politechnika gdańska - Pomeranian Digital Library

POLITECHNIKA GDAŃSKA
WYDZIAŁ MECHANICZNY
KATEDRA TECHNIKI CIEPLNEJ
Rozprawa doktorska
Weihong Yang
EXPERIMENTAL AND MATHEMATICAL MODELING
STUDY OF GAS COMBUSTION UNDER THE HIGHTEMPERATURE AND OXYGEN-DEFICIENT OXIDIZERS
Promotor pracy:
Prof. dr hab. inż. Jan Stąsiek
Gdańsk 2008
Pragnę
złożyć
serdeczne
podziękowanie
promotorowi Profesorowi Janowi Stąsiekowi za
życzliwą pomoc, konsultancje oraz ogromną
cierpliwość podczas realizacji pracy.
Dziękuję również Profesorowi Włodziemierzowi
Błasiakowi za umożliwienie pobytu naukowego w
Royal Institute of Technology (KTH), Sztokholm,
Szwecja i uczestniczenia w pracach badawczych
realizowanych w laboratoriach Katedry Wymiany
Ciepła i Pieców Przemysłowych (Division of Heat
and Furnace Technology).
TABLE OF CONTENTS
List of Figures
iv
List of tables
viii
Nomenclature
ix
1.
1
2.
3.
Introduction
1.1 Motivation
1
1.2 Thesis
4
1.3 Objective
4
1.4 Scope of the work
4
State-of-the-art
6
2.1 History of the HiTAC
6
2.2 Study of single jet flame
9
2.3 Study of a semi-industrial furnace with HiTAC burners
14
2.4 Flame volume, length and entrainment
17
2.5 NO emissions
17
Methodology
20
3.1 Experimental facility
20
3.1.1 Single jet experimental facility
20
3.1.2 Classification of the HiTAC
22
3.1.3 The HiTAC test furnace
26
3.1.4 Measurement program
30
3.2 Mathematical model for HiTAC
31
3.2.1 turbulent combustion model
31
3.2.2 NOx modes
33
3.2.2.1 Thermal NO
34
3.2.2.2 Prompt-NO
35
3.2.2.3 N2O –intermediate Nox
36
3.2.2.4 NO reburning
38
3.2.2.5 NO source term and turbulence-chemistry interaction
39
3.2.3 Other models
40
i
3.3 Developed concept for description of HiTAC characteristics
4.
5.
6.
7.
40
Cross-flow single jet flame study
45
4.1 Study of mathematical modelling
45
4.2 Study of Mathematical models
49
4.3 Mean Residence Time and Flame Peak Temperature
52
4.4 Gas Temperature Uniformity Ratio
53
4.5 Conclusions
55
Flame volume and length study in co-flow single jet flame
56
5.1 Flame appearance
57
5.2 Effect of oxygen concentration
57
5.3 Effect of oxidizer temperature
63
5.4 Effect of fuel temperature
64
5.5 Effect of fuel flow rates
66
5.6 Effect of fuel nozzle diameter
67
5.7 Scaling analysis
68
5.8 conclusion
72
Flame entrainment study in co-flow single jet flame
73
6.1 Study Cases
74
6.2 Effects of the oxygen concentration of the oxidizer on entrainment
76
6.3 Effects of the temperature of the oxidizer on entrainment
83
6.4 Effects of the fuel flux on entrainment
84
6.5. Effects of the buoyancy
85
6.6. Global field behaviour of the entrainment
90
6.7 Conclusions
94
Semi-industrial furnace with HiTAC burner study
95
7.1 Experimental measured and verification of mathematical modelling for HiTAC
furnace with one-flame burner
95
7.1.1 Energy Balance
95
7.1.2 Temperature Field
97
7.1.3 Gas Species
98
7.1.4 Features of Combustion and Flow of HiTAC
101
7.1.5 Heat Transfer Elevation
108
7.1.6 Simulation of a Moving Slab
110
7.2 Study of the HiTAC furnace with a twin-flame HiTAC burner
111
ii
7.2.1 Experimental and verification of the modelling prediction Heat transfer
elevation
7.2.2 Effect of flame configurations
114
7.2.3 Effect of excess air ratio
117
7.2.4 Effect of Fuel/Air injection momentum ratio NO emission
119
7.3. OPTIMAL DESIGN OF A HiTAC FURNACE
120
7.3.1 Flame Entrainment Ratio
120
7.3.2 Optimal Design of HiTAC Furnace
124
7. 4 CONCLUSIONS
8.
111
NOx formation and destruction mechanism study
129
131
8.1 Experimental and verification of NO emissions models
131
8.2 Effects of NO model on the NO formations
134
8.3 NO emissions From Coflow Gas Jet Combustion Study
139
8.4 Conclusion
143
9.
Thermodynamics investigation of the HiTAC combustion
144
10.
Mainly Conclusion of the thesis
149
11.
References
151
iii
List of Figures
Figure 3.1 Schematic of the combustion chamber for single fuel jet test furnace (a)Cross
flow (b) Co flow
21
Figure 3.2 Scheme of one-flame HRS
24
Figure 3.3 Scheme of two-flame HRS
25
Figure 3.4 Two-flame regenerator system firing configuration for uniformity temperature
profile in the furnace (F means Fuel, A means air and F.G. means flue gas) (a)
single-side firing configuration (b) stagger firing configuration(c) counter firing
configuration
25
Figure 3.5 HiTAC test furnace and burner (a) HiTAC test furnace at KTH, (b)
Configuration of HiTAC test furnace with one-flame HRS
Figure 3.6 Configuration of the HiTAC test furnace with inlet arrangements
27
28
Figure 3.7 Configuration of HiTAC test furnace with two-flame HRS (a) Two-flame HRS
(b) Top view of HiTAC test furnace and test positions
28
Figure 3.8 Top view of HiTAC test furnace and test positions
29
Figure 4.1 3D-computational domain and meshes of HTAC test furnace
46
Figure 4.2 Predicted temperature profiles for 10% oxygen in the air preheated upto 1041K
and at fuel inlet temperature equal to 473 K. (a) PDF model; (b)EBU model
47
Figure 4.3 Predicted distributions of oxidation mixture ratio for 10% oxygen in the air
preheated upto 1041K and at fuel inlet temperature equal to 473 K. (The same
scale as Figure 4.2)
47
Figure 4.4 Predicted gas temperature profiles for 10% oxygen in the air preheated upto
1041K and at fuel inlet temperature equal to 288 K. A=4.0, (b) A=2.0, (c) A=1.0
48
Figure 4.5 Predicted distributions of the oxidation mixture ratio for 10% oxygen in the air
preheated upto 1041 K and at fuel inlet temperature equal to 288 K. (a) A = 4.0,
(b) A = 2.0, (c) A = 1.0
Figure 4.6 Rf versus oxygen concentration for various air and fuel temperatures (K)
49
50
Figure 4.7 Rf versus preheated air temperature for oxygen concentration and fuel
temperatures (K)
51
Figure 4.8 Mean residence time (s) versus oxygen concentration and combustion air and
fuel temperature (K)
52
iv
Figure 4.9 Peak flame temperature (Tmax) versus oxygen concentration for various
combustion air and fuel temperature(K)
53
Figure 4.10 Temperature uniformity ratio, Rtu versus oxygen concentration for various air
and fuel temperatures(K)
54
Figure 5.1 Flame appearance for the LPG jet co-flowing vertically with hot and oxygen
depleted flue gas. Fuel nozzle diameter 0.5 mm. Fuel jet velocity equals to 25.5
m/s. Flue gas velocity equals to 0.98 m/s
57
Figure 5.2 Flame apparence versus oxygen concentration To = 1173 K, dF = 5.E-4 m, QF =
0.01 g/s, TF =299K
58
Figure 5.3 Predicted flame shape and size for different oxygen concentration (To = 1173 K,
dF = 5.E-4 m, QF = 0.01 g/s, TF =299K)
59
Figure 5.4 Length and volume of HiTAC flame versus oxygen concentration.(To =1173 K.
dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
60
Figure 5.5 Predicted temperature profiles for different oxygen concentration.(To =1173 K.
dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
61
Figure 5.6 Flame maximum temperatures versus oxygen concentration.( To =1173K. dF =
5.E-4 m. QF = 0.01 g/s. TF =299K)
62
Figure 5.7 Flame maximum temperatures versus air temperature.( [O2]=12.8%. dF = 5.E4 m. QF = 0.01 g/s. TF =299K)
62
Figure 5.8 Length and volume of the HiTAC flame versus oxidizer temperature.
([O2]=12.8%. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
64
Figure 5.9 Length and volume of the HiTAC flame versus fuel temperature. [O2]=10%. dF
= 5.E-4 m. QF = 0.01 g/s. TO =1173K
65
Figure 5.10 Length and volume of the HiTAC flame versus fuel firing rate.(TO =1173K. dF
= 5.E-4 m.)
66
Figure 5.11 Length and volume of the HiTAC flame at different cases ( [O2]=10%. dF =
5.E-4 m. TO =1173K)
67
Figure 5.12 HiTAC Flame length and volume versus diameter of fuel nozzle. ([O2]=10%.
TO =1173K. TF=299K QF =0.01 g/s)
68
Figure 5.13 HiTAC Flame length and volume correlated with the flame Froude number
70
Figure 6.1 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different oxygen concentrations in the oxidizer
76
v
Figure 6.2 Entrainment rate change versus the oxygen concentration and the characteristic
temperature ratio
(T f / To ) 0.5
81
Figure 6.3 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different preheated temperature of the oxidizer
83
Figure 6.4 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different preheated temperature of the oxidizer
85
Figure 6.5 The local influence of buoyancy as a function of the axial distance normalize by
a) (x/d*), and b) the flame height in log scale for difference cases
89
Figure 6.6 Dimensionless entrainment rates v.s. ratio of axial distance from the nozzle tip to
flame height
91
Figure 7.1 Temperature distribution on the side wall of the furnace at x = 0.8 m, and y = 0.3m
98
Figure 7.2 Predicted and measured O2 profiles in the furnace x = 0, z = 0.3 m (b) x = 0, z =
0.6m (c) x = 0, z = 1.2 m
99
Figure 7.3 Predicted and measured CO profiles in the furnace (a) x = 0, z = 0.3 m , (b) x =
0, z = 0.6 m, (c) x = 0, z = 1.2 m, (d) x = 0, z = 2.15 m
100
Figure 7.4 Predicted velocity vectors at a cross section through the fuel and one of the air
inlets in HiTAC mode [m/s]
102
Figure 7.5 Predicted temperature profile at a cross section through the fuel and one of the
air inlets at HiTAC mode [K]
103
Figure 7.6 Predicted and measured flame shapes and volumes shown by the oxidation
mixture ratio RO=0.99 (a) Predicted (b) Estimated by experimental data
105
Figure 7.7 Predicted flame shape Conventional burner
107
Figure 7.8 Predications of heat flux absorbed by the charge along central line on the surface
of sink Case 0: Test furnace with a one-burner HRS without any charge or heat
sink. Case 1: Test furnace with conventional turbulent jet flame without any
charge or heat sink. Case 2: Test furnace with a one-burner HRS and with a
stationary heat sink whose surface temperature is equal to 20OC and constant,
Case 3: Test furnace with conventional turbulent jet flame with a stationary heat
sink whose surface temperature is equal to 20OC and constant.Case 4: Test
furnace with a one-burner HRS with a moving steel slab which initial surface
temperature is equal to 20OC.
vi
109
Figure 7.9 Predicted temperature distribution of moving slab with one-flame HRS
111
Figure 7.10 Predicted and measured CO profiles in the upper part of the furnace (a) x=0,
z=1.7m, (b) x=-0.475 m, z=1.4m, (c) x=0., z=0.75m
113
Figure 7.11 Predicted and measured O2 profiles in the upper part of the furnace at x=-0.475
m, z=1.4m
114
Figure 7.12 Temperature profiles (K) at various firing locations (a) Counter mode (b)
Single-side mode, (c) Stagger mode
114
Figure 7.13 Flame shapes shown by the oxidation mixture ratio RO=0.99 for difference
firing modes (a) Counter mode (b) Single-side mode (c) Stagger mode
116
Figure 7.14 Effect of firing configurations on NO emissions
117
Figure 7.15 NOx emission vs excess air ration during the combustion of counter mode
118
Figure 7.16 NO variations with the ratio of the fuel/air injection momentum
120
Figure 7.17 Flame entrainment ratio vs. excess air ratio (ALFA)
121
Figure 7.18 Excess air ratio vs. flame occupation coefficient and peak temperature
122
Figure 7.19 Effect of fuel temperature and flame locations on flame entrainment ratio
123
Figure 7.20 The scheme of structure of the flame zone and recirculation in an enclosed
combustion chamber
125
Figure 7.21 Distributions of Temperature and Flame entrainment ratio along the direction
of the furnace width
128
Figure 8.1 Effects of excess air ratio on NO emission
132
Figure 8.2 Effect of excess air ratio on maximum temperature
133
Figure 8.3 Predicted NO contours (mass%) at the cross-section through fuel and one air
nozzle in the test furnace at excess air ratio (λ) equal to 1.04 and 1.09 with and
without N2O-route
136
Figure 8.4 Predicted NO contours (mass%) at the cross-section through fuel and one air
nozzle in the test furnace at excess air ratio (λ) equal to 1.15 and 1.25with and
without N2O-route
137
Figure 8.5 Temperature distribution at the cross-section through fuel and one air nozzle in
the test furnace at different excess air ratio (λ)
138
vii
Figure 8.6 Effect of oxygen concentration on NO emissions at an air temperature of 1173K
and fuel temperature of 299K for gas jet combustion in the co-flow test
141
Figure 8.7 Effect of preheated air temperature on NO emissions at a fuel temperature of
299K for the case of gas jet combustion in the co-flow test
142
Figure 8.8 Effect of fuel flowrate on NO emission at an air temperature of 1173K and a fuel
temperature of 299K for gas jet combustion in the co-flow test
142
Figure 9.1 Enthalpy-temperature diagram for CH4 mixture with different oxygen level at
the stoichiometric
145
Figure 9.2 Excess enthalpy in log versus inlet temperature at different oxygen concentration
for a methane/oxygen/nitrogen mixture
146
viii
List of Tables
Table 3.1: Arrhenius kinetic coefficients used for nitrous oxide mechanism reactions
Units:cal,mol,cm3,sec
37
Table 6.1 Values of Variables of Numerical Studies
74
Table 6.2 Summary of the results for present study and previous works
78
Table 7.1 Fuel characteristic and burner operating conditions
96
Table 7.2 Energy balance from measurements and from modelling data (Reference
temperature: 298K)
97
ix
Nomenclature
A - empirical coefficient
Acr - criteria area of flue gas recirculation
Ar - area of flue gas recirculation
B - empirical coefficient
C - pre-exponential factor of chemical reaction
Ce - entrainment coefficient
CHiTAC - Richardson correction ratio for the HiTAC combustion
Cp - specific heat, kJ/(kg.K)
D - diameter, m
d* - equivalent diameter defined
ds - source diameter, m
Ea - activation energy, kJ/mol
Fb - the buoyancy force, N
Fm - the initial inertia force of the fuel nozzle, N.
Fr - Froude number
fs - the stoichiometry
g - gravitational acceleration, m/s2.
k - turbulent kinetic energy, m2/s2
ka - absorption coefficient of flue gas
L - flame length, m
L* - dimensionless flame length
M - molecular weight (kgmol-1)
m - mass fraction
M molecular weight (kgmol-1)
m* - non-dimension entrainment mass flux
m0 - initial jet mass flux,
me - the jet mass flux,
n - stoichiometric coefficient (number of moles)
qc
- local heat release of difference fuel species, (kW/m3)
Qf - the heat release in flame zone, (kW).
qFHR - average Flame Heat releasing ( kW/m3)
R - the universal gas constant, kj/kmol.K
x
r - the density weighted velocity ratio between the jet and coflow, (( ρ F u F2 ) /( ρou o2 )) 0.5
R - universal gas constant, (Jkmol-1K-1)
r - radial distance to the flame centreline within the flame, m
Re - Reynolds number
REBU - fuel consumption rate for eddy break up model (kgm-3s-1)
Rent - entrainment rate
Rfe - flame entrainment ratio
Rflame - radial distance from the flame centreline to the flame board, m.
RFOC - furnace occupation coefficient
RHiTAC - rate of high-temperature and oxygen deficient (HiTCA) combustion, kg/m3/s
Ri - Richardson number
RKIN - Arrhenius reaction rate (kgm-3s-1)
RKIN - Arrhenius reaction rates, kg/m3/s
Ro - oxidation mixture ratio
Ro - oxidation mixture ratio
Rtu - furnace temperature uniformity ratio
S - stochiometric air to fuel mass ratio,
T - temperature (K)
ui - velocity in the i cell number, m/s
v
- velocity
Vf
- flame
volume (m3).
VF - furnace volume ( m3)
vj - the species rate exponents,
x - the downstream distance from the virtual origin, m.
Z - horizontal distance from burner face, (m)
β
- temperature exponent
δ - jet width, m
ξ - the nondimensional streamwise coordinate
ξ’ - correction of the nondimensional streamwise coordinate
ξHiTAC - correction nondimensional streamwise coordinate for HiTAC combustion
ρ - density, kg/m3
υ - the specie rate exponent
u - average velocity, m/s
T - average temperature in furnace (K)
xi
ε - turbulent kinetic energy dissipation,m2/s3
[O2] - oxygen concentration, %
∏ ( x ) - denotes the product of all xj.
j
β - temperature exponent of chemical reaction
ΔT f - the characteristic temperature rise resulting from combustion, K
Subscripts
0 - initial state
∞ - ambient
a - air
ad - adiabatic temperature
c - species
f - flame
F - fuel
F - Furnace
Fl - fuel
i
- calculation cell number
Max - maximum values
O - oxygen
O, o - oxidizer
out - final state
P - product
r - recirculation flue gas
s - source
stoic - stoichiometry
xii
Abstract
The combustion proceeds in an atmosphere of low oxygen concentration, as well as at high
temperatures of the oxidizer, mostly above the auto ignition temperature of the fuel, spread
out many new features, such as significantly higher flame stability at all fuel-air (including
very lean mixtures) a larger flame volume, a uniform temperature distributions, higher heat
transfer, low NOx emission. These features have been demonstrated in a lot of practice
applications, either with air as oxidizers combing a modern regenerative system, or with
pure oxygen as oxidizers.
A comprehensive view of these works has been provided, in regard to the fundamental
differences in the thermal, chemical and fluid dynamics characteristics of the flame.
However, since the high temperature and low oxygen deficient are the unique parameters
differ from conversations combustion, there is little knowledge of the quantitatively effects
of them on the flame properties.
The generally objectives of this work are theoretical analysis and development of new
combustion concepts. In particular,
•
Quantitative study of the flame properties, attention is focused on the volume,
length, and entrainment of an ‘invisible’ flame.
•
NOx formation and destruction mechanisms during new combustion conditions.
The investigation has been performed using a single fuel jet flame facility with cross and
co-flowing, and a semi-industrial furnace equipped with High-temperature Air
Combustion-HiTAC
burners.
Experimental,
numerical
and
theoretical
analyzing
investigations are adopted.
In this work, a ‘chemical’ flame volume and ‘chemical’ flame length were used to describe
this ‘invisible flame’. Results from single jet flame study show that:
•
Flame length increases with either the decrease of oxygen content, or increase of
oxidizer temperature, or decrease of fuel temperature. Furthermore, the flame length is
independent of the fuel flow rate and the diameter of the fuel nozzle for the studied
cases.
xiii
•
Flame volume increases either with the decrease of oxygen content and increase of
oxidizer temperature, or with the reduction of fuel temperature, or with the increasing in
fuel firing rate. Flame volume depends very much on the oxygen concentration in the
oxidizer.
