5.1 Temperature-time profile of steady state operation . . . . . . . . . . . . . 50
6.1 Pressure drop data for Run 2 . . . . . . . . . . . . . . . . . . . . . . . . 60
6.2 Pressure drop data for Run 3 . . . . . . . . . . . . . . . . . . . . . . . . 61
6.3 Axial temperature profiles . . . . . . . . . . . . . . . . . . . . . . . . . . 63
6.4 Viscosity-temperature curve comparison . . . . . . . . . . . . . . . . . . 66
6.5 Geometry for calculating pressure drop . . . . . . . . . . . . . . . . . . . 67
viii
Chapter 1
Introduction
1.1 Jet Fuel and Aircraft Thermal Management
Jet fuel is the source of power for many types of aircraft, from small turboprops to
heavy jet aircraft. Jet fuel is a middle-distillate product of the crude oil refinery process.
The most common commercial jet fuel in North America are Jet A and Jet A-1, which
are kerosene-type jet fuels. JP-8 is the primary fuel of the U.S. Air Force [1], and it is
essentially Jet A with additives added to improve lubricity, and thermal stability. These
fuels are composed of many different types of compounds, but primarily of hydrocarbons
with carbon numbers from C8 to C16 [1], with the most prevalent being C11 [2]. Other
compounds, such as dissolved oxygen, metals, sulfur are present in the fuel in minute
amounts, but play a significant role in the thermal stability of the fuel.
In addition to powering the aircraft, one very important function of jet fuel is its
ability to be used as a heat sink. Many aircraft systems generate large quantities of heat
and require cooling [3]. On modern aircraft, these cooling demands are ever increasing.
For example, modern engines operate at very high combustion temperatures to increase
efficiency, and their lubrication systems require cooling as well [4]. In addition, there are
ever more avionics and electronic equipment that generate an excessive amount of waste
1
Chapter 1. Introduction 2
heat that must be removed [5].
Using the fuel system as a heat sink offers several advantages as compared to another
cooling method, which uses engine bleed air [6] or ram air [7] as the cooling medium. First,
jet fuel has a greater heat sink capacity under a wider range of operating conditions than
air cooling. The disadvantage to using air as a coolant is that as flight speeds increase,
the stagnation temperature of the air rises as well, thus decreasing the effectiveness of the
heat sink. Jet fuel, on the other hand, is not as restricted by operating conditions. Fuel
temperatures in aircraft fuel tanks stay relatively cool, ranging from −40 °F to 120 °F
(−40 °C to 49 °C) [8]. Another advantage of fuel cooling is that it is weight-saving. Air
cooling requires heavy equipment to be installed, and may incur drag penalties [5]. With
jet fuel, the cooling infrastructure can be designed into an aircraft’s fuel system and does
not require a separate system, reducing weight penalties. Yet another advantage to a fuel
cooling is that during the cooling process, the jet fuel is heated, resulting in an elevated
fuel temperature entering the combustion chamber, which leads to more efficient fuel
burn and reduced fuel consumption [5].
1.2 Jet Fuel Thermal Stability
While jet fuel is an effective cooling medium, its heat sink capabilities are limited,
primarily by the property known as thermal stability. When fuel is heated to high
temperatures, which is referred to as being thermally stressed, chemical reactions take
place in the fuel that break it down and form solid precipitates that are eventually
deposited onto the walls of the fuel lines in a process referred to as coking. If left
unchecked, this could not only result in severe damage to engine components, but also
reduce the effectiveness of heat transfer in the heat exchangers between the fuel and
engine components [8, p. 2-3].
Many factors affect the thermal stability of jet fuel and coking, such as the dissolved
Chapter 1. Introduction 3
oxygen content, temperature, pressure, chemical composition of the fuel [9], and the
temperature, material and finish of the wetted surface [10]. One of the most important
factors, dissolved oxygen content, is responsible for the autoxidation mechanism. It has
been found in many studies that by removing the dissolved oxygen from the fuel, coking
due to autoxidation can be reduced significantly [11–13]. In laboratory settings, nitrogen
sparging has been used to deoxygenate the fuel for studying the effects of dissolved
oxygen on coking [13,14]. Various other methods have been investigated for this purpose
in a U.S. Air Force study in 1988 [15], which included chemical and molecular means of
deoxygenation. However, while effective for laboratory purposes, these methods are not
practical for large quantities of fuel and real fuel systems. On-board fuel deoxygenation
systems have been studied, such as a membrane diffusion method proposed by Spadaccini
and Huang [11].
Removing dissolved oxygen from the fuel is one way to improve the fuel’s thermal
stability. Other methods have also been investigated, such as the addition of additives
to the fuel. An example of this is the U.S. Air Force’s JP-8+100 program. JP-8 fuel is
widely used in the U.S. military, and its temperature thermal stability limit is 300-325 °F
(149-163 °C) [16]. This program developed an additive that increased this temperature
limits by 100 °F to 425 °F (218 °C), and reduced deposits by 50-95% [2].
Thermal stability is an important limiting factor on the ability of the fuel as a heat
sink, and ignoring its importance can lead to consequences that include fuel line clogging
and degraded engine performance. Therefore, research in this area has been extensive,
with the goal of better understanding all aspects of thermal stability, in order to improve
the designs of fuel systems, as well as the fuels themselves, and mitigate the adverse
effects of thermal instability.
Chapter 2
Project Definition
2.1 Fuel Injector Coking
Jet fuel thermal stability and coking affect all parts of an aircraft’s fuel system. One
particular area of interest is how deposits form in fuel injector nozzle passages. In fact,
fuel injector coking has been identified as the most widespread problem in jet fuel thermal
stability [8]. The injector nozzle passages differ from the rest of the fuel system in two
major ways. First, fuel injector passages are small in diameter and short in length as
compared to the tubing in the rest of the system. The injector nozzle diameters can
be as small as 0.01 inches (in.) (0.254 mm) [17]. Second, because injector nozzles are
located inside the combustion chamber, they are subject to much higher temperatures.
The combustion temperatures will increase the wetted wall temperatures of the nozzles
to greater than 250 °C (482 °F) [3]. At these temperatures, significant coking is observed
in standard engine operations. Coking can block the passages and affect spray patterns,
thus affecting combustion performance.
The goal of present project is to design and build an apparatus to simulate the
geometry and conditions that are seen by injector nozzles in a gas turbine engine, to
study the effect of various factors on coking in the injector nozzles. The project will
4
Chapter 2. Project Definition 5
combine experimental results with those of a concurrent numerical study of the flow
and chemical reactions to better understand the formation of deposits in fuel injector
nozzles. This thesis will present the experimental component of the project, and consists
of the development, design and validation of the experimental apparatus. Pressure drop
measurements will be used as the primary means of characterizing the amount of deposits
and blockage of the passages, and preliminary experiments were performed in order to
ensure valid and useful measurements can be obtained from the experimental apparatus.
The long term goal of the project is to study a more comprehensive set of factors, such
as fuel composition, metal surface material, and fuel storage times. Furthermore, results
from the numerical simulation study will be combined with those of the experimental
study to eventually produce a set of correlations that will be useful as a tool in gas
turbine design.
2.2 General Experimental Approach
In order to study and characterize the coking of the fuel injector nozzles, the experi-
mental apparatus must be able to simulate the fuel system on an actual aircraft. Previous
thermal stability research at the University of Toronto Institute for Aerospace Studies
(UTIAS) was done with a single tube dynamic flow apparatus with a test section of
approximately 1 m in length [13]. This was accomplished in a bench-top apparatus that
used a tube furnace as the heat source. While this type of experiment provided valuable
information on the general coking characteristics of jet fuels, it was not designed to be
representative of any specific part of an aircraft’s fuel system.
On an aircraft, as the fuel travels from the fuel tanks to the combustion chamber,
it flows through several heat exchangers that expose the fuel to different temperature
ranges [8]. The residence times of the fuel in these various heat exchangers also vary
greatly, depending on several factors such as the size of the aircraft, the location of the
Chapter 2. Project Definition 6
fuel tanks, the geometry and arrangement of the fuel lines, and the operating condition of
the aircraft. Because the focus of the present project is coking in fuel injector passages, it
is important to ensure that the fuel flowing through the passages have conditions that are
representative of real-world fuel systems and operating conditions. Furthermore, since
the fuel is exposed to different temperatures, the experimental apparatus must be able to
maintain different temperatures at different parts of the fuel flow path. Table 2.1 provides
a summary of residence time and temperatures that are present in various components in
the fuel flow path on modern aircraft. The components include pumps, heat exchangers
and fuel nozzles. It can be seen that aside from the fuel nozzles, engine components
such as the engine oil cooler expose the fuel to the highest temperatures. The data from
Table 2.1 was taken from [8], and represents typical data for idle or descent conditions.
Because of the low power setting during these conditions, fuel flow rates are relatively
low compared to those during the takeoff and cruising conditions. Thus, descent and idle
present favourable conditions for fuel system coking and is of great interest in thermal
stability research.
Chapter 2. Project Definition 7
Tab
le2.
1:C
omm
erci
alai
rcra
ftfu
elen
vir
onm
ent,
flig
ht
idle
(des
cent)
condit
ions.
Dat
aex
trac
ted
from
Chap
ter
1,T
able
2of
[8].
Fuel
Syst
em
Sin
gle
Pass
Resi
dence
Fuel
Tem
pera
ture
,P
ress
ure
,psi
aS
urf
ace
Tem
pera
ture
,C
om
ponent
Tim
e,
s°F
(°C
)°F
(°C
)
Fuel
tank
5×
103-5×
104
−40
-120
(−40
-49)
2.5-
3.5
−50
-130
(−46
-54)
Tan
kb
oos
tpum
p15
-30
123-
128
(51-
53)
25-3
5F
uel
Engi
ne
1st
stag
epum
p13
-32
127-
150
(53-
56)
75-1
00F
uel
Engi
ne
oil
cool
er0.
7-1.
424
5-32
0(1
18-1
60)
75-4
0020
0-32
0(9
3-16
0)
Engi
ne
gear
pum
p0.
5-0.
624
5-32
0(1
18-1
60)
350-
400
Fuel
Fuel
filt
er0.
5-1.
024
5-32
0(1
18-1
60)
75-4
00F
uel
Mai
nen
gine
contr
ol1.
1-5.
024
5-32
0(1
18-1
60)
350-
400
Fuel
Con
trol
serv
os1.
0-20
245-
320
(118
-160
)35
0-40
0F
uel
Bypas
sre
circ
ula
tion
...
245-
320
(118
-160
)75
-100
Fuel
Gen
erat
oroi
lco
oler
10-1
524
5-32
0(1
18-1
60)
40-1
0032
0(1
60)
Fuel
noz
zles
0.15
-0.8
300-
400
(150
-200
)40
-100
450-
500
(232
-288
)
Chapter 2. Project Definition 8
However, short of constructing an exact replica of a fuel system, it is difficult and
impractical to build a test rig to simulate all the different temperatures seen in the fuel
system. Therefore, for the present project, the complex and varied temperatures and
residence times in the many fuel system components were simplified. The “fuel system”
is simplified and represented by two sections in the fuel flow path: the “preheating
section” and “test section”. The preheating section represents the heat exchangers that
the fuel flows through before it reaches the injector. The test section is a fuel passage
designed to simulate the fuel injector. The temperatures in each of these two sections are
to be controlled independently therefore two temperature and time scales are identified
that are to be the primary parameters of interest:
• Tin - The temperature to which the fuel is stressed before it reaches the test section,
also known as test section inlet temperature;
• Twall - The wetted wall temperature in the injector nozzle passage, or test section.
The fuel (at Tpreheat) will be exposed to this temperature as it passes through the
test section.
In addition, the two temperatures give rise to two time scales that are important in
the experiments:
• tpreheat - The duration that the fuel is subjected to thermal stress before it reaches
the injection passage. This will correspond to the time that the fuel spends in the
heat exchangers before reaching the combustor. This will also be referred to as
residence time, or preheat time.
• ttest - The amount of time the fuel injector passage is subjected to flow of thermally
stressed fuel or test time. This will correspond to the duration that the engine is
in operation. The injector passage, or test section, will have a wetted inside wall
temperature maintained at Twall.
Chapter 2. Project Definition 9
These two time and temperature scales will be controlled independently in the ex-
perimental apparatus, along with the pressure and flow rate of the fuel, as well as the
diameter and length of the fuel injector passage (the test section).
Chapter 3
Background and Literature Review
3.1 Thermal Stability of Jet Fuels
The literature on jet fuel thermal stability is extensive, and many experiments have
been performed to characterize many aspects of the subject. This section will summarize
some important findings in previous research. The production of insoluble deposits in
jet fuels under thermal stress is primarily the result of two mechanisms: autoxidation
and pyrolysis. The former is dominant from fuel temperatures of about 150 °C to 350 °C
(302 °F to 662 °F) [18], and the latter is the dominant process from 300 °C (572 °F) and
higher [19]. These temperature boundaries are not absolute; slight variations have been
reported by different research efforts.
