CHEMICAL KINETIC AND THERMAL NUMERICAL ......Abstract CHEMICAL KINETIC AND THERMAL NUMERICAL SIMULATION OF COKING PROCESS OF JET FUELS IN THIN NOZZLE SECTIONS WITHIN AUTOXIDATION TEMPERATURE

CHEMICAL KINETIC AND THERMAL NUMERICALSIMULATION OF COKING PROCESS OF JET FUELS INTHIN NOZZLE SECTIONS WITHIN AUTOXIDATION

TEMPERATURE REGIME

by

Zhao (John) Liu

A thesis submitted in conformity with the requirementsfor the degree of Master of Science

Graduate Department of University of Toronto Institute of AerospaceStudies

University of Toronto

c© Copyright 2014 by Zhao (John) Liu

Abstract

CHEMICAL KINETIC AND THERMAL NUMERICAL SIMULATION OF COKING

PROCESS OF JET FUELS IN THIN NOZZLE SECTIONS WITHIN

AUTOXIDATION TEMPERATURE REGIME

Zhao (John) Liu

Master of Science

Graduate Department of University of Toronto Institute of Aerospace Studies

University of Toronto

2014

The autoxidation and deposition process of air saturated Jet A-1 was simulated for

injection nozzle-like test sections with preheated fuel in the range of autoxidation regime.

The model also includes a preheating section to represent upstream condition of injection

nozzles. 2-D computational fluid dynamics along with a pseudo-detailed chemical kinetics

model was used to model deposition. Velocity and temperature profiles can significantly

influence the chemical reactions. Due to temperature differences, local concentrations of

species varied throughout the preheating and test sections. Deposition rate in the test

section is dependent on both upstream and local conditions. Higher temperatures in the

test section resulted in increased deposition rate. To determine the deposition rate at

a certain location, the entire history of the upstream conditions or the exact chemical

composition at the location of interest is necessary.

ii

Acknowledgements

Studying at UTIAS had been an illuminating experience. First I would like to thank Dr.

Omer Gulder. The guidance he had given me during my study at UTIAS is invaluable.

Thanks to Pratt & Whitney Canada for providing the funding and valuable feedback of

the project.

I also want to give thanks to all of my colleagues; especially to Owen Wong, who provided

the previous experiential results and test rig; to Jason Liang, who designed and built the

new experimental rig to accompany this study, and Frank Yuen, for helping with the

construction of the rig and advice on experimental and numerical studies.

Lastly, I am thankful for my parents and my friends, for being there with me for all the

ups and lows in the past few years.

iii

Contents

1 Introduction 1

1.1 Aviation Fuel Heat Sink Utilization in Aircraft . . . . . . . . . . . . . . . 1

1.2 Motivation of Thermal Stability Research . . . . . . . . . . . . . . . . . . 2

1.2.1 Degeneration of Aviation Fuel at Elevated Temperatures . . . . . 2

1.2.2 Future Trend of Aircraft Development . . . . . . . . . . . . . . . 3

1.2.3 Previous Thermal Stability Studies . . . . . . . . . . . . . . . . . 3

1.2.4 Coking in Fuel Injection Nozzles . . . . . . . . . . . . . . . . . . . 4

1.3 Objective . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2 Literature Review 6

2.1 The Mechanism and Regimes of Jet Fuel Degeneration . . . . . . . . . . 6

2.2 Physical Effects on Autoxidation . . . . . . . . . . . . . . . . . . . . . . 7

2.2.1 Effect of Temperature on Autoxidation . . . . . . . . . . . . . . . 7

2.2.2 Effect of Pressure . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.3 Effect of Oxygen on Autoxidation . . . . . . . . . . . . . . . . . . . . . . 11

2.4 Chemical Effects on Autoxidation . . . . . . . . . . . . . . . . . . . . . . 12

2.4.1 Hydrocarbon Content and Autoxidation . . . . . . . . . . . . . . 12

2.4.2 Effect of Heteroatomic Species . . . . . . . . . . . . . . . . . . . . 15

2.5 Chemical Kinetics of Autoxidation . . . . . . . . . . . . . . . . . . . . . 16

2.5.1 Autoxidation of Paraffins . . . . . . . . . . . . . . . . . . . . . . . 16

2.6 Model and Simulation of Autoxidation and Deposition of Jet Fuels . . . . 18

2.6.1 General Schemes of Autoxidation Deposition . . . . . . . . . . . . 18

2.6.2 Global Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

2.6.3 Pseudo-Detailed Chemistry Model . . . . . . . . . . . . . . . . . . 26

3 Experimental and Numerical Setup 32

3.1 Previous UTIAS Single Tube Heat Exchanger Apparatus . . . . . . . . . 32

iv

3.2 Previous Experimental Setup . . . . . . . . . . . . . . . . . . . . . . . . 32

3.3 Design of New Experimental Apparatus . . . . . . . . . . . . . . . . . . . 36

3.3.1 Test Section and Test Section Heating System . . . . . . . . . . . 36

3.3.2 Preheating System . . . . . . . . . . . . . . . . . . . . . . . . . . 38

3.4 Data Acquisition/Analysis Techniques . . . . . . . . . . . . . . . . . . . 39

3.4.1 Pressure Drop and Deposit Formation . . . . . . . . . . . . . . . 39

3.5 Experimental Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . 40

3.5.1 Experimental Parameters . . . . . . . . . . . . . . . . . . . . . . 40

3.5.2 Experimental Procedures . . . . . . . . . . . . . . . . . . . . . . . 41

3.6 Simulation of Jet Fuel Autoxidation . . . . . . . . . . . . . . . . . . . . . 41

3.6.1 Chemical Kinetics Model . . . . . . . . . . . . . . . . . . . . . . . 41

3.6.2 Initial Conditions of Simulations . . . . . . . . . . . . . . . . . . . 44

3.6.3 Computational Fluid Dynamics . . . . . . . . . . . . . . . . . . . 45

3.6.4 Simulations for the Deposition Studies . . . . . . . . . . . . . . . 49

4 Results and Discussion 51

4.1 Simulations of Temperature and Species Concentration in Preheating Section 51

4.1.1 Temperature Profiles . . . . . . . . . . . . . . . . . . . . . . . . . 51

4.1.2 Oxygen Profile . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

4.1.3 Hydroperoxide (ROOH) . . . . . . . . . . . . . . . . . . . . . . . 56

4.1.4 Velocity Profile and Radial Diffusion Effects . . . . . . . . . . . . 57

4.2 Simulations of Deposition in Test Section . . . . . . . . . . . . . . . . . . 61

4.3 Comparison with Experimental Results . . . . . . . . . . . . . . . . . . . 63

5 Conclusions and Recommendations 66

5.1 Suggestion for Future Work . . . . . . . . . . . . . . . . . . . . . . . . . 67

Bibliography 68

v

List of Tables

2.1 Critical properties of current jet fuels and n-paraffins . . . . . . . . . . . 11

2.2 Paraffin autoxidation mechanisms. . . . . . . . . . . . . . . . . . . . . . 18

2.3 Global thermal-oxidative deposition mechanism from Katta and Roquemore 24

2.4 Revised global thermal-oxidative deposition mechanism from Ervin and

Williams . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

2.5 Pseudo-detailed mechanism for liquid phase jet fuel autoxidation by Zabar-

nick et. al. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29

3.1 Commercial aircraft fuel enviroment, flight idle (descent) conditions . . . 37

3.2 Summary of test section dimensions. . . . . . . . . . . . . . . . . . . . . 38

3.3 Summary of fuel temperature experiments. . . . . . . . . . . . . . . . . 40

3.4 Summary of wetted wall temperature experiments. . . . . . . . . . . . . 40

3.5 Revised pseudo-detailed chemical kinetics used for simulation . . . . . . . 42

3.6 Deposition reactions used in simulation . . . . . . . . . . . . . . . . . . . 43

3.7 One step wall reactions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44

3.8 Initial concentration of species used in simulation of jet A-1 fuel. . . . . 45

3.9 Simulations for study of deposit formation in test sections . . . . . . . . 49

vi

List of Figures

2.1 The regimes of jet fuel deposition . . . . . . . . . . . . . . . . . . . . . . 7

2.2 Deposition rate as a function of oxygen concentration . . . . . . . . . . . 9

2.3 Oxygen level and depsition rate in JFTOT experiment . . . . . . . . . . 13

2.4 Taylor’s two-step model of autoxidative deposition process. . . . . . . . 19

2.5 Pathways of deposit formation . . . . . . . . . . . . . . . . . . . . . . . . 20

2.6 Major reaction paths of autoxidation in pseudo-detailed mechanism . . . 29

3.1 Thermal stability test rig developed at UTIAS (old) . . . . . . . . . . . . 33

3.2 Schematic of modified rig for thermal stability studies . . . . . . . . . . . 35

3.3 Assembly and cross section of the test section and heating block . . . . . 36

3.4 Major physical components modeled in the simulation. . . . . . . . . . . 45

3.5 Grid setup of simulation. . . . . . . . . . . . . . . . . . . . . . . . . . . 46

4.1 Comparison between simulated and measured temperature for 112 ◦C (224◦F) and 188 ◦C (371 ◦) at the outlet of the preheating section. . . . . . . 52

4.2 Temperature profile along the length of preheating section at three differ-

ent radial locations (at r = 0, 1/2 r and r) for various preheating temper-

atures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 55

4.3 Oxygen concentration profile along the length of preheating section at

three different radial locations (at r = 0, 1/2 r and r) for various preheating

temperatures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 58

4.4 ROOH concentration profile along the length of preheating section at three

different radial locations (at r = 0, 1/2 r and r) for various preheating

temperatures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59

4.5 Reynolds number and average velocity profile for 140 ◦C (284 ◦F) and 260◦C (500 ◦F), S1 and S4 respectively . . . . . . . . . . . . . . . . . . . . 60

4.6 Deposition rate at test section outer wall temperature of 266 ◦C (510 ◦F)

with various preheating temperature. . . . . . . . . . . . . . . . . . . . 62

vii

4.7 Deposition rate at test section with test section heater temperatures of

266◦ and 185◦ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

4.8 Deposition rate at test section corresponding to experimental runs. . . . 65

viii

Chapter 1

Introduction

1.1 Aviation Fuel Heat Sink Utilization in Aircraft

Aircraft thermal management is one of the key drivers in commercial and military aircraft

design. An aircraft produces considerable amount of heat during operation. Sources of

these heat include electronics, hydraulics, aerodynamics and combustion systems. Heat

generated in the aircraft must be removed to maintain systems and components within

desired operation temperatures. An aircraft utilizes a variety of cooling technology,

including mechanical cooling systems, forced air cooling, and use of heat sinks in thermal-

mechanical systems.

Jet fuel is the preferred heat sink for engine systems, hydraulic systems and elec-

tronics in modern aircraft [1,2]. Fuel-cooled systems are more compact due to the higher

density, specific heat and thermal conductivity of jet fuel. They also avoid the loss of

efficiency due to the need to divert air streams in air-cooling system. Air and mechanical

cooling systems are also heavier, more complex and larger in volume [3,4]. They include

many components such as air bleeds, ejector systems and mechanically driven fans that

are not necessary in a fuel-cooled system. Lastly, jet fuel is able to retain thermal energy

thus benefiting the engine thermodynamic cycle.

1

Chapter 1. Introduction 2

Fuel-cooled systems can be further divided into direct or indirect systems. In direct

cooling systems, heat exchangers are integrated into the high temperature components

so that heat is directly transferred to fuel in the fuel tanks or fuel entering the engine.

On the other hand, in indirect systems, ram air or compressor bleed air are cooled by

fuel and in turn used to cool other components.

1.2 Motivation of Thermal Stability Research

1.2.1 Degeneration of Aviation Fuel at Elevated Temperatures

Despite the advantages of using jet fuel as heat sink, high fuel temperatures introduce

new problems. At elevated temperatures, chemical reactions occurring within the fuel

may result in the formation of solid carbon deposits. There are two types of reaction

regimes responsible for the production of solid deposits. The first of which requires the

presence of dissolved oxygen and begins at a relatively lower temperature of around

150 ◦C (302 ◦F), known as autoxidation [5–7]. At around 400 ◦C (752 ◦F), a different

set of reaction mechanisms called pyrolysis takes place without the presence of oxygen.

Pyrolysis reactions involve the break-down and recombination of hydrocarbon chains

resulting in the formation of solid deposits [7, 8].

The presence of solids in fuel systems can adversely affect aircraft performances.

Deposit build up along the wall of fuel lines can alter flow characteristics and decrease

heat exchange efficiency in heat exchangers [9, 10]. Narrow passages such as valves and

injection nozzles are susceptible to deposit build-up and blockage. Blockages within the

fuel line may significantly increase the power required from the pumps to maintain flow

rate [11,12]. Blockage near the fuel injection nozzles can alter fuel spray patterns, caus-

ing change in flame characteristics which in turn may result in damage to combustion

components or even flame out. To avoid the production of carbon deposits, maximum

bulk fuel temperature and wetted wall temperature in aircraft are limited by fuel speci-


fications. United States Air Force (USAF) reports state the maximum fuel temperature

for primary jet fuels (JP-4, JP-5, Jet A, Jet A-1) allowed to reach before entering the

combustor is 163 ◦C (325 ◦F) [6]. Near the injection nozzle, wetted-wall temperature

may be significantly higher. For commercial aircraft, temperatures may reach 232 ◦C -

287 ◦C (450 ◦F - 547 ◦F) although the residence time is relatively short [6].

1.2.2 Future Trend of Aircraft Development

Thermal management of aircraft faces increasing challenges with the advance of aircraft

design. Modern turbine engines are more efficient. They operate at higher compression

ratios and lower fuel consumption. As a result, temperatures in combustion chambers

and turbines are increased while fuel flow rates are decreased [4]. Thus, the overall cooling

capacity available from fuel is decreased. Another trend in aircraft design is the increasing

usage of electrical power. Although advanced avionics uses less power, the increased

usage of electronics in modern aircraft by far outpace the increase in efficiency [2]. As

an example, the Boeing 777 introduced in the 1990s used about 200 kW of on-board

electric power, while the newly introduced 787 Dreamliner is expected to require close to

1 MW [13].

1.2.3 Previous Thermal Stability Studies

The trend in engine and electronics development means that future aircraft will have

to deal with higher heat load and a more complex thermal environment. A robust and

effective thermal management system that is well integrated with propulsion and power

system is crucial. There is also a growing interest in elevating the cooling capability

of jet fuels. To achieve these goals, a precise knowledge of the physical processes and

mechanisms of deposit production is necessary. Over the past decades, there have been

numerous studies on the thermal stability and coking characteristics of jet fuels. However,

deposit formation is an extremely complex process that involves several levels of chemistry


which may also be significantly affected by operating conditions such as temperature,

pressure, flow characteristics and materials that come into contact with the fuel.

There are currently two areas of focus of research in order to raise the cooling

capacity of jet fuels. The first is the development of advanced, thermally stable jet fuels

such as the JP-8+100 and JP-900 [14, 15]. The other is the fundamental research on

chemical kinetics and operation conditions with the aid of computer simulations to more

accurately predict deposition behavior in flow systems.