•
Influences of high temperature and low oxygen concentration in the oxidizer on the
flame Froude number, Frf were examined. Regimes of momentum- or buoyancycontrol, were determined on the assumption that oxidizer temperature and oxygen
concentration are changeable. A simple correlation of the ‘flame’ length and volume
with flow parameters has been derived in terms of a flame Froude number for
momentum-buoyancy transition jet flame under the HiTAC condition. The criteria
constants of the dimensionless flame volume V* and the dimensionless flame length L*
to assess momentum– or buoyancy–control flame are given.
Additionally, the entrainments of this ‘invisible’ flame have been numerically and
theoretical studied. Conclusions are:
•
The uniformity of the heat release in reacting jets has strong effect on the flame
entrainment. More uniform the heat release, larger the entrainment. The effect of heat
release reduces the entrainment in the near field of the reacting jets with the same
factor of the characteristic ratio (Tf/To)0.5.
•
The entrainment increases as the oxygen concentration is decreased. Furthermore, the
entrainment is independent of the fuel flow rate and the preheated temperature of the
oxidizer for the investigated temperature range (1073-1573K).
•
The effect of the oxygen concentration and preheated temperature of the oxidizer on
buoyancy was examined. A correction Richardson coordinate, where the effect of the
oxygen concentration (stoichiometric ratio) is included, was derived to describe the
local influence of buoyancy force along the chemical flame length under the high
temperature and oxygen deficient oxidizer condition. It can be concluded that the
buoyancy force increases with the reduction of the oxygen concentration in the
oxidizer.
•
The global behaviour of the entrainment was revealed. The entrainment of jet flames
can be identified as two regimes: (a) the near field where entrainment coefficient is
positive; and (b) the far field where entrainment coefficient is negative. Corrections of
entrainment rates were derived in terms of a Frf number for momentum-buoyancy
xiv
transition jet flame under the high temperature and low oxygen concentration oxidizer
condition. Furthermore, the maximum entrainments along the flame length are
estimated
Further on, the benefits of HiTAC technology are quantitatively demonstrated by
mathematical models. These benefits are: lower peak temperature, larger flame volume,
more uniform thermal field, lower local firing rate, higher heat transfer, higher energy
utilizing efficiency and lower combustion noise.
NOx formation and destruction during this new combustion phenomenon has been studied
numerically. It was found that the NO formation via N2O mechanism may be important.
The approximate percentage of NO production by the nitrous oxide according to the
Zeldovich and prompt mechanism varies from 5:95 at 10% oxygen concentration to 95:5 at
5% oxygen concentration.
Finally, a thermodynamic analysis of combustion process shows that oxyfuel combustion is
able to increases the available energy of the flameless combustion, thus higher combustion
intensity. Additionally, the flameless oxyfuel burner does not need preheating the oxidizer,
this extend the concept of the HiTAC.
xv
1 Introduction
1.1 Motivation
According to the Kyoto Protocol on Global Climate changes as the third conference of
parties (COP3) to the United Nationals Frame conversion that developed countries should
reduce their total emissions of greenhouse gases by at least 5% from the level of 1990
between the years 2008 and 2012. The European Union and its implementation recently
decided to set a target of 20% CO2-reduction, and 20% renewable energy using.
Fossil fuel, such as gas and oil is used in the various fields such as steel production,
petroleum, electronics and machinery and the energy saving on these industrial processes
will result in big effect for reducing global warming gases. In 2007 the fossil fuels
constituted 80% of the global energy supply (35% petroleum + 24% coal + 21% natural gas)
[1]. Further on, emission of NOx (NO,NO2, N2O, N2O4, N2O5,etc), which is known to be
responsible for the destruction of ozone layer in the upper atmosphere, is another
environment problem when fossil fuel is used. Therefore the energy saving and
environment protection technologies should be developed and then applied as soon as
possible in the practical fields.
The most efficient saving energy technique for the combustion of fossil or derived fuels is
either preheating combustion air /fuel by recovering flue gas heat or reducing the flue gas
volume (oxygen enhancement in air ), or combining both of them. However, directly used
of these methods will lead to high combustion temperature, thus a higher NOx emission
High temperature air combustion-HiTAC technology can simultaneously reduce CO2 and
nitric oxide emissions and reduce energy consumption. It has been now well recognized as
the most advance combustion technique to achieve all requirements.
The essence of the HiTAC is fuel combustion under the condition of high-temperature and
oxygen and deficient atmosphere. A lower oxygen concentration during the combustion
process can depress the flame temperature even at a higher preheated oxidizer /fuel, and a
higher temperature of oxidizer can sustain the flame stability. The necessary temperature of
the oxidizer can be decided according to the oxygen concentration in the oxidizer.
1
It has been now well recognized that HiTAC has advantages as:
•
Energy savings
•
CO2 reduction
•
NOx reduction
•
Uniformity of temperature profile
•
Enhancement of heat transfer
•
Lower combustion noise
In the practical applications, the high-temperature of oxidizer can be obtained using a
modern regenerative system, or mixing of cold oxidizer (air/oxygen) and the hot flue gas by
a higher injection velocity of the oxidizer. For the fore method, the combustion oxidizer can
be preheated by regenerative system directly, and for the latter, the oxidizer can be heated
by stronger flue gas internal recirculation. Both technological solutions can be obtained
high-preheated oxidizer temperature before combustion occurs. The temperature of the
preheated oxidizer at the point before combustion occurs should be higher than the autoignition temperature of used fuel. The low oxygen concentration in the oxidizer can be
obtained either by an external exhaust gas recirculation, or by an internal flue gas
recirculation obtained by a high velocity injection of oxidizer and/or fuel.
Gases and liquid fuel combustion with modern regenerative type of heat exchangers has
been demonstrated in industrial furnaces, for example, reheating furnaces, heat treatment
furnaces, melting furnace etc, for energy savings (about 30%), reduce pollutant’s emission,
including CO2(about 30%), and reduce equipment size (about 25%) [2]
Additionally, for oxygen enhancement or pure oxygen combustion, this kind of
phenomenon is achieved by supersonic injection of oxygen, which forms a strong internal
flue gas recirculation. This technology is called as ‘’flameless oxyfuel’’. Dramatically fuel
energy savings have been successfully demonstrated in different types of furnaces along
with uniform thermal field and pollution.
High- temperature air combustion technology has already been applied in steel industrial,
reported by Yasuda [3], Suzukawa [4] and Mori [5]. In all these projects very high fuel
savings (sometimes as high as 60%), reduction of NOx (around 50%) and production
2
increase (by 20 -50 %) was achieved. Base on these factors, apart from steel industry,
HiTAC were applied to melt aluminum or to incinerate odour, vapor gases for example in
pulp and paper industry.
Industrial applications of this so called “flameless” oxyfuel combustion are documented [610] and prove enormous future potential of this technology in various thermal processes for
example soaking pits, box-, walking beam-, catenary furnaces [10].
This new combustion proceeds in an atmosphere of low oxygen concentration, as well as at
high temperatures of the oxidizer, mostly above the auto ignition temperature of the fuel.
The high-temperature (above the fuel’s auto ignition temperature) and oxygen deficient
atmosphere are unique characteristics that differ from any other combustion. A more
accurate definition would be that combustion is spread out in a large volume, which some
refer to as ‘volumetric combustion’, resulting in uniform and lower temperature of such
flame. The essence of these larger flame volume and lower flame temperature is the low
temperature increase during the combustion process. This makes a combustion chamber
more like a well stirred reactor. Again, the effect of heat release on the combustion
characteristics is less in the case of a low temperature increase. For example, larger flame
volume and gas entrainments are found during high temperature and oxygen deficient
conditions.
This works firstly concerns the studying of some thermal and physical properties of gaseous
fuel flames, especial focus on the quantitative of ‘invislabe’ flame volume, length, and
flame entrainment for both cross and co-flow of single jet flame. Further on, the flame
properties in a semi-industrial HiTAC furnace with modern ceramic honeycomb
regenerative burner have been studied. These concepts were used to describe the
characteristics of the HiTAC in a practical application, which provides help for optimal
design of a HiTAC furnace and burners. As well, the benefits of HiTAC technology are
quantitatively demonstrated. Further on, the formation machines of NO during the HiTAC
has been studied. Finally, a thermodynamic analysis basing on the 2nd law of this new
combustion process will be performed in order to understand its limitation/drawbacks, the
essence of the HiTAC, and the potential extension of this new combustion.
3
1.2 Thesis
This thesis work will verify the fundamental phenomena of the gas combustion under high
temperature and oxygen deficient atmosphere, including larger and invisible flame, and
check the NOx formation and destruction mechanism. Furthermore, demonstrate the high
energy efficiency, low NOx emission in a semi-industrial furnace.
1.3 Objective
The generally objectives of this work are theoretical analysis and development of new
combustion concepts. In particular,
•
Quantitative study of the flame physical properties, attention is focused on the flame
volume, flame length, and flame entrainment.
•
NOx formation and destruction mechanisms during new combustion conditions.
1.4 Scope of the work
The investigation has been performed using a single fuel jet flame with cross and co-flow
during high-temperature and oxygen deficient atmosphere, and a semi-industrial furnace
equipped with HiTAC burners. Experimental, numerical and theoretical analyzing
investigations are adopted.
Firstly, basing on the single jet flame facilities, the concepts to description of the ’invisible
flame’ physical properties will be performed. The volume, length of the ‘invisible’ flame
will be quantitative studied with the unique characteristic parameters of the HiTAC, i.e.
high temperature and low oxygen concentrations. NO formation and destruction
mechanisms during this new combustion phenomenon will be investigated.
Secondly, a semi-industrial furnace equipped with HiTAC burners will be used for
researching. The benefits of the HiTAC technology will be quantitatively demonstrated by
mathematical models verified by experimental data. The developed concepts in above
works will be used to this HiTAC application to understand the combustion in a practice
furnace.
4
Finally, a thermodynamic analysis basing on the 2nd law of this new combustion process
will be performed in order to understand its limitation/drawbacks, the essence of the
HiTAC, and the potential extension of this new combustion, for example from air-fuel
combustion to oxyfuel combustion.
5
2 State-of-the-art
2.1 History of HiTAC
Weinberg [11] found the fact that preheating an ultra-lean mixture using the heat
recirculation method can achieve stable combustion and expand the flammability limits of
the mixture. He proposed the concept of additional enthalpy combustion, or ‘excess
enthalpy combustion’, which can be defined as energy contained in the hot exhaust gas is
recirculated back to the inlet air either internally or externally.
Further efforts were made by British Gas and later at Hotwork International, UK [12] to use
this concept for the larger scale industrial furnace, to develop both recuperative and
regenerative burners. However, the problem of higher NOx emissions was met when a
higher preheating air temperature is achieved.
Technical breakthrough of recovering waste heat from exhaust gases using honeycomb
regenerative burner for the industrial furnace was made by Mr R. Tanaka of NFK and
Hasegawa T etc his cooperators [2, 13, 14, 15]. In this concept, a modern regenerative
material is used to recover waste heat and the temperature of combustion air can be only 50
o
C-100 oC lower than the exhaust gases temperature. They reduced NOx emission at so
high preheating air temperature (over 1000 oC) using a novel idea of a low oxygen
concentration. Since the combustion occurs during the condition of high-temperature and
low oxygen, novel combustion phenomena appear. Since the combustion is preheating in
very high level, it was first named as ‘high-preheating air combustion’, or HPAC. Late on,
this was called high-temperature air combustion, i.e. HTAC or HiTAC. Further more,
Gupta [15] etc gave the definition of the HiTAC as ‘the maximum temperature of the
reaction zone is held constant by recirculating oxygen reduced combustion products back
into the inlet air. The combustion air is preheated to temperature in excess of 1000 oC’.
Another reason to use HiTAC is due to high-preheated air temperature upto 1300 oC using
modern regenerative honeycomb heat exchanger to recover waste heat from exhaust gas. In
this way, the utilization energy efficiency can be achieved very high level. There is another
name for this high temperature combustion technology- HiTCOT. This novel combustion
was also pioneering trial and industrial works carried out by Tokyo Gas with a technology
of Directly Fuel Injection, called as DFI.
6
The researching of this heat recirculation technology was also conduced in European
countries, for example in Germany. Wunning [16] proposed an idea of FLameless
OXidation for low thermal NOx emission. In this work a FLOX was proposed, and defined
as FLOX, FLameless Oxidation with high recirculation rates preheating oxidizer to a
o
o
minimum of 650 C with combustion chamber temperatures greater than 800 C.
It can be seen that the essence of this new combustion is that oxidize is diluted before
combustion occurs. During this new combustion atmosphere, the color of ‘flame’ can be
different, for example, green, blue, bluish depending on the dilutant used for oxidants, and
degree of diluting, and different the fuels. In some certainly condition, this is no ‘visible
flame’, and is referred ‘flameless’ Oxidation /combustion/ by Wunning [16], or ‘colorless
combustion’ by Gupta etc [17].
MILD is another name of this technology. For example, in the work of Daily [18], he used
this combustion as MILD, and defined as Moderate and Intense low Oxygen Diffusion
Combustion, Uses exhaust gas recirculation to raise the oxidizer temperature and dilute the
oxygen concentration to maintain a low flame temperature. Cavaliere et al [19] further
studied this technology, and also explained this MILD combustion technology in the way as
the temperature increase due to the reaction is lower or milder than the temperature of the
preheated reactants.
Diluted Combustion is another name for this technology, for example, Milani etc [20]
described this technology as Diluted Combustion here, dilution means that fuel and oxidizer
are mixed ‘locally’ with a ballast of inert gases before they react so that the oxygen
concentration in the reactants is substantially reduced with respect to the 21% of the
standard oxidizing air.
Another definition would be that combustion is spread out in a large volume, which some
scientists refer to as “volumetric combustion”, resulting in uniform and lower temperature
of such flame. The essence of these larger flame volume and lower flame temperature is
the low temperature rise during the combustion process. This is the fact that a low
temperature increase makes a combustion chamber more similar as a well-stirred-reactor.
7
Again, the effect of heat release on the combustion characteristic is less at the case of a low
temperature rise. For example, larger flame volume and entrainments are found during the
high-temperature and oxygen deficiency condition. This concept was developed to used to
the other fields, such as boiler, rotary kiln etc [21].
Blasiak etc [21] also named this combustion as a ‘quasi-isothermal combustion’ since this
new combustion has a lower excess enthalpy, which firstly extends this idea to the
combustion with pure oxygen.
Further, innovation of this combustion phenomenon was extended to combustion with
oxygen or oxygen enrichment [21]. When this new combustion is realized with air, it is
generally called as ‘flameless-airfuel’. When pure oxygen or oxygen enrichment is used
instead of air as the oxidizer, this is called as ‘flameless oxyfuel’.
Some of study of oil [2, 22, 23] and coal [2, 23, 24] has been conduced with hightemperature air in order to achieve all the benefits obtained from gaseous fuel combustion.
They are still a lot of ongoing working on this. It has been understood that high-temperature
and low oxygen concentration are the unique characteristic than that of other combustion
technology.
The use of high temperature oxidizer has been further developed for the biomass
gasification processing. In this application, referred to as High Temperature Agent
Gasification (HiTAG), a highly preheated gasifying media with temperature as high as
1273 K, provides additional energy into the gasification process, which enhances the
thermal decomposition of the gasified solids. HiTAG has significant advantages in the
gasification of low-rank biomass fuels and traditionally unusable waste streams like for
example as sludge. It can also operate efficiently on a wide range of feedstock [25, 26, 27,
28] and the HiTAG gasifier system can be built extremely compact with an atmospheric
pressure, lowering component costs.
Since the characteristics of the Flameless Oxidization, such as very stable combustion with
low NOx levels, it has been attractive people to use this idea to high pressure combustor,
for example the combustor of gas turbines. Major works in this field can be found in the
8
proceedings of Flamelesss Combustion for gas turbine, in Lund, Sweden, June 21, 2005.
[29].
A lot works have been done since 1990. Here, a basic literatures related to current work
will be given in next section, focusing on the study of a single jet gas-fuel and semiindustrial industries application.
2.2 Study of Single jet flame
Hasegawa et al,(1997) [14] performed studying with the NKF test facility, which designed
and fabricated by the Nippon Furnace Kogyo Kaisha, NFK, consists of two combustion
chambers, each fitted with ceramic honeycomb regenerator, controllers for flow and
switching sequence. The high temperature combustion air, produced from the regenerative
test facility, and a gas jet can then be injected into this preheated oxidizer. In their work, a
mixture of air, N2 and CO2 was as the oxidizer. They reported that the single jet gas flame
flammability does not decrease with decreasing of oxygen concentration when the oxidizer
temperature is sufficiently high. For example, the limiting temperature was 900oC for the
LPG fuel. They also found that the HTAC flame had much better uniformity of temperature
and the temperature fluctuations went from 197 oC under the conventional conditions down
to below 5 oC for HTAC. They also reported that the flame with 3% O2 in the oxidizer and
preheated to 1010 oC had very low visible radiation and a much large volume. The color of
flames was reported to be blue at conventional conditions and ‘Bluish-green’ or green at
HTAC firings. They found that CO2 had a stronger effect on NOx reduction then if N2 was
used to dilute the oxidizer.
Kishimoto et al (1997) [30] performed experiments on a similar NFK HTAC test facility.
The natural gas was chosen as fuel, and N2 as dilutant and took chemiluminescence images.
They found that the flame seemed to be stabilized at a large distance from the injection
nozzle. When diluting the oxidizer that the dilution made the flame unstable, increasing the
temperature made the flame stable again. They also performed experiments on a similar
NFK HTAC test facility. The natural gas was chosen as fuel, and N2 as dilutant and took
chemiluminescence images. The reported that the flame color changed from violet to blue
and then to green. Their explanation was the decreasing CH emission and the increasing C2
emission
9
Amagai et al (1995)[31], made an experiment in stead-state with a electrically heat oxidizer
into which they injected a co-flowing single fuel jet of propane. Both gas jet diffusion and
premixed flames were checked. The results reported that both the diffusion and premixed
flame were well stabilized when the temperature was over the auto-ignition temperature.
This work also demonstrated that the flame length of the laminar diffusion and the
premixed flames decrease with increasing temperature of the oxidizer temperature. On the
other hand, the length of turbulent diffusion increases with flame-laminarization due to high
temperature.
Kitagawa et. al.(1998) [32] used a spectra-video camera to capture the fluctuations of LPG
flame in oxidizer diluted with N2 from NFK HTAC test facility. The also measured the
temperature profile in the combustion chamber with thermocouple. They calculated the
vibration temperature in 2-D images of the flame using the two C2 bands. Very uniform
temperature profile was found at high preheatly and low oxygen concentration, which
coincided with the lowest temperature fluctuations.
Uniformity and homogenization of combustion of methane and propane with highly
preheated air (1000ºC) have been studied by Isiguro et al. (1998)[33]. They obtained
images of OH, CH and C2 emissions for a number of experimental conditions that differed
in the air preheat level and oxygen content of the air. The investigators concluded that the
increase in air temperature results in a decrease of flame temperature gradients
(homogenization of flames).
Furthermore the higher the combustion air temperature the lower the flame fluctuations are.
Plessing et al. (1998) [34] used laser-induced predissociative fluorescence and Rayleigh
thermometry to examine flameless oxidation at laboratory scale. They observed that the
flameless oxidation occurring in the well-stirred reactor regime. The OH concentration in
the combustion zones of flameless oxidation is lower than in non-preheated undiluted
turbulent premixed flames.
Soon afterwards studies [35-36] have been undertook research on the effect of combustion
air temperature and oxygen concentration on flame color, visibility and thermal emission
spectra. Both methane and propane were used as the fuel. The spectroscopy measurements
10
checked that there were large emissions CH and C2 as well as other intermediate species for
propane flame. But in the methane flame they did not measure such significant emissions.