In this chapter, some of the fundamental concepts of jet fuel thermal stability are
presented, especially those that are most relevant to the present project. Some of the
past experimental studies, their methodologies and results are also surveyed. More com-
prehensive reviews of the literature on jet fuel thermal stability, as well as more detailed
analyses of the chemistry and mechanisms behind deposit formation, were given by Ha-
zlett [8], Watkinson [19], and Wong [13].
10
Chapter 3. Background and Literature Review 11
3.2 Mechanics and Chemistry of Thermal Stability
3.2.1 Autoxidation Mechanism
The autoxidation mechanism of deposit formation is attributed to many factors such
as fuel composition, presence of heteroatomic species such as sulfur, and most impor-
tantly, dissolved oxygen. Hazlett in [8] provides an overview of the paraffin oxidation
mechanism, listed in Table 3.1. In this model, hydrocarbons in the fuel reacts with dis-
solved oxygen to form hydroperoxides, which act as a precursor to deposit formation. It
has also been shown that a fuel that oxidizes more easily tends to form less deposit. In
other words, a fuel that oxidizes easily is more thermally stable. This inverse relationship
was reported in [20].
Table 3.1: Overview of paraffin autoxidation mechanism [8].
Initiation R-H + X −→ R ·+XH
Propagation R ·+O2 −→ ROO·ROO ·+R-H −→ ROOH + R·
Chain Termination ROO ·+ROO· −→ ROH + R′COR′′ + O2
ROO ·+R· −→ ROORR ·+R· −→ R-R
Another theory of jet fuel autoxidation is based on soluble macromolecular oxidatively
reactive species (SMORS). Hardy and Wechter proposed this concept in [21], in a study of
deposit formation in long-term storage of diesel fuel. The theory is based on polar species
such as phenols as precursor species, which undergo oxidation reactions and grow in
molecular weight to form SMORS. These reactions occur until the molecular weights are
high enough that solid precipitates are produced [22]. Since then, the SMORS mechanism
has been studied further. Beaver et. al. [22] proposed that the SMORS mechanism also
applies to jet fuel thermal stability.
Chapter 3. Background and Literature Review 12
3.2.2 Pyrolysis Mechanism
The pyrolysis of long-chain hydrocarbons becomes the dominant process at higher
fuel temperatures from around 300 °C (572 °F). At this temperature, the dissolved oxy-
gen in the fuel will have been consumed and there is little contribution from autoxidative
deposits. The formation of deposits due to pyrolysis is attributed to polyaromatic hy-
drocarbons formed as a result of the cracking of alkanes to form cycloalkanes. Andreson
et. al. showed that the aromatic content in the solid deposits consisted of 6- to 7-ring
structures [18]. In the present project, the pyrolysis mechanism was not studied.
3.3 Factors that Affect Thermal Stability
3.3.1 Dissolved Oxygen
Dissolved oxygen content in the fuel have been shown to have a significant effect
on deposit formation. Air-saturated jet fuel has a nominal 70 ppmv of dissolved oxy-
gen, and different levels of dissolved oxygen can be achieved by the process of nitrogen
sparging [13]. In general, deposit formation has been found to decrease significantly in
deoxygenated fuel [8,12]. In the autoxidation process, dissolved oxygen is consumed with
zeroth order and first order kinetics in oxygen. One study has found that the process
of depletion of oxygen starts as a zeroth order process, and then becomes first order as
dissolved oxygen decreases to around 20% of air-saturated dissolved oxygen level [23].
These studies have modelled the depletion of oxygen quite accurately with a bimolec-
ular reaction and Arrhenius rate constant [24]. In some studies, it was found that the
availability of dissolved oxygen is linearly related to total deposit formation [12]. In a
study by Ervin and Williams, different dissolved oxygen levels in thermally stressed fuel
were investigated; the authors found that there is an increase of deposits with decreased
oxygen consumption, and that maximum deposits are produced for a “least favourable”
Chapter 3. Background and Literature Review 13
dissolved oxygen concentration [14].
3.3.2 Temperature
Temperature is an important parameter in jet fuel deposit formation. Deposit for-
mation regimes are defined according to temperature. The autoxidation regime starts
at a bulk fuel temperature of around 150 °C to 350 °C (302 °F to 662 °F), and exhibits
a decline of deposition between 300 °C to 450 °C (572 °F to 842 °F). This is known
as the transition regime, in which the dissolved oxygen in the fuel has been completely
consumed [25] and pyrolytic processes begin to dominate the deposition process. This
deposit dependence on temperature is schematically represented in Figure 3.1. The for-
mation of deposits in a flowing experiment also depends on the wall temperature, as it
influences the transport of insolubles to the wall, as is found in [14]. This study showed
that the amount of deposition in a flowing experiment has a greater dependence on wall
temperature than on the bulk fuel temperature, as the fuel near the walls is heated at a
greater rate.
Figure 3.1: Temperature regimes of jet fuel deposit formation [25].
Chapter 3. Background and Literature Review 14
3.3.3 Flow Rate and Residence Time
Several studies examined the effect of velocity and pressure on deposit formation.
Spadaccini et. al. [25] investigated the effects on deposit formation with respect to space
velocity, which is the ratio of the volumetric flow rate to the total test passage volume.
The work found that as the space velocity increased, the carbon deposits increased lin-
early. The work also compared the effects of Reynolds number and residence time on the
deposits, concluding that species transport is an important factor. As the flow becomes
turbulent (corresponding to higher flow rates) deposit formation is not limited by species
diffusion, thus deposition increased with flow rate [25]. At lower Reynolds numbers (cor-
responding to lower flow rates), where the flow in the test section is laminar, diffusion
is a controlling factor in deposition on the wall and longer residence times will result in
more deposits [23].
3.4 Past Experimental Studies and Apparatuses
3.4.1 Dynamic Flow Tests
In the literature, a large number of aviation fuel stress tests are dynamic flow tests,
and these have been conducted in single-pass heat exchangers. This type of apparatus
typically relies on a length of heated tubing to collect deposits, while fuel flows through
the tube under steady state conditions. The heat can be provided by a furnace or a
heated copper block. An example of this is the Phoenix rig developed by Heneghan
et. al. at the University of Dayton [24]. Spadaccini et. al. used a multiple-tube test
rig configuration in their research [25]. This apparatus allowed for tests for 5 different
conditions to be conducted simultaneously. Yet another example of a dynamic flow test
apparatus is the near-isothermal flowing test rig (NIFTR) used by Jones and Balster
in [26]. This apparatus can be configured such that the fuel passes through the heater
Chapter 3. Background and Literature Review 15
twice, for increased residence time.
The previous experimental apparatus at the University of Toronto Institute for Aerospace
Studies was a single-pass flow test apparatus, using a furnace as the heat source. This
apparatus will be discussed in more detail in Section 3.6.
3.4.2 Fuel Injector Studies
Dynamic flow tests provide good platforms for studying a variety of factors affecting
jet fuel thermal stability. However, they often do not provide conditions that are repre-
sentative of actual engine operating conditions. Fuel system simulators that can replicate
engine operating conditions and geometries were developed for testing the effects of jet
fuel thermal stability on various engine components. These types of apparatuses are
necessary for investigating coking in fuel nozzles.
One effort in studying fuel nozzle coking behaviour was done by Bullock et. al. in [4],
in which a fuel system simulator, called the Aviation Fuel Thermal Stability Test Unit,
was developed. The apparatus was modular and included several component simulators,
such as fuel pumps, filters, and injectors. In the fuel injector module, a length of tubing
simulating the geometry and flow conditions of the fuel injector nozzle was heated by
constant power radio-frequency heater, to simulate heating by the hot combustion air.
The constant power heating ensured constant heat flux, and allowed for coking buildup
to be monitored and inferred by temperature rise due to the insulating effect of carbon
deposits.
Hazlett surveyed several studies on injector nozzle testing using nozzle components
located in a hot gas stream that simulates the combustion environment [8, pp.6-7,25-
27]. These studies tested actual engine fuel injector nozzles, and found that performance
of the nozzles is adversely affected by coking buildup, and is primarily exhibited in a
reduction of flow through the injector fuel passage.
Chapter 3. Background and Literature Review 16
3.5 Analytical Techniques
Analytical techniques for jet fuel thermal stability studies can be classified into two
types. The first type are techniques that analyze and characterize the carbon deposits on
the tube surfaces, and the second type are techniques that focus on the chemical changes
in the bulk fuel as a result of thermal stressing. Some examples of these techniques are
presented in this section.
3.5.1 Temperature Programmed Oxidation
Temperature programmed oxidation (TPO) is the most commonly used technique for
directly determining the amount of carbon deposit buildup in metal tubes. Also known
as carbon burn-off, it is a method of coke removal in aircraft fuel systems [27]. It has
been used in many studies in jet fuel thermal stability to characterize the deposits formed
in the fuel passage [7, 22, 25, 28–30]. For this purpose, the oxidation of solid carbon to
produce carbon dioxide is used:
C(solid) + O2heat−−→ CO2
In this procedure, the fuel from the thermally stressed metal test section is dried, and
is placed in a constant flow of oxygen, usually in a furnace. The test section is heated
gradually. As the temperature rises, the solid carbon deposits will oxidize to produce car-
bon dioxide. This carbon dioxide is then measured with a CO2 analyzer to obtain a CO2
profile in temperature. Different carbon structures will oxidize at different temperatures,
therefore obtaining the carbon dioxide concentration as a function of temperature will
enable characterization of the carbon deposits. A reference on the different morpholo-
gies of carbon deposits and their approximate burning temperatures is provided in [29].
After complete oxidation of the deposits, the CO2-temperature profile can be integrated
to obtain the total mass of carbon deposits in the test section. Carbon burn-off can
also be used to construct an axial deposition profile along the test section by cutting the
Chapter 3. Background and Literature Review 17
test section into smaller segments and performing the carbon burn-off procedure on each
segment.
In order for carbon burn-off to be an effective analytical method, very accurate mea-
surements of CO2 concentrations are required. In past UTIAS studies, carbon burn-off
has been effective as a qualitative measurement [13]. The potential for carbon burn-off
to provide accurate quantitative data was limited by the sensitivity of the CO2 analyzer.
The current CO2 analyzer at UTIAS has an accuracy of the greater of ±100 ppm or 5%
of reading. For low levels of CO2 which are typical in the current application, this error is
on the same order as the measurements. Therefore, without more sensitive and accurate
instrumentation, carbon burn-off measurements can only provide qualitative insights on
the carbon deposits. An accurate carbon determinator such as the Leco RC612 or the
Eltra SC-800 is required for this purpose, and was unavailable due to budget constraints.
Therefore, carbon burn-off was not used in this thesis, and will be implemented in future
work in this project.
3.5.2 Pressure Drop Measurement
Another method to determine the amount of deposits is to measure the change in
total pressure loss in the laminar flow across a test section, caused by a reduction in
cross-sectional area due to accumulated carbon deposits.
In an incompressible flow, according to the Bernoulli equation, total pressure is the
sum of dynamic pressure and static pressure:
p0 = ps +1
2ρv2 (3.1)
where p0 is the total pressure, ps is the static pressure, v is the flow velocity, ρ is the fluid
density, q = 12ρv2 is the dynamic pressure, and variation of height is assumed to be zero.
In an inviscid flow, the total pressure is constant at any two points in the flow. If the
Chapter 3. Background and Literature Review 18
pipe has constant cross sectional area, then the velocity, and thus the dynamic pressure,
are constant. This leads to equal static pressures measured at any two points in the flow
to be equal. However, in a real pipe flow where friction is a factor, there is a loss of
total pressure, or head loss, along a length of piping. If constant flow rate is assumed,
there will be no change in the flow velocity or dynamic pressure. As a result, there will
be a measured static pressure difference across a length of pipe due to the loss of total
pressure. Total pressure loss of laminar flow in a circular pipe of constant diameter can
be calculated with the Hagen-Poiseuille law [31]:
Q =πR4∆P
8µL(3.2)
where Q is the volumetric flow rate, R is the radius of the pipe, µ is the dynamic viscosity
of the fluid, and L is the length of the section of pipe, ∆P is the pressure drop across L.