1.2.4 Coking in Fuel Injection Nozzles

Coking in the injection nozzle is of particular interest since it is one of the most com-

mon problem in aircraft related to thermal stability. Jet fuel is exposed to the highest

temperatures in the injection nozzles due to being close to the combustion chamber. In

many cases, surface temperature of the injection nozzle can even be higher than the limit

specified in fuel specifications. In addition, injection nozzles passages are much thinner

compared to other jet fuel passages. The inner diameter of injection nozzle can be as

small 0.01 in.(0.254 mm ) [16]. The small diameter and high temperatures make injection

nozzles much more susceptible to deposit formation and accumulation.

1.3 Objective

The objective of this project is to develop a numerical and experimental framework

to study jet fuel deposition process in injection nozzles under conditions similar to a

realistic aircraft fuel delivery system. The project consists of an experimental part and

a numerical part. A small-scale test rig capable of stressing the fuel at conditions that

resemble passing through the nozzle section in an aircraft is developed. This thesis will

present the numerical portion of the project. The simulation will utilize the most detailed

autoxidation deposition chemistry model to date to investigate the influence of chemical


and physical factors on deposition rate such as temperature, geometry, flow rate and fuel

chemical composition. The numerical simulation will be based on the experimental setup

presented in detail by Liang [17]. This study will focus on the autoxidation regime since

fuel in an aircraft fuel system is mostly exposed to temperatures within this regime.

Chapter 2

Literature Review

2.1 The Mechanism and Regimes of Jet Fuel Degen-

eration

The thermal stability of aviation fuel has a strong dependence on temperature as it dic-

tates the mechanisms that produce solid carbon deposits. The degeneration process of

jet fuel is divided into two categories based on the underlying chemistry: autoxidation

and pyrolysis. Study of thermal stability is further broken down into 3 regimes based on

temperature. Autoxidation occurs at relatively lower temperatures ranging from 150 ◦C

(302 ◦F) to 260 ◦C (500 ◦F) and pyrolysis occurs at temperatures above 400 ◦C (752 ◦F).

The temperature range from 260 ◦C (500 ◦F) to 400 ◦C (752 ◦F) is considered a transition

regime in which both autoxidation and pyrolysis may take place [8, 18]. The chemistry

taking place with respect to temperature is outlined in Figure 2.1. In the autoxidation

regime, oxygen content significantly affects the deposit formation process. Autoxidation

rate increases with temperature until the upper limit of autoxidation regime [3, 19, 20].

From then on deposit formation due to autoxidation decreases with temperature until the

pyrolytic regime. If the oxygen dissolved in the fuel is eliminated the deposit formed by

autoxidation reaction is reduced or suppressed. Past the autoxidation regime hydrocar-

6

Chapter 2. Literature Review 7

bons begin to decompose through pyrolysis reactions, known as hydrocarbon cracking,

which are heavily dependent on temperature, fuel composition and time exposed at high

temperatures [21].

Figure 2.1: Oxidative deposition involving oxygen ceases upon oxygen depletion. Astemperature further increases, cracking of hydrocarbon chains occur to initiate depositionformation in the pyrolytic regime. Free radical reactions are key in both autoxidationand pyrolytic reactions [5].

It is important to note that fuel degeneration does not necessarily lead to deposit

formation [22]. Both autoxidation and pyrolysis reactions may occur in liquid phase

within the bulk fuel. In fact, deposit formation is often regarded as a side branch of jet fuel

degeneration because only a relatively small portion of the fuel participates in reactions

that ultimately form solid deposit [6]. Fuel degeneration refers to all the chemical changes

within the fuel at elevated temperatures. However, solid deposition is one of the most

important aspects of jet fuel degeneration due to the negative effects it may bring to the

operation of an aircraft.

2.2 Physical Effects on Autoxidation

2.2.1 Effect of Temperature on Autoxidation

Temperature is the most important parameter in the study of jet fuel autoxidation. Ear-

lier global deposition models were based on the assumption that the amount of deposition

on metal surfaces is proportional to the amount of heat transferred [5,23]. However, later


studies have shown the effect of temperature on deposition rate to be much more com-

plex. Deposit formation occurs later in the autoxidation process. It is also possible for

autoxidation to occur without forming insolubles on either the wall or in bulk fluid un-

der certain conditions [22]. It should be pointed out that increasing the temperature

does in fact increase the rate of autoxidation reactions in liquid phase. This is observed

from increased oxygen consumption rate with respect to increased temperature in many

studies. However, faster autoxidation does not necessarily correlate to increased deposit

formation for many types of fuels. Observations show deposit formation occurs mainly

during oxygen (O2) depletion. Therefore, the effect of temperature on the coupling of

liquid phase and deposition formation reactions can be very complex.

Experimental work by Jones and Balster [24] using the near-isothermal flowing test

rig (NIFTR) with a single-pass heat exchanger has shown an inverse relationship between

the wall temperature and the amount of both wall and bulk insolubles (Figure 2.2).

This is the case for a large variety of jet fuels (Jet-A, JP-5, JP-8, JPTS). Moreover,

the experiments have shown that temperature has a more profound effect on deposition

in less thermally stable fuels. For thermally stable fuels, the effect of temperature on

deposition is relatively smaller. The explanation proposed for this inverse relationship

revolves around the most important species in the autoxidation process, hydroperoxides

(ROOH). ROOH is one of the least stable product in the mechanism. It is responsible

for carrying the chain propagation reactions and the formation of precursor species. At

higher temperatures, ROOH dissociates when coming in contact with hot stainless-steel

surfaces [25], thus reducing deposition. The explanation of more significant influence of

temperature on less thermally stable fuels revolves around heteratomic species, which

will be discussed in detail in the next section.

Most thermal stability studies are carried out under isothermal condition. However,

evidence have shown that it is possible for deposit to form in cooled regions, even an

order of magnitude greater than that of the heated region under certain conditions [26].


(a) Bulk/Surface deposit at 185 ◦C (b) Bulk/Surface deposit at 225 ◦C

Figure 2.2: Experiement of Balster and Jones [24] with NIFTR single pass heat exchanger.The apparatus consists of a test section stainless-steel tube (0.318 cm o.d, 0.216 cm i.d.)Bulk insolubles were collected by filters. Surface deposits are measured with carbonburnoff methods. At 225 ◦C (437 ◦F), most fuel types produce significantly less deposits.This inverse dependence of temperature and deposit amount is most apparent with fuelsthat have lower thermal stability.

Certain deposit forming precursors are believed to be soluble at higher temperatures,

and only starts to precipitate from the bulk fuel onto the wall once reaching a certain

critical temperature. In addition, testing with various additive packages with both heated

and cooled regions suggested that thermal stability additives may be less effective to

suppress deposit formation in cooled regions than heated regions. The implication of the

complex effect of temperature on autoxidation and deposit formation is that in order to

accurately predict deposition process and amount, the entire history of the flow conditions

(temperature and flow rate) of the fuel is necessary.

2.2.2 Effect of Pressure

Pressure may alter the characteristics of the flow resulting in changes in Reynolds number,

mass transfer and residence time. The effect of pressure on coking rate is drastically

different between autoxidation and pyrolytic regimes. Within the autoxidation regime,


pressure has a relatively small effect. Thus, pressure should have a minor impact on the

autoxidation deposit formation [27]. From the limited studies related to pressure effects

in the autoxidation regime, most data show no change or less deposition in air saturated

and deoxygenated fuels as pressure of the system is increased [27–29].

Pressure has a more erratic impact on deposit formation at higher temperatures.

At temperatures of approximately 370 ◦C (698 ◦F), fuels begin to reach critical tem-

peratures. If the pressure of the system is sufficiently low, fuel may exist in a partially

vaporized state [30], known as supercritical state. At supercritical conditions, pressure

can significantly alter the state of the fluid and change the residence times. Supercritical

jet fuel has a much higher solubility, thus may minimize the deposition in supercritical

sections while increase the deposition in downstream cooler sections. A list of critical

properties for selected jet fuels are shown in Table 2.1. Commercial Jet A (C11.6H22.0)

and Jet A-1 fuels have critical temperatures and pressures similar to JP-5 and JP-8.

Critical temperatures in Table 2.1 are calculated from the specific gravity and boiling

curve of the fuels by the method outlined in ASTM D2889. The complex state of the

liquid/vapor mixture induces turbulence to significantly improve heat and mass transfer.

In various studies with temperatures above autoxidation regime, no simple correlation

between deposit formation and pressure is observed. Depending on the fuel type and

degree of deoxygenation, deposit amount is observed to decrease, remain constant and in

some cases increase. Supercritical jet fuel deposition characteristics is beyond the scope

of this study. The maximum fuel temperature investigated in this study is 260 ◦C (500

◦F), well below the critical temperature. Pressure is expected to have small effect on the

results.


Table 2.1: Critical properties of current jet fuels and n-paraffins [30].

Fuel Type Tcrit,◦F (◦C) Pcrit, psia (atm)

JP-4 C8.5H16.9 550 - 690 (288 - 365) 500 - 400 (34.0 - 27.2)JP-5 C11.9H22.2 720 - 780 (382 - 415) 300 - 280 (20.4 - 19.0)JP-7 C12.1H24.4 760 (405) 270 (18.4)JP-8 C10.9H20.9 700 - 760 (370 - 405) 350 - 275 (23.8 - 18.7)

n−C8H18 564 (296) 365 (24.8)n−C12H26 726 (386) 263 (17.9)n−C16H34 836 (477) 200 (13.6)

2.3 Effect of Oxygen on Autoxidation

In the autoxidation regime, temperature, fuel composition and the amount of oxygen

present are the most significant parameters to determine the amount and rate of deposit

formation [19]. Dissolved oxygen in fuel reacts with fuel molecules to form alkylperox-

ides [31]. These species in turn form precursor species that ultimately lead to deposit

formation. It is widely believed that reducing the amount of dissolved oxygen in the fuel

will lead to reduced amount of deposit. The effect of deoxygenation and the case of zero

oxygen concentration has been studied extensively in the past. Results have universally

shown improved thermal stability after removing oxygen for a variety of commercially

available and experimental jet fuels [32,33]. Figure 2.3 shows the effect of dissolved oxygen

on deposition in a Shell aircraft fuel system simulator. The research reported a significant

decrease of deposition only when oxygen level is reduced below 30 ppm [34]. Some studies

reported an even lower oxygen level is needed to significantly reduce deposition rate [19].

The extent of suppression of deposits by deoxygenation in these studies was found to

vary considerably across fuel types. Evidence suggest these variations are caused by the

trace components containing sulfur, nitrogen and oxygen [19]. These trace species were

thought to behave chemically different in deoxygenated environments [35,36]. Numerous

onboard fuel deoxygenation methods, such as membrane separation and chemical based

systems are currently under investigation [3, 37,38].


The amount of dissolved oxygen in air-saturated jet fuel is dependent on the fuel

composition [39]. This may explain the large variation of reported value of oxygen con-

tent, ranging from 45 ppm to 70 ppm by mass fraction, in air saturated jet fuels. Due

to the strong dependence of deposition rate on oxygen content, the total amount of de-

posit obtained by complete conversion from oxygen present can be used as a measure of

the thermal stability of the fuel [40]. In a realistic aircraft fuel system, jet fuel maybe

be exposed to varying atmospheric conditions during landing, takeoff and cruising. In

addition, as jet fuel flows through the complex geometric and thermal enviroment of the

fuel system, only a portion of oxygen may react due to high flow rate and low residence

time through certain high temperature sections. Under conditions of incomplete oxygen

consumption, the relationship between oxygen level, oxygen consumed and deposition

rate can be extremely complex, especially when coupled with temperature [26]. How-

ever, most relevant studies focus on complete oxygen consumption and complete oxygen

removal.

Despite the fact that deoxygenation may decrease overall deposition, some experi-

mental evidence suggest deoxygenation may alter the deposition characteristic in a flow

system. For example, Ervin et. al. [26] reported in a study with a series of heated and

cooled sections, fuel deoxygenation may sometimes increase surface fouling under certain

conditions. Morphological studies of the deposits also suggest that the chemical and

physical process of deposit formation may be different in partially deoxygenated and air

saturated fuels.

2.4 Chemical Effects on Autoxidation

2.4.1 Hydrocarbon Content and Autoxidation

Jet fuel is comprised of approximately 98% or more of hydrocarbons. Although only

an extremely small portion of hydrocarbons within the fuel is ultimately converted into


Figure 2.3: The jet fuel thermal oxidation tester (JFTOT) experiment performed by AeroPropulsion Laboratory showed significant reduction of deposition rate via deoxygenationoccurred when oxygen level is below 30 ppm by weight [6, p. 101]. mil2 is square mil,one millionth of a square inch. BTU is British thermal unit, equivalent of about 1055 J.

insolubles via autoxidation, the chemical makeup of the hydrocarbons defines the base

for fuel autoxidation chemistry [6]. Most jet fuels are composed of mainly alkanes and

cycloalkanes [41]. Olefin content of up to 5% is allowed by many fuel specifications.

Aromatic hydrocarbons concentrations are limited to 20% or 25% depending on fuel

specifications [6].

Many studies have investigated the stability behaviour of pure hydrocarbons and

blends of pure hydrocarbons. Some earlier general findings related to paraffins are listed

below [42–44]:


• Deposit formation for n-alkanes moderately decreases as molecular size increases.

• Branched alkanes produce more deposit relative to n-alkanes.

• Olefins exhibits tendency to increase deposit formation. (Experiment carried out

at 10 wt% olefin blend, which is significantly higher than allowed amount in jet

fuel specifications.)

• Cycloalkanes (e.g. decalin) modestly inhibit the deposit rate of decane.

The effect of aromatics on autoxidation and deposition is very complex. Simple

one-ring benzenes have nominal effects on autoxidation rate and deposition [6, 45]. As

the molecular size and number of ring structures increase, autoxidation rate can be sig-

nificantly increased or decreased depending on the specific type of aromatics. Tetralins,

which may be produced by mild hydrotreatment of fuels containing naphthalene, signif-

icantly increases autoxidation rate by rapid consumption of oxygen and increasing the

radical pool [22, 31, 37]. The products from tetralin autoxidation readily forms solid de-

posits in the autoxidation regime. This effect is opposite of the general behaviour of

faster oxidation leading to less deposit formation exhibited by most fuels [46]. Indene

also significantly increases deposition in a paraffin blend [44], although unlike the case of

tetralin, autoxidation rate is slow. On the other hand, some aromatics such as fluorene

reduce autoxidation rate and deposition significantly [42]. These effects from aromatics

are most prevalent at lower temperatures in the autoxidation regime.