However, they noticed that there were much less intermediate species emissions for the
methane flame.
Bolz (1998) [36] and Gupta (1999) [37] further investigated the influences on HiTAC
volume and lift-off (standoff) distances with direct flame photography. In this work, N2 was
as diluted oxidizer and fuel was LPG. They found that the flame color varied from yellow,
at 21 %O2 and 1100 oC, to yellow-blue flame when decreasing the oxygen concentration.
And at the HiTAC firing, the flame was found to be blue-green. At 900 oC, however, the
flame changed from yellow to blue, without any green color. Gupta et al, [37] further
investigated found that no yellow color in flames with less than 950 oC and low oxygen
concentration in the oxidizer. They concluded that the effect of oxygen concentration and
oxidizer temperature on flame color be of use when designing applications which need
certain radiant heat flux.
Mochida and Hasegawa (2000) [38] have been developing a flame visualization technique
based on the luminescence intensity ratio of C2 and CH radicals.
Lille et al. (2000) [39] built an experimental facility for studying fuel jets immersing into a
cross-flowing high temperature air stream and their preliminary findings have been
reported. In this present, the fuel was propane and N2 was used as dilutant. The high-speed
photography, direct and schlieren color visualization were used to record images of flames.
They conducted that a lower oxygen concentration increases the flame size. Air temperature
has an opposite effect of decreasing the flame size but not in the same proportions as an
increase in oxygen content. The flame visibility decreases with decreasing oxygen
concentration and increases with air temperature. The flame color changes from bright
white/yellow to blue/green/yellow and is primarily influenced by oxygen concentration.
Lille(2000)’s work depicted that a lower oxygen concentration increase the lift-off distance,
and for high velocity fuel jet increases the lift-off distance.
Lille et al (2002) [40] further investigated fuel jet combustion under the condition of high
temperature and oxygen deficient air. In this work, the jet of fuel was co-axially injected
into high temperature exhaust gases generated by means of a gas burner fired with Gasol
11
(mainly propane). They found that the oxygen concentration in the oxidizer have a
substantial effect on flame size, luminosity, color, visibility. The flame became first bluish
and then non-visible at sufficiently low oxygen concentration in the oxidizer.
In the co-flow study of Fujimori et al.(1997 [41],1998 [42])) and Sato et al (1997)[43], they
demonstrated that the reduction of NOx emission was largest for flames with high liftoff
distance, and diluting the fuel with N2 had a considerable effect on the NOx reduction. The
same results were found for air dilution.
Blasiak (2001) [44] reported that addition of N2 to methane results in lower emission of NO
and CO, as well as higher fuel injection velocity.
Rota et al [45] presented an experiment furnace in which partially premised fuel was coaxially injected into a preheated and diluted air stream. The investigations gave that the
NOx was decrease when run under HTAC conditions.
Using NFK HTAC test stand, Mochida, et al.(1995) [46] investigated the flame radiation by
means of CCD camera. The results showed than the relative intensity of HTAC flame is
much lower, approximately one-third, and is longer than a conventional firing. They
explained that the radical particles, which produce a more luminous flame, was formed in
conventional firing comparing to the HTAC flame.
Yuan et al.(1999) [47] investigated numerically the temperature distribution, soot formation
and NO emission under the condition of the highly temperature and deficient oxygen air.
They carried out the numerical model for NFK test furnace. In this work PDF model was
adopted to simulate the turbulent diffusion combustion using the equilibrium chemistry
method for chemical reactions. The soot formation was simulated using a two-step Tesner
model and both thermal-NO and prompt NO were calculated. They shown that the effect of
dilution of air stream on the decrease of maximum flame temperature various as
CO2>N2>He because the difference of heat capacity and/or difference of mixing processes
between fuel and diluted air. The influence of diluents on the NO emission has the same
tendency as on the temperature.
12
Dong [48] performed the mathematical modeling of single jet under the highly preheated
air combustion, the finite rate/eddy dissipation model and mixture fraction/ PDF are used
and compared each other, general characteristics of flame were conducted. Results show
that the mixture fraction/PDF model is better than the finite rate/eddy dissipation model for
predicting of the thermal fields of the turbulent jet flames, and thus can give out reasonable
prediction of NOx formation.
Dong(2000) [49] furthermore numerical studied gas single jet under high temperature air
combustion with advance flow model. In this work, the large eddy simulation (LES) and
Reynold Stress Model(RSM) together with the finite rate/eddy dissipation reaction model
were adopted. The results showed that these models gave a small difference in the near
field predicted of flow, in contrary to the empirical constant as for example Cs in LES
model which has significant influence on the predictions.
Yang (2001) [50] conducted whole studying of in highly preheated and oxygen deficient air.
Combustion of a single jet of propane in a cross-flowing stream of preheated and oxygen
deficient air is analyzed.
In the book on HiTAC (2000) [2], the characteristics of HTAC combustion numerical
calculating were analyzed. HTAC firing is controlled both by chemical kinetics and by
mixing. Then, the mixed fraction/PDF model and Eddy-Break-Up model are not suitable
for HTAC, since the chemical reactions of both of them are assumed as infinitely fast,
furthermore, PDF model was experientially obtained based on normal combustion reaction,
which implicitly infer the use of ambient air. Thus, the eddy dissipation concept with multisteps reaction of fuel is an interesting suggestion for calculating HTAC at present.
Furthermore, the single gas jet combustion under the condition of high temperature and
deficient air was carried out. The combustion models were investigated and comparing.
These models include one-step global reaction model, Jones’s four-step reaction model and
Srivatsa’s four-step reaction model. The results show that Srivatsa’s model was better
among the three model concerning the flame-lifted height and maximum flame temperature.
Mörtberg [51] recently has used Particle Image Velocimetry-PIV to obtain information on
the flow dynamics of a fuel jet injected into a crossflow of oxidizer at either a normal
temperature or a very high temperature. Light emission spectroscopy was used to collect
13
information on time-averaged radical distribution in the combustion jet. Jet turbulence,
time-averaged velocity distribution, fuel jet mixing, the distribution of radicals such as CH,
OH, and C2 and flame photographs were investigated. The results showed delayed mixing
and combustion under high-tem rapture low oxygen concentration condition. The
combustion air preheats temperature and oxygen concentrations were found to have a
significant effect on the burning, fuel-jet behaviour.
2.3 Study of a semi-industrial furnace with HiTAC burners
In order to understand the characteristics of this novel combustion technology, HiTAC has
been modeled by a single jet gas combustion in a low-oxygen partial pressure and hightemperature air environment. However progress in the use of different applications of
HiTAC technology has increased the need for more information and data for furnace and
process designers. In particular, it is very important to specify optimal conditions for
installation of HiTAC in industrial furnaces. For these reasons, studies are performed at a
larger scale where at least one set of regenerative burners systems is installed.
The advantages and efficiency of HiTAC technology have been investigated and tested for
semi-industrial and industrial-scale atmospheric burners by some researchers. For example,
the main characteristics of flameless oxidation as promising high efficiency and low NOx
combustion technology have been described by Wünning [16]. Performances of the
recently developed High-cycle-Regenerative burner System- HRS have been evaluated by
Suzukawa et al. [53]. Two pairs of burners were tested using a mixture of by-product gases
from steel plant, with a lower heating value of 11.5 MJ/m3. The capacity of each twoburner system was 0.93 MW. The honeycomb regenerator exhibited very good performance,
with preheated-air temperature being close to the furnace exhaust gas temperature, and very
low levels of NOx emission recorded.
The performances of two different HRSs operating on natural gas were examined by
Quinqueneau et al. [54, 55]. Ranges of burner operation and levels of CO and NOx
emissions were measured for both one and two-burner systems. Both systems were tested at
the same capacity of 200 kW.
14
The International Flame Research Foundation-IFRF has carried out a series of semiindustrial-scale tests to identify the principal characteristics of the combustion process
using high-temperature air. The fuel used in these tests included natural gas [56], oils [57]
and one type of coal [58]. The behaviours of this technology were demonstrated through
measuring flow field, in-furnace temperatures and heat fluxes. These experiments
confirmed that HiTAC technology is an extremely attractive technology for efficient and
environmentally favourable combustion. A review of above works has been undertaker by
Weber [59]. Furthermore, the most important discoveries relating to ’at the burner’ heat
recovery methods and NOx reduction methodology have also been summarised and
reviewed by Weber (with respect to combustion air of temperature in excess of 1000oC is
used) [60].
The effects that changing the size of test furnace chamber on the performance of the oneflame HRS have been investigated by Szabo et al. [61], who also performed in-flame
measurements of for example O2, CO and temperature, and also studied the effects of firing
rate, fuel type, NO emissions and heat transfer. Measurements of composition and heat
fields have been performed inside a semi-industrial test furnace equipped with a 200kW
HiTAC burner at KTH [62, 63, 64 ]. These investigations into the performances of one and
two-flame burners used Liquefied-Petroleum-Gas-PLPG fuel. The results of these were also
compared with a conventional high-velocity-jet burner.
In Parallel with these semi-industrial tests, numerical models of the furnace have to be
developed and verified. A numerical simulation in an industrial slab reheat furnace with
two-flame HRS was performed by Ishii [65] and Hino [66] and respective co-researchers.
These results showed that the ratio of the air-to-fuel injection velocities has a strong
influence on the rate of NOx production in the furnace. However, they found that NOx
models used currently might require improvement in order to properly describe NOx
formation under “low-temperature” conditions.
An unsteady flamelet model was employed to calculate the mild combustion mode in a
laboratory-scale combustor by Coelho and fellow researchers [67]. Their results show that
the steady flamelet library was unable to describe the formation of NO since this is a
chemically slow process, which is sensitive to transient effect the unsteady flamelet model
was able to predict the correct order of magnitude of NO emissions.
15
A steady state simulation of a one-flame burner with 200kW capacity has been performed
by Pasenti et al. [68]. The numerical results showed good agreement for total radiative heat
flux between measured and predicted values, but NOx emissions were substantially
underpredicted. However, the trend of NOx emissions on the furnace exiting temperature
was correctly represented in the predictions.
IFRF tests using natural gas combustion with preheated air were modelled and reported by
Dong [49]. The semi-industrial test furnace used was equipped with one burner operated
under steady state conditions. The high-temperature regenerator was replaced by a precombustor, where natural gas was burned with air under lean conditions. Experiment data
from this test furnace were used by Orsino et al.[69] to assess the abilities of exitedcombustion models to predict the characteristics of this new combustion technology. The
combustion models tested included the eddy-break-up model, the eddy-dissipation-concept
model, and PDF/mixture fraction model, had all correctly reproduced the characteristics of
the high temperature air combustion with the exception of the small region located within
the natural gas jet. Based on the above IFRF furnace, a recent simulation predicting the
combustion characteristics with an emphasis on NOx formation and destruction of nitrogen
oxides was performed by Mancini et al. [70]. They identified the location of regions of NOx
formation. They concluded that the predicted values were substantially lower than the
measured emissions for low preheated levels.
The differences in heat transfer and combustion features between conventional highvelocity-turbulent jet flame and HiTAC flame have been numerically investigated by Yang
et al [71]. The influence of stationary and permanent heat sinks on furnace heat transfer in
these two different types of burner systems was also investigated and compared. The
authors concluded that the HiTAC technology is more sensitive to heat sink than
conventional combustion (high velocity turbulent jet flame combustion).
These publications focus on the evaluation of HRS performance in relation to heat recovery
efficiency, heat flux evaluation and NOx emissions. Before utilizing HiTAC in various
applications, flame properties have to be known as these determine the flame stability, and
the designing of the burner and the chamber.
16
Rafidi Nabil [72] has done the thermodynamic study and of the new combustion technology.
His analyses display the possibilities of reducing thermodynamic irreversibility of
combustion by considering an oxygen-diluted combustion process that utilizes both gasand/or heat-recirculation. Additionally, the results showed that an oxygen-diluted
combustion system that utilizes oxygen as an oxidizer, in place of air, results in higher 1st
and 2nd law efficiencies. Furthermore, Mathematical models of heat regenerator’s heat
transfer were performed for maximized heat recovery.
2.4 Flame volume, length and entrainment
It is known that the diffusion flame length can be generalized as the function of fuel jet
momentum and furnace temperature by means of the Froude number, Frf [73, 74, 75].
The entrainment of the jets is the key technique solution for the industrial applications of
this novel combustion technology. Jet entrainment is the radial inward flux of ambient fluid
drawn into a jet. The entrainment into a turbulent jet was investigated in a number of earlier
studies [76-90]. For the free nonreacting turbulent jets, Ricou and Spalding [76] gave the
general, well-know expression where Ce =0.32 is the entrainment coefficient. In reacting
jets, the entrainment behavior is less straightforward due to heat release and buoyancy.
Important quantitative measurements and correlations of the entrainment rate based on this
equation induced by turbulent gas diffusion flames are further investigated by [77-89]. For
example, the constant Ce has been corrected for the cases of reacting jets by works [78, 79].
In order to get a general expression of the entrainment ratio for reacting jets, efforts have
been made by works [82, 84, 85]. The influences of the heat release and/or the buoyancy
force on the jet entrainment rate were investigated as well [78, 79].
2.5 NO emissions
The HiTAC flame has also proved to have many features that are superior to conventional
turbulent diffusion flames. These major features include a larger flame volume, a more
uniform temperature filed and a less luminous, sometimes invisible flame. NOx emissions
are also kept at a very low level. This is because during the HiTAC condition, the large
17
quantities of recirculation of combustion products are entrained into the fresh reactants
before combustion thus higher peak temperature is lack. As a result, thermal NOx is
suppressed and much of NO maybe form mainly by mechanisms that are insignificant in
most conventional combustors. A more details examination of the NO formation process in
HiTAC conditions is needed in order to present a clear background to the continuously
developing range of Nitrogen Oxides (NOx) control techniques and equipping HiTAC in
real industrial furnaces.
Although full chemical equilibriums model of NOx emission can offer more detail
information, simplified chemistry models of NOx formation and destruction are more
reasonably and realistic for predictions of an industrial-scale combustion chamber because
of cost-efficient at this stage. A number of works for predicting the NO formation in
furnaces equipped with HiTAC burners have been shown [68, 70, 71]. In these works, the
existing NO models were examined. The generally conclusion can be drawn as the existing
NOx models are able to predict the trend of the NOx emissions, the thermal NOx formation
is predominant and the NOx reburning mechanism is of little importance in which low
hydrocarbon concentration fuel is used[70]. Additionally, NO predictions are sensitive to
the form of PDF used to take into account the effect of turbulence-chemistry interactions
[7], and the use of the partial equilibrium or equilibrium O radical approaches to determine
the O radical concentration has a strong influence on the NO predictions [2]. However, the
predictions of NO emission are lower than the measured values [2, 65 66, 91, 92].
Therefore, present NOx models might require improvement to describe properly the NOx
formation under ‘low-temperature’ condition [2, 65 66, 91, 92].
The N2O-intermediate mechanism is important in fuel-lean or low temperature condition.
The N2O-intermediate NOx mechanism was proposed first by Malte and Pratt [93] for NO
formation from molecular nitrogen (N2) via nitrous oxide (N2O). This model was further
extended by some researchers [94-96]. The study [94] points out that under favourable
circumstances, this mechanism may contribute to as much as 90% of the NOx formed in
combustion process. The relevance of NOx formation in gas-phase from N2O has been
observed indirectly, and theoretically speculated for a number of combustion systems by a
number of different researchers [93, 96-99]. However, there is no any mathematical model
available of reduction mechanism for the calculation of NO formation via N2O-intermediate
18
as argued in [99] that the lack of a proper model for the nitrous oxide mechanism has not
allowed an exact calculation of NO concentration for HiTAC.
19
3 Methodology
3.1 Experimental faculties
3.1.1 Single Jet flame Experimental facilities
The combustion of a propane gas jet in the laboratory furnace was studied. Both cross-flow
and co-flow of fuel and air nozzles were considered.
The schematic of the combustion chamber are shown in Figure 3.1. The combustion air for
the first test case is preheated by an electrical heater, and diluted by nitrogen. The fuel and
air is injected with a cross-flow arrangement as shown in Figure 3.1 (a). The combustion air
for the second test case is preheated and diluted by a flue gas generator. In this case, the
composition of the oxidizer is close to what can be found in a real industrial furnace. The
fuel jet was injected in a co-flow arrangement to the main flow of the hot flue gases (Figure
3.1(b).
The fuel studied in this work is liquefied Petroleum Gas-LPG. The components (vol%) are:
CH4=0.02, C2H6=0.95, C3H8 =98.35, C4H10 =0.67.
The ranges for the parameters studied are as follows:
1. Flow rate of the fuel was varied from 4.33×10-6 kg/s (0.13 nl/min) to 1.73×10-5 kg/s
(0.53 nl/min),
2. Flow rate of the oxidizer was from 1.1kg/s to 1.3 kg/s,
3. Fuel preheat temperature was varied in the range between 388 K and 1073 K,
4. Air preheat temperature was in the range from 1041 K to 1573 K,
5. Oxygen concentration in the preheated air was varied from 2% to 23.2% (mass%).
20
Data logger for
temperatures
Upper probe
Yellow color
Flame zone
Temperature
measurements
Temperature
measurements
Blue color
Flame zone
Lower probe
Lift-off
Distance
[mm]
Fuel injection
Nozzle
(0,5 - 0,9 mm)
Thermocouple
K - type
Thermocouple
S - type
Flue gas analysis
(NOx, CO, CO2, O2)
Combustion gases
Propane fuel
(0,2 - 0,4 nlitres/min)
Water
cooling
0,2
0,28
0,75
0,3
Fuel
Diluted and
Preheated Air
(a)
Diluted and
Preheated
Air
Fuel
(b)
Figure 3.1 Schematic of the combustion chamber for single fuel jet test furnace
(a) Cross flow (b) Coflow
21
3.1.2 Classification of HiTAC burners
Combustion with high-temperature preheated air diluted by hot-combustion-production has
become increasingly attractive in industrial furnaces over recent years. This technology,
when applied together with a modern regenerative system, offers significantly increased
energy efficiency, very low CO, CO2 and NOx emissions and high quality of the product at
increased production rate.
For its industrial application, fuel nozzles and combustion air nozzles are arranged on the
burner at a certain distance from each other, fuel and high temperature air are injected
directly into the furnace at a high velocity. Because of the entrainment of these injection
moments, the in-furnace gas in the zone near the burner is thoroughly mixed and its partial
pressure of oxygen is lowered. The combustion stability of fuel directly injection into this
zone of oxygen at low partial pressure is possible if the air preheated temperature exceeds
the auto ignition temperature of the fuel. In the industrial furnace, the combustion air can be
obtained a temperature of 800-1350oC, and only lowers 50oC as exhaust flue gas by
recovered from exhaust flue gas improved through the very effective recovery of waste heat
by a high performance heat exchanger, for example, a regenerative heat exchanger switched
in the high cycle, which can recover as much as 90% of the waste heat. Thus, a large energy
saving is achieved. Meantime, the temperature of fuel and air are raised well above the
auto-ignition temperature of the most of gas fuel, which means that the conditions for flame
stabilization are very favourable.
It would be very helpful to classify the application of high- temperature air combustion
(HiTAC) technology with modern regenerator burner before the study is begun.
Due to the design of the HiTAC burners to control the mixing of the fuel and air jets with
the furnace gases, the chemical reaction rate in these burners is lower than in conventional
combustion. The typical characteristics of the HTAC technology, is its capacity to generate
larger flame volumes than conventional combustion, which results in an increased heat
transfer. Additionally, the more homogeneous reaction associated with HiTAC technology
also implies that the heat release in the reaction zone is more widespread, leading to a more
even and moderate temperature rise. Consequently, the emission of NOx can be kept low.