Rearranging 3.2 to solve for ∆P results in Equation 3.3:
∆P =8µQL
πR4(3.3)
From this result, it is apparent that the total pressure drop is inversely proportional
to the 4th power of the pipe diameter. This relationship can be more practically applied
to measuring the coking blockage in a section of small diameter tubing simulating an
injector nozzle. Coking leads to a reduction in tube inner cross sectional area, thus
reducing the radius R. If ∆Pi and ∆Pf are initial and final pressure drops measured over
one experiment, respectively, and Ri and Rf are the initial and final radii, respectively,
then Equation 3.3 can be used to obtain the following relation:
∆Pi
∆Pf
=
(Rf
Ri
)4
(3.4)
By measuring the pressure drop before and after the experiment, Equation 3.4 can be
applied to infer an “average” reduction in radius, and therefore a thickness of the layer
Chapter 3. Background and Literature Review 19
of deposits can be calculated. It is important to note that this calculated thickness is
only an estimated average. In reality, pressure drop measurement cannot determine how
the deposit is distributed along the length of the test section, as the deposition may not
be uniform [32]. One way to gain a better understanding of the axial deposit profile is to
use carbon burn-off in conjunction with pressure drop measurements. Not only can an
axial deposit profile be obtained, but an estimate of the average density of the deposit
can also be calculated.
The implementation of this measurement in the experimental apparatus is discussed
in Section 4.6, and the preliminary results are presented in Section 6.3.
3.5.3 Spectroscopic Techniques for Chemical Analysis
In jet fuel thermal stability research, spectroscopic techniques can be used to analyze
and quantify changes to the chemical composition of the fuel due to thermal stressing.
UV-visible spectroscopy includes absorbance and fluorescence spectroscopic analysis of
chemical species, and Commodo et. al. showed that these methods are sensitive to
the chemical changes that occur as a result of autoxidative reactions of jet fuel thermal
stressing [33]. Furthermore, they used 3-dimensional UV-visible fluorescence to show the
high-temperature formation of polycyclic aromatic hydrocarbons (PAH) in the stressed
fuel [34]. Li et. al. quantified the hydroperoxides formed during thermal stressing using
absorption spectra [35].
In the present project, chemical changes due to thermal stress is not a primary focus.
However, long term goals of the project include investigation of the effects of fuel storage
duration and temperature on the chemical composition of the fuel, and in turn on the
thermal stability and tendencies of deposit formation.
Chapter 3. Background and Literature Review 20
3.6 Jet Fuel Thermal Stability Research at UTIAS
The University of Toronto Institute for Aerospace Studies (UTIAS) has been con-
ducting research in jet fuel thermal stability. The Advanced Fuel Research Laboratory is
equipped with a bench-top single-tube coking apparatus, capable of conducting dynamic
flow tests.
A Teledyne Isco 500D syringe pump, operating with constant continuous flow, was
used to deliver fuel through a tube that is heated by a tube furnace. The test section
consisted of 1/8 in. (3.175 mm) outer diameter type 316 stainless steel tubing. The
furnace has a heated section of 36 in. (91 cm) length. The syringe pump is able to
operate at precise flow rates with only minor pressure fluctuations. The pressure of
the system is regulated by a back pressure regulator, downstream of the furnace. Fuel
flow temperatures are measured by K-type thermocouples placed before and after the
test section tubing, as well as at other locations where it is necessary to monitor fuel
temperature.
Carbon burn-off, or temperature programmed oxidation, was used to analyze the
amount of deposits collected during a test. UV-visible absorption and fluorescence spec-
troscopy, and electrospray ionization mass spectroscopy were used to analyze chemical
changes to the fuel, such as the growth of polycyclic aromatic hydrocarbons [33,34], and
changes in polar species [36]. Three-dimensional fluorescence spectra were also used to
assess the level of fuel degradation due to thermal stressing [37]. The experimental ap-
paratus was designed and developed by Wong and is shown in Figure 3.2. Details on its
development was reported in [13].
Chapter 3. Background and Literature Review 21
Chapter 3. Experimental Apparatus and Analytical Techniques 36
Notes:
- Drawing is not to scale
- Pipe threads are all NPT
Teledyne
Isco
A500
Dual
Syringe
Pump
Kodiak
RC022 Recirculating Chiller
Fuel Cooling Coil
Pump
Controller
To Waste Thermcraft
Split-Hinged Three-Zone
Tube Furnace
0.5 µm
Filter
Back-
Pressure
Regulator
Waste Fuel
Tank
Fresh Fuel
Tank
To
Surroundings
To Sample
Collection
Dissolved
Oxygen
Sensor
Pressurized
Nitrogen Gas
Pressurized
Oxygen Gas
To
Surroundings
140 µm Filter
Legend
FlowmeterBonnet
Needle Valve
Hydrocarbon
Trap
Thermocouple
Probe
3-Way
Switching
Valve
Pressure
Gauge
Pressure
Regulator
Figure 3.2: Detailed layout of dynamic flow setup. Jet A-1 started in fresh fuel tank that doubled as
the gas sparging vessel, and was pushed through the furnace, cooling coil, back-pressure regulator by
the syringe pump. The 3-way switching valves only allow flow between the bottom port and only one of
the side ports at any given time.
chemical reducing agents, an oxygen-diffusive membrane (Figure 3.4), or by nitrogen
sparging [4]. Molecular sieve adsorbents are impractical because a large surface area
sieve is required to reduce the dissolved oxygen mass fraction from 70 × 10−6 (70 ppm
by mass) to a single digit values [35]. Chemical reducing agents or oxygen scavengers are
generally unstable, cannot be exposed to air, and may react with the fuel. A proprietary
oxygen-diffusive membrane was developed by UTRC [4] that is capable of deoxygenating
a fuel with a dissolved oxygen mass fraction of 70 × 10−6 (70 ppm by mass) down to
1×10−6 (1 ppm by mass). The diffusion process was driven by the oxygen partial pressure
difference between the fuel’s dissolved oxygen and the absence of oxygen in the nitrogen
carrier gas or vacuum on the other side of the membrane. The oxygen permeability of the
membrane was linearly proportional to the pressure difference and was also affected by
Figure 3.2: Schematic diagram of the single tube dynamic flow test apparatus [38].
Chapter 4
Experimental Apparatus
In developing the experimental apparatus for this project, some components from the
previous single-tube dynamic flow test apparatus at UTIAS (Figure 3.2) were utilized.
Because of the different objectives of the current project, the existing apparatus could not
be used without modification. Therefore, new components were added which replaced
some of the existing components, in order to accommodate the requirements of the present
project. Some key differences between the new apparatus and the previous apparatus
are summarized as follows:
• The previous apparatus did not have provisions to control the two temperature and
time scales as described in Section 2.2. Its furnace was the only heating element,
which heated the test section of approximately 1 m in length. The new apparatus
has two heating elements; the first to preheat the fuel to achieve the test section
inlet temperature Tin, the second to heat the test section to provide a controlled
wetted wall temperature Twall.
• The test sections in the new apparatus are significantly different from those in the
previous apparatus. Since the purpose is to study coking in a passage that simulates
a fuel injector nozzle, the test section is much shorter and narrower than those in
the previous apparatus.
22
Chapter 4. Experimental Apparatus 23
• The method of heating the test section is different. Whereas in the previous ap-
paratus the tube furnace was used, a new design with a nozzle band heater and a
brass heating block is employed.
4.1 Overview of the Experimental Apparatus
The new experimental apparatus is shown in Figure 4.1, and Figure 4.2 shows a
schematic diagram of the major components of the experimental apparatus, which consist
of the following:
• Fuel pump
• Fuel preheater
• Test section and heater
• Fuel cooler
• Supply and waste fuel tanks
The fuel is pumped from a stainless steel supply tank by a syringe pump, operating
at a constant flow rate. The fuel is pumped through 1/4 in. (6.35 mm) outer diameter
stainless steel tubing which is submerged in a preheating oil bath. In the oil bath, the
fuel is heated to the desired test section inlet temperature, Tin. Tin is measured by a
K-type probe thermocouple, located just upstream from the test section. The fuel then
passes through the test section, several designs of which are possible (Section 4.4.2).
For the preliminary experiments reported in this thesis, the test section was 1/8 in.
(3.18 mm) outer diameter, 0.027 in. (0.686 mm) inner diameter type 304 stainless steel
tubing. The test section is heated to the desired inner wetted wall temperature Twall, by
a brass block with a diameter of 1.5 in. (38.1 mm) and a length of 2 in. (50.8 mm),
which is clamped around the test section by a nozzle band heater that has a maximum
temperature of 649 °C (1200 °F). Four thermocouples are inserted through the brass
block, contacting and measuring the temperature of the outer surface of the test section
Chapter 4. Experimental Apparatus 24
Figure 4.1: A photograph of the experimental apparatus.
tubing. Downstream of the test section, the fuel exit temperature Tout is measured in
the same way as Tin.
The pressure drop ∆P across the test section is measured at the same locations as
Tin and Tout, by means of two union cross fittings and 1/8 in. (3.18 mm) outer diameter
impulse lines. The pressure drop is measured with a differential pressure transducer of
1 psid (pounds per square inch difference) (6.89 kPa) range. To prevent as much heat loss
as possible, all components that are not heated between the exit of the oil bath and the
Tout measurement location are insulated with several layers of ceramic strip insulation,
wrapped tightly around the tubing and fittings and secured with aluminum foil tape
(Figure 4.3).
Further downstream, the fuel passes through a sintered stainless steel filter that has
0.5 µm porosity to ensure that no solid particles pass through and contaminate the
Chapter 4. Experimental Apparatus 25
Syringe
Pump,
Flow Rate
Controller
Back Pressure
Regulator
Oil Bath
Preheating Tubing
Insulation
Differential
Pressure
Transducer3-way
valve
Thermocouples
Filter
Co
olin
g
Co
il
Flow
Test Section Tubing
Test Section Heater
Tts,1 Tts,2 Tts,3 Tts,4Tin Tout
3-way
valve
N2
Tank
Supply
TankWaste
Tank
Figure 4.2: Schematic overview of the major components of the experimental apparatus.
Figure 4.3: The insulation applied to prevent heat loss from non-heated components.Ceramic strip insulation is wrapped tightly around the tubing and fittings, and is securedwith aluminum foil tape.
sensitive components further downstream. After the filter, the fuel passes through a
heat exchanger to rapidly cool down the hot fuel to quench deposit-forming reactions. A
cooling coil submerged in recirculating cold water, supplied by a recirculating chiller, is
used for this purpose. At the exit of the cooling coil, the fuel is at or below the ambient
Chapter 4. Experimental Apparatus 26
temperature. A back pressure regulator is located downstream of the cooling coil, which
controls the pressure of the system. After the fuel exits the regulator, it is directed to a
waste fuel tank, which is identical to the supply tank.
Pressurized nitrogen is used to purge the entire system of fuel during test rig shut-
down. A pressurized nitrogen tank is connected to the system using a 3-way valve located
between the syringe pump and the oil bath. A check valve is installed at the 3-way valve
to prevent any back flow of fuel into the nitrogen lines. Figure 4.4 shows the nitrogen
purging system.
Check
valve
N2
tank
Fuel from
syringe pump
3-way
Switching
valve
To oil bath,
test section
Figure 4.4: The nitrogen purging system in the experimental apparatus. The solid lineindicates the the flow path of the nitrogen when the 3-way switching valve is set to thepurging mode. Note the check valve that is located at the 3-way valve to prevent backflow of fuel into the nitrogen line.
The syringe pump, supply tank and the oil bath are mounted on a bench that is
assembled from square aluminum profiles and high density polyethylene (HDPE) bench
tops, and all components downstream of the oil bath, with the exception of the recircu-
lating chiller, are mounted vertically on an instrumentation rack. The major components
of the apparatus are described in more detail in the following sections.
Chapter 4. Experimental Apparatus 27
4.2 Fuel Pump and Back Pressure Regulator
4.2.1 Syringe Pumps
The fuel is pumped through the system by a syringe pump system. The A500 dual
syringe pump system purchased from Teledyne Isco was used for this purpose. The system
consists of two 500D syringe pumps connected together through a single controller. The
pumps are run in continuous flow mode, with automatic pump switching and refilling
made possible by a valve system which is operated by compressed air and controlled by
the pump controller. This system allows for smooth switch-over from one pump to the
other, with minimal pressure fluctuations. The pump can operate in constant pressure
mode and constant flow rate mode, and allows for precise and constant flow rate control.
In the present project, the pumps were always operated in constant flow rate mode since
all experiments required a constant flow rate.
4.2.2 Back Pressure Regulator
With the syringe pumps operating in constant flow rate mode, pressure of the system
is regulated through a back pressure regulator installed downstream of the test section,
just before the fuel enters the waste tank. The regulator is supplied by Swagelok (part
number KPB1N0G425P20000).
