design and development of an experimental apparatus to ... · abstract design and development of an...

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


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.


Tab

le2.

1:C

omm

erci

alai

rcra

ftfu

elen

vir

onm

ent,

flig

ht

idle

(des

cent)

condit

ions.

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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

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)75

-100

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Gen

erat

oroi

lco

oler

10-1

524

5-32

0(1

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60)

40-1

0032

0(1

60)

Fuel

noz

zles

0.15

-0.8

300-

400

(150

-200

)40

-100

450-

500

(232

-288

)


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.


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.


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”


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].


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


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.


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


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


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


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.


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. 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


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


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


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.


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


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.


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.).


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


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


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


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.


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


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


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


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


(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


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


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.


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.


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.


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.


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).


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


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


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


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.


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


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-


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


• 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


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


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-


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.


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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


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


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.


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-


- 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


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


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


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.


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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