22
Results from previously published studies of HiTAC applications in industrial furnaces,
demonstrate that high-temperature air can be obtained using a regenerative heat exchanger.
In this heat exchanger, which can take the form of ceramic balls or honeycomb, the heat is
periodically stored and withdrawn from the heat storage material,. Hot furnace gas and cold
combustion air flows alternatively in the regenerator in contact with the regenerator
material. The furnace gas transmits heat to the storage material, during the heating period.
The combustion air absorbs heat from the storage material, cooling down the regenerator
during the cooling period.
To maintain a continuous operation at least two regenerators are required. Continuous
operation is achieved in the model regenerator by switching periodically from hot, furnace
gas to cold air in the regenerator with a short switching interval. It was found that a shorter
switching time results in a more efficient waste-heat recovery rate and a more uniform and
high temperature [52]. When more than one pair of burners are used, different firing
configurations can be used to obtain a better heating performance, that is desired in special
applications [5]. The switching interval used, can be varied from 4-5 seconds [52] to 60
seconds, and correspond to known applications of HiTAC, that are referred to as Highcycle Regenerative System or HRS.
The two main HRS solutions currently in use feature either one or two-flame burner
systems. A one-flame HRS is characterized by a single flame created by one fuel nozzle
surrounded by air inlets and flue gas outlets [16, 52]. The scheme is shown in Figure 3.2.
The single flame develops along the axis of the fuel-jet nozzle during the cooling and
heating periods. Fuel is supplied continuously through the same nozzle. In this way a single
flame can be formed with a permanent position. This position remains almost unchanged
between heating and cooling periods, as the regenerators are located around the fuel jet
nozzle.
23
Fresh Air
F.G.
F.
Flue gas
F.G.
H.A.
Figure 3.2 Scheme of one-flame HRS
In a two-flame HRS, there are two separated high-cycle regenerative burners. This scheme
is shown in Figure 3.3. The two burners are located in the walls of the furnace and work in
pairs by a set of valves that change the direction of the air and the flue-gases according to
the required switching time. Normally there are a few pairs of burners working together.
Each burner has a preheated air outlet located centrally, and two fuel nozzles located
laterally. When the hot furnace gas passes out through the regenerator of one of burners
(heating period), the fuel nozzle of this burner is closed, and the combustion air and flame
are switched off. During this time another burner operates in the combustion mode, or
cooling period of the regenerator. That means the air is preheated via cooling of the
regenerator, and both fuel nozzles of the fired burner are on and two flames can be created.
In this way, the flame can be shifted from one burner to another in accordance with the
switching time between the heating and cooling periods of the regenerator.
24
F.
H.A
F.
F.G.
Fresh Air
Flue Gas
Figure 3.3 Scheme of two-flame HRS
F
A
F
F
A
F
F
F.G.
A
F
F.G.
F.G.
(a)
F.G.
F.G.
F
A
F
A
F
F
(b)
F.G.
F
A
F
(c)
Figure 3.4 Two-flame regenerator system firing configuration for uniformity
temperature profile in the furnace (F means Fuel, A means air and F.G. means flue
gas)
(a) single-side firing configuration (b) stagger firing configuration(c) counter firing
configuration
To discharge uniformly heated product, the furnace itself must have uniform temperature
profile. The thermal profiles control available with two-flame HRS basing on composition
25
of burners with different switch time. Firing configurations that provides temperature
control across the width of the furnace through three basic flame configurations (Figure 3.4)
named as a) Single-side firing Mode, b) Counter Mode, and c), Stagger Mode were
investigated numerical.
3.1.3 The HiTAC Test Furnace
The HiTAC test used in this study is equipped with modern regenerative burner system.
The outside dimensions of the furnace body are 3.5×2.2×2.2 m. Figure 3.5 represents the
configuration of the HiTAC test furnace at KTH. Four tubes with an external diameter of
0.11m each and cooled with air have been installed horizontally in each corner of the
furnace to remove heat from the combustion chamber. On the opposite side of the furnace
to the face, there are two flue gas ducts of 0.11 m external diameter for removing hot flue
gases from the furnace. The walls of the test furnace consist of two layers: an outer steel
cover 5.0×10-3 m thick, and an inner layer of fibrous ceramic insulation 0.3 m thick. The
inner volume of the combustion chamber is 7.2 m3. There are also a number of openings in
the furnace body for measurements and observations inside the combustion chamber.
26
(a)
(b)
Figure 3.5 HiTAC test furnace
(a)HiTAC test furnace at KTH, (b)Configuration of HiTAC test furnace with one-flame
HRS.
27
22
23
11
Air-nozzle
Exhaust flue nozzle
D=220
350
Figure 3.6 Configuration of the HiTAC test furnace with inlet arrangements
20
140
D 3.7
92.5
(a)
(b)
Figure 3.7 Configuration of HiTAC test furnace with two-flame HRS
(a) Two-flame HRS
(b) Top view of HiTAC test furnace and test positions
28
C
11
Two Flame
D
8
6
4
12
9
13
10
3
2
One
Flame
1
5
A
Two Flame
7
B
Figure 3.8 Top view of HiTAC test furnace and test positions
The furnace is designed such that two different HRS can be used. The first HRS is attached
to the front of the furnace. It is a so-called one-flame system and has a thermal capacity of
200 kW.. Figure 3.5 (b) represents the computational domain of the HiTAC test furnace.
Figure 3.6 gives a basically inlet configuration of an one-flame HiTAC burner The other
system which is composed of two pairs of HRS is installed on the left and right sides of the
furnace as shown in Figure 3.7.
The 200 kW one-flame HRS with honeycomb regenerative burners was the first used in this
project. The ceramic honeycomb regenerators, through which the exaust gas and
combustion air are vented, are an intergral part of the burner body. Figure 3.6 shows the
dimensions of the burner and locations of fuel and air injection ports. There are 12
regenerators in total, working in pairs and organised into two groups separated by intervals.
80% of flue gases are vented through the burner outlets, which is sufficient to preheat the
combsution air for the desired fuel. The remainder of the exhaust gases flow out from the
furance through the chimney located on the rear wall of the furnace.
29
The second calculated case assumes that the furnace is equipped with four high-cycle
regenerative burners with capacity of 100 kW each. The burners (four burners marked A, B,
C, D) are placed on the sidewalls of the furnace as it is shown in Figure 3.7(a). Each burner
consists of one injection port for combustion air and two nozzles for fuel injection.
Combustion air and fuel are injected separately. The combustion air injection port is located
in the centre of the burner. The fuel nozzles are placed in the same plane on both sides of
the combustion air port as shown in Figure 3.7 (b). This type of regenerator allows
preheating of combustion air up to 1537 K.
The fuel used in the study was LPG with a flow rate of 7.7 Nm3/h for all the studied cases.
The composition of the fuel used was 0.22% CH4, 0.95% C2H6, 98.35% C3H8, and 0.67%
C4H10. The air flux was around 200Nm3/s
3. 1.4 Measurement program for the semi-industrial furnace
For flue gas composition in furnace measurements, a water-cooled gas-sampling probe was
inserted for taking samples at various points inside the furnace. The probe was mounted at
the top of the furnace on the traversing system and can be inserted down in the furnace to
1.2m from the ceiling. The movements of the probe tip in x, y and z coordinated can be
carried out by a computer controlled traversing system.
The probe quenches the sample and reduces the temperature down to 140 oC in a very short
time and to prevent further processing of the sample. In order to avoid wash out of some
species such as NO2, the sample was immediately cooled down to the below ambient
temperature by means of a cooler and condensate remover. However some dissociation
products, even if present in the furnace, will not be detected when using conventional
cooled sampling probes since they will recombine faster than the probe can quench the
reaction.
3 thermocouples were fitted with the furnace left sidewall (x=0.8m and y=−0.3m). These
thermocouples are applied to measure furnace wall temperature.
30
The measurements of the flue gas parameters were made at 13 fixed locations on the
horizontal plane as shown in Figure 3.8. The measurements of species were performed for
11 horizontal planes at the vertical distances with the interval 0.05m.
The uncertainties associated with the probe positions are ± 10 mm in the x-and z-directions,
and ± 2 mm in the y-direction. The average uncertainty of measured parameters in flue gas
according to the micro-GC measurement, it is composed mainly of 98% propane, 0.9%
ethane, and 0.8% butane, with a lower heating value of 93.2 MJ/Nm3.
3.2 MATHEMATICAL MODEL FOR HiTAC
3.2.1 Turbulent Combustion Model
For HiTAC, the incoming preheated air is diluted with combustion products that are
recirculated inside the furnace before the preheated and diluted air jet makes contact with
the fuel. The temperature of the air is normally higher than the autoignition temperature of
the fuel. The combustion can take place immediately after air and fuel are mixed. However,
this rate of combustion is slower due to the lower partial pressure of oxygen in the
combustion air. The zone of chemical reaction tends to be larger, which is quite different
from conventional combustion. Studies, [2] have shown that the characteristic time of
kinetics and turbulence are comparable, (Damköhle number≈1) and the two are coupled
with each other. Therefore, the combustion rate is controlled by both chemical kinetics and
by the turbulent-mixing. As a consequence, the combustion model that is based on the
assumption that mixing–is–burned is not suitable for predicting HiTAC.
To be able to accurately simulate HiTAC, when using the full reaction mechanism it is
indispensable to consider all the intermediates. However, a practical simulation of an
industrial furnace including a three dimensional flow with full reaction mechanism is far
beyond the capability of present computers. Therefore, the most realistic solution would be
to adopt a set of greatly simplified reaction mechanisms covering some intermediates.
31
Attention should be paid to the rate constant of a reaction during simulation of the HiTAC.
These constants are commonly obtained for normal combustion using ambient temperature
air. The same problem exists when a full reaction mechanism is used, even for the
elementary reactions, since the accuracy of all the associated rate constants has not been
confirmed. Therefore, the constants in models have to be optimized on the assumption that
air temperature and oxygen concentration are variable. The combustion model used for
HiTAC simulation must be a model capable of expressing precise reaction rates in a hightemperature and low oxygen partial pressure atmosphere.
In this study, the combustion model involves both chemical-kinetic and turbulent-mixing
based models. This entails evaluation of both rates locally and then taking the slower of the
two as the controlling rate according to the following:
RHiTAC = − min[ REBU , RKIN
]
(3.1)
where REBU is the corresponding turbulence-controlled rate, determined from the eddybreak-up model [101], and RKIN is the kinetic rate.
Considering the used fuel in the trial whose composition is listed above, the reactions were
given by:
CH4+1.5O2 Æ CO+2H2O
(R3.1)
C2H6+2.5O2 Æ 2CO+3H2O
(R3.2)
C3H8+3.5O2 Æ 3CO+4H2O
(R3.3)
C4H10+4.5O2 Æ 4CO+5H2O
(R3.4)
CO +0.5O2 Æ CO2
(R3.5)
The kinetically controlled reaction rate of the fuel RKIN is defined as:
R KIN = CM F T β ∏ (
allj
ρm j
Mj
v
) j e − Ea / RT (3.2)
32
Here
∏ ( x ) denotes
j
the product of all xj , vj are the species rate exponents, Mj the
molecular weights, mj is mass fraction of the species, C, β and Ea are the pre-exponential
factor, temperature exponent and activation energy for the reaction, respectively. R is the
universal gas constant, T is temperature. All kinetic rates were taken from Reference [102].
In the context of the Magnussen and Hjertager model [101], the kinetic rates are
deliberately set very high so that turbulent mixing is guaranteed to be the controlling rate.
Mathematically, these statements translate into the following equation:
R EBU = −
ρε
⎡
m
m ⎤
A min ⎢m F , O , B P ⎥
k
sO
sP ⎦
⎣
(3.3)
Where, sO = nO M O / nF M F , s P = n P M P / n F M F , A and B are empirical coefficients.
In order to decide on a suitable combustion model for HiTAC, besides the eddy-break-up
model, a PDF-mixture fraction model with chemical equilibrium [113]was also used.
The mixture is assumed to obey the ideal gas law. The viscosity, thermal conductivity and
specific heat of the mixture have been computed from the properties of individual species,
and are all functions of temperature
3.2.2 NOx models
The amount of NO is small and the time scale for NOx reactions is larger than the time
scales for the turbulent mixing process and the combustion of hydrocarbons that control the
heat-releasing reactions. Hence, therefore, it is possible to assume that the reactions
involved in the NO chemistry can be decoupled from the main combustion reaction
mechanism. In this work, four different mechanisms have been identified for the formation
and destruction of NO, i.e.., thermal NO, prompt NO, NO reburning and N2O- intermediate
mechanism.
33
3.2.2.1 Thermal NO
The formation of thermal NO is determined by the following three extended Zeldovich
mechanism[113]:
k ,k
1b
O + N 2 ←⎯1f⎯
⎯
→ NO + N ;
(R3.6)
,k
k
2f
2b
⎯⎯
→ NO + O .
N + O2 ←⎯
(R3.7)
k ,k
3f
3b
N + OH ←⎯
⎯
⎯
→ NO + H
(R3.8)
Based on the quasi-stead state assumption for N radical concentration, the net rate of NO
formation via the foregoing reaction can be determined by:
d [NO ]
=
dt T _ NO
1
k1b [NO ]
1+
k 2 f [O2 ] + k 3 f [OH ]
×
(3.4)
⎡
⎤
2k1b
(k 2b [O ][NO ]) + k 3b [H ][NO ]⎥
⎢2k1 f [O ][N 2 ] −
k 2 f [O2 ] + k 3 f [OH ]
⎢⎣
⎥⎦
Where
k i f (b ) = Ai T Bi exp(−Ci / T )
Where, T is temperature, K. The reaction constants, Ai,
Bi and Ci , were taken from Baulch et al. [103].
All species concentrations used in Eq.[3.4] and (R3.6-R3.8), except for O, H and OH
radical concentration, are from the main combustion simulation. In this work, the
concentration of H atom is set to zero. The concentrations [O] and [OH] can be calculated
from the thermodynamic equilibrium as following [104]:
[O] = 3.97 × 10
5
× exp(−31090 / T )
T
[O2 ]1 / 2
mol/m3
(3.5)
34
[OH ] =
1
−2
6.24 × 10 × T
0.057
× exp(8600 / T )
[O2 ]1 / 2 [H 2 O]1 / 2
mol/m3
(3.6)
3.2.2.2 Prompt-NO formation mechanism
For gaseous fuels, De Soete [105] proposed a roughly estimated chemical reaction rate
appropriate for the prompt-NO formation mechanism. In this rate, Missaghi et al. [106]
have included a prompt factor to extend the expression to natural gas combustion. In terms
of concentrations it reads, in g_mole/cm3s,
E
d [NO ]
M 1+b
b
= f pr C 1+b [O2 ] [N 2 ][Fuel ]T exp(− a )
dt P _ NO
RT mol/cm3s (3.7)
ρ
where M stands for the mixture molecular weight and ρ is the mixture density. The oxygen
power b is related to oxygen mole fraction in the flame,
⎧1.0[O 2] ≤ 4.1 × 10 −3
⎪
⎪− 3.95 − 0.9 ln[O2 ]
b=⎨
⎪− 0.35 − 0.1ln[O2 ]
⎪0
⎩
[O2 ] ≤ 4.1 × 10 −3
4.1 × 10 −3 ≤ [O2 ] ≤ 1.11 × 10 − 2
1.11 × 10 − 2 ≤ [O2 ] ≤ 0.03
[O2 ] ≥ 0.03
(3.8)
For C3H8 the constants C and Ea take the following values :
C=6.4×106
Ea= 49.65×103
s-1
cal/mole
(3.9)
(3.10)
The prompt-factor f pr is calculated as follows [105]:
35
f pr = 4.75 + 0.0819 N c − 23.2
1
λO
2
+ 32.0
1
λ
2
− 12.2
O2
1
λ3
(3.11)
O2
where NC is the number of carbon atoms in the hydrocarbon and λ O 2 the local
stoichiometry.
3.2.2.3 N2O-intermediate NOx
The emphasis of this study in on the nitrous oxide mechanism as following [108]:
N2 +O + M ←
⎯→ N 2 O + M
(R3.9)
N 2O + O ⎯
⎯→ N 2 + O2
(R3.10)
N 2O + H ⎯
⎯→ N 2 + OH
(R3.11)
N 2 O + OH ←
⎯→ N 2 + HO2
(R3.12)
N 2O + O ⎯
⎯→ NO + NO
(R3.13)
N 2O + H ⎯
⎯→ NO + NH
(R3.14)
The N2Oforms by reaction (R3.9) and is destroyed by reactions (R3.10) – (R3.14). Portions
of this N2O are converted to NO by reactions (R3.13) and (R3.14). The kinetics rates used
are listed in Table 3.1.
36
Table 3.1: Arrhenius kinetic coefficients used for nitrous oxide mechanism reactions
Units:cal,mol,cm3,sec
KKIN=ATbexp(-Ea/RT)
Reaction
R9
A
b
Ea
4.7e12
0
17614
Reference
CECR[109]
M body collision efficiencies:
H2=2, O2=0.4, Ch4=2,CO2=2, CO=1.5, H2O=6,other
GRI 3.0[110]
species=1.
R10
1.4e12
0
10810
GRI 3.0[110]
R11
3.87e14
0
18880
GRI 3.0[110]
R12
2.0e12
0
21060
GRI 3.0[110]
R13
2.9e13
0
23150
GRI 3.0[110]
R14
1.385e17
-0.5
33729
GRI 3.0[110]
It is noteworthy that all reactions rates depend strongly on O, OH and H radial
concentrations. Therefore, the prediction model of these radicals will be sensitive for the
prediction of NO emission.
The effect of turbulence on N2O formation and destruction shall be considered. The EddyDissipation-Conception model [2] has achieved some degree of success in HiTAC is. Here,
it is extensively [111] used in the calculation of the chemical reaction rate of N2O-route NO
model. The N2O concentration, which accounted by its production (R3.9) and its
reconversion to N2 (R3.10)-(R3.12), has been calculated with this finite rate-eddy
dissipation assumption. According to the model, the reaction rate of NO is given by:
R EBU _ NO _ i = −
⎡
mj ⎤
Aebu min ⎢mi , ⎥
k
s j ⎦⎥
⎣⎢
ρε
(3.12)
where, ρ is density of mixture, kgm-3, ε is turbulence kinetic energy dissipation rate,m-2s-3,
k is turbulence kinetic energy, m-2s-2. In the terminology of above, mi is mass fraction of N2
37
and N2O for (R3.9) and (R3.10)-(R3.14), respectively. mj is mass fraction of O for (R3.9),
(R3.10) and (R3.13). It is mass fraction of H for (R3.11) and (R3.14). For (R3.12), it is
mass fraction of OH. s j = ni M i / n j M j and n is the stoichiometric coefficient (number of
moles) and M is the molecular weight. Aebu is empirical coefficients with value 4.
Then, the reaction rate for reaction (R3.9)-(R3.14) entails evaluating chemical-kinetic and
turbulent-mixing rates and taking the slowest as controlling,
[
R N 2O _ NO _ i = − min R EBU _ NO _ i , R KIN _ i
]
kg/m3s
(3.13)
where REBU_NO_i is the corresponding turbulence-controlled rate, determined from Equ. 3.12,
and RKIN_i is the kinetics rates obtained from table 1.
Furthermore, the N2O transport equation has been solved adding only reaction rates of
(R3.9)--(R3.12) as chemical sources.
The NO concentrations which come from (R3.13) and (R3.14) was implemented to add as a
source term to the NO transport equation.
The similar N2O-intermediate model has been implemented by Tobacco [111] in order to
assess the effects of this NOx formation pathway with increasing pressure.