4.3 Fuel Preheater
The purpose of the fuel preheater is to simulate the heat exchangers that the jet
fuel passes through on an aircraft. In an aircraft’s fuel system, jet fuel is used as a
coolant for engine oil, gear pumps and other critical components [8]. Throughout this
process, heat is added to the fuel, raising its temperature. The addition of heat in a real
fuel system depends heavily not only on the engine operating conditions and the fuel
Chapter 4. Experimental Apparatus 28
flow rate, but also on the size of the aircraft and the actual design of the fuel delivery
system. In some designs, fuel does not simply pass through the heat exchangers and
go on to the combustor; it may be recirculated back to the feed tank in cases where
cooling demands are high but fuel flow rate to the engine is low. An example of this
is the Concorde supersonic aircraft in cruising flight [3], where the high speeds lead to
aerodynamic heating but the engine power demand is low, thus requiring a relatively low
flow rate compared to take off and other power-intensive situations.
Since the addition of heat to the fuel on an aircraft is not a linear process, and given
the wide variety of fuel system designs, it is impossible to design a test apparatus to
generally simulate the fuel heating in all fuel systems. However, “reference” operating
conditions which will provide useful design data can be obtained. To achieve these
reference conditions, the fuel preheater should be able to maintain a constant and uniform
fuel temperature throughout the preheating process.
A few options were considered in the fuel preheater design. First, commercially
available heat exchangers were considered. Heat exchangers are an effective method of
heating the fuel to a desired exit temperature [39], but since they are complete packaged
units, there is no control over the preheat residence time tpreheat. The existing furnace
was also considered for use as a preheater, but because its heated chamber is only a 1
in. (25.4 mm) diameter and 36 in. (92 cm) long tube, the preheat tubing is limited to
that size, and therefore the preheat residence time is also limited. The third and chosen
option was an oil bath. Its heated space is a rectangular box which offers much more
flexibility on the length and arrangement of the preheat tubes, and therefore allows for
more flexible control of the preheat time. Oil baths also have very stable and uniform
temperature fields due to the high heat capacity of the bath fluid.
Chapter 4. Experimental Apparatus 29
4.3.1 Oil Bath for Preheating
The design chosen for preheating in the current apparatus involves heating the fuel
in an oil bath. Tubing will be submerged in this temperature-controlled bath, with
thermocouples placed at the entrance to the test section, to ensure that the fuel is at the
required temperature Tin at the test section inlet.
The oil bath that was chosen was the Memmert model ONE 45 (Figure 4.5). In
this oil bath, the bath fluid is heated by resistance heaters on the bottom surface and
two vertical surfaces of the bath, and uniform temperature is maintained by natural
convection. Its maximum operating temperature is 200 °C (392 °F), and has an internal
usable capacity of 45 L (12 U.S. gal.).
Figure 4.5: The Memmert ONE 45 oil bath. Maximum operating temperature is 200 °C(392 °F), and usable heated space is 45 L (12 US gal.).
Chapter 4. Experimental Apparatus 30
4.3.2 Silicone Bath Fluid
There are several criteria to consider when choosing a bath fluid for fuel heating.
First, the fluid must be able to operate at the required temperatures. In this case, since
the maximum operating temperature of the oil bath is 200 °C (392 °F), the bath fluid had
to be able to operate at that temperature for extended periods of time. Silicone fluids, or
silicone oils, were chosen for this purpose. Silicone fluids are an inorganic, silicone-based
oil that can be produced with a wide range of properties [40]. An important property to
consider when choosing a bath fluid is its gel time, which describes a fluid’s tendency to
polymerize at high temperatures over time. If the fluid’s gel time is shorter than desired,
the fluid may polymerize and form a gel, possibly doubling its volume [40].
The fluid chosen was the DPDM-400 high temperature silicone bath fluid, purchased
from Clearco Products. This fluid was designed specifically for high-temperature bath
applications, and has a upper operating temperature limit of 250 °C (482 °F). While
other fluids that were considered also had this operating range, the gel time of this fluid
is virtually unlimited at within the specified temperature limits, which is an essential
property in the long-term testing of the current application.
Preheating Tubing
The preheating tubing that is submerged in the oil bath is 1/4 in. (6.35 mm) outer
diameter, 0.18 in. (4.57 mm) inner diameter type 316 stainless steel tubing. The inner
diameter of the preheating tubing was chosen to be much larger than that of the test
section tubing, such that the loss of total pressure between the syringe pump and the
test section are minimal and constant as compared to that in the test section. This will
allow for a more stable fuel pressure at the test section.
Because of the large heated space of the oil bath, the arrangement of the preheating
tubing is very flexible. For shorter preheating times, a simple length of tubing can be
installed in the system as the preheating tubing. If longer preheating times are desired, a
Chapter 4. Experimental Apparatus 31
system of modular coiled tubing was manufactured. These can be installed in any config-
uration to produce different lengths of preheating tubing. In the preliminary experiments
reported in this thesis, a 1 m (39 in.) length of tubing was used for preheating, and was
bent to fit inside the oil bath.
4.4 Test Section
4.4.1 Test Section Design Drivers
It was required to design the test section to simulate the geometry and operating con-
ditions of a small gas turbine combustor fuel injector nozzle. In a gas turbine combustor,
high combustion temperatures can increase the injector’s wetted wall temperatures up
to 204.4 - 232.2 °C (400 - 450 °F). The size of the fuel injector passages are very small
compared to the fuel lines, with diameters on the order of 0.01 in. (0.254 mm).
Geometry and Operating Conditions
The test section in the rig must have the required geometry and be able to achieve the
wetted wall temperatures as specified in the test matrix supplied by Pratt & Whitney
Canada. The original test matrix specified a variety of tube diameters from 0.009 in. to
0.150 in. (0.23 mm to 3.81 mm), and wetted wall temperature from 93.3 °C to 315.6 °C
(200 °F to 600 °F). Even though this test matrix would eventually be simplified to cover
only a few of these items, as will be shown in Chapter 5, the test section had to be capable
of reaching all of the specified conditions in the original test matrix. The test matrix
did not specify the material of the test section tubing, therefore the effect of different
materials on coking was outside the scope of the present study.
Coke Deposit Analysis
Deposit analysis was another key design driver in the test section design. Such small
Chapter 4. Experimental Apparatus 32
diameters can potentially be difficult to analyze. Since the primary analysis method is
pressure drop measurement, the test section must be designed to accommodate it. In
addition, the test section was to be designed to support carbon burn-off measurements.
4.4.2 Test Section Designs
To satisfy the above requirements, several designs were considered for the test section’s
fuel injector passage, two of which were chosen for experiments. These will be detailed
in this section.
Commercially Available Tubing
This design was used in the preliminary experiments described in this thesis. In this
design, commercially available stainless steel tubing was to be used as the test section.
Stainless steel tubing are specified by outer diameter and wall thickness, and by its
material. While it is recognized that surface deposition is affected by surface treating
[10, 41] and material differences, investigating this factor was outside the scope of the
current study. Rather, material consistency within a certain set of experiments was more
important.
Seamless stainless steel tubing with outer diameters (O.D.) of 1/16 in. (1.59 mm)
and 1/8 in. (3.18 mm) were purchased with various wall thicknesses, to achieve a range
of inner diameters. Table 4.1 summarizes the tube sizes. In the preliminary experiments
described in this thesis, 1/8 in. (3.18 mm) outer diameter, 0.027 in (0.686 mm) inner
diameter type 304 stainless steel tubing was used as the test sections.
1/4 in. Tubing With Flow Restriction
In this design, the same 1/4 in. outer diameter tubing used in the preheater tubing
is continued throughout the test section, but a constriction in the inside diameter is
inserted as the test section. Figure 4.6 shows an example of this design. This type
of flow path is not commercially available in tubing stock, therefore requires custom
Chapter 4. Experimental Apparatus 33
Table 4.1: Tube sizes available with commercial stainless steel tubing.
Outer Diameter, Inner Diameter, Wall Thickness,in. (mm) in. (mm) in. (mm)
1/16 (1.59)0.0225 (0.572) 0.020 (0.508)0.0345 (0.876) 0.014 (0.356)
1/8 (3.18)0.0270 (0.686) 0.049 (1.245)0.0550 (1.397) 0.035 (0.889)0.0690 (1.753) 0.028 (0.711)
manufacturing. However, this was considered as the design that will most realistically
simulate the geometry and operating environment that the injector is exposed to in the
combustor. While this design was not used in the preliminary experiments as described
in the present work, it will be implemented in the future.
Figure 4.6: An example of the 1/4 in. O.D. tubing and EDM-drilled constriction inserttest section. Shown is the 0.009 in. (0.2286 mm) hole diameter and 0.2 in. (5.08 mm)long fuel injector passage. Both the outer tubing and the insert are made of type 316stainless steel. The insert piece is chamfered 45° to provide a finite transition from thepreheat tubing to the injector passage.
Chapter 4. Experimental Apparatus 34
This test section consists of two parts: the outer tubing and a constriction insert.
The outer tubing is a roughly 3 in. (7.62 cm) long section of 1/4 in. outer diameter
standard stainless steel tubing, and the constriction insert is a drilled-through piece of
stainless steel rod that is press-fit into the outer tubing and welded in place. Because
it requires custom manufacturing, this design offers the best flexibility in terms of the
geometry of the test section. The challenges in producing these test sections were two-
fold; first, producing a smooth reduction in diameter in the constriction, while ensuring
that there are no gaps between the insert and the outer tubing; second, precisely drilling
the very small holes that will be the fuel injector passages which are as small as 0.009 in.
(0.2286 mm). The smallest available conventional drill bit, the #80 bit, has 0.0135 in.
(0.343 mm) diameter, which is not small enough for the required sizes. Also, conventional
drilling may leave burrs that may be difficult to clean up after drilling such small holes.
The solution is to use electrical discharge machining (EDM) to drill the fuel passage
holes. This was done by the machine shop of the University of Toronto Department of
Mechanical and Industrial Engineering. The EDM process consists of an electrode (the
“drill bit”), an electrically conductive work piece, a dielectric fluid, and electrical power.
The EDM process removes material from the work piece by creating a high-voltage spark
across a small gap between the electrode and the work piece, thus vaporizing minute
amounts of both the work piece and the electrode. After vaporization, the removed
material re-solidifies and is carried away by the dielectric fluid, which also acts as cooling
material for the work piece. Electrical discharge machining is comprehensively described
by Jameson in [42].
EDM drilling is a subtype of electrical discharge machining, and it uses an electrode
wire passing through a wire guide to remove material from the work piece, thus “drilling”
a hole. During the drilling process, the electrode wire is consumed as well. EDM allows
very precise holes to be drilled in the work piece, without creating any burrs. However,
there are limitations. Since the electrode wires are very thin, and due to the fact that
Chapter 4. Experimental Apparatus 35
the wire must be extended past the wire guide as the hole is drilled deeper, it is difficult
for the wire to stay aligned with the drilling axis past a certain hole depth [42]. The
Agie Charmilles Drill 11 EDM drill used by the Mechanical Engineering machine shop
is able to drill holes as small as 0.006 in. (0.1524 mm) to depths of up to 200 times the
electrode diameter. This is sufficient to cover a wide range of injector passage diameters.
4.5 Test Section Heater
Since wetted wall temperatures of up to 204.4 - 232.2 °C (400 - 450 °F) are required,
the heater of the test section must be able to maintain this temperature for an extended
period of time. In past thermal stability research, there have been several methods used
to heat the test section to desired temperatures.
Furnaces were used in a number of experimental setups, such as the previous test rig
at UTIAS and also at the University of Dayton Research Institute [41]. Use of the existing
furnace to heat the test section was considered, but it was found to be impractical. Since
the test section is very short compared to the furnace’s heated chamber, it would have
been difficult to control the exact heated boundaries. Also, placing the entire test section
assembly in the heated chamber will heat the fittings as well, and past experience has
shown that high temperatures affect the fittings, and may cause leaks when the fittings
are reassembled.
In a NASA study by Faith et. al., the test section was heated by passing current
directly through the metal of the test section tubing [6]. The electrical power is held
constant to maintain a constant heat flux, rather than a constant temperature. Ther-
mocouples were cemented along the length of the test section to measure the outer wall
temperatures. This constant heat flux allows the monitoring of temperature increase
due to carbon deposit buildup, since carbon deposits have a thermal insulation effect [4].
This method was not used in the current study due to safety concerns from passing large
Chapter 4. Experimental Apparatus 36
amounts of current through exposed tubing.
In another test rig at the University of Dayton, a heated copper block clamped around
the test section tubing was used to provide even heating to the test section. In this
application, the copper heated block was a cylinder of 460 mm (18.1 in.) length and
76 mm (3 in.) diameter, and is able to maintain the test section outer wall temperature
to up to 497 °C (927 °F) [24]. Thermocouples welded to the test section tubing enabled
measurement of the axial temperature gradient.