The ability to inhibit autoxidation shown by some aromatics is attributed to their

weakest C-H bonds with bond strength of 343 kJ/mol (82 kcal/mol) or less [6]. These aro-

matics behave like antioxidants, some categorize them as secondary antioxidants within

fuels [47,48]. The radicals formed from these types of aromatics via hydrogen abstraction

are relatively stable. In comparison, reactions with said radicals with oxygen is signifi-

cantly slower than alkyl radicals. However, it should be noted that reduction of oxidation

rate does not necessarily lead to less deposition.

Characteristics of aromatic species are also used to explain the discrepancy be-


tween autoxidation and deposition characteristic of jet-fuel/paraffin blends and straight

run/hydrotreated blends. In hydrotreated fuels, olefins are saturated and trace metal

species are eliminated. Aromatic content may also be reduced depending on the extent

of hydrotreatment. Naturally occurring nitrogen, sulfur, and oxygen containing het-

eroatom species are also reduced [48]. From these respects compositional changes caused

by hydrotreatment are similar to dilution of straight run fuels with paraffins. Studies

have shown that only a 9:1 blend of hydrotreated fuel with straight run, respectively,

will result in a more thermally stable fuel at 185◦C [47]. However, studies with jet fuel

and paraffinic/cycloparaffinic solvent blends have shown slower autoxidation rates across

all blend ratios [48]. The major difference between a hydrotreated blend and a paraf-

finic/cycloparaffinic blend is that the latter is signifcantly lower in aromatic content.

Thus, the reduced aromatic content in paraffin blends was considered the source of the

discrepancy.

2.4.2 Effect of Heteroatomic Species

Heteroatomic compounds containing sulfur, nitrogen and oxygen occur naturally in jet

fuels. Naturally occurring heteroatomic species are known to contribute to deposit forma-

tion [6, 35, 49]. Heteroatomic species slow down autoxidation by consuming free radicals

without further propagation of the reaction chain. However, the products from these

reactions can ultimately lead to deposit formation [47].

Sulfur containing compounds are the most abundant in jet fuels. Content of sulfur

species is limited to 0.3% in most commercial jet fuels [6]. The amount of nitrogen

and oxygen containing species is not limited in fuel specifications. However, they are

assumed to be present in low quantities in petroleum-derived jet fuels. Fuel processing

technologies such as hydrotreatment removes most of the natural heteroatomic species.

Such fuels are generally more thermally stable than straight-run fuels.

The major sulfur containing compounds in jet fuels are mercaptans (thiols), sul-


phides and thiophenes [50]. Studies with individual sulfur species showed increased de-

posit formation as thiols and most sulfides content are increased. Thiophenes produced

only moderate increase in deposits in comparison [35, 51, 52]. In commercial jet fuels,

good correlations can be observed between deposition rate and total sulfur content of the

fuel [6]. Studies on nitrogen and oxygenated species are not as detailed as sulfur com-

pounds. The level of nitrogen containing species in fuel is varied, commonly existing in

the forms of anilines, indoles, carbazoles, quinolines and pyridines. However, it is usually

much lower than oxygen and sulfur containing species. In general, increasing nitrogen

containing species decreases jet fuel stability [6].

2.5 Chemical Kinetics of Autoxidation

2.5.1 Autoxidation of Paraffins

Autoxidation refers to the fuel/oxygen catalytic chemical reaction process that involves

free radical chains. The autoxidation of paraffinic hydrocarbons is relatively well under-

stood. Many jet fuel autoxidation deposition models developed recently are based on

the oxidation of paraffins to avoid the complexity of modeling large number of different

hydrocarbons in jet fuels. The basic autoxidation scheme including the chain initiation,

propagation, and termination reactions are as follows [22]:

Initiation:

RH −−→ R · (2.1a)

ROOH −−→ RO · + OH · (2.1b)

2 ROOH −−→ RO2 · + RO · + H2O (2.1c)

Propagation:

R · + O2 −−→ ROO · (2.2)


RO2 · + RH −−→ ROOH + R · (hydroperoxide) (2.3a)

RO2 · + RH −−→ ROOR · ( ≡ R · )(polyperoxide) (2.3b)

RO · + RH −−→ radicals, products (2.4)

Termination:

ROO · + ROO · −−→ products (2.5)

ROO · + R · −−→ products (2.6)

R · + R · −−→ products (2.7)

Reactions 2.1a, 2.2, 2.3 and 2.5 - 2.7 are known as the basic autoxidation scheme

(BAS). The BAS and an alternative simplified version are listed in Table 2.2. In these

reactions, R denotes hydrocarbon chains, H and O represent hydrogen and oxygen atoms

respectively. Hydrocarbons (RH), hydroperoxides (ROOH) and oxygen are naturally oc-

curring species in jet fuels. The dissociation of hydrocarbons in the initiation reaction

may be caused by thermal decomposition or reaction with chemical initiators such as

metal ions, or ultraviolet light. In oxygen rich conditions the hydrocarbon radicals (R·)

promptly react with oxygen to form peroxy radicals (ROO·) in reaction 2.2. Propagation

continues with reactions 2.3a and 2.3b, forming hydroperoxides (ROOH) and polyper-

oxides (ROOR·( ≡ R · )). Hydroperoxides are the central species of autoxidation. Once

ROOH is introduced into the reaction chain, autoxidation is significantly catalyzed with

the dissociation of ROOH (2.1b, 2.1c). ROOH may also be present in fuel as a result of

prolonged storage, thus significantly decreasing the initial rapid autoxidation delay (rad-

ical build-up process). The unimolecular reaction 2.1b generally occurs at low ROOH

concentration whereas the bimolecular reaction 2.1c is likely to occur at higher ROOH

concentrations. The dominant termination reactions depend mostly on the amount of

oxygen present. High oxygen levels result in an abundance of alkyl peroxy radicals, thus


termination occurs mainly via reaction 2.7. If oxygen level is low, less peroxy radicals

are produced from reaction 2.2. Thus, the dominant radical species is a mixture of

alkoxy RO· and alkyl radicals R·. Under these conditions, reactions 2.6 and 2.7 are more

significant.

Table 2.2: Paraffin autoxidation mechanisms.

Hazlett [6] Balster and Jones [19]Initiation: Initiation:RH + X −−→ R·+ XH RH + O2 −−→ R·Propagation: ROOH −−→ RO·+ HO·R·+ O2 −−→ ROO· Propagation:ROO·+ RH −−→ ROOH + R· R·+ O2 −−→ ROO·Termination: ROO·+ RH −−→ ROOH + R·ROO·+ ROO· −−→ ROH + RCOR + O2 Termination:ROO·+ R· −−→ ROOR ROO·+ ROO· −−→ nonradical ProductsR·+ R· −−→ R−R

2.6 Model and Simulation of Autoxidation and De-

position of Jet Fuels

2.6.1 General Schemes of Autoxidation Deposition

The previous sections have outlined the key experimental observations regarding autox-

idation deposition. Any mechanism proposed for autoxidation deposit formation must

address these observations. In summary, it is proposed that an autoxidation deposit

formation mechanism must meet the following criteria [6, 53]:

1. dissolved oxygen initiates the process;

2. heteroatom-containing molecules should play an important role;

3. only a small amount of the fuel is involved in the deposit formation process;

4. the mechanism must account for the inverse relationship between the ease of oxi-

dation and the formation of deposits.


Point 4 is relatively new development in the study of autoxidative deposition stud-

ies (Hardy et. al. [54], Heneghan et. al. [53], and Jones [25]) therefore is not addressed

in mechanisms developed earlier. These earlier mechanisms include Bol’shakov’s the-

ory published in 1970 and described in [53], Clark and Smith’s two-step mechanism in

1988 [6, p. 92], and Taylor and Frankenfeld’s mechanism in 1986 [55]. The mechanism

proposed by Taylor [35] provided the most comprehensive description of the gross chemi-

cal and physical process (Figure 2.4). The mechanism involves two-step oxidation process

followed by physical deposition processes. The chemical and physical processes resemble

many of the more recently proposed models.

Figure 2.4: Taylor’s two-step model of autoxidative deposition process. The processincludes chemical and physical processes [35].

Surface deposits originate from two sources based on physical processes: direct

formation at the wall from precursor species or adhesion/accumulation of bulk insolubles

(Figure 2.5) [22]. In both cases the transfer of species and the characteristics of metal


surface are important in the deposition process.

Figure 2.5: Pathways of deposit formation [22].

The effects of surface on autoxidative deposition are not well understood. The

nature of the effect can be chemical or physical. Metal surfaces can influence surface

reactions via catalytic effects. The material of the surface is observed to significantly

influence deposition rate and characteristic [6, 56]. For example, deposition formation

rate within stainless steel tubes treated by adding a silica-based layer (Silcosteel) is

observed to be slower than untreated stainless steel tubes [57]. The silica-based layer

prevents direct contact of fuel and nickel containing stainless steel which is regarded as

more reactive. The variation of deposition rate due to material can also be physical in

nature, since smoother surfaces tend to inhibit deposition [25,58]. The effect of surfaces

is further complicated by the fact that the surface in contact with the fuel changes in

characteristic as deposit accumulates. Low initial deposition rate is often seen in treated,

clean and smooth surfaces. However, as test duration lengthens, the overall effect of

different metal surfaces becomes less substantial.

Recent studies have identified many key species (oxygen, hydroperoxides and polar

species) that influence deposition process. In addition, wall reactions are increasingly

common in recently developed chemistry models [23, 59]. Therefore knowing the species

and temperature distributions is essential in accurately modeling the thermal-oxidative

deposition process. Transport of reactants is mainly influenced by flow velocity and


geometry. The velocity, or flow rate for a system determines whether the flow is laminar

or turbulent. The flow regime in turn determines temperature and reactant distributions

within the test section. Accurate representation of the flow field of the system has been

becoming increasingly important with the development of more detailed chemical kinetics

models for thermal-oxidative deposition.

2.6.2 Global Models

The success of the simulation depends on the development of an accurate chemical kinet-

ics model. In a flowing system, the complex geometries and flow characteristics can be

accurately modeled with the current progress in computational fluid dynamics. On the

other hand, development of chemistry models has been challenging due to the lack of fun-

damental understanding of the complex nature of autoxidative deposition process. This

section will provide an overview of the global chemistry models developed in relatively

early studies. Early global models make the assumption of the amount of deposition be-

ing directly correlated to the amount of heat transferred or dissolved oxygen consumed.

The physical and chemical processes were combined into one or two steps [5, 23]:

Fuel −−→ Deposits (2.8)

dD

dt= αA2a exp

−E2a

RT(2.9)

Reaction 2.8 is a simple reaction that combines the autoxidation process and de-

position process in one step. Equation 2.9 represents the deposition rate at the surface.

D represents the deposit amount, α is a constant that may be a function of velocity, E2a

is the activation energy, A2a is the Arrhenius constant and finally T is the temperature.

Due to the over-simplifications of the mechanism, details of the physical and chemical

processes are masked. Thus, the model assumes the quantity of carbon deposited on wall


surfaces to be related to physical parameters such as the amount of heat transferred in

the duration of operation. A two step model was later proposed and applied successfully

to a number of heated-tube experiments by Giovanetti [60]:

Fuel + O2 −−→ Products (2.10)

Products −−→ Deposits (2.11)

The two step mechanism separates the autoxidation reaction and deposit reactions.

In the autoxidation reaction represented by Reaction 2.10, intermediate precursor prod-

ucts are produced. These precursor species are believed to include oxygenated species

such as hydroperoxides, alcohols, ketone and carbon monoxide. The oxygen consumption

rate and deposition rate are represented in equation 2.12 and 2.13 respectively.

−d[O2]

dt= k1[RH][O2]

n (2.12)

d[Deposits]

dt= k2[Products]n (2.13)

In equation 2.12, the hydrocarbons in fuel is represented collectively as RH. The rate

constants k1 and k2 are Arrhenius type and each one is dependent on a pre-exponential

factor and an activation energy term. The order of reaction is given by n, which is de-

pendent on operation conditions or concentration of certain species. In this mechanism,

oxygen consumption is tightly coupled to deposition rate. However, this is only true

when R· and ROO· are high in concentration. As discussed earlier in Section 2.3, oxy-

gen consumption rate does not directly correlate to deposit formation rate in jet fuels.

The significant amount of antioxidants present in the jet fuel limit the radical pool in

autoxidation reactions, thus producing a lag between oxygen consumption and deposit


formation.

In addition to the limitations mentioned above, simple global mechanisms also do

not take mass transfer effects into account. Studies on fuel velocity and deposition rates

suggested that at high temperatures, chemical reactions become limited by transport

[28]. At lower temperatures of the autoxidation regime, surface-catalyzed (heterogeneous)

reactions are dominant in the deposition process. At higher temperatures, homogeneous

reactions in the bulk fuel are dominant [4]. In summary, in order to develop an accurate

simulation of deposition rate to reflect the combined effects of temperature, oxygen and

minor species (such as antioxidants), detailed bulk and surface reactions are necessary.

In an attempt to incorporate effects of mass transfer and deposit thickness, a global

chemistry mechanism consisting of two bulk reactions and two surface reaction was in-

troduced by Katta and Jones [23]:

Bulk:

Fuel + O2 −−→ Precursor (2.14)

Precursor −−→ Soluble (2.15)

Surface:

Fuel + O2 −−→ Deposit (2.16)

Precursor −−→ Deposit (2.17)

The reaction rates of bulk reactions 2.14 and 2.15 are represented by Arrhenius

form. The reaction rates of equation 2.16 and 2.17 are expressed as a function of deposit

thickness in the form of:

Reaction Rate = A1 exp−E1

∆Dep

(2.18)


A1 is the pre-exponential factor and E1 is the activation energy. The term ∆Dep is a

factor in meters depicting the net result of the deposit thickness due to reactions 2.16 and

2.17. The 4-step model is further developed into a 9-step model consisting of eight species

(Table 2.3). The 9-step model includes radical species formed from reaction between fuel

and dissolved O2. Unlike the previous mechanism, bulk-insoluble and deposit-forming

precursors are produced in different pathways.

Table 2.3: Global thermal-oxidative deposition mechanism from Katta and Roquemore[61].

Reaction Number Reactions Pre-exponential factor A Activation Energy(mol, L, s) (kcal/mol)

Bulk Reactions:R1 F + O2 −−→ ROO· 2.5× 1013 32R2 ROO·+ F −−→ solubles 1× 104 10R3 ROO·+ Fs −−→ P 8× 109 15R4 ROO·+ F −−→ Dbulk 200 10R5 P + F −−→ solubles 3.2× 1012 30R6 Dbulk + F −−→ 2 Dbulk 1× 10−3 0Wall Reactions:R7 O2 + F −−→ P 5.2× 10−3 12R8 P −−→ Dwall 260 17R9 Dbulk −−→ Dwall 0.8 10

In the mechanism listed in Table 2.3, fuel is represented by F, precursor by P and

deposit on the surface by Dwall. A zero-order reaction was used for O2 consumption rate

in reaction R1. When O2 reaches concentration of 10 ppm, a first-order reaction is used.