3.2.2.4 NO reburning mechanism
Under fuel rich conditions, NO can be reduced by CHi radicals. The reactions involved in
this global mechanism are[113]:
where,
k i = Ai T Bi exp(−Ci / T )
k1
NO + C x H y ⎯⎯→
HCN + ...
(R3.15)
k2
HCN + O2 ⎯⎯→
NO + ...
(R3.16)
k3
HCN + NO ⎯⎯→
N 2 + ...
(R3.17)
The reaction constants, Ai, Bi and Ci , are taken from
Bowman [112]. The rate of depletion for NO is given by:
38
d [NO ]
= k1 [CH ][NO ] + k 2 [CH 2 ][NO ] + k 3 [CH 3 ][NO ]
dt
(3.14)
3.2.2.5 NO source term and turbulence-chemistry interaction
The NO source term due to the formation and destruction of the thermal NO, prompt NO,
N2O route and NO reburning can be calculated as:
⎧ d [NO ]T _ NO d [NO ]P _ NO d [NO ]N 2O _ NO d [NO ]R _ NO ⎫
S NO = M NO ⎨
+
+
−
⎬
dt
dt
dt
dt
⎩
⎭
(3.15)
The interaction between turbulent flow and chemistry is taken into account through the
transport equation of NO chemical species. Thermal NO is strongly temperature dependent
and so it may be argued that the variation of temperature has a great impact on NO
production that species concentration fluctuation. The mean NO concentration formatted
via thermal NO is computed on the basis of the single-variable probability density function
(pdf) model. The Beta-function PDF is used in this study since it is widely used in turbulent
combustion simulations.
The mean NO concentration is calculated by solving its transport equation based on the
flow field as combustion solution from the combustion simulations.
ρui
∂YNO
∂Y
∂
( ρ D NO ) + S NO
=
∂xi
∂xi
∂xi
(3.16)
where, ρ is mean density, u is mean velocity in the i direction, and YNO is the mean mass
fraction of NO. The mean source term, S NO , is determined from above NO mechanisms.
39
3.2.3 Other Models Used
The flows in the industrial furnace are turbulent. It follows that the performance predictions
of combustors depend very much on the turbulence model adopted. With the requirement of
a million nodes per cubic millimeter, it is clear that the application of DNS (direct
numerical simulation) for engineering purposes is not practical for application here. LES
(Large eddy simulation) seems to have a bright future, but more research on LES is
required. Dong [49] has carried out a simulation of a single fuel jet flow in hightemperature diluted air combustion. It was found that advanced turbulent models, such as
LES and RSM, gave small differences in the near field when predicting the flow. However,
the empirical constants, for example Cs in the LES model, have a significant influence on
the predictions. This implies that the empirical constants in traditional models must be
adjustable to be able to obtain the best performance for HiTAC simulations.
The k-ε model remains the obvious starting point, especially for diffusion flames for
engineering calculations. It has been verified that it is robust and efficient for most
engineering calculation purposes in the range it can be used. For flows in HiTAC, due to
the larger reaction zone and the similarities to a ‘well-stirred reactor’, the assumption of
non isotropy of the turbulence is weak. Furthermore, the buoyancy in the furnace is
relatively small compared to conventional combustion due to both the preheated air and
lack of a distinct flame zone. Consequently the k-ε model is chosen to be used in this paper.
Additionally, radiation was handled using the discrete transfer method [113]. Radiation
properties of flue gas were assumed to be of the grey body type and were temperature- or
concentration-dependent, or both. Available absorption coefficient model called the
weighted-sum-of-grey-gases model or WSGG was used to determine the absorption
coefficient.
3.3
DEVELOPED
CONCEPTS
FOR
DESCRIBING
HiTAC
CHARACTERISTICS
40
HiTAC has many characteristics that are completely different from conventional
combustion. For example, the HiTAC flame is less visible than the flame from conventional
combustion where there is a high concentration of oxygen by volume (more than or equal to
21%, i.e air). Therefore it is generally accepted, that flame length is not a suitable parameter
for characterizing flame size for HiTAC. Instead, it is necessary to characterize the flame
shape and size using a comprehensive numerical simulation.
To describe the flame under HiTAC conditions, the oxidation mixture ratio is used in this
work. The oxidation mixture ratio allows the combustion progress to be estimated and to be
calculated as the mass fraction of oxygen divided by the mass fraction of oxygen plus the
amount of oxygen needed to achieve complete combustion at any point in the combustion
chamber, as follows:
mO
mO + ∑ s c m F , c
Ro =
c
where
s = nO M O / nF M F
(3.17)
. This ratio has a value of Ro =1 at the air inlet or when the
combustion is completed, and a value of Ro = 0 at the fuel inlet.
To determine flame border through this parameter, the critical Rcr must be given. Eq. 3.17
can be transferred as following:
Ro =
1
1 + ∑ sc
c
m F ,c
(3.18)
mO
For HiTAC technology, the preheated air temperature is above the fuel’s autoignition
temperature, therefore, the fuel flammability limit plays the main role for combustion
stability. Thus, the criteria Rcr can be determined through the flue flammability limit φ in
Eq. 3.19.
Rcr =
1
1 + ∑ scφc
(3.19)
c
41
The lean flammability limits for different fuel species has been used to indicate the outside
border of the flame, and the rich flammability limits of them are used to given the insider
border of the flame. According to Glassman [114], the flammability lean limit of propane in
air and oxygen is 2 % (fuel volume percent), and this value of CO in air is 12%, then RO=
0.99 is assumed to indicate a flame border. Thus the flame volume can be approximate
defined when
0 ≤ RO ≤ 0.99
(3.20)
In order to describe the overall radiation field of a flame, the radiant fraction (frad) is used. It
is defined as the ratio of net radiative heat loss from the flame (Qrad) to the total heat
released during combustion (QF) as follows:
f rad =
Qrad
QF
(3.21)
To evaluate gas temperature field uniformity inside the furnace, a furnace-gas-temperature
uniformity-ratio, Rtu, is defined as follows:
⎛ (T − T ) ⎞
⎟⎟
= ∑ ⎜⎜ i
Rtu
⎝ T
⎠
2
(3.22)
where Ti [K] is the temperature of calculated cell number and T [K] is the average
temperature in the furnace. When Rtu =0, there is no gas temperature gradient inside the
furnace.
For HiTAC flames, to be able to characterize the flame volume in relation to the volume of
the combustion chamber, a dimensionless coefficient called Furnace Flame Occupation
Coefficient (FOC), RFOC is defined as the ratio between the flame and furnace volume:
RFOC =
Vf
VF
(3.23)
42
where, Vf [m3] is flame volume calculated according to the relationship in equation 5, and
VF [m3] is the furnace volume calculated from the geometrical dimensions of the furnace.
Combustion intensity is also a very important parameter for designing the furnace and the
burner. To evaluate quantitatively the chemical reaction intensity in the furnace and
especially in the chemical reaction zone (flame), two parameters are used. One is Flame
Heat Release (FHR), which is defined as the ratio of heat released inside the flame zone
(Qf ) to the flame volume (Vf):
q FHR =
Qf
Vf
(3.25)
where Qf is obtained as follows:
Qf = ∫
Vf
∑ q dV
c
(3.26)
Here qc (kW/m3) is local heat release of different fuel species.
Another parameter used in this work is the ratio between the heat released by the flame
zone (Qf) and total heat released inside the combustion chamber, (QF). It is named Flame
Heat Occupation Coefficient (FHOC), and defined as follows:
RFHOC =
Qf
QF
(3.27)
where, QF is calculated according to:
QF = ∫
VF
∑ q dV
c
(3.28)
Furthermore the entrainment ratio of the nozzle is a good parameter for the description of
the internal recirculation of the flue gas which plays an important role in the HRS systems.
43
Due to the interaction between fuel and air nozzles for one-flame HRS, the entrainment of a
single nozzle is not a suitable parameter for the characterization of the internal flue gas
recirculation. Therefore, the entrainment ratio must include the interactions between the
nozzles. In order to describe the interactions between the fuel and air nozzles, flame
entrainment ratio is more efficient. The flame entrainment ratio (Rfe) is defined as the
following:
R fe =
mf
m0
(3.29)
Here, m0 (kg/s) and mf (kg/s) represent the initial total mass flow rates and mass flow rates
through the cross section of the flame respectively.
44
4 Cross-flow single jet flame study
Combustion of a single fuel jet of propane in a laboratory furnace was studied numerically.
Furnace’s combustion chamber has dimension of 0.16 × 0.2 × 0.28 m. Computational domain
and meshes of HiTAC test furnace can be seen in figure 2. Fuel nozzle is placed on the wall
in a cross-flow to the main flow of oxygen deficient and preheated air. Flow rates of
preheated air and fuel were kept constant and equal to 3.333 × 10-3 m3s-1 and 5.0 × 10-6 m3s-1,
respectively. Fuel preheat temperature was in the range from 288 K to 873 K and air
preheat temperature from 1041 K to 1273 K. Oxygen concentration in the preheated air
varied from 2% (mass) to 18%( mass). Reynolds numbers were 1162 and 3800 for the
preheated air and fuel gas inlet respectively. The blowing ratio as defined in the paper by
Hasselbrink et. al. [117] is 160. At this stage of the study the thermal decomposition of
propane at elevated temperature before outlet from the gas nozzle was not considered.
Adiabatic wall boundary conditions are assumed in the heat transfer model (i.e. zero heat
flux).
4.1 Study of mathematical modelling
The computational domain, representing laboratory test furnace, was divided into 23148
computational cells using unstructured meshes with embedded refinement to refine the fuel
nozzle. Figure 4.1 displays the spatial discretization of the computation domain.
45
Local refined mesh near
the fuel jet
Fig.4.1 3D-computational domain and meshes of HTAC test furnace
Gas temperature profiles and flame zones are shown in Figures 4.2 and 4.3. Figure 4.2
shows that the maximum gas temperature predicted with use of PDF combustion model is
equal to 1952 K but it is equal to 1846 K if the EBU combustion model is employed. Also
areas of maximum gas temperature within the flame are predicted larger when the PDF
model was used. The lower peak temperature seems more realistic when oxygen
concentration is only 10%. Moreover, one can see from Figure 8 that the flame volume as
defined by means of the oxidation mixture ratio was predicted smaller when the PDF model
is used. Also it can be noticed that the flame size and shape predicted by the EBU model is
visually similar to the flame presented at photography (Figure 4.3c) taken. Based on these
facts, it was assumed that the HiTAC flame predictions with the EBU model are more
realistic. Thus, the EBU model was further studied as more applicable to HiTAC flame
modelling.
46
(a) pdf
(b) ebu
Fig.4.2 Predicted temperature profiles for 10% oxygen in the air preheated upto 1041K and
at fuel inlet temperature equal to 473 K.
(a) PDF model; (b)EBU model
(a)
(b)
(c)
(a) PDF model; (b)EBU model, (c) flame
photograph
Fig.4.3 Predicted distributions of oxidation mixture ratio for 10% oxygen in the air
preheated upto 1041K and at fuel inlet temperature equal to 473 K.
Two empirical coefficients, A and B, are involved in the rate of reaction equation (3.3) in
the EBU model. It has often proved to be necessary to adjust the A and B to obtain the good
performance for a particular application. B is taken into account to inhibit reaction where
the temperature is low. However in HiTAC combustion, the air temperature is higher than
47
fuel self-ignition temperature, the influence of empirical B is insignificant, so B is set as
normal constant 0.5. Concerning of A, it is known that the fuel consumption decreases
when A decreases according to reaction equation as the reaction rate will also decreases.
Many experimental works proved that HiTAC combustion rate is slower than traditional
combustion. This indicates that A should be less than the nominal value 4. Thus, simulation
for the following three different values of A that is 4.0, 2.0 and 1.0 was performed.
Figure 4.4 shows that predicted temperature field is influenced by value of the A coefficient.
Lower value of the A coefficient gives smaller temperature gradients in the flame. Clearly it
is seen that area of flame with the highest temperature is reduced when the lower value of
the A coefficient is applied.
(a
(b)
(c)
Fig. 4.4 Predicted gas temperature profiles for 10% oxygen in the air preheated upto 1041K
and at fuel inlet temperature equal to 288 K.
(a) A=4.0, (b) A=2.0, (c) A=1.0
Influence of the A value on the oxidation mixture ratio is shown in Figure 4.5. Numerical
predictions demonstrate that the flame volume increases with the reduction in A value. It
shows also more uniform distribution of the fuel inside the flame when smaller values of
the A are used. It confirms also that the combustion rate is slow according to equation (3.3).
48
(a)
(b)
(c)
Fig. 4.5 Predicted distributions of the oxidation mixture ratio for 10% oxygen in the air
preheated upto 1041 K and at fuel inlet temperature equal to 288 K.
(a) A = 4.0, (b) A = 2.0, (c) A = 1.0
4.2 Flame Volume Ratio, Rf
Consequently the EBU model was used to study influence of the fuel preheat temperature
on the combustion.
Figure 4.6 shows that the flame volume ratio Rf depends very much on oxygen
concentration in the preheated combustion air, temperature of the combustion air and the
fuel inlet temperature. Flame volume ratio always increases when the oxygen concentration
in the preheated air is reduced. This is clearer if oxygen concentration in the preheated air is
below 5%.
49
250
225
200
175
Rf
150
125
100
75
50
25
0
0
2
4
6
8
10
12
Oxygen
Concentration
(mass%)
Ta=1041,TF=288
Ta=1041,TF=573
Ta=1173,TF=573
Ta=1273,TF=288
Ta=1273,TF=573
14
16
18
20
Ta=1041,TF=473
Ta=1173,TF=288
Ta=1173,TF=873
Ta=1273,TF=473
Ta=1273,TF=873
Fig.4.6 Rf versus oxygen concentration for various air and fuel temperatures (K)
Elevating fuel temperature leads to reduction of the flame volume ratio Rf at constant
oxygen concentration. For example when oxygen concentration is equal to 5% and
combustion air temperature is equal to 1273 K, the Rf in case of the fuel inlet temperature
equal to 288K is 6 times larger than in case of the fuel inlet temperature equal to 873 K. At
10% of oxygen concentration, this relation is 5.3, and at 18% of oxygen concentration, it is
equal to 8.7. Reason for this is decrease of the fuel density with increase temperature thus
increase of fuel inlet velocity at constant fuel fluxes. Changes of density and fuel inlet
velocity are proportional to the fuel temperature. Increasing of initial velocity of the fuel jet
improves the mixing between fuel and preheated air what results in flame volume
decreasing. Improvement of mixing is approximately proportional to the fuel inlet velocity.
Therefore, reduction of Rf is also proportional to the fuel inlet temperature.
Figure 4.7 also shows that combustion air temperature has much less significant influence
on the flame volume at constant oxygen concentration and fuel temperature. For the
investigated temperature range (1041 K–1273 K) of the preheated air flame volume was
found almost constant at fixed oxygen concentration and fuel inlet temperature.
50
120
100
Rf
80
60
40
20
0
1030
1080
1130
1180
Preheated air temperature (K)
Oxyg.=5%,TF=288
Oxyg.=10%,TF=288
Oxyg.=10%,TF=473
Oxyg.=5%,TF=573
1230
1280
Oxyg.=10%,TF=573
Oxyg.=18%,TF=573
Oxyg.=5%,TF=873
Fig.4.7 Rf versus preheated air temperature for oxygen concentration and fuel temperatures
(K)
Generally it can be concluded that the largest flames are obtained at the lowest oxygen
concentration in the combustion air and at lowest fuel and preheated air temperatures.
Assuming the combustion air temperature equal to 1273 K and the fuel inlet temperature
equal to 473K the HiTAC flame at oxygen concentration equal to 5 % is more then eight
times larger than the flame at oxygen concentration equal to 18%. Flame volume in case of
2% oxygen concentration and 288 K fuel temperature is 223 times larger than that in case
of 18% oxygen concentration and 873 K fuel temperature. The performed numerical
experiments show that HiTAC process is spread over much larger volume then
conventional turbulent diffusion flame. It confirms that the HiTAC is a large volume
combustion with reduced combustion rate.
From the above, it can be also stated that three primary factors determine flame volume.
These are as follows: 1 - initial momentum of the fuel jet; 2 - oxygen concentration in
combustion air; 3 - ratio of fuel inlet density to ambient gas density, ρ F ρ 0 .
51
4.3 Mean Residence Time and Flame Peak Temperature
Figure 4.8 shows that the mean residence time, τ R increases with decrease of oxygen
concentration. This dependence is almost consistent with change of flame volume ratio, Rf
as showed in Figure 4.6. It confirms primary importance of oxygen concentration and fuel
temperature on mean residence time. Fuel temperature influences the value of τ R , but not
as much as the molar fraction of oxygen. For example, for air temperature equal to 1273K
and fuel temperature equal to 473K, the mean residence time at the case of 5% oxygen
concentration is 6.7 times larger then at the case of 18%. When oxygen concentration is 5%
and combustion air temperature is 1273 K, the mean residence time in case of fuel
temperature equal to 288K is 2.9 times larger than that in case of fuel temperature equal to
873 K. These values are somewhat less than changes of Rf at the same conditions as the
flame density compensates. It can be concluded that the mean residence time increases with
reduction of oxygen concentration as well as with decrease of fuel inlet temperature.
0,14
0,12
0,1
0,08
0,06
0,04
0,02
0
0
2
4
6
8
10
12
14
16
18
20
Oxygen Concentration (mass%)
Ta=1041,TF=288
Ta=1173,TF=573
Ta=1273,TF=473
Ta=1273,TF=873
Ta=1173,TF=288
Ta=1173,TF=873
Ta=1273,TF=573
Fig.4.8 Mean residence time (s) versus oxygen concentration and combustion air and fuel
temperature (K)
52
The mean residence time also slightly increases with decrease of the preheated air
temperature. For example, assuming oxygen content equal to 5% and fuel temperature
equal to 473K, the changes of local residence time is 1.12 times for combustion air
temperature from 1273K down to 1041K. This is because flame density increases with
reduction of combustion air temperature.
It is obvious as showed in Figure 4.9 that for any preheated air and fuel temperature, flame
peak temperature approaches maximum for the highest oxygen concentration, and
minimum for lowest oxygen concentration. At any preheated air temperature the peak
temperatures fall slightly for constant oxygen concentration and at reduced fuel inlet
temperature.
2500
Tmax , K
2300
2100
1900
1700
1500
0
2
4
6
8
10
12
14
16
18
20
Oxygen Concentration (mass%)
Ta=1173,TF=288
Ta=1173,TF=873
Ta=1273,TF=473
Ta=1173,TF=573
Ta=1273,TF=288
Ta=1273,TF=873
Fig.4.9 Peak flame temperature (Tmax) versus oxygen concentration for various combustion
air and fuel temperature(K)
4.4 Gas Temperature Uniformity Ratio, Rtu
53
Gas temperature uniformity ratio, Rtu versus oxygen concentration for various air and fuel
temperatures is shown in Fig.4.10. The Rtu decreases with reduction of oxygen
concentration, which indicates the more uniform temperature field at lower oxygen content.
Rtu in the case of oxygen concentration equal to 5 % and at fuel inlet temperature equal to
288 K is two times lower than in the case of oxygen concentration equal to 18%. Increase
of temperature field uniformity results from the reduction of flame peak temperature and
from increasing of the flame volume Rf at lower oxygen concentrations.