4.5.1 Brass Heating Block
For the current project, a compact heater was desired because of the small size of the
test sections, therefore the heated metal block approach was chosen. However, instead
of copper, brass was used because it was more readily available. Brass, being an alloy of
copper and zinc, has slightly lower thermal conductivity than pure copper. However, it
is still more thermally conductive than stainless steel by a factor on the order of 10 [43],
which allows for faster heating times. The heater block assembly is a cylindrical block of
1.5 in. (38.1 mm) diameter and 2 in. (50.8 mm) length, thus providing a heated length
of 2 in. A hole with the diameter of the test section tubing (either 1/8 in. or 1/4 in.) was
drilled through the centre of the block. The block was machined in two halves, which
are clamped around the test section by a nozzle band heater (see Section 4.5.2). Four
1/8 in. (3.175 mm) holes are drilled through one of the halves such that 4 thermocouples
can be inserted through the block to physically contact the test section outer wall. The
thermocouple holes are equally spaced, with the end holes as close to the edge as possible
as allowed by the machining processes. Diagrams of the test section assembly, with the
1/8 in. (3.175 mm) outer diameter, 0.027 in. (0.686 mm) inner diameter commercially
available stainless steel tubing as the test section, are shown in Figure 4.7.
The thermocouples used to measure the test section outer wall temperatures are type
K probe thermocouples, clad in 1/8 in. (3.175 mm) diameter stainless steel sheaths. A
Chapter 4. Experimental Apparatus 37
mechanism was designed to ensure that the thermocouples make physical contact with
the outer surface of the test section tubing. This mechanism uses compression fittings in
conjunction with small precision springs to press the thermocouple probes down on to the
outer surface of the test section tubing. The mechanism is shown and explained in Figure
4.8. The fittings are attached to the brass block by silver soldering, and all manufacturing
and machining was done by the University of Toronto Mechanical Engineering machine
shop.
4.5.2 Band Heater and Temperature Controller
Heat is provided to the brass block by a 300 W band heater commonly used in heating
nozzles in injection molding machines. These heaters are compact and are simple to set
up. The heater is purchased from McMaster-Carr, and has a maximum temperature out-
put of 649 °C (1200 °F). The band heater is controlled by an Omega CN7523 temperature
controller. A solid-state relay is used to power the heater circuit. The wiring diagram
for the heater and controller is shown in Figure 4.9. The thermocouple that provides
the controller feedback is a dual-element thermocouple which also provides temperature
data for the data acquisition system which will be described in Section 4.7.
4.6 Pressure Drop Measurements
The primary means of monitoring deposit buildup is the measurement of the total
pressure drop across the test section. As discussed in Section 3.5.2, the pressure drop is
inversely proportional to the 4th power of the passage diameter.
To measure the pressure drop, two static pressure taps are installed, one upstream
and one downstream of the test section (Figure 4.10). Since the measured pressure drop
is the loss in total pressure, the pressure taps must be at locations of equal cross sectional
area, to ensure equal dynamic pressure at both locations. The flow cross sectional areas
Chapter 4. Experimental Apparatus 38
(a) (b)
Figure 4.7: (a) Drawing showing the complete test section assembly, with all thermocou-ples and the test section tubing inserted. Note that only 2 in. (50.8 mm) of the 3.25in. (82.55 mm) test section is heated. The unheated segments at the ends are necessaryfor the fittings, and are insulated. (b) Drawing showing internal structure of the brassheater block. 4 holes of 1/8 in. (3.18 mm) diameter are drilled through the top half ofthe block and the 4 thermocouple tubes are silver soldered to the brass piece.
at the two locations must also be equal, since the constriction of the test section will
create a narrow flow cross sectional area immediately upstream and downstream of the
contraction. The pressure drop, if measured at locations of flow constrictions, will have
a component of static pressure difference in addition to the loss of total pressure [44].
A differential pressure transducer is used to measure the pressure drop across the test
section. Initially, a transducer of 0 - 200 psid (0 - 1.38 MPa) range was used. However, it
Chapter 4. Experimental Apparatus 39
Nut
Spring
Washer
Ferrule and
set screw
(fixed to probe)
Thermocouple
probe
Test section
outer wall
Fitting
soldered to
brass block
Figure 4.8: The thermocouple contact mechanism in the brass heater block. The ferruleis held fixed to the thermocouple probe by a hex set screw. A spring and washer reston top of the ferrule. The thermocouple is inserted into the fitting through the brassblock. As the nut is threaded on to the fitting, it compresses the spring which providesthe downward pressure on the ferrule, which keeps the thermocouple probe in contactwith the test section outer wall.
was found that it was not sensitive enough to measure the pressure drop across the test
section. Any measured signal was well within the uncertainty of the transducer, therefore
it produced no useful data. To remedy this, a 0 - 1 psid (0 - 6.89 kPa) transducer was
used instead.
The pressure transducer used is a wet/wet differential pressure transducer with 1 psid
(6.89 kPa) range, and is purchased from Omega (model number MMDWU001V5P3A0T1A1).
It has a 0 - 5 V DC output, and can be directly measured by the data acquisition system
(see Section 4.7). In addition to the data acquisition system, the output of the transducer
Chapter 4. Experimental Apparatus 40
Solid State Relay
+ -
Heater
300 W
2.5 A
9 10+ -
-
+4
6120 V AC
Thermocouple
Temperature
Controller
1
2
Switch
Fuse
Figure 4.9: Wiring diagram for the test section heater, with controller terminal num-bers shown. Solid State Relay: Omega SSRL240DC10; Temperature Controller: OmegaCN7523; Thermocouple: K-type dual element, Omega SCASS-125U-6-DUAL; Switch:DPST, 10 A @ 125 V AC; Fuse: 10 A.
is also simultaneously monitored with a process meter, which also provides the excitation
voltage for the transducer.
The transducer is set up with overpressure protection by means of a 3-way valve.
During normal operation, the valve is switched so that the high pressure port on the
transducer is connected to the upstream pressure tap. However, in high flow situations,
such as nitrogen purging, the pressure drop may be well above the 1 psid range. In this
case, the 3 way valve is switched to connect the high and low pressure ports together,
such that the pressure differential is equalized. Figure 4.11 schematically shows these
two configurations.
Chapter 4. Experimental Apparatus 41
Pressure tap
locations
Test section
Thermocouple
probes
¼” union
cross fitting
Flow
Figure 4.10: The placement of the static pressure taps for the pressure drop measurement.The taps are located sufficiently away from the test section such that the flow crosssectional area is not affected by the constriction of the test section.
Differential
pressure
transducer
Test section
3-way
valve
Flow
Hi Lo
(a) Normal operation
Differential
pressure
transducer
Test section
3-way
valve
Flow
Hi Lo
(b) High-flow operation
Figure 4.11: The two configurations of the impulse lines of the differential pressuretransducer. The solid lines indicate the high pressure lines and the dashed lines indicatethe low pressure lines. (a) shows normal operation, where the high and low pressureports are connected to their respective pressure taps. (b) shows the configuration whenhigh flow rates result in a differential pressure greater than the transducer’s range. Inthis case, both the high and low pressure ports are connected to the downstream pressuretap, thus equalizing the pressure difference.
Chapter 4. Experimental Apparatus 42
4.7 Data Acquisition
A computerized data acquisition (DAQ) system is used to collect all relevant data
from the experiments. The following variables are monitored and recorded by a data
acquisition system:
• Temperature upstream and downstream of the test section, Tin and Tout, respec-
tively
• Temperature profile of the test section outer wall, provided by the 4 thermocouples
in the brass block, Tts,1 through Tts,4 (Figure 4.2)
• Pressure drop across the test section, ∆P , provided by the differential pressure
transducer
A National Instruments USB-6210 multifunction data acquisition device is used to
collect the pressure transducer data, which has a 0-5 V DC output signal. Rather than
purchasing the LabVIEW software, the MATLAB Data Acquisition Toolbox was used to
provide the interface between the National Instruments DAQ device and the computer.
The data from the thermocouples are collected with an 8-channel thermocouple data
logger, supplied by Omega (model number OM-CP-OCTTEMP2000).
4.8 Other Components
Sintered Stainless Steel Filter
To ensure that no solid particles are allowed into the cooling coil and back pressure
regulator, a 0.5 µm sintered stainless steel filter is installed immediately downstream of
the Tout measurement location. Without the filter, solid precipitates may build up in
the custom manufactured cooling coil or the back pressure regulator, which are costly to
replace. In comparison, the filter elements can be easily and inexpensively replaced on a
regular basis.
Chapter 4. Experimental Apparatus 43
Cooling Coil and Recirculating Chiller
To quench deposit-forming reactions after the heated sections, the fuel passes through
a cooling heat exchanger, which is a coiled section of 1/8 in. (3.18 mm) outer diameter
tubing, encased in an 1.5 in. (38.1 mm) outer diameter tubing. The cooling medium
is cold water, which is supplied by a Lytron Kodiak recirculating chiller (model number
RC022J03BE2). This chiller can supply water at 10 °C (50 °F), for effective heat removal.
The fuel and the water flow in the same direction, such that the hot fuel is exposed to
the cold water as soon as possible. The coil and chiller system is described in more detail
in Section 3.2.4 of [13].
Supply and Waste Fuel Tanks
The supply and waste fuel tanks are mirror-finish stainless steel pressure tanks pur-
chased from McMaster-Carr. The tanks are of 8 L (2 U.S. gal.) capacity, and are fitted
with female NPT (national pipe thread taper) connection ports for inlet and outlet con-
nections. The supply tank is fitted with a nitrogen sparging system for de-oxygenating
the supply fuel. Experiments with de-oxygenated fuel was not within the scope of this
thesis, therefore this system was not used. The system is described in more detail in
Chapter 3 of [13].
Chapter 5
Experimental Methodology and
Procedures
5.1 Overview
For the initial part of the experimental work, the scope of the experiments is restricted
to vary only the two temperature scales, fuel test section inlet temperature Tin and test
section wetted wall temperature Twall. These temperatures were defined in an updated
test matrix, shown in Tables 5.1 and 5.2, provided by Pratt & Whitney Canada. All other
experimental parameters are to be kept constant. For each test point in the matrix, an
experiment, or a test run, is performed. Data is collected throughout the test run and
is analyzed afterwards. The same test procedures were carried out in the preliminary
experiments reported in this thesis. The test run procedure, as well as the data collection
and analysis procedures, are described in this chapter.
5.2 Numerical Simulation
As stated in Section 2.1, the project has a concurrent numerical simulation com-
ponent. The objective of this component is to combine computational fluid dynamics
44
Chapter 5. Experimental Methodology and Procedures 45
Table 5.1: The revised Pratt & Whitney Canada test matrix for fuel temperature (Tin)tests.
Fuel Temperature Tests
Fuel Temperature, Wetted Wall Fuel Pressure, Flow Rate,°F (°C) Temperature, °F (°C) psig (MPa) pph (mL/min)
150 (65.6) 450 (232.2) 100 (0.69) 2 (20.408)200 (93.3) 450 (232.2) 100 (0.69) 2 (20.408)250 (121.1) 450 (232.2) 100 (0.69) 2 (20.408)300 (148.9) 450 (232.2) 100 (0.69) 2 (20.408)325 (162.8) 450 (232.2) 100 (0.69) 2 (20.408)
Table 5.2: The revised Pratt & Whitney Canada test matrix for wetted wall temperature(Twall) tests.
Wetted Wall Temperature Tests
Fuel Temperature, Wetted Wall Fuel Pressure, Flow Rate,°F (°C) Temperature, °F (°C) psig (MPa) pph (mL/min)
250 (121.1) 275 (135.0) 100 (0.69) 2 (20.408)250 (121.1) 300 (148.9) 100 (0.69) 2 (20.408)250 (121.1) 325 (162.8) 100 (0.69) 2 (20.408)250 (121.1) 340 (171.1) 100 (0.69) 2 (20.408)250 (121.1) 355 (179.4) 100 (0.69) 2 (20.408)250 (121.1) 370 (187.8) 100 (0.69) 2 (20.408)250 (121.1) 385 (196.1) 100 (0.69) 2 (20.408)250 (121.1) 400 (204.4) 100 (0.69) 2 (20.408)250 (121.1) 425 (218.3) 100 (0.69) 2 (20.408)250 (121.1) 450 (232.2) 100 (0.69) 2 (20.408)
(CFD) and pseudo-detailed chemistry models to simulate the processes in deposit for-
mation. The CFD and heat transfer simulations are helpful in estimating the heat losses
and temperature profiles within the experimental apparatus, thus providing a guidelines
for temperature settings for the various heaters in the apparatus.
Chapter 5. Experimental Methodology and Procedures 46
5.3 Description of Experimental Procedures
Each test run will be conducted in sessions of 5 hour in length. Usually one session
is performed each day, for a total of three sessions, or 15 hours for each run. This
scheme was chosen for practical reasons, since it was unrealistic to run coking tests for
an extended period of time without supervision. The NASA technical report by Faith et.
al. in 1971 adopted a similar scheme, in which runs of 20 or 100 hours in total length were
conducted in 5- or 10-hour sessions, with a complete shutdown between each session [6].