The rate of deposition at the wall is expressed as a function of shear stress at the fuel-

deposit interface and the sticking coefficient representing the probability of P and Dbulk

to collide with and bind to the surface. The rate for Dwall is expressed as

d[Dwall]

dt=

S

τawk3[Oxidation Products]n (2.19)

In this equation, k3 is the rate constant, S is the sticking coefficient representing the


possibility of particles P and Dbulk colliding and sticking to the surface, τw is the wall shear

stress. The model provided better predictions of deposition compared to the previous

models with the use of more detailed chemical reactions and more realistic physical

processes. The 9-step global model is further developed to reflect deposition characteristic

in cooled regions by Ervin et. al. [26]. The modified global kinetic mechanism is outlined

in Table 2.4.

Table 2.4: Revised global thermal-oxidative deposition mechanism from Ervin andWilliams [26] .

Reaction Number Reactions Pre-exponential factor A Activation Energy(mol, L, s) (kcal/mol)

Bulk Reactions:1a F + O2 −−→ ROO· 4.2× 1013 322a ROO·+ F −−→ solubles 1× 104 103a ROO·+ Fs −−→ P 8× 109 154a ROO·+ F −−→ Dbulk 100 105a P + F −−→ P1 1× 106 156a P1 + F −−→ Dbulk 0.3 07a P1 + F −−→ solubles 1× 104 30Wall Reactions:8a O2 + F −−→ P 5.2× 10−3 129a P −−→ Dwall 260 1710a P1 −−→ Dwallc 1.0 1011a Dbulkc −−→ Dwallc 0.08

The mechanism introduces 2 new reactions and 3 new species P1, Dwallc and Dbulkc.

P1 is the precursor species responsible for forming deposit in the cooled regions. Dwallc

is wall deposit formed only when fuel temperature is cooled to a certain critical level.

Dbulkc are insolubles within the bulk fuel formed from bulk species that are soluble in

higher temperature but become insoluble at lower temperatures.

The global chemistry models outlined thus far have provided limited success in

predicting deposit formation from various studies. However, species used in global mech-

anism do not represent real species. Hence, the reactions does not represent the actual

chemistry of jet fuel autoxidation and deposition. The use of the global models requires


calibrations with experimental results, making the models fuel and system dependent.

Therefore, such models can not be effectively used for correlating results from different

experiments or in practical applications.

2.6.3 Pseudo-Detailed Chemistry Model

While the relatively few reactions from global mechanism are often inadequate to accu-

rately reflect autoxidation and deposition processes, a detailed kinetic mechanism may

include hundreds of reactions to describe each species and reactions associated with them.

Thus, a detailed chemical approach combined with computational fluid dynamics is not

practical for the simulation of thermal stability behaviour of jet fuel. Recent devel-

opments of autoxidation chemistry model focus on representing groups of species with

similar roles in autoxidation. Thus far, the most prevalent of such chemical models is

the pseudo-detailed model proposed by Zabarnick [59] and the soluble marcromolecular

oxidatively reactive speices (SMORS) mechanism.

Pseudo-Detailed Autoxidative Deposition Mechanism

The pseudo-detailed mechanism includes radical chain reactions and antioxidant reac-

tions. It is proposed in this mechanism that naturally occurring molecules with weakly

bonded hydrogen (antioxidants) are responsible for oxidation resistance and deposit for-

mation.

The mechanism revolves around radical chain reactions and the various ways of

chain breaking reactions related to antioxidants. A simplified autoxidation chain reaction

model is used (similar to those in section 2.5.1) to describe the formation of free radicals

as a result of hydrocarbon fuel oxidation:

Initiation:

Formation of R · (2.20)


Propagation:

R · + O2 −−→ RO · (2.21)

RO2 · + RH −−→ RO2H + R · (2.22)

Termination:

RO2 · + RO2 · −−→ products (2.23)

Reaction 2.21 proceeds with no activation energy while reaction 2.22 has a signif-

icant activation energy for most hydrocarbons. Therefore, in oxygen rich environments,

reaction 2.22 is the slow step and RO2· is the major radical species. Reaction 2.22 also

determines if a fuel is easily oxidized. Fuels containing hydrocarbons with weak R−H

bonds can significantly lower the activation energy of reaction 2.22. A chain-breaking

antioxidant reaction is introduced to the chain mechanism:

RO2 · + AH −−→ RO2H + A · (2.24)

The antioxidant species is represented by AH. AH represents species with an easy to

abstract hydrogen atom. Antixodants are also often referred to as inhibitors in literature.

Reaction 2.24 competes with reaction 2.22 to remove chain propagating RO2· radicals. In

order for this reaction to suppress oxidation process, the radical A· must not propagate

the reaction chain. It is assumed that the AH bond strength is weaker than the RH bond

strength, thus preventing the regeneration of R· radicals from A· radicals. It is possible,

however, for A· radicals to react with oxygen similar to reactions 2.21 and 2.22:

A · + O2 −−→ AO2 · (2.25)

AO2 · + RH −−→ AO2H + R · (2.26)

It is not clear how much the activation energy is different from reaction 2.26 to


reaction 2.22. It is generally argued that in hydrogen abstraction reactions such as

reaction 2.26, the activation energy depends on the C-H bond strength in the hydrocarbon

molecule RH, instead of the stability of the underlying radical of the peroxy radical A·.

Therefore, it is possible that reaction 2.26 may have similar capability of propagating

the chain. There are also reasons for reaction 2.26 to be slower than 2.22, such as

steric hindrance, which lead to a lower pre-exponential factor (Table 2.5) and other

characteristics related to the A· radical. Termination reactions of the antioxidant reaction

chain consists of two reactions:

AO2 · + AO2 · −−→ products (2.27)

A · + A · −−→ products (2.28)

In reaction 2.27, termination products are proposed as the major oxidative pathway

to deposit formation in this mechanism. Lastly, an additional reaction is added to take

into account that AH may oxidize in its own chain reactions:

AO2 · + AH −−→ AO2H + A · (2.29)

Reactions 2.20 - 2.29 outline the basic features of the mechanism now widely known

as the pseudo-detailed chemical kinetic mechanism. An expanded version was later pro-

posed (Table 2.5). The major additions of the model in Table 2.5 compared to the

previous 9 step model are reactions that occur after dissolved oxygen is consumed and

thermal decomposition of alkyl hydroperoxides.

The 17 reactions in the pseudo-detailed chemical kinetics mechanism include the

most important reactions to determine fuel oxidation behaviour. The most important

species are: Hydrocarbons (RH), dissolved oxygen (O2), peroxy radical inhibitors or


Table 2.5: Pseudo-detailed mechanism for liquid phase jet fuel autoxidation by Zabarnicket. al. [59].

Reaction Reactions Arrhenius A Factor Activation EnergyNumber (mol, l, s) Ea, (kcal/mol)1 I −−→ R· 1× 10−3 02 R·+ O2 −−→ RO2· 3× 109 03 RO2·+ RH −−→ ROOH + R· 3× 109 104 RO2·+ RO2· −−→ Termination 3× 109 05 RO2·+ AH −−→ ROOH + A· 3× 109 56 AO2·+ RH −−→ AO2H + A· 3× 105 107 A·+ O2 −−→ AO2· 3× 109 08 AO2·+ AH −−→ AO2H + A· 3× 109 69 AO2·+ AO2· −−→ Products 3× 109 010 R·+ R· −−→ R2 3× 109 011 ROOH −−→ RO·+ ·OH 1× 1015 4212 RO· −−→ RH −−→ ROH + R· 3× 109 1013 RO· −−→ R · prime + Carbonyl 1× 1016 1514 ·OH + RH −−→ H2O + R· 3× 109 1015 RO·+ RO· −−→ Termination 3× 109 016 R · prime + RH −−→ Alkane + R· 3× 109 1017 ROOH + SH −−→ Products 3× 109 16

Figure 2.6: Major reaction pathways of jet fuel autoxidation and the effects of SH andAH species [62].

antioxidants (AH), hydroperoxide decomposers (SH), and hydroperoxides (ROOH). RH

represents the bulk fuel and is assumed to have the chemical properties of a straight-chain


alkane. R · prime is a smaller alkyl radical created by the unimolecular reaction 13. The

major reaction pathways are shown in Figure 2.6:

• Autoxidation reactions are driven by radicals. R1 - R4 represent the creation of

radical and fuel oxidation. The cycle begins with a poorly understood process that

produces a hydrocarbon radical R·. The R· rapidly reacts with dissolved oxygen to

produce a peroxy radical RO2·. The RO2· radical can extract a hydrogen atom from

the fuel RH forming a hydroperoxide (ROOH) and regenerating the R· radicals. R3

is believed to be much faster than R1, therefore after the cycle is started, R3 is

negligible.

• The propagation of R· can be slowed by antioxidant (polar) species AH. Polar

species can occur naturally in a fuel or can be added to the fuel. AH affects

autoxidation by intercepting the peroxy radical (R5), thus slowing the hydrocarbon

radical propagation. AH is also believed to contribute to surface deposition. Overall

it is observed that AH species tend to suppress liquid phase oxidation rate and

increase deposition rate. The antioxidant reactions are represented by R5 - R9.

• Hydroperoxides (ROOH) are important species in fuel autoxidation. They are

created and destroyed during autoxidation and increase the radical pool (R9 - R16)

in the process. Kinetic analysis have shown that at 185 ◦C (365 ◦F) decomposition

of ROOH can create a radical pool that is sufficient to initiate the autoxidation

chain.

• The presence of SH species may also decrease oxidation rate. SH decomposes hy-

droperoxides through a route without producing radical species (R17), thus limiting

the radical pool size. SH includes a subset of sulfur containing species in jet fu-

els. In addition to affecting oxidation, there is also evidence that reactive sulfur

containing species promote surface deposit rate.


SMORS Oxidative Deposition Mechanisms

The SMORS mechanism originates from mechanism of oxidative deposit formation

in middle distillates such as diesel fuels. According to SMORS mechanism, subsequent

reactions following rapid oxidation formed precursor species with molecular weight in

the range of 600 - 1000 dalton (Da). These species are referred to as SMORS and are

soluble until molecular weight are increased to certain threshold. Details of the SMORS

mechanism can be found in [63–66]. The scheme will not be discussed in detail due to

the limited use of the mechanism currently in simulation studies.

Chapter 3

Experimental and Numerical Setup

3.1 Previous UTIAS Single Tube Heat Exchanger

Apparatus

A single tube heat exchanger test rig was developed at the University of Toronto Institute

of Aerospace Studies (UTIAS) for jet fuel thermal stability studies [66]. A schematic of

the rig is provided in Figure 3.1. The major components of the test rig include a dual

syringe pump (Teledyne Isco A500) and a 3 zone tube furnace (Thermcraft) capable

of providing a heated environment of up to 900 ◦C (1652 ◦F) for a 36 in. (91 cm)

long section. Additional components include a fuel storage vessel, a gas sparging vessel,

dissolved oxygen sensor and cooling coils. For details regarding to the specifications of

the rig, refer to Wong [66]. Additional studies at UTIAS by Commodo et. al. related to

jet fuel autoxidation and the previous experiential rig can be found in [67–69].

3.2 Previous Experimental Setup

The previous experimental setup was discussed in detail by Wong [66]. The experimen-

tal rig used test sections with dimensions of 36 in. (91 cm) in length and 0.069 in. (1.8

32

Chapter 3. Experimental and Numerical Setup 33

Figure 3.1: Schematic of the thermal stability test rig developed at UTIAS by Wong [66].Jet fuel can be sparged with nitrogen in the 8 L stainless steel fuel storage tank ifnecessary. A Broadley oxygen sensor measures the fuel storage tank oxygen level. Thefuel pump flow accuracy of ±0.5% of set flow rate and also an accuracy of ±0.5% forset pressure. The test sections used were Swagelok 0.125 in. (3.18 mm) outer diameter,0.069 in(1.8 mm) inner diameter 316 stainless steel tubes. Previous experiments wereconducted at temperatures ranging from 120 ◦C (248 ◦F) to 400 ◦C (752 ◦F), flow ratesof 1 ml/min, 5 ml/min and 10 ml/min, and pressures of 600 psig and 1200 psig.


mm) inner diameter. The test section is uniformly heated by the furnace. It is assumed

that all heat transfer and chemical reactions take place within the test section. The test

section design is similar to many thermal stability studies. It is a good representation

of heat exchange sections of a fuel delivery system. The current study aims to model

the deposition characteristics of fuel injection nozzles. Injection nozzles in aircraft com-

bustion chambers have much smaller diameter than typical fuel lines (in the order of 0.1

mm). Jet fuel in the nozzles have a higher flow speed than in pipelines and are exposed

to higher wetted wall temperatures. Residence time in nozzles is also shorter in com-

parison. In addition, prior to reaching the injection nozzles, jet fuel pass through heat

exchangers that expose jet fuel to various temperatures. Table 3.1 outlines the residence

time and temperature jet fuel is exposed to in relevant aircraft components. It can be

seen that prior to reaching the injection nozzle, jet fuel may be exposed to temperature

sufficiently high for autoxidation reactions to take place. Therefore, upstream conditions

must be captured in order to accurately predict deposition in fuel injection nozzles. In

order to properly reflect these unique characteristics of an injection nozzle, modifications

to the previous apparatus was necessary. A schematic of the modified design is shown in

Figure 3.2.

A preheating section was added to more accurately reflect the fuel condition prior to

entering injection nozzles. The test section was also redesigned. Other major components

of the test rig including the fuel delivery and waste fuel cooling system are identical to

the previous set up outlined in [66]. Details of all components of the new test rig can

be found in [17]. The following sections will present the design of the major components

that are relevant to numerical simulation.


Figure 3.2: Schematic of modified rig for thermal stability studies [17].


3.3 Design of New Experimental Apparatus

3.3.1 Test Section and Test Section Heating System

In order to provide an evenly distributed temperature profile around the test section, a

new test section and heating system was designed. Figure 3.3 shows the cross section

and component assembly of the test section and heating block. The major components

include a brass cylindrical block and a band heater that surrounds the brass block.

Figure 3.3: Assembly (a) and cross section (b) of the test section and heating block [17].

The 2 in. (50.8 mm) test sections that were used are commercially available stainless


Tab

le3.

1:C

omm

erci

alai

rcra

ftfu

elen

vir

omen

t,flig

ht

idle

(des

cent)

condit

ions.

Dat

aex

trac

ted

from

Chap

ter

1,T

able

2of

[6].

Fuel

Syst

emC

omp

onen

tSin

gle

Pas

sR

esid

ence

Fuel

Tem

per

ature

,P

ress

ure

,Surf

ace

Tem

per

ature

,T

ime,

s◦ F

(◦C

)psi

a◦ F

(◦C

)F

uel

tank

5×

103-5×

103

-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

firs

tst

age

pum

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)E

ngi

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

Cru

ise

serv

os1.