40
36
Rtu
32
28
24
20
16
12
0
2
4
6
8
10
12
14
Oxygen
Concentration
(mass%)
Ta=1041,TF=288
Ta=1041,TF=578
Ta=1273,TF=473
Ta=1273,TF=873
16
18
20
Ta=1041,TF=473
Ta=1273,TF=288
Ta=1273,TF=573
Ta=1173,TF=288
Fig.4.10 Temperature uniformity ratio, Rtu versus oxygen concentration for various air and
fuel temperatures(K)
Influence of the fuel inlet temperature on the temperature uniformity ratio Rtu is difficult to
detect. In general it was noticed that at lower oxygen concentration the influence of fuel
inlet temperature is more distinct. At high oxygen levels, for example at 18% oxygen
concentration, the magnitudes of Rtu at different fuel and air temperatures are very close to
each other. It is apparent that the average temperature T and the temperature of ith
computational cell, Ti in test furnace increase when the fuel temperature increases.
54
4.5 Conclusions
1) Concept of the flame volume ratio, Rf introduced in this work was used to describe the
flame volume changes during High Temperature Air Combustion.
2) It was showed that High Temperature Air Combustion is spread over much larger
volume then conventional combustion.
3) Flame volume was found almost constant for the investigated temperature range (1041
K–1273 K) of the preheated air at fixed oxygen concentration and fuel inlet temperature
4) Concept of gas temperature uniformity ratio Rtu introduced in this work was used to
characterise the temperature uniformity changes inside the flame.
5) Mean residence time of fuel gas parcels inside the flame volume increases with
reduction of oxygen concentration as well as with decrease of fuel inlet temperature and
slightly increases with decrease of the preheated air temperature.
6) Increase of the fuel inlet temperature results in smaller flame, shorter mean residence
time, smaller temperature peaks and lower formation of NO.
55
5 Flame Length and volume study in co-flow single jet
flame
The studies of global flame features showed the flame color to change from yellow to blue
to bluish-green to green, and in some cases hybrid and purple color flame was also
observed. Under certain conditions flameless or colorless oxidation of the fuel has also
been demonstrated. It is generally accepted that HiTAC flame length is not enough for
characterizing flame properties of high temperature air combustion because of its poor
visibility and big chemical reaction volume. Quantitative numerical studies were performed
in work, where the HiTAC flame volume was defined to demonstrate the HiTAC flame
properties and its changes depending on process parameters were studied.
It is known that the diffusion flame length can be generalized as the function of fuel jet
momentum and furnace temperature by means of the Froude number, Frf [116]. However,
there is little knowledge available about the influences of high temperature combustion air
and low oxygen concentration on the flame length. It has proved that the temperature
profile in combustion chamber is more uniform in the HiTAC condition although the
combustion air is preheated to very high temperature. This implies that the buoyant force in
furnace is smaller. Consequently, the Frf is larger than that in the case of conventional
combustion with the same initial jet momentum. Therefore, the constants of the flame Frf
used to assess momentum– or buoyancy–control have to be optimized on the assumption
that air temperature and oxygen concentration are changeable. The present study intends to
fill this gap. The “flame” is determined in this paper according to the flammability limits of
the combustible gases in the combustion chamber during the HiTAC condition. The HiTAC
flame length and volume in co-axial flow with high-temperature and low-oxygen
concentration oxidizer were numerically studied. The “flame” length and volume describe
the physical properties of the chemical reaction zone in the combustion chamber.
Consequently, they are referred to as a “chemical” flame length or height [117] and volume,
to distinguish them from luminous flame lengths and volumes determined based on visual
observations. The flame length and volume in the latter part of this work mean chemical
flame length and volume. The HiTAC “flame” length and volume in coaxial flow with a
high temperature and low-oxygen-concentration oxidizer were studied. The studied
56
parameters included oxygen concentration in oxidizer, fuel and oxidizer temperatures,
firing rate, and diameter of fuel nozzle. The HiTAC flame volume was also expected to be
synthesized as a function of these various parameters.
5.1 Flame appearance
Gas jet flame co-axial with flow of flue gases also at various oxygen molar fractions in the
oxidizer is shown in Figure 5.1. Combustion of fuel gas in hot and oxygen deficient flue
gases appeared to be very stable and complete even at very low oxygen concentration.
(a) 10% O2
(b) 12.8% O2
(c) 13.3% O2
(d) 16.8% O2
(e) 21% O2
Figure 5.1 Flame appearance for the LPG jet co-flowing vertically with hot and oxygen
depleted flue gas. Fuel nozzle diameter 0.5 mm. Fuel jet velocity equals to 25.5 m/s. Flue
gas velocity equals to 0.98 m/s [40].
This experiment also shows that reduced oxygen concentration increases the flame size,
lift-off distance and decreases luminosity and visibility. Flame becomes first bluish and
next non-visible. It is important to stress that the substantial differences were noticed
particularly in phenomena of heat transfer by radiation. Flame formed by burning fuel gas
jet in low oxygen content flue gases has higher emissivity what results also in much higher
total radiative heat flux.
5.2 Effect of oxygen concentration
57
Gas jet flame co-axial with flow of flue gases also at various oxygen molar fractions in the
oxidizer is shown in Figure 5.2. In this figure, the flame pictures were taken by a 60×60
mm Hasselblad camera through the window opened in the front wall of the combustion
chamber. This experiment also showed that reduced oxygen concentration increases the
flame size, lift-off distance and decreases luminosity and visibility. Flame becomes first
bluish and next non-visible. Experimental flame lengths were obtained by visual
determination as shown in Figure 5.2. For the cases of [O2] equal to 16.8, 12.8 and 11.0,
they were 0.25 m, 0.30 m and 0.33 m, respectively.
(b) O2=12.8 %
Length
Length
Length
(a) O2=11.0%
(c) O2=16.8%
Figure 5.2 Flame apparence versus oxygen concentration
To = 1173 K, dF = 5.E-4 m, QF = 0.01 g/s, TF =299K
Figure 5.3 shows the flame shape and size defined according to Eq. 1. Consequently
calculated flame length and volume of LPG flames were showed in Figure 19. Comparison
with the experimental results, predicted values are larger. This is because of the difference
of definitions for both predicted and measured. However, the changing trend was similar
and the agreement can be accepted.
58
(a) 5% O2
(b) 10% O2
(c) 12.8% O2
(d) 15.8% O2
(e) 21% O2
RO
Figure 5.3 Predicted flame shape and size for different oxygen concentration
(To = 1173 K, dF = 5.E-4 m, QF = 0.01 g/s, TF =299K)
From Figure 5.4 one can further see that drop in oxygen concentration in the oxidizer
increases the flame length as well as the flame volume. The influence of oxygen
concentration on flame volume is more visible. When the oxygen concentration varies from
21% down to 5%, the flame volume increases 13.2 times, but the flame is barely 2.4 times
longer. It is obvious because the volume is proportional to the cubic of the length if the
shape is similar. The rising trend in flame volume is clear if the oxygen concentration is
below 10%. The possibly explanation is that although the chemical reaction between fuel
and oxidizer is very fast due to the high oxidizer temperature, because of deficiency of
oxygen concentration in the combustion regime, the combustion can not complete in small
volume and the flame needs to disperse in order to complete oxidizing reaction of
hydrocarbon until the combustibles are consumed. Therefore, this combustion phenomenon
is very similar as Well Stirred Reactor (WSR). These behaviours offer further quantitative
evidence for ‘volumetric combustion’ instead of ‘flame front combustion’ for HiTAC.
59
1,0
1,2E-03
0,8
Experimental Length
9,0E-04
Flame Volume
0,6
y = 0,1072x -0,5674
0,4
6,0E-04
3,0E-04
0,2
Flame Volume, m3
Flame Length, m
Flame Length
y = 3E-06x -1,7052
0,0
0,0E+00
0
0,05
0,1
0,15
0,2
0,25
O2 content in oxidizer(vol%)
Figure 5.4 Length and volume of HiTAC flame versus oxygen concentration.
(To =1173 K. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
Figure 5.5 gives the temperature distribution on the central vertical cross section versus the
oxygen concentration in the oxidizer. It is obvious that the flame peak temperature
approaches maximum for the highest oxygen concentration, and minimum for lowest
oxygen concentration. This implies that temperature distribution tends to be more uniform
as the oxygen concentration decreases. The heat release in such conditions is also
distributed, leading to a dispersed and moderate temperature rise. It can be concluded that
the lower the oxygen concentration in the oxidizer, the bigger the reaction zone. To see the
relationship between oxygen concentration and the maximum temperature in the furnace,
these values were plotted in Figure 5.6. The peak temperature decreases linearly with the
reduction of oxygen concentration.
60
(a) 5% Oxy.
(b) 10% Oxy.
(c) 12.8% Oxy. (d) 15.8% Oxy.
(e) 21% Oxy.
T, K
Figure 5.5 Predicted temperature profiles for different oxygen concentration.
(To =1173 K. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
Changes of the peak temperature with change of oxidizer temperature were shown in Figure
5.7. However, these changes are weak in comparison to its changes with oxygen content in
the oxidizer.
61
Flame maximum temperature, K
3000
Tair=1173K
2500
2000
1500
1000
0
0,05
0,1
0,15
0,2
Oxygen concentration, vol%
0,25
Figure 5.6 Flame maximum temperatures versus oxygen concentration.
( To =1173K. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
Flame maximum
temperature, K
3000
Oxygen=12.8(vol%)
2500
2000
y = 0,5892x + 1312,8
1500
1000
1000
1200
1400
1600
Preheated air temperature, K
1800
Figure 5.7 Flame maximum temperatures versus air temperature.
( [O2]=12.8%. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
62
These results were consistent with experimental results presented in work [15, 17, 100] and
theoretical analysis results in works [2]. Although in these cases, fuel jet in a cross-flow of
oxidizer was investigated.
Maximum temperature of flame, (Tf,max) of common fuel during preheated and diluted air
combustion can be estimated using the following relationship:
T f ,max = To +
Tad
21 /[O2 ] + 1
(5.1)
5.3 Effect of oxidizer temperature
Figure 5.8 shows influences of oxidizer temperature on the HiTAC flame length and
volume at 12.8% of oxygen concentration. The Reynolds number of the fuel jet was 3316.
The results show that the flame length is only slightly higher for higher temperature of
oxidizer. This is because, as the inlet temperature is increased, mass flux by turbulent
mixing is reduced due to decrease of air density. Meanwhile, the small expansion of the
burnt gas does not suppress the turbulent mixing. These effects are considered to
compensate each other. The same results have been experimentally observed by Fujimori
and co-workers [41].
The results also show that increase in preheated temperature of oxidizer increases the flame
volume by about 40%. This is because the low air density at high temperature causes less
diffusion of oxidant into the fuel flow. Less penetration of air into the fuel jet results in
flame volume increase.
63
3,0E-04
y = 2E-07x - 9E-06
0,8
2,4E-04
0,6
1,8E-04
y = 1E-04x + 0,2352
0,4
1,2E-04
Flame Length
0,2
6,0E-05
Flame Volume, m3
Flame Length, m
1,0
Flame Volume
0,0
0,0E+00
1000
1200
1400
1600
1800
Air preheated temperature, K
Figure 5.8 Length and volume of the HiTAC flame versus oxidizer temperature.
([O2]=12.8%. dF = 5.E-4 m. QF = 0.01 g/s. TF =299K)
5.4 Effect of fuel temperature
Figure 5.9 shows influence of the fuel temperature on the HiTAC flame length and volume
at 10% of oxygen concentration. In these cases, in order to simplify the numerical
calculations, the pyrolysis of fuel has not been taken into account. It can be seen that
increase of fuel temperature leads to decrease of both flame length and volume. Flame
volume in the case of 299 K fuel temperature is 2.9 times larger than that in the case of
1073 K. The reason for this is that decrease of the fuel density results in increase of fuel
inlet velocity at constant fuel flow rate, and changes in density and fuel inlet velocity are
proportional to the fuel temperature. Increasing initial velocity of the fuel jet improves the
mixing between fuel and oxidizer, which results in flame volume decrease.
64
1,0
4,0E-04
Flame Length, m
Flame Volume
3,0E-04
y = 3E-06x -1,7052
2,0E-04
0,6
0,4
0,2
1,0E-04
y = -0,0002x + 0,441
0,0
Flame Volume, m3
Flame Length
0,8
0,0E+00
273
473
673
873
1073
Fuel preheated temperature, K
Figure 5.9 Length and volume of the HiTAC flame versus fuel temperature. [O2]=10%. dF
= 5.E-4 m. QF = 0.01 g/s. TO =1173K
For flame length, although the trend is similar, the effect of fuel temperature is not so
pronounced. Only 1.45 times the flame length is increased. The shape and size of flame at
different fuel temperatures were shown in Figure 5.10. It can be seen that the changes in
flame volume comes from not only flame’s length, but also flame’s diameter. This can
explain the above behaviours. The function of the flame length to the fuel temperature is
linear, and the flame volume is power function of the fuel temperature.
65
QF =0.01 g/s
QF =0.01 g/s
QF =0.04 g/s
TF=299 K
TF=1073 K
TF=299 K
(a)
(b)
RO
(c)
Figure 5.10 Length and volume of the HiTAC flame versus fuel firing rate.
(TO =1173K. dF = 5.E-4 m.)
5.5 Effect of fuel flow rates
Figure 5.11 depicts the influence of fuel flow rate on flame length and volume versus
oxygen concentration. The fuel mass flow rates are 0.01, 0.02 and 0.04 g/s respectively.
Flame length was found to be nearly constant for each oxygen concentration and is almost
independent of the fuel flow rate. This may be because the studied cases were in the range
of turbulent flame, the Reynolds numbers of fuel nozzle were 3316, 6631 and 13263,
respectively.
The flame volume increases with increase in fuel flow rate. The shape and size of flame at
different fuel flow rates are given (a) and (c) in Figure 5.11. It is obvious that the diameter
of the flame increases with the increase of fuel flow rate. The increasing trend is higher at
lower oxygen concentration. For example when oxygen concentration is equal to 5%, flame
volume in the case of fuel flow rate equal to 0.8 kW is 2.3 times larger than in the case of
66
the fuel flow rate equal to 0.2 kW. At 10% oxygen concentration, this relationship is 1.8,
and at 15.8% oxygen concentration, it is equal to 1.6. This is the fact that the mean
residence time increases with either decrease of the oxygen concentration, or increase of
0,9
3,E-03
0,6
2,E-03
0,3
1,E-03
0,0
0,E+00
1,0
0,0
0,2
0,4
0,6
0,8
Volum e, m 3
Length, m
fuel flow rate.
Fuel firing rate, kW
L. at 15.8%[O2]
L. at 10%[O2]
L.at 5%[O2]
V. at 15.8%[O2]
V. at 10%[O2]
V.at 5%[O2]
Figure 5.11 Length and volume of the HiTAC flame at different cases
( [O2]=10%. dF = 5.E-4 m. TO =1173K)
5.6 Effect of fuel nozzle diameter
The influences of diameter of fuel nozzle on the HiTAC flame length and volume at the
constant fuel flow rate were showed in Figure 5.12. The fuel velocity at the fuel nozzle
outlet was kept constant. The increases of the fuel nozzle diameter were obtained by
elevated fuel temperature. The influence of the fuel nozzle diameter on flame length and
67
volume can be negligible for the studied cases. This is because the changes of mixing
between fuel and the oxidizer are small for the studied cases because of the constant
velocities of fuel and air. It can be argued that the combustion with high temperature and
low oxygen concentration oxidizer is diffusion combustion although it is ‘volumetric
combustion’ instead of ‘flame front’ combustion.
4,E-04
3,E-04
0,6
2,E-04
0,3
Length
Volume
0,0
1,E-04
0,E+00
Fla m e Vo lum e, m 3
Flam e Length, m
0,9
2,E-04 4,E-04 6,E-04 8,E-04 1,E-03
Diameter of fuel nozzle, m
Figure 5.12 HiTAC Flame length and volume versus diameter of fuel nozzle.
([O2]=10%. TO =1173K. TF=299K QF =0.01 g/s)
5.7 Scaling analysis
Generally, Froude number is defined to establish momentum-controlled and buoyancycontrolled regimes. For turbulent jet diffusion flames, the following definition of flame
Froude number is used [73]:
68
Fr f =
v F f s3 / 2
⎡ ΔT f
⎤
ρ
( F )1 / 4 ⎢
gd F ⎥
ρ∞
⎣ T∞
⎦
1/ 2
(5.2)
where ΔT f was estimated using ΔT f = T f ,max − To , and
fs =
1
(5.3)
(mo / m F ) stoic + 1
Flame length and volume were plotted against the flame Froude number, Frf, as the
abscissa and a dimensionless flame length L* and V* as the ordinates for flame length and
flame volume, respectively in Figure 5.13.
L* =
[d
Lf fs
F
( ρ F / ρ o )1 / 2
]
(5.4)
and
V* =
[
V f f s2
L f × d F ( ρ F / ρ o )1 / 2
]
2
(5.5)
Using the dimensionless groups defined above, two regimes can be identified: a buoyancydominated regime that is correlated by the expression:
69
30
3
L*
V*
20
L*
V*
2
10
1
0
0
1
4
7
Frf
10
Figure 5.13 HiTAC Flame length and volume correlated with the flame Froude number
L* =
8.22 Fr f0.4
(1 + 0.07 Fr f2 ) 0.2
Fr f < 3
(5.6)
and a momentum-dominated regime by the expression:
L*=11
Fr f ≥ 3
(5.7)
Comparing with Delichatsios’ correlation [73] obtained for conventional combustion,
L* =
13.5Fr f0.4
(1 + 0.07 Fr f2 ) 0.2
L*=23
Fr f ≥ 5
Fr f < 5
(5.8)
(5.9)
70
The criteria constants, L* and Fr f , in the obtained correlation Equ.5.6 and Equ.5.7 are
smaller comparing that in Equ.5.8 and Equ.5.9, respectively. This can be understood from
their definitions by Equation 5.4 and 5.5, respectively. When the temperature of the
oxidizer is increased and the oxygen concentration is decreased, both fs and ρo are
decreased. Therefore, L* could be greatly decreased. The criteria constants vary from 23 to
11 at momentum-dominated regime, and the constant at buoyancy-dominated regime varies
from 13.5 to 8.22.
Influences of oxygen concentration and temperature in oxidizer on Fr f can also be directly
analysed as well if ΔT / T∞ in Equation 5.2 is substituted by Equation 5.1. Generally, the
decrease of oxygen concentration leads to decrease of Fr f . However, the increase of
oxidizer temperature causes the increases in Fr f . The result of both of these compensative
effects leads to a small decrease of Fr f . Thus, the criteria Fr number separating buoyancy
or momentum-dominated regime varies from 5 to 3.
More should be said that there are other similar correlations ( for example in work [75] ) to
that of Delichatsios have been proposed for normal combustion conditions. In these works,
various investigators use different definitions and experimental techniques to determine
flame lengths. However, the effects of oxygen concentration and temperature on these
correlations should be similar as the above work.
The predicted data of flame volume collapsed to a single linear curve for buoyancydominated regime as:
V * = 0.17 Fr f + 0.15
Fr f < 3
(5.10)
And a momentum-dominated regime where the dimensionless flame volume V* is constant:
V*=0.7
Fr f ≥ 3
(5.11)
71
5.8 Conclusions
Influences of combustion with high temperature and low oxygen concentration oxidizer on
flame length and volume were numerically investigated. The following conclusions can be
drawn:
Flame length increases with either the decrease of oxygen content, or increase of oxidizer
temperature, or decrease of fuel temperature. Furthermore, the flame length is independent
of the fuel flow rate and the diameter of the fuel nozzle.
Flame volume increases with either the decrease of oxygen content and increase of oxidizer
temperature, or the reduction of fuel temperature, or the increasing in fuel firing rate. Flame
volume depends very much on the oxygen concentration in the oxidizer.
Influences of high temperature and low oxygen concentration in the oxidizer on the flame
Froude number, Frf were examined. Regimes of momentum- or buoyancy-control, were
determined on the assumption that oxidizer temperature and oxygen concentration are
changeable.