The total number of hours for each session will also depend on the coking charac-
teristics for each fuel. Longer experiments may be required for more stable fuels. To
determine the number of sessions required for each set of test runs, an initial trial run
will be performed, and repeated to ensure consistency of the procedures.
5.3.1 Preparation and Set Up
At the beginning of each session, the supply tank is filled with fuel. Also, the waste
tank is emptied with a plastic hand-held fuel pump into a larger (20 L, 5.3 gal.) waste
fuel container. Once this is done, power is turned on to the oil bath, equipment rack,
and subsequently the test section heater and process meter, which supplies the power to
and monitors the data from the pressure transducer. The fume hood is also powered on
to ensure proper ventilation, as there may be fumes from the heated silicone bath fluid.
Before powering on the syringe pump, compressed air must be supplied to the air
valve system necessary for continuous flow. An air line connected to the building’s
compressed air supply is connected to the air valve system. A pressure regulator controls
the air pressure from the supply line. For the air valve system, the pressure is set at
90 - 100 psig (0.62 - 0.68 MPa).
Chapter 5. Experimental Methodology and Procedures 47
5.3.2 Run Procedure
Each session includes three segments:
• Heat-up, during which the oil bath and the test section are heated to the desired
temperature set points. The duration of this segment will vary depending on the
set temperatures;
• Steady-state run, which is the 5-hour timed segment, during which oil bath and
test section temperatures are essentially at steady state and deposits are collected
in the test section. The temperatures from the thermocouples and the pressure
drop measured by the pressure transducer are recorded during this segment;
• Shutdown, which includes purging the system of fuel with compressed nitrogen,
cooling down of the test section heater, and refilling of the syringe pump cylinders.
5.3.3 Heat-Up
The heat-up phase consists of bringing the oil bath and the test section heater up to
their required temperature. The oil bath temperature is set such that the test section inlet
temperature, Tin, matches the test conditions specified in the test matrix. The process of
heating the bath fluid from room temperature to the desired preheat temperature can take
as long as 1.5 hours. The DPDM-400 bath fluid has a heat capacity of 0.35 cal ·g−1·◦C−1,
or 1470 J ·kg−1 ·K−1 at 25 °C (77 °F), and there are 42 L (11 U.S. gal.) of the fluid in the
bath, equivalent to approximately 44.5 kg. The electrical power rating of the oil bath is
2800 W during heating, and assuming perfect efficiency and a constant specific heat, the
total time needed to heat the bath fluid from a room temperature of 25 °C (77 °F) to
170 °C (338 °F) is approximately 1 hour. Factoring in heat losses, the heating time can
take as long as 1.5 hours, at approximately 1.6 °C (2.9 °F) per minute.
During this heating process, no fuel is pumped through the system. Because the
heat up process takes up to 1.5 hours, the preheat tubes are exposed to autoxidation
Chapter 5. Experimental Methodology and Procedures 48
temperatures for non-negligible lengths of time. Therefore, fuel flow is only started
during the last few minutes of the heat-up phase, or when the oil bath temperature
reaches within 40 - 30 °C (72 - 54 °F) below the set point, as described below. A very
low flow of nitrogen is maintained in the entire system, to remove the oxygen that is
present in the air which can oxidize and thus remove any deposits that may already have
been collected in the test section. The pressure of the nitrogen flow will vary depending
on the degree to which the stainless steel filter has been clogged, but has to be at least
20 psig (0.14 MPa) to crack the check valve, as explained in Section 4.1.
As the oil bath temperature reaches within 40 - 30 °C (72 - 54 °F) below the set point,
fuel flow is started, and the system is pressurized with the back pressure regulator. Heat
is applied to the test section at this point. The fuel flow is started at this point because
the heating time of the test section is much shorter than that for the oil bath. Heating
up the test section takes approximately 10 to 20 minutes, depending on the desired test
section temperature set point. During test section heat-up, visual checks for leaks are
performed at all fittings.
Oil Bath Temperature Setting
Due to heat losses to the ambient environment, the temperature set point for the oil
bath is set such that the test section inlet temperature Tin is controlled at the desired
temperature. As shown in Figure 5.1a, there is an approximately 25 °C (45 °F) tempera-
ture drop between the oil bath exit and the test section inlet, despite the insulation that
is applied to reduce heat loss. The fuel temperature exiting the oil bath is assumed to
be at the oil bath temperature, since the length of the preheating tubing is sufficient to
heat the fuel up to the oil bath temperature. Table 5.3 lists the oil bath temperature
settings for various test section inlet temperatures.
Test Section Heater Temperature Setting
The four thermocouples for the test section are set up such that they contact the outer
Chapter 5. Experimental Methodology and Procedures 49
Table 5.3: Oil bath temperature settings for various test section inlet temperatures.
Test Section Inlet Oil Bath TemperatureTemperature Tin, °C (°F) Set Point, °C (°F)
162.8 (325) 188.5 (371.3)
121.1 (250) 144.8 (292.6)
93.3 (200) 112.6 (234.7)
wall of the test section tubing, as described in Section 4.5.1. This will only provide a
measurement of the outer wall temperature, not the wetted wall temperature Twall, or
inner wall temperature. Since the fuel enters the test section at a lower temperature than
the specified Twall, there exists a temperature gradient between the inner and outer walls
of the test section tubing. Due to this gradient, the test section heater is set at a higher
temperature than the desired wetted wall temperature. This correction is obtained from
a numerical simulation of the temperature field in the test section tubing.
5.3.4 Steady State Operation
The steady state phase, or the timed phase of the session, begins when the temper-
atures are within a few degrees of the set points. This method was chosen because the
temperature rise as a function of time slows down significantly as the temperatures ap-
proach the set points. This behaviour is dictated by the proportional-integral-differential
(PID) control parameters set in the temperature controllers. If the temperature is al-
lowed to reach the set point before the steady state phase begins, the system will have
already been running at near the set point temperature for a significant amount of time,
allowing deposits to build up. Therefore, it is more reasonable to define a temperature
above which it can be assumed to be sufficiently close to the steady state set point. In
Chapter 5. Experimental Methodology and Procedures 50
this thesis, this temperature is defined as at 10 °C (18 °F) below the desired Tin, and is
kept consistent throughout each run. Since the temperatures will keep rising after the
steady state phase begins, all temperatures are recorded. Figure 5.1 shows an example
of the temperature-time profiles of the oil bath and test section.
0 5 1 0 1 5 2 0 2 5 3 01 5 5
1 6 0
1 6 5
1 7 0
1 7 5
1 8 0
1 8 5
1 9 0
���
����
��� �
�
T i m e ( m i n )
O i l B a t h T e s t S e c t i o n I n l e t T i n T e s t S e c t i o n O u t l e t T o u t
(a) Oil bath and test section inlet/outlet
0 5 1 0 1 5 2 0 2 5 3 0
2 5 5
2 6 0
2 6 5
2 7 0
T e s t S e c t i o n O u t e r W a l l���
����
��� �
�
T i m e ( m i n )(b) Test section wetted wall
Figure 5.1: The temperature-time profiles of the first 30 minutes of steady-state operationin a typical session. In this case, the conditions are Tin = 162.8 °C (325 °F), Twall = 260 °C(500 °F). Note that timing starts before the temperatures reach steady state, which occursapproximately 20-25 minutes after timing starts. (a) shows the oil bath temperature andtest section inlet and outlet temperatures, Tin and Tout respectively. Note that Tout ishigher than Tin due to test section heat addition, and that there is a roughly 25 °C(45 °F) temperature loss between the oil bath exit and test section inlet. (b) shows thetest section outer wall temperature from the first thermocouple on the test section. Notethat this outer wall temperature is set higher than the desired wetted wall temperatureTwall.
Test Section Steady State Temperature Profile
At steady state, the heated portion of the test section tubing is at nearly isothermal
conditions, with a variation of up to 2 °C (3.6 °F). This was made possible with improved
insulation as shown in Figure 4.3 that insulates the unheated sides of the brass heater
block.
Chapter 5. Experimental Methodology and Procedures 51
System Back Pressure Adjustment
During the steady state phase, the system pressure is constantly adjusted using the
back pressure regulator. The regulator is adjusted such that the pressure at the syringe
pump is held constant, rather than the pressure at the back pressure regulator. This
is done to ensure that the pressure at the test section is held as constant as possible.
Throughout the session, the stressed fuel passing through the filter downstream of the test
section will leave deposits and increase the pressure drop across the filter. To maintain a
constant flow rate, the syringe pump will increase its pressure, thereby increasing the total
pressure drop from the pump to the regulator. If the back pressure regulator pressure is
kept constant, then the pressure at the test section will increase, and it was found that
there is a dependence of pressure drop on the static pressure. Since the preheat tubing
between the pump to the test section has much larger diameter than that of the test
section, the pressure drop through it is small and relatively constant. Thus, keeping the
pressure constant at the syringe pump will ensure that the pressure of the test section is
constant, thus giving a more reliable measurement of the pressure drop.
5.3.5 Shutdown and Purging
At the end of the timed portion, the shutdown procedure is carried out. During
shutdown, it is important to minimize the amount of thermal soakback coking that will
occur in the test section as a result of residual heat left in the heater. Pratt & Whitney
Canada has identified thermal soakback as a significant contributor to fuel line coking
in aircraft, due to the high temperature of the fuel system after engine shutdown and
the presence of stationary fuel left in the fuel lines. Soakback coking will be studied as
a long term objective of this project, but it is beyond the scope of this thesis. However,
it is important to avoid soakback coking in the test section in the current apparatus.
The preheat tubing in the oil bath has a much longer thermal soakback period because
of the high heat capacity of the bath fluid. However, the coking in the large-diameter
Chapter 5. Experimental Methodology and Procedures 52
preheat tubing is not of major concern in the present work. The following procedures
are followed to minimize soakback coking in the test section during the shutdown phase.
Nitrogen Purging
Fuel is purged immediately from the test section after the timed session ends. At
the end of a timed session, the oil bath and test section heaters are turned off, and
the pressures at the two ports of the differential pressure transducer is equalized by
switching the valve to the configuration shown in Figure 4.11b. At this point, the system
is gradually de-pressurized. Fuel flow is stopped and nitrogen purging is started as
soon as the fuel is at atmospheric pressure. Nitrogen purging pressure is set at at least
50 psig (0.34 MPa).
Test Section Cooling
During nitrogen purging, the test section is cooled down rapidly using compressed
air impingement. This forced convection significantly speeds up the cooling of the test
section, to further minimize soakback coking. The air is directed at the test section heater
until the test section thermocouple readings are below 80 °C (176 °F) to ensure that the
wetted wall temperature is below the autoxidation coking temperature threshold.
5.3.6 Data Collection
As discussed in Section 4.7, the data is collected with automated data acquisition
systems. The Data Acquisition Toolbox in MATLAB was used to record the pressure
drop data, and the Omega OM-CP-OCTTEMP2000 thermocouple data logger is used to
collect temperature data.
Pressure Drop Data
One data point is recorded per minute for the pressure drop data. Each data point
is the average over a 2-second recording at 1000 Hz. This will ensure that any high-
Chapter 5. Experimental Methodology and Procedures 53
frequency fluctuation in the input voltage is damped out. The pressure drop across the
test section varies over periods of hours, and it is not necessary to record data points at
shorter intervals.
Temperature Data
Temperature data is recorded at the same 1-minute interval as the pressure drop data.
The data from 6 thermocouples are recorded:
• The four thermocouples contacting the outer wall of the test section tubing, Tts,1 -
Tts,4
• The test section inlet and outlet temperatures, Tin and Tout
Chapter 6
Results and Discussion
Preliminary experiments were performed with the apparatus, for the purpose of test-
ing the apparatus to ensure proper functionality and that measured data are valid and
significant. From these preliminary results, it was shown that pressure drop measurement
is an effective method of measuring the carbon deposit build-up in the test section.
6.1 Fuel Batch and Jet Fuel Thermal Oxidation Tester
(JFTOT) Results
The fuel used in the preliminary experiments was Jet A-1 fuel, supplied by Shell
Canada, and was delivered in July 2009 [13, p.49]. This was the same batch (“Batch 2” as
referred to in [13,38]) that was used in previous UTIAS jet fuel thermal stability research.