0-20

245-

320

(118

-160

)35

0-40

0F

uel

Bypas

sre

circ

ula

tion

-24

5-32

0(1

18-1

60)

75-1

00F

uel

Gen

erat

oroi

lco

oler

10-1

524

5-32

0(1

18-1

60)

40-1

0032

0(1

60)

Fuel

noz

zles

0.15

-0.8

300-

400

(150

-200

)40

-100

450-

500

(232

-288

)


steel tubes. Since surface material affects the rate of autoxidation and deposition, it is

important to consistently use test sections of the same material. The test sections are 2

in. (50.8 mm) long with two different outer diameters (O.D.), 1/8 in. (3.18 mm) and 1/16

in. (1.59 mm). The wall thickness of the test sections were varied in order to produce a

range of inner diameters (I.D.). The range of test section dimensions are summarized in

Table 3.2.

Table 3.2: Summary of test section dimensions.

Outer Diameter in. (mm) Inner Diameter in. (mm) Wall Thickness 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)

The test section was clamped by a cylindrical block with inner diameters matching the

outer diameter of the test sections. The brass block is 1.5 in. (38.1 mm) thick and

2 in. (50.8 mm) long. On top of the brass block four 1/8 in. (3.18 mm) holes were

drilled. Thermocouples were inserted into these locations to measure the temperatures

along the test section outer wall that was directly in contact with the heating block.

Another thermocouple was placed directly upstream of the test section to measure the

temperature of fuel entering the test section. Test section heating was provided by a 300

W band heater. The heater is tightly wrapped around the copper brass block. The heater

is capable of reaching a maximum temperature of 649 ◦C (1200 ◦F). Output temperature

is controlled by temperature controller (Omega CN7523).

3.3.2 Preheating System

Fuel preheating was provided by an oil bath (Memmert ONE 45). The oil bath was filled

with high temperature silicone bath fluid (DPDM-400). The system is able to maintain

a maximum temperature of 200 ◦C (392 ◦F). Stainless steel tubing was submerged in the


silicon fluid for jet fuel to pass through. The tubing has an outer diameter of 1/4 in.

(6.46 mm) and an inner diameter of 0.18 in. (4.57 mm). Given the set flow rate of the

system, the residence time of fuel within in preheating section can be varied by changing

the length of tubing submerged inside the fluid.

3.4 Data Acquisition/Analysis Techniques

The data extracted from the experimental setup were test section inlet/outlet temper-

atures and pressure differential across the test section. Computerized data acquisition

system was used to collect the data. For details on these components, refer to [17].

3.4.1 Pressure Drop and Deposit Formation

An in situ measurement method (pressure drop measurement) was used in the experi-

mental setup instead of the carbon burnoff method employed in the previous study by

Wong [66]. This method was used due to the difficulty of using the burn-off method on

a small test section. The pressure differential method does not provide a direct measure-

ment of the amount of deposit as does the burn-off method. The pressure difference was

used in the formula that describes the pressure loss of a laminar flow in a circular pipe

which is known as Hagen-Poiseuille law:

∆P =8µQL

πR4(3.1)

In this equation, 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 interest, ∆P is

the pressure drop across L. The amount of deposit can be translated to an equivalent

blockage radius which can be used to infer deposition amount.


3.5 Experimental Methodology

3.5.1 Experimental Parameters

Experiments are to be conducted with two varying parameters: the temperatures exposed

to the preheating section and the wetted wall temperature of the test section. Planned

experiments with various values of the fuel temperatures and wetted wall temperatures

are listed in Table 3.3 and Table 3.4. All other physical parameters were kept constant.

The back pressure (as at the position of the pressure regulator) was kept constant at 132

psig (9.10 bar). The fuel flow rate used for all experiments was 20.408 ml/min (approx.

2 lb/hr).

Table 3.3: Summary of fuel temperature experiments.

Fuel Temperature, Wetted Wall Temperature, Fuel Pressure, Flow Rate◦F (◦C) ◦F (◦C) psig (bar) pph (ml/min)

150 (65.5) 450 (232.2) 100 (6.89) 2 (20.408)200 (93.3) 450 (232.2) 100 (6.89) 2 (20.408)250 (121.1) 450 (232.2) 100 (6.89) 2 (20.408)300 (148.9) 450 (232.2) 100 (6.89) 2 (20.408)325 (162.8) 450 (232.2) 100 (6.89) 2 (20.408)

Table 3.4: Summary of wetted wall temperature experiments.

Fuel Temperature, Wetted Wall Temperature, Fuel Pressure, Flow Rate◦F (◦C) ◦F (◦C) psig (bar) pph (ml/min)

250 (121.1) 375 (135.0) 100 (6.89) 2 (20.408)250 (121.1) 300 (148.9) 100 (6.89) 2 (20.408)250 (121.1) 325 (162.8) 100 (6.89) 2 (20.408)250 (121.1) 340 (171.1) 100 (6.89) 2 (20.408)250 (121.1) 355 (179.4) 100 (6.89) 2 (20.408)250 (121.1) 370 (187.8) 100 (6.89) 2 (20.408)250 (121.1) 385 (196.1) 100 (6.89) 2 (20.408)250 (121.1) 400 (204.4) 100 (6.89) 2 (20.408)250 (121.1) 425 (218.3) 100 (6.89) 2 (20.408)250 (121.1) 450 (232.2) 100 (6.89) 2 (20.408)


3.5.2 Experimental Procedures

Each test was conducted for a total of 15 hours. All tests were split into 5 hour sessions

due to practical reasons. Complete shut-down and warm-up process was carried out for

each session.

In each session, the oil bath and test section heating block were warmed up to the

desired temperature. No fuel were passed through the system in the warm-up process

to avoid potential reactions since the warm-up process can take an extended period

of time. During warm-up, nitrogen was passed through the system to remove air and

oxygen residing in the system. Once the oil bath and test section heating block reached

the desired temperature, fuel was pumped through the system from the fuel reservoir

tank. Timed experiment started after the system reached steady state and necessary

adjustments were made to maintain the desired temperatures in the preheating and test

sections.

At the end of the timed test session, fuel flow was immediately stopped and the

system purged with nitrogen. The test sections and heating block were cooled down

quickly by compressed air to avoid possible soak back effects.

3.6 Simulation of Jet Fuel Autoxidation

The numerical simulation was designed to model the physical and chemical processes

within the experimental apparatus. The following sections will present the major aspects

of the chemical and numerical setup of the simulation.

3.6.1 Chemical Kinetics Model

The chemical kinetics model used in the numerical analysis is known as the pseudo-

detailed chemical kinetics model [70]. This formulation is in between the detailed chem-

istry and global chemistry in terms of complexity. The model has seen reasonable success


in modeling thermal oxidation of jet fuel. It should be noted that the pseudo-detailed

chemical kinetics mechanism was initially more focused on liquid phase oxidation when

proposed by Zabarnick in 1993 [59]. A few recent efforts, from [71, 72], have looked into

the modifications of the mechanism to better suit solid deposition modelling. The 18

step mechanism is outlined in Table 3.5. The mechanism is a modified version of the

mechanism described in section 2.6.3 (Table 2.5). The 16 step mechanism outlined in

section 2.6.3 focuses on liquid phase reactions and are all assumed to occur in the bulk

fuel. The mechanism in Table 3.5 includes an additional ROOH decomposition reaction

that is assumed to occur only in the bulk fuel near the surface (R18). This reaction is

assumed to be catalyzed by active metal surface, thus having a higher rate of reaction

compared to R11.

Table 3.5: Revised Pseudo-detailed chemical kinetics used for simulation. [71]

Number Reaction Number ReactionR1 I −−→ R· R10 R·+ R· −−→ R2

R2 R·+ O2 −−→ RO2· R11 ROOH −−→ RO·+ ·OHR3 RO2·+ RH −−→ ROOH + R· R12 RO·+ RH −−→ ROH + R·R4 RO2·+ RO2· −−→ termination R13 RO· −−→ R · prime + carbonylR5 RO2·+ AH −−→ ROOH + A· R14 ·OH + RH −−→ H2O + R·R6 AO2·+ RH −−→ AO2H + R· R15 RO·+ RO· −−→ terminationR7 A·+ O2 −−→ AO2· R16 R · prime + RH −−→ alkane + R·R8 AO2·+ AH −−→ AO2H + A· R17 ROOH + SH −−→ productsR9 AO2·+ AO2· −−→ products R18 ROOH −−→ RO·+ ·OH

The species are identified in section 2.6 and will not be repeated here. R1 - R17

are reactions occurring in the bulk fuel. The rate constant for each reaction is expressed

as:

k = A · exp(− Ea

RT) (3.2)

In equation 3.2, A is the pre-exponential factor, Ea is the activation energy, R is the

universal gas constant, and T is the temperature. For R18, the decomposition rate of


hydroperoxide (ROOH) at the wall is given as:

−d[ROOH]

dt= A · exp(− Ea

RT)[ROOH]α (3.3)

The α in equation 3.3 is the order of reaction. This value, in conjunction with the

activation energy of R3 was calibrated to best represent the experimental measurements

of oxidation level in [71]. Such information was not obtained from the our current ex-

perimental setup. Therefore, we were unable to carry out calibration specifically for our

experimental apparatus. The values of α, A, and Ea for R18 and Ea value for R3 used

in this simulation are borrowed from [71], where α = 0.8.

The surface deposition mechanism used in the simulations is shown in Table 3.6,

proposed by Kuprinick et. al. [62]. In this deposition mechanism, it is hypothesized

that deposit results from products formed from autoxidation reactions involving peroxy

radical inhibitors AH. This hypothesis explains the tendency of faster oxidation fuel to

form less deposits and slower oxidizing fuel to form more deposits. It also addresses the

role of polar species in the formation of deposits.

Table 3.6: Deposition reactions used in simulation. [71]

Number Reaction A (mol, L, s) Ea (kcal mol−1)R19 ProductsAH −−→ solubles 1× 109 0R20 ProductsAH −−→ insolubles 3.8× 1010 6.5R21 insolubles −−→ deposits 3× 103 16.3

Reactions R19 and R20 occur in the bulk fuel while R21 occurs at the wall. AH

reaction products from autoxidation form both soluble and insoluble precursor species.

The reactions of insoluble products at the wall ultimately lead to wall deposit. Deposit

formations reactions R19 - R21 are global reactions that simulate numerous poorly un-

derstood reactions and processes. Therefore, unlike autoxidation reactions outlined in

Table 3.5, R19 - R21 do not represent actual chemical reactions, and the parameters in

Table 3.6 have no chemical significance. The 3-step deposition mechanism in Table 3.6


allows the calibration of the location and magnitude of peak deposition. The location of

peak deposition is primarily based on the oxidation rate and wall reaction rate, while the

magnitude depend on the competition for ProductsAH. This mechanism is more realistic

than other one step models proposed as deposit formation mechanism for the pseudo-

detailed chemical mechanism. This mechanism have seen relative success in predicting

the location of maximum deposit and magnitude of maximum deposition in stable fu-

els while one step models are more accurate in predicting deposition in less stable fuels.

Since the Jet A-1 fuel used in the simulation have low polar and sulfur content, the 3 step

model outlined in Table 3.6 should be a better choice for the deposition characteristics

of the fuel and was used instead of the one step models. The one step models are listed

in Table 3.7 for reference.

Table 3.7: One step wall reactions [71]

Number Reaction Reaction Type1 ProductsAH −−→ solubles Wall2 ProductsAH −−→ insolubles Wall

3.6.2 Initial Conditions of Simulations

In the pseudo-detailed mechanism, different types of fuels are represented by different

initial concentrations of AH, SH and ROOH. For the case of Jet A-1 fuel, the initial

concentrations used in the simulation are listed in Table 3.8. All other species were

assumed to be 0 initially. The values were borrowed from [59]. The concentration of

RH was estimated from the molarity of n-dodecane liquid at room temperature. The

dissolved oxygen concentrations was estimated based on a variety of jet fuel properties.

At room temperature of 20 ◦C - 25 ◦C (68 ◦F - 77 ◦F) and 1 atm, dissolved oxygen

level was estimated to be 79 ppm by weight (1.8 × 10−3 M, mole of solute per liter of

solution, when converted to concentration). A small amount of species I was included to

initiate the chain reaction. The value of SH and AH can be varied to reflect properties


of different fuels. A typical value of AH was used in this case for Jet A-1 fuel.

Table 3.8: Initial concentration of species used in simulation of jet A-1 fuel.

Species Initial Concentration (M)I 1.0× 10−7

O2 1.8× 10−3

RH 4.4AH 7.5× 10−5

all other 0

3.6.3 Computational Fluid Dynamics

In order to investigate the effects of temperature, fluid dynamics and metal surfaces, the

flow and autoxidation chemistry throughout the apparatus need to be simulated. For sim-

plification, only the preheating section and test sections were included in the simulation

(Figure 3.4). The preheating section and the test section were modeled as simple closed

metal tubes. Joints and connections between the preheating and test sections (such as

connections to the pressure transducer and thermocouples) were neglected. Thus, the

fuel flow was assumed to proceed from preheating sections directly into test section. A

constant wall temperature was prescribed at the outer diameter of the preheating and

test sections. Simple conduction was assumed within the metal wall of the tubes and

brass block.

Figure 3.4: Major physical components modeled in the simulation.


The flow inside the preheating and test section was assumed to be axisymmetric,

therefore cylindrical coordinate system was employed. The flow was also assumed to be

laminar (Reynolds number less than 2300), this may not be true for experiments with

smaller cross sections, high flow rates and high temperatures. This will be discussed

further in later sections.

Computational Grids

Two-dimensional grids were used in the simulations in this study. Figure 3.5 shows the

grid system used in the simulations. Due to the axisymmetric nature of the flow, only

half of the tube was modeled with r = 0 being the center of the tube. In similar studies on

simulations of autoxidation reactions in metal tubes, a grid size of 5×10−3 m is typically

used [5]. Further refinement of the grid sizes showed negligible changes in accuracy in

solutions. Similar order of cell resolution was used in this study. However, instead of

using grids of the same size throughout the test section, finer grids were used at the inlet.

Such grid system was used because the most reactive part of the flow are portions with

high temperature or high temperature gradient, which occurs at the wall boundary and

inlet at the wall boundary. In sections such as further downstream and away from the

inlet and inner walls, reactions were expected to be more uniform, thus larger grid size

was used to speed up calculation time.

Figure 3.5: Grid setup of simulation.


Velocity Field

The flow was assumed to be fully developed at the entrance. For laminar flow of with no

slip boundary condition, the flow velocity is parabolic with the form:

u = umax[1− (r

rw)2] (3.4)

umax is the maximum axial velocity of the fuel. In fully developed laminar pipe flow, the

value of umax is twice the value of the bulk average velocity, whereas the average velocity

was calculated from the mass flow rate and cross sectional area. The velocity profile is

important to the autoxidation process. The fuel near the center flows at a faster speed

which results in a shorter residence time. Fuel near the boundary flows at slower speed

thus have a longer residence time. The velocity profile is expected to create a radial

concentration profile, making both radial and axial diffusion of species important in the

overall rate of reaction. The laminar tubular flow is generally a good approximation

until Reynolds number reaches 2300. For high Reynolds numbers the laminar flow may

transition to turbulent flow downstream of the inlet.