A simple correlation of the ‘flame’ length and volume with flow parameters has been
derived in terms of a flame Froude number for momentum-buoyancy transition jet flame
under the HiTAC condition. The criteria constants of the dimensionless flame volume V*
and the dimensionless flame length L* to assess momentum– or buoyancy–control flame
are given.
72
6 Flame Entrainments study in coflow single jet flame
The entrainment of the jets is the key technique solution for the industrial applications of
this novel combustion technology. However, little knowledge is available about the
influences of preheated temperature and oxygen concentration on the jet-flame entrainment
rate. When combustion occurs in the condition of high-temperature and oxygen deficient
oxidizer, reaction zones are relatively widely distributed to yield a somewhat widespread
and mild heat release, and hence a uniform temperature distribution. This also implies that
the buoyancy force is decreased. The effects of the mild heat release and less buoyancy
force on the entrainment could be of fundamental interest.
In the previous work, a ‘chemical’ flame is used to describe this less luminosity (or
invisible) chemical reaction zone, which is coincident with the internal zone in the
outerside border flame. The flame’s outerside border is determined according to the
flammability limits of the combustible gases in the combustion chamber in term terms of
the oxidation mixture ratio. The length and volume of the ‘chemical’ flame zone are used to
describe the physical properties of the chemical reaction zone in the combustion chamber,
which have been systemic studied in the previous work. In the context of this paper, the
flame means a chemical flame zone.
In this work, studies were performed to understand the entrainment into a ‘chemical’ flame
zone induced by a turbulent jet flame using the high-temperature and oxygen deficient
oxidizer. The combustion with High-Temperature (above the fuel’s auto ignition
temperature) and Oxygen Deficient atmosphere is called as HiTAC combustion hereinafter,
since those two parameters are unique characteristic differing from any other combustion.
A flame entrainment rate is proposed to describe the entrainment induced by a jet flame.
The effect of the preheat temperature and the oxygen concentration in the oxidizer, heat
release and buoyancy on the entrainment rate is investigated. A correction Richardson
coordinate, where the effect of the oxygen concentration (stoichiometric ratio) is included,
was derived to describe the local influence of buoyancy force along the chemical flame
length under the HiTAC condition. The global behaviour of the entrainment was revealed.
Corrections of entrainment rates were derived in terms of an Frf number for momentumbuoyancy transition jet flame under the HiTAC condition.
73
6.1 Study Cases
Variables chosen for numerical and theoretical studies were as follows: flow rate and
temperature of fuel, oxygen concentration and temperature of the oxidizer, and diameter of
the fuel nozzle (Table 6.1). The compositions of the oxidizer (O2, CO2, H2O and N2) were
obtained according to the level of oxygen based on the chemical balance of the
conventional burner that was used to produce the oxidizer.
Table 6.1 Values of Variables of Numerical Studies
Cas
[O2],
To,
Tf,
Cp1,
uF,,
u o,
es
mass%
K
K
kJ/(kg.K)
m/s
m/s
dF, m
Reo
Lf2, m
105Ris
Frf,
(wet)
1
5
1173
299
1.3030
25
0.98
5.0E-4
3316
0.625
1.89
0.67
2
10.2
1173
299
1.2653
25
0.98
5.0E-4
3316
0.400
1.90
1.37
3
13.4
1173
299
1.2446
25
0.98
5.0E-4
3316
0.350
1.90
1.83
4
16.9
1173
299
1.2215
25
0.98
5.0E-4
3316
0.313
1.91
2.29
5
23.2
1173
299
1.1790
25
0.98
5.0E-4
3316
0.263
1.92
3.16
6
13.4
1073
299
1.2237
25
0.98
5.0E-4
3316
0.338
1.82
1.73
7
13.4
1373
299
1.2373
25
0.98
5.0E-4
3316
0.375
2.06
1.99
8
13.4
1473
299
1.2228
25
0.98
5.0E-4
3316
0.375
2.13
2.09
9
13.4
1573
299
1.1919
25
0.98
5.0E-4
3316
0.388
2.20
2.18
10
10.2
1173
710
1.2653
25
0.98
7.7E-4
2146
0.425
1.89
1.36
11
10.2
1173
1121
1.2653
25
0.98
9.7E-4
1708
0.438
1.89
1.33
12
16.9
1173
299
1.2215
50
0.98
5.0E-4
6631
0.325
0.48
4.35
13
16.9
1173
299
1.2215
100
0.98
5.0E-4
13263
0.313
0.12
8.49
1
2
Cp is calculated according to the compositions of the oxidizer.
Lf was obtained from previous work.
Studied temperatures of the oxidizer are greater than 1000K, which is higher than the
autoignition temperature of most of fossil fuels, and it is in the range of HiTAC conditions.
The coflow speed is about 4% of the jet speeded for most of studied cases. Additionally,
since for all studied cases, this velocity was kept as constant, thus its influence was ignored.
In addition, the axial distance, x, is counted from the virtual origin from the nozzle.
74
There is no need for a pilot flame to stabilize the flame under HiTAC conditions. This is
very important for the studies of the effect of the heat release on the entrainment. Although
the mass flow rate of the pilot is very small, the heat release from the pilot occupies a large
portion if, for example, the hydrogen is used as the pilot fuel. The heat-flux ratio between
the jet and pilot is only 22% even the pilot-to-jet mass flux ration is less than 1.2%, as
analysed in the work [90]. Therefore, in the HiTAC condition, the studies of the effect of
the heat release on the entrainment especially in the near field will have a higher accuracy.
Flame length was calculated as the distance between the end of the fuel nozzle and the axial
location of the oxidation mixture ratio equal to 0.99. The lift-off distance was negligible.
There will be no any error for the calculation of flame length.
In this work, a flame entrainment rate, Rent , is proposed. It is defined as a ratio of the mass
flow rates through the cross section of the flame (me) to the initial jet mass rates (m0), and
be expressed as:
Rent = me m0
(6.1)
Here,
d * = d F ( ρ F / ρ o ) 0.5
(6.2)
The mass flux, me, can be calculated by the following integral of velocity and density
profiles for whole section of the chemical reaction zone.
me = ∫
R flame
0
2 ρu πrdr = ∑ ρ i u i Ai
(6.3)
i is the calculation cell number.
This definition of this flame entrainment rate can be used to estimate the flame
entrainments by a group of jet flames, which can be interference between the individual jet
flames. In this paper, only one jet flame is studied. Consequently, the results from this work
can be compared with that from the previous published works, which only single jet
75
entrainment was studied. In order to compare with the available data from the literatures, it
is deducted as the well-know expression:
Rent = me m0 = C e ( ρ o ρ F ) 0.5 x d F = C e x / d *
(6.4)
6.2 Effects of the oxygen concentration of the oxidizer on
entrainment
The entrainment rate for the burning region of jet flames with different oxygen
concentration in the oxidizer are presented in Figure 6.1.
120
Case1
Case2
Case3
Case4
Case 5
me/m0
80
40
Chemical Flame length
0
0
100
200
300
400
500
600
x/d*
Figure 6.1 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different oxygen concentrations in the oxidizer
It can be seen that the curves show that the trend is similar at different oxygen levels. The
entrainment increases linearly for the first half of the flame length, reaching a maximum
value, and the decrease of the entrainment occurs over the last third of the flame length.
76
This global various trend might be explained by the nature of the jet. From previous studies
[88], jets in co-flow are known to show a jet-like behaviour in the near field, but a wakelike behaviour in the far field. For a fully developed self-similar reacting jet, the
entrainment coefficient is a constant. For a fully developed wake with a small excess
velocity, the entrainment is achieved by the encroachment of the boundary on the
surrounding fluid, and should decrease according to self-similarity. This can be observed in
the jet-like region (from the nozzle tip to the location of the maximum value of the
entrainment rate) in Figure 6.1, where the entrainment rate increases, and the wake-like
region (from the location of the maximum value of the entrainment rate to the end of the
chemical flame), where the entrainment decreases, and the entrainment coefficient has a
negative value. There are few reports in the literature about this phenomenon for the
reacting jet. One of the most important reasons is the used definition of the flame border. In
this work, the chemical flame border is separated even at the tip of the flame. At far
downstream locations in the reacting jet, the distribution becomes smaller than the free jet.
Therefore, a reduction of the entrainment is possible.
It is worthy to note that both the near and far fields of the jet are of importance. In the near
field, a significant portion of the mixing process occurs, which affects the flame
stabilization, and in the far field, the flame length is defined. In particularly, the location of
the maximum entrainment decides the optimal design of the furnace and the location of the
burners. In this part, these behaviours of the entrainments are discussed firstly, and the
location of the maximum entrainment will be discussed later.
The average entrainment coefficients Ce from the nozzle tip to the locations before the
maximum entrainment for Case 1, 2, 3, 4, and 5 were summarized in the Table 6.2.
The entrainment coefficient Ce was determined to be 0.32 for momentum driven free jet by
Ricou and Spalding [76]. Becker and Yamazaki [78] reported Ce =0.16 for a momentum
driven reacting jet. Han and Mungal reported [82] Ce=0.13 for a reacting free jet. A
summary of the results from the present study and previous reported work are listed in
Table 6.2.
77
Table 6.2 Summary of the results for present study and previous works
Reo
[O2],
Flame
uF, m/s
uo,m/s
Lf/dF
Ce
mass%
Case 1
3316
5
Yes
25
0.98
1250
0.28
Case 2
3316
10.2
Yes
25
0.98
800
0.20
Case 3
3316
13.4
Yes
25
0.98
700
0.17
Case 4
3316
16.9
Yes
25
0.98
625
0.15
Case 5
3316
23.2
Yes
25
0.98
525
0.14
R&S [76]
>25 000
23.2
No
*
0
*
0.32
Muniz
&Mungal
10 000
23.2
Yes
36
0.45
100
0.12a
&Mungal
37 500
23.2
Yes
134
2.0
152
0.076a
23.2
Yes
0.16b
23.2
Yes
0.13(1-e-0.036(r-1))
[85]
Muniz
[85]
Becker
&
Yamazaki [78]
Han &Mungal
[90]
a
Effect of ρi′ and ui′ is neglected. Ce could be underestimated by around 20%.
b
For the entrainment coefficient at small ξ.
One major difference in the present work and other reported results is the conditions of the
oxidizer. Here, high temperature and oxygen deficient oxidizer was used. Another
difference in the present and other results is the definition of the entrainment, which the
flame entrainment is used in this work rather than the jet entrainment in other works.
It can be observed that the entrainment rate for the studied cases are in the range of
0.14 ≤ C e ≤ 0.28 . They are smaller than those of non reacting jets, but larger than those of
reacting jets under normal combustion conditions. It is been known from previous studies
[78, 82, 85], that heat release in reacting jets reduces the entrainment. This means that the
entrainment coefficient for the reacting jets should be lower than 0.32. This principle is also
valid by flames under the HiTAC condition.
Table 6.2 shows that the maximum entrainment rate for the studied cases occurs at the
lowest oxygen concentrations. With the reduction of the oxygen concentration, the
78
entrainment increases. For example, the flame at the case of 5% oxygen concentration
experiences a 50% increase in the entrainment than at 21% oxygen concentration. This
trend is consistence with the variation of the HiTAC flame volume or temperature
uniformity. It has been shown, that a larger chemical reaction zone exits when the
combustion occurs in a high temperature and low oxygen concentration oxidizer. Lower the
oxygen concentration in the oxidizer, larger the chemical reaction zone and consequently a
more uniform the heat release. Therefore, these features might be used to explain the effect
of the oxygen concentration on the entrainment. i.e., a larger heat release uniformity, results
in a larger entrainment. In order to understand the present results, further analysis of the
effect of heat release on the entrainment rate is necessary to understand the above results.
In the absence of buoyancy, the momentum flux of a turbulent jet, which is defined as the
momentum crossing a plane perpendicular to the jet axis, remains constant with axial
distance and is equal to the initial jet momentum flux. When ambient fluid is entrained into
a turbulent jet, the mass flux of the jet increases and the velocity of the jet decreases due to
the conservation of the momentum. It has been shown [87], that flames have higher
centreline velocities and concentrations than those of non reacting jets at the same axial
location with the same exit conditions. It can be assumed that the total mass flux is smaller
in a flame than in a non reacting jet with the same exit conditions. Therefore, the mass flux
by entrainment is less if the initial mass fluxes are the same. The increase in centreline
velocities and concentrations is caused by the heat release. It can be concluded that a
uniform heat release rate will have a less effect on these parameters, thus a larger
entrainment rate. It has been demonstrated [2,14-20] that the concentration and temperature
profiles are more uniform under the condition of high-temperature and oxygen deficient
oxidizer than those in the conventional condition. Furthermore, a recent study proved [51]
that the centreline velocities are smaller when the oxygen concentration in the oxidizer
decreases during the HiTAC condition. These features implicates that the entrainment
during the HiTAC condition will be larger than under conventional conditions. At the
extreme, when the oxygen concentration in the oxidizer is equal to 0, the most uniform heat
release is achieved, and there will be no effect by the heat release. Additionally, lower
centreline velocities at the lower oxygen concentration lead to a larger the flame widths,
which have proved in [2, 14-20]. Consistent with this, the entrainment at the low oxygen
concentration increases.
79
The fluctuations of the velocities and temperatures also affect the entrainment rate. These
intensities of the fluctuations vary along the flame length. It is known that there exists some
effect that increases along the flame length, that enhances the large-scale structures and
fluctuations in a reacting jet [82]. Becker and Yamazaki [84] identified this effect as
buoyancy, which adds additional momentum to the jet. This will be further discussed later.
Finally, one possible reason for the increase in the entrainment rate in HiTAC combustion
than that in the normal cases is the use of the different definition of the entrainment. Here,
the flame entrainment is used rather than the jet entrainment, which was adopted by most of
researchers. In HiTAC combustion, the flame width increases with the reduction of the
oxygen concentration [2, 14-18].This contributes the increase of the entrainment as well.
One effect the reduced entrainment has on the flame is an increase of the chemical flame
volume. Consequently, an estimate of the entrainment increase caused by lower oxygen
concentration in the oxidiser can be obtained from the flame volume growth rate, because
the flame volume is dependent on the oxygen concentration [2,14-18]. Here, the effect of
oxygen concentration on the entrainment rate has been shown in Figure 6.2a. It can be seen
that the entrainment increases when the oxygen concentration in the oxidizer reduces. This
is clearer if the oxygen concentration in the oxidizer is below 10%. This trend is consistent
with changes of the chemical flame volume versus the oxygen concentration in the oxidizer.
Moreover, this trend becomes weaker when the oxygen concentration is approach to 21% in
the oxidizer.
80
Entrainment coefficient, Ce
0,4
0,3
0,2
0,1
0
0
5
10
15
20
25
Oxygen Concentration in the oxidizer, %
Entrainment coefficient, Ce
a)
0,4
0,3
0,2
0,1
y = -0,6306x + 1,0099
0
1
1,1
1,2
1,3
1,4
1,5
(Tf/To)**0.5
b)
Figure 6.2 Entrainment rate change versus the oxygen concentration and the characteristic
temperature ratio (T f / To ) 0.5 .
81
The uniformity of the heat release can be simply represented by the characteristic
temperature ratio, (Tf/To). Then, the effect of uniformity of the heat release on the
entrainment can be further quantitative analyzed.
It is very obvious that the highly preheating temperature and oxygen concentration of the
oxidizer plays an important role on the flame temperature during the HiTAC combustion.
Consequently, the HiTAC’s flame temperature can be featured when it is expressed as a
function of the preheated temperature of the oxidizer. The characteristic temperature ratio
can be estimated through the relationship from previous work as following.
T f ,max = To +
Tad
21
+1
[O2 ]
(6.5)
and use Tf,max instead Tf, then, the characteristic temperature ratio can be obtained:
Tf
To
=
Tad
1
+1
To 21
+1
[O2 ]
(6.6)
Here Tad is the adiabatic temperature of propane at the stoichiometric combustion with air
(79% N2 and 21% O2).
The entrainment rate in this work is plotted as a function of the characteristic temperature
ratio resulting from combustion, (T f / To ) 0.5 in Figure 6.2 (b). It can be seen that the
entrainment rate changes with the same factor of the (T f / To ) 0.5 when (T f / To ) 0.5 is less
than 1.4.
It is interesting to note that Han [82] argued that in the near field of reacting jets, the effect
of heat release on the entrainment can be estimated by the ratio of (ρo/ρa)0.5. The ratio of
(ρo/ρa)0.5 can be approximated by (Tf/To)0.5 for a constant pressure condition. This result is
consistent with our presented data.
82
To summarize of these discussions, the conclusion can be made gives on the hand that the
entrainment rate during high-temperature and oxygen deficient conditions will be larger
than under conventional combustion. The lower the oxygen concentration in the oxidizer is,
more uniform the heat release rate is, thus larger the entrainment rate. The effect of heat
release reduces the entrainment in the near field of the reacting jets with the same factor of
the characteristic ratio (Tf/To)0.5.
6.3 Effects of the temperature of the oxidizer on entrainment
The preheated temperature of the oxidizer was increased when the other conditions were
kept constant for the studied cases (case 3, 6-9). Preheat temperature of the oxidizer has
small effect on the entrainment for studied cases, as shown in Figure 31
60
Case 3
Case 7
Case 9
Case 6
Case 8
50
100
me/m0
40
20
0
0
150
200
250
300
x/d*
Figure 6.3 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different preheated temperature of the oxidizer
83
The studied cases have the same entrainment ratio up to x/d*<50. This trend can be
extended to the point of the maximum entrainment ratio. At the flame tip, a slightly higher
entrainment takes place at lower preheat temperatures.
The momentum of the fuel jet for these cases was maintained constant. As the inlet
temperature of the oxidizer is increased, the mass flux by turbulent mixing is reduced due
to the decrease of air density. Meanwhile, the small expansion of the burnt gases does not
suppress the turbulent mixing. These effects are considered to compensate each other. As a
result, the entrainment is maintained almost constant. In particular, up to x/d*<50, it can be
seen that this is the jet developing regime. The same momentum will cause the same
entrainment. At the flame tip region, located in the buoyancy –control zone, the coherent
structures of the flame are easy to penetrate the ambient gases.
This result can also quantitatively explained by using the above relationship:
Re ∝ (T f / To ) 0.5 . When the preheat temperature increases for the same level oxygen
concentration, the flame temperature increases as well. Thus, the characteristic temperature
ratio has a small variation, and this variation becomes smaller when the oxidizer’s
temperature is increased. Consequently, the effect of the preheating temperature of the
oxidizer is smaller, especial in the momentum controlled zone (near field). In the far field,
it might be in the regime of the buoyant force- control zone.
6.4 Effects of the fuel flux on entrainment
For the studied cases 4, 12 and 13, the fuel fluxes were increased through an increase of the
fuel velocities, while the other conditions were kept constant. The effects of the fuel flux on
the entrainment were shown in Figure 6.4. The increase in the fuel fluxes leads to a slight
increase in the flame temperature, the density of the flue gases is decreased, thus the low
entrainment. Additionally, the diameter of the flame obviously increases as the fuel flow
rate increases which increases the mass entrainment. These factors compensate each other
and achieve a constant mass entrainment.
84
60
Case4
Case 12
Case 13
me/m0
40
20
0
0
50
100
150
200
250
300
x/d*
Figure 6.4 The variation of the entrainment rate along the axial direction from fuel nozzle
tip at different preheated temperature of the oxidizer
6.5. Effects of the buoyancy
Becker and Yamazaki [78] have established a criterion by which the effects of buoyancy in
a flame can be quantified. The parameter of interest is a nondimensional stream wise
coordinate, ξ
ξ ≡ Ris1 / 3
x
ds
(6.7)
where, Ris is the source Richardson number, and ds is source diameter. For a uniform exit
velocity, uo, the source Richardson number and diameter become: Ris = gds/uo2, where g is
acceleration due to gravity. ξ contains the integrated effect of the buoyancy along the jet
trajectory. Buoyancy forces are considered to be negligible when ξ< 1. This coordinate is
widely used by other researchers. [82, 85] etc.