The jet fuel thermal oxidation tester (JFTOT) was used to test the thermal stability
properties of this batch of fuel [38]. The JFTOT is a standardized test (ASTM D 3241)
to rate the thermal stability of a particular batch of jet fuel. The JFTOT is described
in detail in [8, pp. 16-19]. In the test, heated fuel is flowed through a test section tube
at specified temperature, pressure, for a specified length of time. At the end of the test,
the fuel is rated to a pass or fail grade based on the following two criteria:
54
Chapter 6. Results and Discussion 55
1. Tube deposit rating. This is a visual rating of the deposits on a test section tube.
The used test section tube is compared against a colour standard to determine a
rating of 0, 1, 2, 3, or 4, with 0 being “no visible deposits”. For a “pass” rating, a
visual rating of less than 3 is required.
2. Filter pressure drop. As the stressed fuel is flowed through the test system, a filter
collects deposits and pressure drop builds up across the filter, and this pressure
drop is measured at the end of the test to grade the fuel. For a “pass” rating, the
pressure drop cannot exceed 3.3 kPa (25 mmHg).
JFTOT tests were performed on the current fuel by CanmetENERGY and was re-
ported in [38]. The fuel had a visual deposit rating of 0 (no visible deposits) and a filter
pressure drop of 0 mmHg. While these are passing results for this particular batch of
fuel, it also means that it has a lesser tendency to produce deposits. This was confirmed
by carbon burn-off results reported in [38], in which the current batch produced much
less deposit than a previous batch.
6.2 Apparatus Verification Experiment Conditions
To verify the apparatus operation, two sets of preliminary experiments, or “Run 2”
and “Run 3” as they will be referred to in this thesis, were performed. Each run consisted
of three 5-hour sessions. The experiments followed the procedures described in Chapter 5.
These two sets of experiments only varied in the test section fuel inlet temperature Tin,
while all other parameters remained identical for both sets of experiments. Using the test
matrix in Table 5.1 as a guideline, the experimental conditions for the two experiments
were chosen as follows:
• Test section inlet temperature: Tin = 162.8 °C (325 °F) for Run 2, 93.3 °C (200 °F)
for Run 3
Chapter 6. Results and Discussion 56
• Wetted wall temperature: Twall = 260 °C (500 °F)
• Fuel preheat temperature: as shown in Table 5.3.
• Fuel preheat residence time: tpreheat = 50 s
• Fuel pressure: P = 132 psig (0.91 MPa) at fuel pump
• Flow rate: Q = 20.408 mL/min, or approximately 2 pounds per hour
Temperatures
The purpose of choosing the two Tin as described above was to measure the effects of
two extremes of the bulk fuel temperature on deposit formation. 93.3 °C (200 °F) is on
the lower temperature limit for autoxidative deposit formation, while 260 °C (325 °F) is
near the high end of the fuel temperature spectrum in an aircraft’s fuel system, as shown
in Table 2.1.
The test section wetted wall temperature Twall used in the apparatus verification
experiments is higher than the maximum Twall of 232.2 °C (450 °F) as specified in the
test matrices (Tables 5.1 and 5.2). Because the fuel was rated to be very stable by the
JFTOT, it was decided to set Twall at a higher temperature than specified to ensure
deposit formation.
Pressure
The system pressure was kept constant throughout the experiments. The fuel was kept
at an elevated pressure above the atmospheric pressure, in order to ensure that the fuel
stayed within the liquid phase as it was exposed to the high temperatures of the test
section [45]. It is important to note that the pressure as measured at the syringe pump,
as opposed to that measured at the back pressure regulator, was used as the system
pressure. This was done to ensure that the fuel at the test section was consistently at
the same pressure, since the relatively large diameter tubing upstream of the test section
contributed very little pressure drop between the syringe pump and the test section. It
was not desirable to maintain the pressure constant at the back pressure regulator, as the
Chapter 6. Results and Discussion 57
pressure drop across the filter, which is present between the test section and regulator,
increased significantly over the duration of the tests. This would increase the pressure
at the test section which would change the pressure drop measurements.
Flow Rate
The flow rate of 20.408 mL/min was chosen as specified by Pratt & Whitney Canada
and results in a test section fuel velocity of 0.92 m/s (3.02 ft/s). With the fuel temper-
atures of Run 2, this equated to a Reynolds number of 1670. Since the temperatures of
Run 2 are the highest in the test matrix, this is the maximum Reynolds number that
will be seen at this flow rate. For lower temperature, viscosity increases, making the
Reynolds number lower. This placed all the test conditions in the laminar flow regime,
and made the the Hagen-Poiseuille equation for pressure drop valid, as will be discussed
in Section 6.4.
6.3 Pressure Drop Results
Despite the stable fuel batch, the preliminary experiments produced measurable pres-
sure drop that increased during the high temperature runs. The pressure drop results
from the two preliminary experiments were recorded over the duration of the experiment
and are shown in Figures 6.1 and 6.2 for Runs 2 and 3, respectively. The break between
the 5-hour sessions are shown by vertical grid lines.
Run 2 Results
Over the entirety of Run 2, a pressure drop increase that is outside the error range of
the pressure transducer was clearly measured. The Tin for this run was 162.8 °C (325 °F)
and the Twall was 260 °C (500 °F). The accuracy of the pressure transducer is ±0.08% of
its full measurement range, and given that its range is 1 psid (6.90 kPa), this is equivalent
to ±0.8× 10−3 psid (5.5× 10−3 kPa). In the data, it can be seen that at the beginning
Chapter 6. Results and Discussion 58
of each run, as the temperature of the system was still reaching the steady state, the
pressure drop decreased until the steady state was reached. This dependence of pressure
drop on temperature will be discussed in more detail in Section 6.4.
During the steady state, as carbon deposits began to build up, the pressure drop was
seen to climb steadily with time. In the results, a periodic fluctuation was observed even
during the steady state operation. This was caused by the pressure fluctuations in the
system that were a result of the switching between the two syringe pumps. This switch
is a necessary action of the pump’s continuous flow system. Despite these fluctuations,
the pressure drop was clearly seen to increase over the entire experiment.
From Figure 6.1, it was determined that the final pressure drop ∆Pf was 0.157 psid
and the initial pressure drop ∆Pi was 0.144 psid, or ∆Pf/∆Pi = 1.09. Using Equation 3.4,
the ratio of final radius to initial radius was determined to be Rf/Ri = 0.98, which is a
2% reduction. Using the initial clean tube inside diameter of 0.027 in. (0.686 mm), an
average deposit thickness can be calculated to be 7.3 µm (0.29×10−3 in.). For comparison
purposes, Faith et. al. have reported deposit thicknesses of 130 - 250 µm (0.005 - 0.01 in.)
for cases with oxygenated fuel, and thicknesses of less than 25 µm (0.001 in.) for de-
oxygenated fuel [6]. Such comparisons must be made with caution, as the experimental
conditions in [6], including experimental setup, test duration, temperature and fuel prop-
erties, were different from the present experiments. Furthermore, the calculated deposit
thickness presented here is only a rough estimate, and provides no information on how
the deposit is distributed on the tube surface. As mentioned in Section 3.5.2, carbon
burn-off should be used in conjunction with the pressure drop measurements to provide
more information on the carbon deposits.
Run 3 Results
The experimental conditions for Run 3 were identical to those of Run 2, except for a
lower test section inlet temperature Tin of 93.3 °C (200 °F), which was below the autoxi-
Chapter 6. Results and Discussion 59
dation temperature range in most research efforts [19]. This experiment was performed
to provide a control condition to verify that the pressure drop increase that was measured
in Run 2 was indeed due to the carbon deposits formed under high fuel temperatures.
From Figure 6.2, the pressure drop only increased from about 0.162 psid to 0.165 psid,
for a ∆Pf/∆Pi ratio of 1.02, as compared to 1.09 from Run 2 at higher fuel temperature.
Given that the accuracy of the transducer is ±0.0008 psid, this increase of pressure drop
of 0.003 psid is just outside the accuracy range of the transducer, therefore it cannot
be concluded with certainty that any appreciable amount of deposit was accumulated in
Run 3.
It also must be noted that the starting pressure drop in Run 3 was significantly
higher than that of Run 2, due to the difference in fuel temperature. This dependency
of pressure drop on temperature will be discussed in more detail in following sections.
With these two results, it can be concluded that the pressure drop can be used
effectively to provide a qualitative trend of deposit formation.
Chapter 6. Results and Discussion 60
01
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Chapter 6. Results and Discussion 61
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Chapter 6. Results and Discussion 62
6.3.1 Test Section Temperature Measurement and Profiles
As the fuel flowed through the test section, it was heated by the test section heater,
and its temperature increased. Because the residence time of the fuel in the 2 in. (50.8 mm)
heated segment was only 0.055 seconds, the average bulk fuel temperature did not in-
crease significantly. Numerical simulations were used to provide insight into the axial
temperature profiles at various locations in the flow. The data for Runs 2 and 3 are
shown in Figure 5.1. The fuel temperatures at the centre line and near the wetted wall
are shown, as well as the average bulk fuel temperature. Only the temperature near the
wetted wall increased significantly to a level close to the wetted wall temperature. As a
result, the average bulk fuel temperature increased moderately.
The simulated temperature at the end of the heated segment was higher than the
measured Tout (Table 6.1). This was expected and could be attributed to the heat losses
that occur in the unheated sections, even when insulated.
The simulated average bulk temperature at 0 mm in Run 3 (Figure 6.3b) did not
match the measured Tin of 93.3 °C (200 °F), whereas the two quantities matched well
in Run 2 (Figure 6.3a). In the simulations, the fuel preheat temperature was used to
simulate the chemical reactions that occur in the preheating tubing more accurately.
Since the heat losses between the preheating section and test section can vary depending
on the insulation quality, the test section inlet temperature was not simulated exactly.
This discrepancy will be addressed in future work.
6.4 Pressure Drop Dependence on Temperature
6.4.1 Viscosity as a Function of Temperature
The pressure drop measurement, as shown earlier, has a strong dependence on the
bulk fuel temperature. As was explained in Section 4.6, the total pressure loss across a
Chapter 6. Results and Discussion 63
0 1 0 2 0 3 0 4 0 5 01 6 0
1 8 0
2 0 0
2 2 0
2 4 0
2 6 0
R u n 2 T e m p e r a t u r e a t w e t t e d w a l l T e m p e r a t u r e a t c e n t r e l i n e A v e r a g e b u l k f u e l t e m p e r a t u r e
��
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��
A x i a l P o s i t i o n ( m m )(a) Run 2
0 1 0 2 0 3 0 4 0 5 01 0 0
1 2 0
1 4 0
1 6 0
1 8 0
2 0 0
2 2 0
2 4 0
R u n 3 T e m p e r a t u r e a t w e t t e d w a l l T e m p e r a t u r e a t c e n t e r l i n e A v e r a g e b u l k f u e l t e m p e r a t u r e
���
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A x i a l P o s i t i o n ( m m )(b) Run 3
Figure 6.3: The axial temperature profiles for Runs 2 and 3 obtained from numerical sim-ulations. The temperature profile is shown for the 2 in. (50.8 mm) heated segment of thetest section. It can be seen that centre line temperature stays constant, while the tem-perature near the wetted wall increases significantly. The bulk average fuel temperatureincreases moderately.
Table 6.1: Measured steady state Tin and Tout for the two preliminary experi-ments, Run 2 and Run 3.
Temperature, °C (°F)
Tin Tout
Run 2 162.8 (325) 169.2 (336.6)
Run 3 93.3 (200) 120.6 (249.1)
pipe of circular cross section is linearly dependent on the fluid’s dynamic viscosity µ, as
shown in Equation 3.3, which is repeated below for convenience:
∆P =8µQL
πR4
Chapter 6. Results and Discussion 64
The dynamic viscosity is in turn a function of the fluid’s temperature, and it can
vary significantly across a wide range of temperatures. Unlike gases, for which estimates
of viscosity can be estimated from theoretical formulations [46], liquid viscosity data
is largely empirical. Having a good estimate of how the viscosity of the fuel changes
with fuel temperature will result in a better understanding of the physics behind the
pressure drop measurements, and also provide more insight on the accuracy and validity
of the quantitative measurements. To this end, empirical data from handbooks and a
semi-empirical formulation were used to validate the measurements.
6.4.2 Handbook Viscosity Data
Vargaftik provides the dynamic viscosity data for a kerosene fuel designated as T-
1 at different temperatures [47]. The Coordinating Research Council (CRC) [48] also
provides data for several properties for a variety of aircraft and rocket fuels, including
Jet A-1. In this handbook, the kinematic viscosity ν is provided, as well as the density
ρ. Since Equation 3.3 uses the dynamic viscosity, it must be calculated using the relation
µ = ρν. A curve fit was performed on the density and kinematic viscosity data from the
CRC handbook, and the resulting curves were multiplied to produce an estimate of the
dynamic viscosity dependence on temperature of Jet A-1.