Governing Equations

Because of the assumptions on the velocity field, the only governing equations left to

solve are the energy balance equations and the species balance equations. The equations

and boundary conditions are:

ρCp(∂T

∂t+ u

∂T

∂z) =

∂K

∂z

∂T

∂z+K

∂2T

∂z2+∂K

∂r

∂T

∂r+K

1

r

∂T

∂r+K

∂2T

∂r2(3.5)

T (r, z = 0) = T0 (3.6)

∂T

∂r(r = 0, z) = 0 (3.7)


∂Ci∂t

+ u∂Ci∂z

=∂D

∂z

∂Ci∂z

+D∂2Ci∂z2

+∂D

∂r

∂Ci∂r

+D1

r

∂Ci∂r

+D∂2Ci∂r2

(3.8)

Ci(r, z = 0) = Ci,0 (3.9)

∂Ci∂r

(r = 0, z) = 0 (3.10)

∂Ci∂r

(r = rwall, z) = 0 (3.11)

Ci represents the concentration of each species in Table 3.5. There are a total of 23

species. Ci,0 and T0 are the initial species concentration and temperature, respectively,

for the fuel at the inlet of the preheating and test section. Equations 3.6 and 3.7 follows

from the axisymmetric conditions. In equations 3.5 and 3.8, K (thermal conductivity

coefficient, Wm−1K−1) and D (binary diffusion coefficient, m2/s) were assumed to be

variables. The values were interpolated from existing data according to the local condi-

tions.

In the present numerical setup, the method of lines was used to solve the partial

differential equations. The spatial derivatives with respect r and z were discretized. The

resulting ODE with respect to time was solved by Huen’s predictor-corrector method

until steady state was reached. The system of rate equations obtained from the chemical

kinetics mechanism was solved with the ode15s package included in MATLAB for each

time step. The energy equation was uncoupled from the species equations. This is due

to the fact that energy change due to chemical reactions was not considered. Therefore,

energy and species equations were solved separately.

Boundary Conditions

The boundary conditions are shown in Figure 3.5. The concentration of species at the

inlet are specified in Table 3.8. The temperature of fuel was set to room temperature of


22 ◦C (71.6 ◦F). The bulk average temperature and species concentration was calculated

at the outlet of the preheating section and used as the input to the test section. Oil bath

temperature and heater band temperature were prescribed as the outer wall temperature

of preheating tube and brass heating block respectively.

3.6.4 Simulations for the Deposition Studies

Table 3.9: Simulations for study of deposit formation in test sections

Simulation Number Pre-heating Section Test SectionTemperature, ◦C (◦F) Temperature, ◦C (◦F)

S1 140 (284) 266 (510)S2 180 (356) 266 (510)S3 220 (428) 266 (510)S4 260 (500) 266 (510)S5 260 (500) 185 (365)

Corresponds to experimental runsS6 188.5 (371.3) 266 (510)S7 112.6 (224.7) 266 (510)

At the time of writing this report, the experimental apparatus was in the final stage

of design and operational testing. Therefore, experiments outlined in Table 3.3 and

Table 3.4 were not carried out. The numerical model was used to perform simulations

that corresponds to experimental runs done with the rig [17]. Additional simulations

were design independent of the apparatus as a parametric study of preheating and test

section temperatures.

To study the effect of temperature in the preheating sections, four simulations (S1

- S4) were carried out over the range of autoxidation temperature regime from 140 ◦C

(284 ◦F) to 260 ◦C (500 ◦F). Upper limit of 260 ◦C (500 ◦F) was used so that the

temperature did not tap into the transition regime to complicate the interpretation of

results. The test section heater band temperature for all four simulations were set at 266

◦C (510.8 ◦F) which was close to the high temperature expected in a fuel nozzle. The


ideal temperature of the wetted wall was 260 ◦C (500 ◦F) at the test section. 266 ◦C

(510.8 ◦F) was used to account for the heat loss in the heat block. Simulation S5 used

test section temperature of 185 ◦C (365 ◦F) and preheating section temperature of 260

◦C (500 ◦F). S4 and S5 differs in test section temperature and will provide insight into

the effect of test section temperature. Simulations S1 - S5 serves as parametric studies.

Two additional simulations were carried in accordance with the experiments in [17].

Temperatures profiles were compared with experimental results to check for accuracy of

the simulation.

Chapter 4

Results and Discussion

The fuel used for numerical model and experiment was Jet A-1 fuel. The experimental

fuel was supplied by Shell Canada and was delivered on July 2009 [66]. The same batch

of fuel was used in previous jet fuel thermal stability research at UTIAS. The jet fuel

thermal oxidation tester (JFTOT) was used to examine the thermal stability of the fuel.

Details on the JFTOT test can be found in [6]. The test showed the particular batch of

fuel has a very low tendency to produce deposits [17].

4.1 Simulations of Temperature and Species Con-

centration in Preheating Section

4.1.1 Temperature Profiles

Temperature has a strong influence on autoxidation process, which in turn affects au-

toxidation rate. To predict the deposition rates, the simulation should be capable of

capturing the characteristics of the fluid dynamics and heat transfer of the experimental

system with reasonable accuracy. Figure 4.1 shows the measured preheating section outlet

temperature (located between the test section and preheating section) and the simulated

average bulk temperature profiles in the preheating section for preheating temperature

51

Chapter 4. Results and Discussion 52

of 188 ◦C (371 ◦F) and 112 ◦C (224 ◦F).

Figure 4.1: Comparison between simulated and measured temperature for 112 ◦C (224◦F) and 188 ◦C (371 ◦) at the outlet of the preheating section.

The simulated bulk temperature at the exit for 188 ◦C (371 ◦F) agrees well with

the measured temperature within about 2% error. The measured temperature was lower

than the simulated temperature for preheating temperature for the case of 112 ◦C (224

◦F) with a larger error of about 14%. The thermocouple was located closer to the inlet

of the test section than the outlet of the preheating section. The test section was at

a much higher temperature of 266 ◦C (510 ◦F) at the heater band. A combination of

heat loss to the air and heat transfer from the test section may account for some of the

difference. It is also possible that the simple flow model used for this simulation did


not adequately model the temperature distribution of the flow. In all simulations, the

flow was expected to be laminar since the Reynolds numbers of the simulation conditions

were well below 2300. However, some studies have shown that using turbulent solutions

for temperatures at the exit of tube test sections produced more accurate results than

laminar solutions [73]. Due to the limited experimental results at this time, we were not

able to determine the source of the error.

The temperature profiles along the length of the preheating section for various

preheating (oil bath) temperatures, S1-S4, are shown in Figure 4.2. The temperature

change throughout the preheating section followed similar trend for each case. At the

wall (r), the fuel rapidly increased in temperature upon entering the inlet. After reaching

about 65% of the prescribed outer wall temperature (oil bath temperature), the rates

of increase became gradual but continued to increase along the rest of the preheating

section. Temperatures at half radius (1/2 r) and centerline (r = 0) also increased along

the length of the preheating section at a relatively higher rate in the first third of the

preheating section. The centerline temperatures were almost identical to the half radius

temperatures. The bulk average temperatures were much closer to the centerline and

half radius temperatures for each case. The centerline temperatures and half radius

temperatures are significantly lower than the fluid temperature at the wall. Thus, the

thin layer of slower flowing fluid near the wall was expected to be significantly more

reactive than the bulk fuel. However, this thin layer of the fluid adjacent to the wall is

only a small portion of the fuel flowing through the system.

4.1.2 Oxygen Profile

The consumption of oxygen is central to the autoxidation and deposition process. The

concentrations of oxygen along the length of the preheating section for simulations S1

- S4 are shown in Figure 4.3. The percentages of initial oxygen level remaining at the

outlet were 31.8%, 47.2% and 96.1% for preheating temperatures of 260 ◦C (500 ◦F),


220 ◦C (428 ◦F) and 180 ◦C (356 ◦F) respectively. Negligible amount of oxygen was

consumed in the 140 ◦C (284 ◦F) simulation. The low consumption rates at 140 ◦C (284

◦F) and 180 ◦C (356 ◦F) were expected since these temperatures are at the lower end

of the autoxidation temperature regime and the modeled Jet A-1 fuel is considered a

relatively stable fuel.

The oxygen concentration profiles along the axis show that there were generally a

slow initial oxidation rate followed by a rapid oxidation rate. This is most evident in

the 260 ◦C (500 ◦F) (Figure 4.3a) and 220◦C (428 ◦F) (Figure 4.3b) cases where oxygen

was consumed completely at certain locations. It can be seen by comparing Figure 4.2

and Figure 4.4 that oxygen consumption rate does not directly correlate to temperature.

This is because the rate of autoxidation reactions is driven by the size of the radical

pool instead of solely by temperature. Rapid oxidation occurs once a substantial radical

pools size was built up. A separate process in addition to the autoxidation reactions is

needed for initial radical pool production before enough radicals are produced through

propagation reactions to sustain the chain reaction. This process overall represented by

R1 in Table 3.5 is poorly understood [59]. Temperature seems to have the most significant

impact on the radical building process. The accelerated reaction at the wall for the 220

◦C (428 ◦F) simulation is significantly delayed compared to the same location for the 260

◦C (500 ◦F) simulation. However, once the accelerated reactions initiated, the oxygen

consumption rate for both cases were roughly the same, both occurred within about 0.4

m of the preheating section. The species AH may also delay the accelerated reaction by

reducing the size of the radical pool. However, the details of the combined effect of AH

concentration and temperature were not explored in this study.

One interesting observation to note is that despite the rapid oxygen consumption

occurring at different axial location for different radial positions, the bulk average oxygen

concentrations appeared to decrease linearly once rapid oxidation took place at a certain

location. This may be useful for future correlation and calibration since oxygen is one of


(a) 260 ◦C (500◦F) oil bath temperature, simu-lation S4

(b) 220 ◦C (428◦F) oil bath temperature, simulation S3

(c) 180 ◦C (356◦F) oil bath temperature, simu-lation S2

(d) 140 ◦C (284◦F) oil bath temperature, simulation S1

Figure 4.2: Temperature profile along the length of preheating section at three differentradial locations (at r = 0, 1/2 r and r) for various preheating temperatures.


the species that can be measured experimentally.

4.1.3 Hydroperoxide (ROOH)

ROOH is another important species in the study of autoxidation of jet fuel. Figure 4.4

shows the simulated ROOH concentrations under various preheating temperatures (S1-

S4) similar to O2 in the previous section. The concentration and production rate of

ROOH is closely coupled with oxygen consumption rate. After an initial induction pe-

riod almost identical to that of oxygen, concentration of ROOH rose sharply in cases

where noticeable oxygen was consumed (S1 and S2). The concentration continued rising

until oxygen was completely consumed. Complete oxygen consumption occurred only for

260 ◦C (500 ◦F) simulation at the tube wall (Figure 4.4a) with a maximum ROOH con-

centration of 1.08×10−3 mol/m3. In this case, ROOH concentration gradually decreased

after the point of complete oxygen consumption due to ROOH dissociation. Dissociation

of ROOH was faster at the wall compared to the bulk fluid due to the catalytic effect

of the metal surface. The dissociation rate at the wall was the combined rates of R11

and R18. Since R18 is significantly faster, the rate of ROOH consumption closer to the

axis (r = 0) was expected to be much slower than close to the wall. This effect was

evident from Figure 4.4a where at the outlet of the preheating section, the concentration

of ROOH at 1/2r had already surpassed the peak of ROOH concentration at the wall. It

is noteworthy to point out that the dissociation of ROOH results in an increase in radical

pool according to R11 and R18 in Table 3.5. However, results from the 260 ◦C (500 ◦F)

simulation indicate that the net effect was faster rate of ROOH dissociation.


4.1.4 Velocity Profile and Radial Diffusion Effects

The characteristics of fluid dynamics in the system is essential to the understanding

the deposition process. Both fluid dynamics and heat transfer significantly influence

the autoxidation and deposition process. However, the influence of fluid dynamics is

poorly understood due to limited systematic studies. Comparisons between the studies

available is also difficult since the physical and chemical properties of fuels and the

specific operational parameters of experimental setups can vary drastically [60, 74, 75].

In general, it is observed that correlation between deposition rate and fuel velocity is

weak for velocities above 0.3 m/s and strong for lower velocities [76]. Figures 4.5a and

Figure 4.5b show the Reynolds number and average velocity in the preheating section

for simulations S1 and S4. The velocity falls well below the suggested 0.3 m/s. The

highest Reynolds number reached for the 260 ◦C (500 ◦F) case at the exit was below 300.

Therefore, all simulations were within the laminar flow regime in the preheating section.

Under these conditions, fuel flow rate, which in turn determines the velocity of the fuel

flow, is expected to have substantial influence on the deposition process.

Due to the laminar nature of the flow, the local residence time of fluid increased

radially from the centerline (r=0) to the wall (r). The maximum speed of the flow, which

was two times the average bulk flow speed, occurred at the centerline, r = 0. The speed

decreased to 0 at the wall surface. Therefore, in addition to being closer to the high

temperature wall surface, the fluid near the wall also have a higher residence time. The

combination of these factors result in a higher temperature thus more reactive layer of

fluid near the wall surface than the centerline.

The overall liquid phase oxidation and deposition rate depend on the relative impor-

tance of physical and chemical factors. In a laminar flow system where speed, temperature

and concentration have a set and stable profile, the influence of physical factors is much

more significant than a system with increased mixing effects such as a turbulent system.

For a less reactive fuel, as a result of either low temperature or chemical make-up, species


(a) 260 ◦C (500 ◦F) oil bath temperature, simula-tion S4

(b) 220 ◦C (428 ◦F) oil bath temperature, simulation S3

(c) 180 ◦C (356 ◦F) oil bath temperature, simulationS2

(d) 140 ◦C (284 ◦F) oil bath temperature, simulation S1

Figure 4.3: Oxygen concentration profile along the length of preheating section at threedifferent radial locations (at r = 0, 1/2 r and r) for various preheating temperatures.


(a) 260 ◦C (500 ◦F) oil bath temperature, simula-tion S4

(b) 220 ◦C (428 ◦F) oil bath temperature, simulation S3

(c) 180 ◦C (356 ◦F) oil bath temperature, simulationS2

(d) 140 ◦C (284 ◦F) oil bath temperature, simulation S1

Figure 4.4: ROOH concentration profile along the length of preheating section at threedifferent radial locations (at r = 0, 1/2 r and r) for various preheating temperatures.


(a) Reynolds Number (b) Velocity Profile

Figure 4.5: Reynolds number and average velocity profile for 140 ◦C (284 ◦F) and 260◦C (500 ◦F), S1 and S4 respectively

transportation via diffusion or secondary motion of the flow has relatively minor influence

on the overall reaction rate. On the other hand, in a more reactive fuel, faster reactions

may be limited by species transportation. Secondary flow motions resulting from sources

such as buoyancy and gravity were not included in this simulation. The radial diffusion

of species were included in the species conservation equations described in Section 3.5.2.

It is clear that the characteristic of the flow is crucial to the determination of species

distribution. Accurate modeling of the flow is also becoming increasingly important

with the introduction of new chemical mechanisms that include separate wall and bulk

fluid reactions. In the previous sections, it is established that, by examining oxygen

and ROOH concentrations within the preheating section, the local reactivity of fuel can

be significantly different due to physical nature of the flow. An aircraft fuel system

or an experimental apparatus may contain sections with various flow characteristics.

Identifying whether the flow is laminar, turbulent or transient and applying appropriate

flow models is crucial to the accuracy of the oxidation and deposition simulation.


4.2 Simulations of Deposition in Test Section

Deposition rate at the wall was simulated in the test section with an outer wall tempera-

ture of 266 ◦C (510 ◦F) for a variety of preheating temperatures (S1-S4). The temperature

of 266 ◦C (510 ◦F) was used so that the wetted-wall temperature was close to the desired

260 ◦C (500 ◦F) when heat loss was taken into account. Detailed simulation conditions

are listed in Section 3.5.5. The wall deposition rate is shown in Figure 4.6. The depo-

sition rate is calculated from the rate constants of chemical reactions and concentration

of species of the cell adjacent to the wall. All deposit produced via R21 (Table 3.6) was

assumed to remain on the tube wall. Studies have shown that as the solid deposits start

to build up on stainless steal tubes, the reaction rate at the wall decreases [58]. In this

study the change in rate due to deposit accumulation was not taken into account. The

calculated rate was assumed to be the steady state deposition rate at the wall surface.

In the pseudo-detailed mechanism, deposits are proposed to be originated from AH

and SH species. The concentration of SH used to model the Jet A-1 fuel was zero in this

study, thus all deposits were formed from products of reactions involving AH. Figure 4.6

shows the deposition rates in the test section were closely dependent on the upstream

preheating section temperature. Deposition rate with preheating temperature of 260 ◦C

(500 ◦F) was roughly twice that of 220 ◦C (428 ◦F) preheating temperature. There was

roughly an order of magnitude drop in deposition rate from 220 ◦C (428 ◦F) to 180 ◦C

(356 ◦F) and from 180 ◦C (356 ◦F) to 140 ◦C (284 ◦F) of preheating temperature. The

deposition rate increased from the inlet to the outlet of the test section for all cases.

Fuel entered the test section at a much higher bulk temperature than room tem-

perature due to the preheating section. Simulations of temperature profiles for S1-S4

in the test section showed that the temperature of fluid at the wall in the test sections

reached a near constant temperature shortly after the inlet. On the other hand, cen-

treline temperature stayed nearly constant and the bulk average temperature increased

moderately. The moderate temperature increase was due to the low residence time of fuel


Figure 4.6: Deposition rate at test section outer wall temperature of 266 ◦C (510 ◦F)with various preheating temperature.

in the test section. The test sections has significantly smaller cross sections compared to

the preheating sections. As a result, the bulk residence time of fuel in the 2 in. (50.8

mm) test section is expected to be only 0.055 seconds.

In numerous flowing rig experiments a peak deposition rate is observed. Such peak

usually occurrs near the location of maximum oxidation rate where dissolved O2 was

depleted [58,62]. A peak in deposition rate was not observed in all test cases in this study.

This was due to the relative short test section compared to those commonly used in other

jet fuel autoxidation studies. In addition, complete oxygen consumption did not occur in

the preheating section for S1-S4. The remaining oxygen from the preheating section is

expected to continue oxidizing throughout the test section to produce deposit precursor


species. The deposits formed from the newly created precursor species can result in

increased deposit formation rate while temperature at the test section wall remained

constant. Results from Figure 4.6 shows that while the test section temperature was set

to be the same for S1-S4, the extent of reaction that took place in the preheating section

is significant to the test sections deposition formation rate.

It is also instructive to investigate the effect of temperature in the test section.

Deposition rates in the test section with heating temperatures of 266 ◦C (510 ◦F) and 185

◦C (365 ◦F) corresponding to simulations S4 and S5 are shown in Figure 4.7. Preheating

temperature for both cases were 260◦C (500 ◦F). The simulated average test section

inlet temperatures for both cases were 215 ◦C (419 ◦F). With test section heated at

266 ◦C (510 ◦F) (S4), fuel continued to increase in temperature along the test section,

with bulk average temperature of 225.6 ◦C (438 ◦F) at the outlet. In the case of test

section being heated at 185 ◦C (365 ◦F) (S5), the bulk average jet fuel temperature

decreased from 215 ◦ (419 ◦F) at the inlet to 191 ◦C (375.8 ◦F) at the outlet of the test

section. As expected, the rate decreased as temperature decreased. From the results

in Figure 4.6 and Figure 4.7 it can be concluded that local deposition rate depends

on both local temperature and upstream temperature. Higher temperature in the test

preheating section and test section both lead to higher deposition rate. Therefore, in

order to determine the deposition rate at a specific location, a complete knowledge of the

upstream and local conditions is necessary.

4.3 Comparison with Experimental Results

Two experiments (S6 and S7) were carried out using the rig developed for this project. In

the S6 and S7 experimental runs, increase in pressure drop at the end of the experiment

was reported to be 0.013 psid (pounds per square inch difference)(91 Pa) and 0.003 psid

(21 Pa) respectively. The rates of deposit formation in the test section simulated for


Figure 4.7: Deposition rate at test section with test section heater temperatures of 266◦

and 185◦, preheating temperatures for both cases are 266 ◦C (510 ◦F).

the two cases are presented in Figure 4.8. The experimental results show that there was

about an order of magnitude drop in pressure increase from 188 ◦C (371 ◦F) to 112 ◦C

(224 ◦F). Computed results depicted in Figure 4.8 shows that the drop in deposition rate

to be much more substantial. The average deposition rates were 4.12× 10−5 mol/s ·m3

for simulation S6 and 1.01× 10−6 mol/s ·m3 for simulation S7. It should be noted that

the pressure transducer has an accuracy of ±0.008 psid. Therefore, since the measured

drop of 0.003 pisd (21 Pa) for the experiment S7 was close to the error range, it cannot be

concluded with certainty whether deposit actually formed inside the test section. In order

to directly compare simulation result and experimental result, blockage from experiment

and rate from simulation need to be converted to actual deposit mass over a period time.

This will pose more challenges since information on physical properties, such as density,

is needed. These information are not readily available due to the limited research on


deposit morphology. Nonetheless, in order to make more meaningful comparisons, more

experimental results are needed.

Figure 4.8: Deposition rate at test section corresponding to experimental runs S6 andS7. Preheating temperatures are 188 ◦C (371 ◦F) (S6) and 112 ◦C (224 ◦F) (S7), testsection temperature are 266 ◦C.

Chapter 5

Conclusions and Recommendations

The autoxidation and deposition processes of air saturated Jet A-1 were simulated in this

thesis work. The simulation modeled the experimental setup consisting of a preheating

section and a test section. A pseudo-detailed chemical kinetics mechanism was used

to model the underlying chemistry. The temperature and species concentrations along

the radial and axial direction were calculated. A fuel flow rate of 20.408 ml/min and

a pressure of 100 psi were used for the inlet of the preheating section. A range of

temperature within the autoxidation regime of fuel degeneration was used to simulate

the autoxidation and deposition of fuel. The dissolved oxygen level of air saturated fuel

was assumed to be 79 ppm by weight. The influence of temperature in the preheating

section and test section on deposition in the test section was investigated. The following

points are the main conclusions of this study:

1. The extent of autoxidation reaction in the preheating section was significant to the

deposition rate in the test section. Higher temperatures in the preheating section

increased the deposition rate in the test section. At lower temperatures, where

little chemical reactions took place in the preheating section, deposition rates were

significantly lower.

2. Oxygen consumption was a key measure of extent of oxidation in the preheating

66

Chapter 5. Conclusions and Recommendations 67

section. Once enough radicals were built up by the initiation reactions, oxygen

was consumed rapidly in the preheating section. Once accelerated reactions were

initiated, oxygen was depleted within a relatively small portion of the preheat-

ing section (about 0.4 m). Temperature ultimately determined if and where the

accelerated reactions occurred within the preheating section.

3. Velocity and temperature profiles can significantly influence the chemical reactions.

Due to the laminar nature of the flow, a radial velocity profile existed within the

preheating and test sections. The difference in local residence time resulted in radial

temperature profiles. Since temperature ultimately determines the reactions and

the rates of which those take place, local concentration of species varied throughout

the preheating and test section. Propagation of species in the radial direction was

observed to initiate, increase or limit the rate of reactions.

4. Deposition rate in the test section was dependent on both upstream and local

conditions. Higher temperatures in the test section increased the deposition rate.

However, the exact rate was also dependent on the upstream condition of the flow.

To determine the deposition rate at a certain location, the entire history of the

upstream conditions or exact composition at the location of interest are necessary.

5.1 Suggestion for Future Work

There are several recommendations for future development of the project. The recom-

mendations are summarized below:

1. Verification with experimental results. The experimental setup used in this study

was not a standard setup used by many other fuel autoxidation studies. Therefore,

no data were available to compare the results from simulations and experiments.

Calibration of the chemical kinetics model is essential. To do so, it is necessary

to get extensive experimental results from the rig in which this numerical study is

Chapter 5. Conclusions and Recommendations 68

based on.

2. Better representation of flow fields. The parabolic velocity field used in this study

to model the flow field is a very simplified representation of a laminar flow. In the

presence of steep temperature gradients, factors such as buoyancy maybe significant

in the flow. Due to the significance of species transport, a better representation of

the fluid dynamics of the system will benefit the accuracy of the simulation.

3. Simulation of complete consumption of oxygen within the preheating section. Oxy-

gen content may significantly alter the deposition characteristics. In this study,

the operating conditions simulated did not allow complete consumption of oxygen

within the preheating section. Complete oxygen consumption may very well occur

in a realistic fuel system. Therefore, investigating the deposition behavior at lower

oxygen levels will benefit simulation accuracy. In future studies, lower initial oxy-

gen level, lower flow-rate or smaller diameter of preheating sections can be used to

ensure complete oxygen consumption.

4. Simulation of other types of fuels. The Jet A-1 fuel used in this study is a relatively

stable fuel. The fuel contained a very low concentration of sulfur species, thus the

SH concentration was not taken into account in this study. The AH concentration is

also relatively low for Jet A-1 fuel. The interaction of AH and SH with other species

in the fuel is the key to determining the characteristics and rate of reactions in the

mechanism. Therefore, the influence of different levels of SH and AH concentration

on deposition rate should be investigated, which can be achieved by using different

types of fuel.

Bibliography

[1] H. Streifinger, “Fuel/oil system thermal management in aircraft turbine engines,” in

RTO AVT Symposium on Design Principles and Methods for Aircraft Gas Turbine

Engines, (Neuilly-sur-Seine, France), pp. 11–15, NATO Research and Technology

Organization, May 1998.

[2] M. Ahlers, “Aircraft thermal management,” in Encyclopedia of Aerospace Engineer-

ing, John Wiley & Sons, Ltd., 2011.

[3] L. J. Spadaccini and H. Huang, “On-Line Fuel Deoxygenation for Coke Suppression,”

Journal of Engineering for Gas Turbines and Power, vol. 125, pp. 686–692, July

2003.

[4] L. J. Spadaccini and H. Huang, “Fuel-cooled thermal management for advanced

aeroengines,” Journal of Engineering for Gas Turbines and Power, vol. 126, pp. 284–

293, April 2004.

[5] S. Butnark, Thermally Stable Coal-Based Jet Fuel: Chemical Composition, Thermal

Stability, Physical Properties and Their Relationships. PhD thesis, Pennsylvania

State University, December 2003.

[6] R. N. Hazlett, Thermal Oxidation Stability of Aviation Turbine Fuels. 1916 Race

Street, Philadelphia, PA, U.S.A. 19103: ASTM International, October 1991.

[7] M. A. Roan, The Effect of Dissolved Oxygen on the Pyrolytic Degradation of Jet

Fuels. PhD thesis, Pennsylvania State University Department of Energy and Geo-

Environmental Engineering, May 2003.

69

Bibliography 70

[8] J. M. Andresen, J. J. Strohm, L. Sun, and C. Song, “Relationship between the For-

mation of Aromatic Compounds and Solid Deposition during Thermal Degradation

of Jet Fuels in the Pyrolytic Regime,” Energy & Fuels, vol. 15, no. 3, pp. 714–723,

2001.

[9] J. Ervin, C. Heneghan, and T. Williams, “Surface Effects on Deposits From Jet

Fuels,” Journal of Engineering for Gas Turbines and Power, vol. 118, pp. 278–285,

April 1996.

[10] J. Watt, A. Evans Jr., and R. Hibbard, “Fouling Characteristics of ASTM Jet A

Fuel When Heated to 700◦F in a Simulated heat Exchanger Tube,” Tech. Rep. TN

D-4958, NASA, Washington, District of Columbia, U.S.A., December 1968.

[11] D. Li, W. Fang, Y. Xing, Y. Guo, and R. Lin, “Spectroscopic Studies on Thermal-

Oxidative Stability of Hydrocarbon Fuels,” Fuel, vol. 87, pp. 3286–3291, November

2008.

[12] G. Hemighaus, T. Boval, J. Bacha, F. Barns, M. Franlin, L. Gibbs, N. Hogue,

J. Jones, D. Lesnini, J. Lind, and J. Morris, “Aviation Fuels Technical Review

(FTR-3),” Tech. Rep., Chevron Corporation, 2006.

[13] W. Gerstler and R. Bunker, “Aircraft engine thermal management: The impact of

aviation electric power demands.” Mechanical Engineering Magazine Online, De-

cember 2008.

[14] S. Heneghan and W. Harrison III, “JP-8+100 Jet Fuel Thermal Stability Addi-

tive,” Journal of Energy Resources Technology, Transactions of the ASME, vol. 118,

pp. 170–179, September 1997.

[15] L. Balster, E. Corporan, M. DeWitt, J. Edwards, J. Ervin, J. Graham, S. Lee, S. Pal,

D. Phelps, L. Rudnick, R. Santoro, H. Schobert, L. Shafer, R. Striebich, Z. West,

G. Wilson, R. Woodward, and S. Zabarnick, “Development of an Advanced, Ther-

mally Stable, Coal-Based Jet Fuel,” Fuel Processing Technology, vol. 89, pp. 364–378,

April 2008.

Bibliography 71

[16] T. Edwards, “Cracking and Deposition Behavior of Supercritical Hydrocarbon Avi-

ation Fuels,” Combustion Science and Technology, vol. 178, pp. 307–334, 2006.

[17] J. Liang, “Design and Development of an Experimental Apparatus to Study Jet Fuel

Coking in Small Gas Turbine Fuel Nozzles,” Master’s thesis, University of Toronto

Institute for Aerospace Studies, Toronto, ON, Canada, 2013.

[18] B. D. Beaver, C. B. Clifford, M. G. Fedak, G. Li, P. S. Iyer, and M. Sobkowiak,

“High Heat Sink Jet Fuels. Part 1. Development of Potential Oxidative and Pyrolytic

Additives for JP-8,” Energy & Fuels, vol. 20, pp. 1639–1646, 2006.