6.4.3 Semi-Empirical Approximation
In addition to the handbook data, a semi-empirical formulation for the temperature
dependency of liquid viscosity on temperature was used. It is given in [46] and proposed
by Andrade [49] (Equation 6.1). It expresses the dynamic viscosity µ as an exponential
function of the inverse of the absolute temperature. However, this relation requires
that the viscosity be known at two temperatures in order to determine the arbitrary
parameters A and B.
Chapter 6. Results and Discussion 65
µ = A exp
(B
T
)(6.1)
It has been observed that the viscosity of a liquid falls on the same curve that can be
shifted to any temperature range. In other words, if the viscosity of the liquid is known
at only one temperature, this curve can be used to estimate the viscosities at all other
temperature as long as it is not higher than the boiling point [46]. This relation is known
as the Lewis-Squires chart and can be expressed as Equation 6.2 [46]:
µ−0.2661 = µ−0.2661T1+T − T1
233(6.2)
where µT1 is the known viscosity of the liquid at a given temperature T1. As stated
previously, this relation can be used to estimate the viscosity when the viscosity is known
at only one temperature, to within 5 - 15% accuracy [46]. Using Equation 6.2 and the
maximum specified viscosity limit of Jet A-1 fuel, an approximate viscosity-temperature
curve can be plotted. The specified maximum viscosity of Jet A-1 in Canada is 8.0 mm2/s
at −20 °C (−4 °F) [50]. Multiplying by the maximum specified density of 840 kg/m3 [50],
a value of 6.72 mPa·s is obtained and used as µT1 in Equation 6.2. The above three
viscosity-temperature curves are plotted in Figure 6.4. It is apparent that the two sets of
handbook data are closely matched. The Lewis-Squires approximation is only accurate
at higher temperatures, but can be made more accurate if the viscosity of the current
batch of fuel is known at a specific temperature.
6.4.4 Effect on Pressure Drop Measurements
There are two ways in which this relationship between viscosity and temperature
affects the pressure drop measurement across the test section. First, the bulk fuel tem-
perature at the inlet to the test section, Tin, will affect the viscosity of the fuel, which in
turn will affect the measured pressure drop according to Equation 3.3. In the present ex-
Chapter 6. Results and Discussion 66
- 2 0 0 2 0 4 0 6 0 8 0 1 0 0 1 2 0 1 4 0 1 6 0 1 8 0 2 0 00 . 00 . 51 . 01 . 52 . 02 . 53 . 03 . 54 . 04 . 55 . 05 . 56 . 06 . 57 . 0
Dyna
mic V
iscos
ity (m
Pa-s)
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Figure 6.4: Comparison between the dynamic viscosity of jet fuel, including handbookdata from Vargaftik (T-1 kerosene fuel) [47], the Coordinating Research Council (Jet A-1) [48], and a calculated curve using the Lewis-Squires approximation for liquid viscosity(Eq. 6.2) [46]. The data from the two handbooks are closely matched, but the Lewis-Squires approximation appears to be more accurate at higher temperatures.
periments, Tin can vary from 65.6 °C (150 °F) to 162.8 °C (325 °F). The difference between
the pressure drop measurements of Run 2 and Run 3 (Figure 6.1 and 6.2 respectively)
demonstrates this effect of temperature on pressure drop.
The second way that viscosity affects the pressure drop is more important to the
validity of the data. Due to the non-isothermal temperature profile of the test section,
as discussed in section 6.3.1, the viscosity of the fuel is also not constant throughout the
test section. The bulk fuel temperature increases as it is being heated by the heater,
resulting in a higher exit temperature Tout than inlet temperature Tin. This increase in
Chapter 6. Results and Discussion 67
temperature corresponds to a decrease in the viscosity. Thus, the measured pressure
drop across the test section is lower than that measured in the case of an isothermal test
section temperature profile.
L
xδx
RQ
Figure 6.5: The geometry for calculating the pressure drop across a test section of con-stant cross-sectional area with constant volumetric flow rate.
To quantify this difference, the change in temperature and viscosity along the test
section must be taken into account. This can be done by dividing the length L of the test
section into small segments of length δx, where x is the coordinate along the length of
the test section. The test section has radius R and fuel flow is at a constant volumetric
flow rate Q, as shown in Figure 6.5. The bulk fuel temperature T is a function of x,
or T = T (x), and the viscosity µ is in turn a function of T , or µ = µ(T ). Thus, the
viscosity can be expressed as µ(x) = µ(T (x)
). Applying Equation 3.3 to a segment from
x to x+ δx, the pressure drop δP across the length δx can be expressed as follows:
δP =8Q
πR4µ(T (x)
)δx (6.3)
If δx is an infinitesimal quantity, Equation 6.3 can be integrated in x from 0 to L, to
obtain the total pressure drop ∆P :
∆P =8Q
πR4
∫ L
0
µ(T (x)
)δx (6.4)
To calculate the pressure drop for the measurements made in this thesis, Equation
Chapter 6. Results and Discussion 68
6.4 is written as a sum of the pressure drop over n discrete lengths ∆x:
∆P =8Q
πR4
n−1∑k=0
µ(T (xk)
)∆x (6.5)
where ∆x = L/n and xk = k∆x. Note that this method uses the left endpoint as the rule
for numerical integration. The pressure drops measured in the test section were calculated
using Equation 6.5 for the three different viscosity-temperature curves (Figure 6.4), and
the results are tabulated in Table 6.2. The results show that the calculated pressure
drops are in the same order or magnitude as the measured values. The differences can
be attributed to the estimation of the axial temperature profiles and the temperature-
viscosity profiles. Nevertheless, these calculations show that the measured pressure drops
are valid and they provide a better understanding of the physical phenomena in the test
section.
Temperature Profile Assumptions
To obtain the results in Table 6.2, an axial temperature profile T (x) must be assumed
for the pressure drop calculations using Equation 6.5, since direct measurement of the
fuel temperature was not possible. The simulated bulk fuel temperature profiles in Figure
6.3 was used for the 2 in. (50.8 mm) heated segment. However, the measured pressure
drop was for the entire 3.25 in (82.55 mm) test section. Therefore, linear profiles were
assumed for the 0.625 in. (15.88 mm) unheated and insulated segments that are on either
side, with the measured Tin and Tout as the end boundaries.
6.5 Sources of Experimental Error
Variations in Procedure
The amount of deposit that is produced for any given session can vary due to slight
variations of the procedure. As mentioned in Section 5.3, the heating procedure is such
Chapter 6. Results and Discussion 69
Table 6.2: Comparison between the measured pressure drop data against pressure dropscalculated from empirical and semi-empirical data from the literature.
Pressure Drop, psid (kPa)
Tin = 93.3 °C (200 °F) Tin = 162.8 °C (325 °F)
Measured Value 0.163 (1.12) 0.144 (0.99)
Vargaftik Handbook [47] 0.311 (2.15) 0.227 (1.57)
CRC Handbook [48] 0.255 (1.76) 0.182 (1.26)
Lewis-Squires Approximation [46] 0.327 (2.25) 0.187 (1.29)
that the timed portion of the experiment session is started before the temperatures of the
system have reached steady state. This is done because the slow heating rates of the brass
block and heaters expose the test section to temperatures that are close to test conditions
for non-negligible periods of time. During the approach to steady state, the timing of the
heating of the test section was determined by trial-and-error. If the test section heating
was started slightly earlier or later for a given session, there would be slightly different
amounts of coking that formed in the test section. This error is difficult to quantify, but
it is minimal since the time approaching steady state is brief compared to the duration of
each session. To minimize this variation in coking deposits, the experimental procedure
must be established precisely for each test condition and followed exactly throughout all
test sessions.
Insulation and Heat Losses
The insulation that is applied around the test section is not always exactly identical for
each experiment, since it has to be replaced for each time a new test section is installed.
Chapter 6. Results and Discussion 70
Initially, the thermocouple that is measuring Tout was not insulated, and in the case
where Tin = 162.8 °C (325 °F) Tin, Tout was measured to be 1 °C (1.8 °F) lower than Tin.
This is clearly unreasonable since heat is added to the fuel in the test section. When
insulation was applied to the Tout thermocouple, Tout was 6 °C (10.8 °F) higher than the
same Tin of 162.8 °C (325 °F). Even with insulation, however, the actual Tout of the test
section would be higher than measured, which can affect the pressure drop calculations
discussed previously.
The quality of the insulation also affected the difference between the measured Tin
and the oil bath set temperature. With poor insulation, the heat losses were greater,
and therefore the oil bath must be set at a higher temperature, which can affect the
deposition in the test section.
Pressure Transducer Accuracy
From the pressure drop measurement results shown in Section 6.3, it can be seen that
the differential pressure transducer was accurate enough, with an accuracy of ±0.08% of
full span, to capture very slight changes in the the pressure drop. However, it was crucial
to ensure that the transducer is correctly zeroed. It was found during the installation
of the transducer that it was sensitive to orientation. When the pressure ports were
oriented vertically, a pressure drop of approximately 0.01 psid was measured. Therefore,
care was taken to ensure horizontal orientation of the pressure transducer by monitoring
the readout and securing the transducer when the readout read zero.
Pressure Fluctuations
As discussed in Section 6.3, the pressure drop across the test section fluctuated peri-
odically due to pump switching that is required for constant flow operation. The mag-
nitude of this fluctuation is approximately 0.003 psid (21 Pa), which is greater than the
transducer’s accuracy. With this fluctuation, an error of ±0.0015 psid (±10 Pa) can be
attached to the pressure transducer measurements.
Chapter 7
Conclusion
An experimental apparatus was designed and constructed to simulate the geometry
and operating conditions of injector fuel nozzles in gas turbine engines, in order to study
how the thermal stability of jet fuel affects the carbon deposit formation, or coking,
in the fuel injector passages. These passages are characterized by their small size and
temperatures that are higher than those of the rest of the fuel system. The experimental
apparatus utilized some basic components from the previous thermal stability test rig
designed by Wong [13], with the design and addition of new heating elements in order to
be able to independently control two time scales and temperature scales, namely tpreheat,
ttest, Twall and Tin.
Pressure drop across the test section, measured by a differential pressure transducer
of 1 psid (6.895 kPa) range and recorded with an automated data acquisition system, was
the primary method of measurement and analysis. Two preliminary experiments were
performed with Jet A-1 fuel at test section inlet bulk fuel temperatures of 93.3 °C (200 °F)
and 325 °F (162.8 °C), with the wetted wall temperature of the test section kept con-
stant at 260 °C (500 °F). It was found that the higher fuel temperature fuel produced
a higher pressure drop increase over the duration of the experiment, and this increase
can be qualitatively attributed to a greater amount of carbon deposits. As expected, the
71
Chapter 7. Conclusion 72
case with the lower fuel temperature produced little measurable pressure drop increase,
suggesting that little deposit was formed.
The pressure drop measurements were verified by applying the Hagen-Poiseuille equa-
tion (Equation 3.2) for pressure drop in laminar flow through a cylindrical tube. Correc-
tions for the viscosity variations with fuel temperature were taken into account, and the
resulting calculated pressure drops were on the same order of magnitude as the measured
data.
The increase of pressure drop over the duration of the experiment was used with the
Hagen-Poiseuille equation to calculate an average deposit thickness of 7.3 µm (0.29 ×
10−3 in.) for the Tin = 162.8 °C (325 °F) case.
7.1 Recommendations and Future Work
The following recommendations are made for improvement to the apparatus itself,
and to improve its readiness for further experimental studies.
1. The insulation around the test section can be improved. Thermocouples may be
added to the tubing under the insulation to more accurately measure the test section
temperature profiles;
2. A new batch of fuel was received that is different from the batch that was used in
the preliminary experiments. The new fuel will be used for the formal experiments
to study the various parameters. It is recommended that the JFTOT tests be
performed on the new fuel to compare it to the current batch of fuel. The fuel’s
viscosity should ideally be measured at various temperatures for more accurate
pressure drop calculations;
3. A carbon determinator instrument, such as the Leco RC-412, is recommended for
the implementation of the carbon burn-off method to accurately determine the
Chapter 7. Conclusion 73
amount of deposits. As mentioned previously, this can be combined with pressure
drop measurements to gain more insight into the deposit distribution along the test
section;
4. Perform preliminary tests for the 1/4-in. tubing with flow constriction test sections
as described in Section 4.4.2. These test sections have smaller diameters and should
have higher pressure drops, therefore a differential pressure transducer of larger
range may be required.
The experimental apparatus will be used to study a variety of different test parameters
in the long-term goals of the project:
1. Perform experiments for a wide range of parameters such as temperature, pressure,
flow rates and test durations to fully characterize coking as a function of these
parameters;
2. Investigate the effects of long-term storage on the thermal stability of the fuel and
how it affects coking;
3. Study other parameters, such as metal surface material, fuel composition and ad-
ditives, and thermal soakback conditions that are important factors in coking;
4. Develop correlations with the data from the numerical simulations to provide a
useful tool in gas turbine design.
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