[19] W. Balster and E. Jones, “Effects of Temperature on Formation of Insolubles in

Aviation Fuels,” Journal of Engineering for Gas Turbines and Power, vol. 120,

pp. 289–2571, April 1998.

[20] S. Darrah, “Jet Fuel Deoxygenation,” Tech. Rep.2081, Air Force Wright Aeronauti-

cal Laboratories (AFWAL), October 1988.

[21] S. Eser, “Mesophase and Pyrolytic Carbon Formation in Aircraft Fuel,” Carbon,

vol. 34, no. 4, pp. 539–547, 1995.

[22] A. Watkinson and D. Wilson, “Chemical Reaction Fouling: A Review,” Experimental

Thermal and Fluid Science, vol. 14, pp. 361–374, May 1997.

[23] V. Katta and E. Jones, “Modeling of Deposition Process in Liquid Fuels,” Combus-

tion Science and Technology, vol. 139, pp. 75–111, 1998.

[24] G. Jones and L. Balster, “Effects of Temperature on Formation of Insolubles in

Aviation Fuels,” Journal of Engineering for Gas Turbines and Power, vol. 120,

pp. 298–293, April 1998.

[25] E. Jones, “Autoxidation of Aviation Fuels in Heated Tubes: Surface Effects,” Energy

& Fuels, vol. 10, pp. 831–836, May 1996.

[26] J. Ervin, S. Zabarnick, and T. Williams, “Global Kinetic Modeling of Aviation

Fuel Fouling in Cooled Regions in a Flowing System,” Industrial & Engineering

Chemistry Research, vol. 35, pp. 4028–4036, 1996.

Bibliography 72

[27] L. J. Spadaccini, D. R. Sobel, and H. Huang, “Deposit Formation and Mitigation

in Aircraft Fuels,” Journal of Engineering for Gas Turbines and Power, vol. 123,

pp. 741–746, October 2001.

[28] P. Marteney and L. Spadaccini, “Thermal decomposition of aircraft fuel,” Journal

of Engineering for Gas Turbines and Power, vol. 108, pp. 648–653, 1986.

[29] W. Taylor, “Deposit Formation from Deoxygenated Hydrocarbons. I. General Fea-

tures,” Industrial and Engineering Chemistry Product Research and Development,

vol. 13, pp. 133–138, 1974.

[30] T. Edwards and S. Zabarnick, “Liquid Fuels and Propellants for Aerospace Propul-

sion: 1903 - 2003,” Industrial and Engineering Chemistry Research, vol. 32, no. 12,

pp. 3117–3122, 1993.

[31] M. Roan and L. Andre, “The effect of fuel composition and dissolved oxygen on de-

posit formation from potential jp-900 basestocks,” Energy & Fuels, vol. 18, pp. 835–

843, March 2004.

[32] S. Zabarnick, “Studies of Jet Fuel Additives Using the Quartz Crystal Microbalance

and Pressure Monitoring at 140 ◦C,” Industrial and Engineering Chemistry Research,

vol. 33, pp. 2771–2777, Nov 1995.

[33] S. Heneghan, C. Martel, T. Williams, and D. Ballal, “Effect of Oxygen and Additives

on the Thermal Stability of Jet Fuels,” Journal of Engineering for Gas Turbines and

Power, vol. 117, pp. 120–124, Jan 1995.

[34] A. Nixon, “Vaporizing and Endothermic Fuels for Advanced Engine Application,”

Quarterly Report No.11, Shell Development Co. on contract with Aero Propulsion

Lab, Wright-Patterson Air Force Base, OH, 1966.

[35] W. Taylor, “Deposit Formation from Deoxygenated hydrocarbons. II. Effect of Trace

Sulfur Compounds,” Industrial and Engineering Chemistry Product Research and

Development, vol. 15, no. 1, pp. 64–68, 1976.

[36] W. Taylor, “Development of High Stability Fuel,” Final Report, Esso R&E Co. on

Bibliography 73

contract with Naval Air Propulsion Test Center, Washington, April 1972.

[37] B. Beaver, R. DeMunshi, S. Henegan, S. Whitare, and P. Neta, “Model Studies Di-

rected at the Development of New Thermal Oxidative Stability Enhancing Additives

for Future Jet Fuels,” Energy & Fuels, vol. 11, no. 2, pp. 396–401, 1997.

[38] B. Beaver, L. Gao, G. Fedak, M, M. Coleman, and M. Sobkowiak, “Model Studies

Examining the Use of Dicyclobexylphenylphosphine to Enhance the Oxidative and

Thermal Stability of Future Jet Fuels,” Energy & Fuels, vol. 16, no. 5, pp. 1134–1140,

2002.

[39] R. Striebich and W. Rubey, “Analytical Method for the Detection of Dissolved

Oxygen,” American Chemical Society, Division of Petroleum Chemistry, Preprints,

vol. 39, pp. 47–56, Feb 1994.

[40] G. Jones and W. Balster, “Surface Fouling in Aviation Fuel: Short- vs Long-term

Isothermal Tests,” Energy & Fuels, vol. 9, pp. 610–615, July-Aug 1995.

[41] T. Edwards, “Surrogate Mixtures to Represent Complex Aviation and Rocket Fu-

els,” Journal of Propulsion and Power, vol. 17, pp. 461–466, March/April 2001.

[42] W. Taylor, “Kinetics of Deposit Formation from Hydrocarbons. Fuel Composition

Studies,” Industrial and Engineering Chemistry, Product Research and Development,

vol. 8, pp. 375–380, 1969.

[43] G. Russell, “Oxidation of Unsaturated Compounds, III. Products of the Reaction of

Indene and Oxygen; Stereochemistry of the Addition of Peroxy Radical and Oxygen

to a Double Bond,” Journal of the American Chemical Society, vol. 8, pp. 1035–1040,

1956.

[44] R. Hazlett, Frontiers of Free Radical Chemistry, vol. 78, ch. Free Radical Reactions

Related to Fuel Research, p. 1035. 1980.

[45] E. Bushueva and I. Bespolov, “Effect of Jet Fuel Hydrocarbon Composition on

Thermal Stability,” Chemistry and Technology of Fuels and Oils, vol. 7, pp. 699–

720, 1971.

Bibliography 74

[46] F. Datschefski, “Role of the JFTOT in Aviation Fuel Stability Research,” Fuel

Science and Technology International, vol. 6, pp. 609–631, 1988.

[47] L. M. Balster, S. Zabarnick, R. C. Striebich, L. M. Shafer, and Z. J. West, “Analysis

of Polar Species in Jet Fuel and Determination of Their Role in Autoxidative Deposit

Formation,” Energy and Fuel, vol. 20, pp. 2564–2571, August 2006.

[48] G. Jones and L. Balster, “Autoxidation of Dilute Jet-Fuel Blends,” Energy & Fuels,

vol. 13, no. 4, pp. 769–802, 1999.

[49] G. Jones and L. Balster, “Interaction of a Synthetic Hindered-Phenol with Natural

Fuel Antioxidants in the Autoxidation of Paraffins,” Energy & Fuels, vol. 14, no. 14,

pp. 640–645, 2000.

[50] Z. Savaya, A. Mohammed, and K. Abbas, “The effect of sulphur compounds on

deposit formation in hydrotreated kerosene,” Fuel, vol. 67, pp. 673–677, May 1988.

[51] D. Kendal, D. Clark, and P. Stevenson, “The influence of polar compounds on the

stability of jet fuel,” in Proceedings of 2nd International Conference on Long-Term

Storage Stabilities of Liquid Fuels, (San Antonion TX), pp. 694–703, October 1986.

[52] D. Clark and L. Smith, “Further studies of the effects of polar compounds on the

thermal stability of jet fuel,” in Proceedings of 3nd International Conference on

Long-Term Storage Stabilities of Liquid Fuels, (Institute of Petroleum, London, Eng-

land), pp. 268–275, November 1988.

[53] S. P. Heneghan and S. Zabarnick, “Oxidation of jet fuels and the formation of

deposits,” Fuel, vol. 73, no. 1, pp. 35–43, 1994.

[54] D. Hardy, E. Beal, and J. Burnett, “The influence of polar compounds on the sta-

bility of jet fuel,” in Proceedings of the 5th International Conference on Stability

Handling Liquid Fuels (N. Giles, ed.), (Department of Energy: Washington, DC),

pp. 260–271, 1992.

[55] W. Taylor and J. Frankenfeld, “Chemistry and mechanism of distillate fuel stability,”

in Proceedings of 2nd International Conference on Long-Term Storage Stabilities

Bibliography 75

of Liquid Fuels (N. Giles, ed.), (Southwest Research Institute, San Antonio, TX),

pp. 260–271, 1986.

[56] O. Altin and S. Eser, “Analysis of Carboneceous Deposits from Thermal Stressing

of a JP-8 Fuel on Superalloy Foils in a Flow Reactor,” Industrial & Engineering

Chemistry Research, vol. 40, pp. 596–603, December 2001.

[57] E. Jones, W. Balster, and J. Pickard, “Surface Fouling in Aviation Fuels: An Isother-

mal Chemical Study,” Journal of Engineering for Gas Turbine Power, vol. 118, no. 2,

pp. 286–291, 1996.

[58] J. Ervin, T. Ward, T. Williams, and J. Bento, “Surface Deposition within Treated

and Untreated Stainless Steel Tube Resulting from Thermal-Oxidative and Pyrolytic

Degradation of Jet Fuel,” Energy & Fuels, vol. 17, pp. 577–586, March 2003.

[59] S. Zabarnick, “Chemical Modelling of Jet Fuel Autoxidation and Antioxidant Chem-

istry,” Industrial & Engineering Chemistry Research, vol. 32, pp. 1012–1017, 1993.

[60] A. Giovanetti, “Long Term Deposit Formation in Aviation Turbine Fuel at Elevated

Temperature,” Tech. Rep. NASA Report CR-19579, United Technologies Research

Center on contract with National Aeronautics and Space Administration, Cleveland,

OH, April 1985.

[61] V. Katta and W. M. Roquemore, “Numerical Method for Simulating Fluid-Dynamic

and Heat Transfer Changes in Jet Engine Injector Feed-Arm Due to Fouling,” Jour-

nal of Thermophysics and Heat Transfer, vol. 7, no. 1, pp. 651–660, 1993.

[62] N. J. Kuprowicz, S. Zabarnick, Z. J. West, and J. S. Ervin, “Use of Measured

Species Class Concentrations with Chemical Kinetic Modeling for the Prediction of

Autoxidation and Deposition of Jet Fuels,” Energy & Fuels, vol. 21, pp. 530–544,

2007.

[63] B. Beaver, C. Gao, and M. Sobkowiak, “On the Mechanism of Formation of Thermal

Oxidative Deposits in Jet Fuels. Are Unified Mechanism Possible for Both Storage

and Thermal Oxidative Deposit Formation for Middle Distillate Fuels,” Energy &

Bibliography 76

Fuels, vol. 19, pp. 1574–1579, 2005.

[64] B. Beaver, C. Gao, and M. Sobkowiak, “Insight into the Mechanisms of Middle Dis-

tillate Fuel Oxidative Degradation. Part 1: On the Role of Phenol, Indole, and Car-

bazole Derivatives in the Thermal Oxidative Stability of Fischer-Tropsch/Petroleum

Jet Fuel Blends,” Energy & Fuels, vol. 23, pp. 2041–2046, 2009.

[65] P. Aksoy, G. O., D. Centiner, D. Fonseca, M. Sobkowiak, S. Falcone-Miller, B. Miller,

and B. Beaver, “Insight into the Mechanisms of Middle Distillate Fuel Oxidative

Degradation. Part 2: On the Relationship between Jet Fuel Thermal Oxidative

Deposit, Soluble Macromolecular Oxidatively Reactive Species, and Smoke Point,”

Energy & Fuels, vol. 23, pp. 2047–2051, 2009.

[66] O. Wong, “Design and Development of and Apparatus to Study Aviation Jet Fuel

Thermal Stability.,” Master’s thesis, University of Toronto Institute for Aerospace

Studies, Toronto, ON, Canada, 2010.

[67] M. Commodo, I. Fabris, O. Wong, C.P.T. Groth, and O.L. Gulder, “Three-

dimensional fluorescence spectra of thermally stressed commercial Jet A-1 aviation

fuel in the autoxidative regime,” Energy & Fuels, vol. 26, pp. 2191–2197, 2012.

[68] M. Commodo, I. Fabris, C.P.T. Groth, and O.L. Gulder,, “Analysis of aviation

fuel thermal oxidative stability by Electrospray Ionization Mass Spectrometry (ESI-

MS),” Energy & Fuels, vol. 25, pp. 2142–2150, 2011.

[69] M. Commodo, O. Wong, I. Fabris, C.P.T. Groth and O.L. Gulder, “Spectroscopic

study of aviation jet fuel thermal oxidative stability,” Energy & Fuels, vol. 24,

pp. 6437–6441, 2010.

[70] S. Zabarnick, “Pseudo-Detailed Chemical Kinetic Modeling of Antioxidant Chem-

istry for Jet Fuel Applications,” Energy & Fuels, vol. 12, pp. 547–553, 1998.

[71] T. Doungthip, J. Ervin, S. Zabarnick, and T. Williams, “Simulation of the Effect

of Metal-Surface Catalysis on the Thermal Oxidation of Jet Fuel,” Energy & Fuels,

vol. 18, pp. 425–437, 2004.

Bibliography 77

[72] J. Ervin, S. Zabarnick, and T. Williams, “One-Dimensional Simulations of Jet Fuel

Thermal-Oxidative Degradation and Deposit Formation Within Cylindrical Pas-

sages,” Journal of Energy Resources Technology, vol. 122, pp. 229–238, 2000.

[73] V. Katta, J. Blust, T. Williams, and C. Martel, “Role of Buoyancy in Fuel-Thermal-

Stability Studies,” Journal of Thermophysics and Heat Transfer, vol. 9, pp. 159–168,

Jan-March 1995.

[74] J. Smith, “Fuel for the SST: Effects of Deposits on Heat Transfer to Aviation

Kerosene,” Industrial and Engineering Chemistry Process Design and Development,

vol. 8, pp. 299–305, 1969.

[75] A. Peat, “Thermal Decomposition of Aviation Fuel,” Tech. Rep. Paper 82-GT-27,

American Society of Mechanical Engineers, New York, 1982.

[76] P. Marteney, “Thermal Decomposition of JP-5 in Long Duration Tests,” Tech. Rep.

Report NAPC-PE-201C, United Technologies Research Center on contract with

Naval Air Propulsion Center, Trenton, NJ, June 1989.

Documents

CHEMICAL KINETIC AND THERMAL NUMERICAL ......Abstract CHEMICAL KINETIC AND THERMAL NUMERICAL SIMULATION OF COKING PROCESS OF JET FUELS IN THIN NOZZLE SECTIONS WITHIN AUTOXIDATION TEMPERATURE