85
It should be noticed that this relationship has not considered the influences of the oxygen
concentration and the air preheated temperature. This is because of an assumption, that ρf is
typically less than 0.5ρo was used when this relationship was derived. This assumption is
reasonable for the combustion conditions without preheated air. However, it will cause a
large error when this assumption is used in combustion under highly preheated and oxygen
deficient conditions in the oxidizer. Therefore, at least a correction of this coordinate should
be added to consider the influences of the preheated temperature and oxygen concentration.
A simply approach is introduced as follows. If we consider a cone-shaped jet with the width,
δ, the ratio of the buoyancy that the entire jet brush feels (Fb)compared to the initial
momentum (Fm)can be written as [82],
Fb
=
Fm
(ρ o − ρ f )g
ρFu
2
F
π
4
π
4
d
δ2
x
3
(6.8)
2
F
Letting δ ∝ x and applying a cubic root to the entire expression, the nondimensional
stream wise coordinate, ξ ' can be wrote as:
ξ ' ≡ Ri01 / 3
ρf
x
(1 −
)
*
ρo
d
(6.9)
here,
Ri0 = gd * / u F2
(6.10)
If we assume ρf is typically less than 0.5ρo, i.e., ( ρ o − ρ f ) → ρ o , we can obtain the
expression as Equation 6.9, which defined by Becker and Yamazaki [78]. Under HiTAC
conditions, this assumption is not reasonable. Further analysis is required.
86
In order to obtain a more concise expression of the nondimensional stream wise coordinate,
ξ’, the relationship between ρf and ρo has to be addressed. Considering the relationship
between gas densities and its temperatures, one has:
ρf
T
∝ o
ρo T f
(6.11)
Combing Equation 6.11 and Equation 6.6, one can obtain:
ρf
∝
ρ o Tad
1
1
+1
To 21
+1
[O2 ]
(6.12)
Thus, Equation 6.12 can be modified as:
ξ HiTOD ≡ c HiTOD Ri01 / 3
x
d*
(6.13)
Here, CHiTAC can be named as the Richardson correction ratio for the HiTAC combustion,
and
c HiTOD ≡
1
To 21
1+
(
+ 1)
Tad [O2 ]
(6.14)
From Equation 6.14 it is easy to find the effects of the assumption, ( ρ o − ρ f ) → ρ o even
for the normally combustion (21% [O2]). In this case, Equation 6.414 can be conversed as:
c ≡
1
2T
1+ o
Tad
(6.15)
87
If the adiabatic flame temperature is significantly larger than the free stream (ambient)
temperature, i.e. c=1, the effects of the stream temperature is neglectable. Otherwise, this
assumption could not be reasonable.
The effect of oxygen concentration and preheated temperature of the oxidizer is considered
in the corrected Richardson number. Here, buoyancy forces are still considered to be
negligible when ξHiTAC< 1.
In Figure 6.5a, ξHiTAC in log scale is plotted as a function of axial distance from the jet
origin for the different caes. In order to understand the effect of the buoyancy forces on the
flame length, ξHiTAC is presented in log scale as a function of the axial distance from the
nozzle, normalized by the flame length, Lf in Figure 6.5b.
Figure 6.5 shows the influence of buoyancy on the HiTAC flame. It can be seen that the
buoyancy force increases with the reduction of the oxygen concentration whilst comparing
with Case 1, 2, 3, 4 and 5. For Case 5 (21%O2), the buoyancy forces are considered
negligible for z/d*< 72. The value for case 1 (5%O2) is 140, which is approximately twice
the length over which Case 5 is non-buoyant. This can be understood as a more uniform
temperature profile provides smaller buoyancy forces. Furthermore, for Case 5, this criteria
location is 0.4 of the flame length, which means from the nozzle tip up to 40% of the flame
length; it is in the momentum-control regime. This value is 30% for Case 1. This implies
that more of the flame’s length locates in the momentum –control regime for higher oxygen
concentration as well.
88
100
Corrected Kesi
Buoyancy non-negligible
1
Case1
Case 3
Case 5
Case 9
Case 11
Case 13
Momentum-driven
Case2
Case4
Case 7
Case 10
Case 12
0.01
0
100
200
300
400
500
600
x/d*
(a)
Corrected Kesi
10
1
0.1
Case1
Case2
Case 3
Case4
Case 5
Case 7
Case 9
Case 10
Case 11
Case 12
Case 13
0.01
0.0
0.2
0.4
0.6
0.8
1.0
x/Lflame
(b)
Figure 6.5 The local influence of buoyancy as a function of the axial distance normalize by
a) (x/d*), and b) the flame height in log scale for difference cases
89
The influence of the preheat temperature of the oxidizer on buoyancy is minor for the
studied cases (Case 3, 7, and 9). The buoyancy force can be considered negligible for 0.35
of the flame length for these cases. This also implies that the effect of the oxidizer
temperature on the temperature field uniformity is small when the oxidizer’s temperature is
higher than 1000K. This is because of the increase of the air preheat temperature also
causes an increase in the flame temperature, the different flue gas densities, which is the
major factor for affecting the buoyancy force, is almost kept constant.
The fuel fluxes have stronger effects on the entrainment. The momentum driven regime has
an increase in the length with the increase of the fuel fluxes. For example, the effects of
buoyancy become non-negligible for x/d* >82 for Case 4. These values for case 12 and 13
are 125 and 200, respectively. It is obvious that the larger the initial monument is, it
maintains a longer momentum control regime, and the Ris decreases as the factor of (uo)2. It
can also be seen from Figure 6 that the larger initial monument causes a longer flame length
expose in the monument control regime. For the case 13, around 80% flame length locates
in the regime of monument - controlling. The values for case 12 and 4 are 45% and 32%
respectively.
6.6. Global field behaviour of the entrainment
We have detail discussed the entrainment behaviour in near field. In this part, the global
field behaviour of the entrainment is further chased.
It was chosen to present the data as a function of the axial distance(x) normalized by the
flame height, Lf, because the general relationships of the flame length have been well
studied. Here, Lf is the chemical flame length can be obtained in previous work. Another
dimensionless parameter, m* is adopted. It is an expression of an entrainment rate
normalized by the stoichiometric requirements:
m* =
me
( S + 1)m F
(6.16)
here,
90
S = (mo / m F ) stoic
(6.17)
Figure 7 shows the relationship of this dimensionless parameters, m*, and the parameter
(x/Lflame). In order to see clear various, the log scale of the distance axial is used. It can be
seen that all the maximum normalized entrainment positions can be collapsed to a single
point. It appears at the location of x / L f ≅ 0.7 . Basing on this point, two regimes can be
identified as a near and a far field.
m*=me/((S+1)mF)
2
1,5
Case 1
Case 5
Case 9
Case 13
Case 2
Case 6
Case 10
Case 3
Case 7
Case 11
Case 4
Case 8
Case 12
1
0,5
0
0,001
0,01
0,1
1
10
x/Lflame
Figure 6.6 Dimensionless entrainment rates v.s. ratio of axial distance from the nozzle tip
to flame height
From Figure 6.6, the mass entrainment of the near field can be expressed as:
me
x
= 1.77
Lf
( S + 1)m F
x / L f ≤ 0.7
(6.18)
and the mass entrainment of the far field can be stated as:
91
me
x
x
= 3.67 − 3.31
= 3.31(1.11 −
)
Lf
Lf
( S + 1)m F
x / L f > 0.7 (6.19)
Obviously, the entrainment coefficient is positive in the near field where and it is negative
in the far field.
It is possible to use available flame length’s relationship to further simply above
relationships. In this work, the relationship of the flame length by mean of the flame Froude
number during the HiTAC condition obtained in previous work is used. Only the
entrainment in near field, x / L f < 0.7 , is discussed in this part.
According to the results from previous work, the flame length during the HiTAC condition
can be decided by the flame Froude number as:
L* = 11
L* =
Fr f ≥ 3
8.22 Fr f0.4
Fr f < 3
(1 + 0.07 Fr f2 ) 0.2
(6.20)
(6.21)
Here,
L* =
[d
Lf fs
F
(ρ F / ρo )
1/ 2
]
=
Lf
*
(6.22)
1/ 2
(6.23)
d ( S + 1)
and
Fr f =
u F f s3 / 2
⎡ ΔT f
⎤
ρ
( F )1 / 4 ⎢
gd F ⎥
ρ∞
⎣ T∞
⎦
fs =
1
(mo / m F ) stoic + 1
(6.24)
92
Then, a simplification of the entrainment in the near field can be obtained through
substituting Equation 6.20-Equation 6.24 into Equation 6.18. They can be expressed as:
me
x
= 0.16 * for Fr f ≥ 3
mF
d
2 0.2
me (1 + 0.07 Fr f )
x
=
0. 4
mF
4.64 Fr f
d*
for Fr f < 3
(6.25)
(6.26)
It can be seen from Equation 6.25, the entrainment coefficient, Ce is equal to 0.16 for a
momentum driven reacting jet ( Fr f ≥ 3 ), which is comparable with the data by Becker and
Yamazaki [78].
These relationships can also be use to compare the entrainment coefficient which were
obtained in the previous part in this paper. For example, Case 12 and 13 belong to the
momentum-dominated regime (see Table 1) because of their Frf numbers are greater than 3.
Then, the entrainment coefficient, Ce, of these two cases is 0.16 according to the Equation
6.25. This value is only 6% larger than that obtained in previous data (Ce=0.15). A good
agreement can be obtained.
It is also very interesting to calculate the maximum mass entrainment, which can decides
the optimal design of the furnace and the location of the burners. Let’s put x / L f = 0.7 into
Equation 6.25, then the maximum entrainment can be estimated as:
me
= 1.24( S + 1)
mF
(6.27)
Comparison the maximum entrainments calculated by Equation 6.27 and the results
presented previous part, the results are also very encouraged.
93
6.7 Conclusions
The influences of combustion using a high-temperature and oxygen deficient oxidizer on
the flame entrainment induced by a turbulent reacting jet are numerically and theoretically
investigated. A flame entrainment ratio is proposed. The results can be summarized as
follows.
1.
The uniformity of the heat release in reacting jets has strong effect on the flame
entrainment. More the uniformity of the heat release, larger the entrainment. The effect
of heat release reduces the entrainment in the near field of the reacting jets with the
same factor of the characteristic ratio (Tf/To)0.5.
2.
The entrainment increases as the oxygen concentration is decreased. Furthermore, the
entrainment is independent of the fuel flow rate and the preheated temperature of the
oxidizer for the investigated temperature range (1073-1573K).
3.
The effect of the oxygen concentration and preheated temperature of the oxidizer on
buoyancy was examined. A correction Richardson coordinate, ξ HiTAC ≡ c HiTAC Ri01 / 3
x
,
d*
where the effect of the oxygen concentration (stoichiometric ratio) is included, was
derived to describe the local influence of buoyancy along the axial distance from the
nozzle under the HiTAC condition. It can be concluded that the buoyancy force
increases with the reduction of the oxygen concentration in the oxidizer.
4.
The global behaviour of the entrainment was revealed. Two regimes for the
entrainment have been identified in jet flames (a) the near field where entrainment
rates are positive; and (b) the far field where entrainment rates are negative.
Corrections of entrainment rates were derived in terms of a Frf number for
momentum-buoyancy transition jet flame under the HiTAC condition. Furthermore,
the maximum entrainments along the flame length are estimated.
94
7 Semi-industrial furnace with HiTAC burners study
7.1 Experimental measured and verification of mathematical
modelling for HiTAC furnace with one-flame burner
Experimental measurement and modelling validation are performed by the means of the
following data:
- Energy balance,
- Wall temperature profiles,
- In-furnace gas species
These are done using tests based on the HiTAC test furnace equipped with one-flame Highcycle Regenerative System (HRS).
7.1.1 Energy Balance
The reference temperature is set at T=298 K, so the sensible heat flow rate at the fuel inlets
was zero. The overall thermal energy (only including the fuels’ chemical energy) input to
the test furnace in this study was 182 kW. The thermal input of the fuel for the simulation
was calculated according to the chemical reaction steps (R3.1-R3.5). The heat of exhaust
after the burner for simulation was approximately calculated as the value of heat of flue gas
through the burner minus combustion sensible heat of the combustion air. The summary of
the energy balance of the semi-industrial furnace at the operation condition is given in
Table 7.1.
A figure of 54.65% of the predicted fuel thermal input passed through the burner, and 83.6
percent of the sensible heat carried by the flue gases through the burner outlets was used to
preheat the combustion air from 300 to 1211 K. This means the thermal efficiency of the
regenerator is 83.6%. This value is about 45.71 % of the total thermal input. This implies
that a very high level of energy utilization efficiency is achieved. Reiterating here, 8.94%
95
of the predicted total fuel thermal input is removed by flue gas through the burner. This
value is higher than measured value of 5.21%. Possible reasons for this could be one or a
combination of the following: heat loss in the burner, or the measurement point being on
the outside of the burner.
The predicted amount of heat taken by the air-cooling tubes occupies 51.67 % of the fuel
thermal input. The heat absorbed by air-cooling tube was measured to be around 54.27% of
the fuel thermal input within the margin of measurement error of 6.44%. Therefore, the
predicted and measured amounts are in a reasonable agreement. The predicted results also
indicated that 97.4% of the heat transferred to the air-cooling tube was due to radiation, and
2.6% was due to convection. Furthermore, the calculation enthalpy of the chimney flue gas
constitutes 9.78% of the fuel thermal input, while the actual value measured was 11.25%.
Again, the agreement is acceptable. The predicted heat loss through the furnace walls
accounts for 29.61% of the thermal input, compared with an actual measurement of
29.49%. Thus, these are also in a good agreement.
Table 7.1 Fuel characteristic and burner operating conditions
HiTAC mode
Fuel flow rate [Nm3/h]
7.7
Fuel inlet temperature, [K]
298
Fuel composition (Mass fraction, [%])
CH4 (0.02), C2H6 (0.95),
C3H8 (98.35) , C4H10 (0.67).
Combustion air flow rate, [Nm3/h]
200
Combustion air temperature,[K]
1211 K
Combustion air composition ([Vol %])
N2 (79), O2 (21)
96
Table 7.2 Energy balance from measurements and from modelling data (Reference
temperature: 298K)
Energy [kW]
1. Fuel power
Input
2. Combustion air sensible heat
Total
3 Heat taken away by the air-cooled
tubes
HiTAC trial measured
184.04
182.0
84.12 (after preheated by the
1.85 (before preheated by
burner)
burner)
268.16
183.85
95.09
99.78
4 Heat of flue gas through burner
100.58
5 Heat of exhaust after burner
16.46
9.57
17.99
20.69
51.27
48.13
3.23
--
Uncounted loss
--
6.06
Total
267.85
183.85
6 Sensible heat of flue gas through
Output
HiTAC CFD prediction
main chimney
7 Heat loss from walls
Radiation heat loss at inlets and
outlets
It is possible to improve the efficiency of the furnace by further increasing the efficiency of
the heat recovery from waste flue gases from the current value of 80% to 100%. The extra
heat recovery from the flue gases can be used to preheat the fuel, which can bring benefits,
such as an even greater reduction in NO emissions. This is very important for combustion
stability when using low and medium caloric value fuels in HiTAC technology as in these
cases, the fuel volume is larger. This method features preheating of both fuel and air, and is
referred to in the field as twin-preheating. As such the method involving preheating of the
combustion air only can be referred to as single-preheating.
7.1.2 Temperature Field
97
Furnace temperatures in the HiTAC test furnace were measured at various positions along
the left-hand-side wall of the test furnace (viewed from the burner). Figure 7.1 presents a
comparison of temperature predictions and measurements showing reasonable agreement,
with a maximum difference of about 10 K with a range of error of 2.8%. These values were
obtained from a stationary thermocouple located on the furnace wall.
1500
Temperature, oC
1400
Measured
Prediction
1300
1200
1100
1000
900
800
0
400
800
1200
1600
Z (mm)
2000
2400
2800
Figure 7.1 Temperature distribution on the side wall of the furnace at x = 0.8 m, and y = 0.3m
7.1.3 Gas Species
Figures 7.2 and 7.3 present the measured and predicted concentrations of O2 and CO across
the furnace chamber at specific vertical distances from the burner face. The x-axis
represents the vertical distance from the centreline of the burner.
From figure 7.2, it can be seen that the calculated O2 levels are in good agreement with that
measured. Both the measured and predicted curves show the same locations and
magnitudes for the maximum and minimum O2 and CO2 concentrations. Furthermore, the
measured CO concentration agrees with that predicted. The predicted CO concentrations
on, or close to, the burner centerline are lower than the measured values, however, the
98
relative difference decreases with increasing distance from the burner. Meanwhile, the
modelling underestimates the fuel consumption in the front part of furnace. From these
data, the conclusion can be drawn that the flame diffusion in furnace is well predicted since
there is good agreement of the locations and magnitudes of the combustibles, including
O2(dry%)
hydrocarbon, and CO and O2.
12
10
8
6
4
2
0
-800
Measured
predicted
-400
0
Y (mm)
400
800
(a)
12
O2 (%dry)
10
8
Measured
predicted
6
4
2
0
-800
-400
0
Y (mm)
400
800
O2 (%dry)
(b)
12
10
8
6
4
2
0
-800
Measured
predicted
-400
0
Y (mm)
400
800
(c)
Figure 7.2 Predicted and measured O2 profiles in the furnace
(a) x = 0, z = 0.3 m (b) x = 0, z = 0.6m (c) x = 0, z = 1.2 m
99
CO (ppm)
16000
12000
M.
P.
8000
4000
0
-800
-400
0
Y (mm)
400
0
Y (m m)
400
0
Y (mm)
400
0
Y (mm)
400
800
(a)
CO (ppm)
25000
20000
15000
M
P
10000
5000
0
-800
-400
800
(b)
CO (ppm)
16000
12000
M.
P.
8000
4000
0
-800
-400
800
(c)
5000
M
P
CO (ppm)
4000
3000
2000
1000
0
-800
-400
800
(d)
Figure 7.3 Predicted and measured CO profiles in the furnace
(a) x = 0, z = 0.3 m , (b) x = 0, z = 0.6 m,
(c) x = 0, z = 1.2 m, (d) x = 0, z = 2.15 m
100
7.1.4 Features of Combustion and Flow of HiTAC
Simulation studies were performed and the differences between the heat transfer and
combustion features between a conventional high velocity turbulent jet flame and HiTAC
flame were found. The influence of a heat sink on furnace heat transfer with these two types
of burner systems was investigated as well.
Vectors of the in-furnace gas velocity for HiTAC with one-flame HRS are shown in Figure
7.4. A cross section through the fuel and one of air inlets is presented in order to clearly
show the flow characteristic of HiTAC. The combustion air was injected into the furnace
with a velocity as high as 100 m/s. This large injection velocity leads to strong internal
recirculation zones (IRZ). On the other hand, flue gas flows to the root of flame on its way
out through the burner located at the root of flame as shown in figure 7.4. Recirculation
allows good mixing of the combustion air with the flue gases before ignition occurs.
Because of this, the HiTAC mode leads to a lower peak temperature and a larger
combustion volume and consequently a lower NO emission. Thus, the HiTAC flame
stability depends on existence of the strong internal recirculation zones.
101
Figure 7.4 Predicted velocity vectors at a cross section through the fuel and one of the air
inlets in HiTAC mode [m/s]
102
The highest temperature zone was found along the central axis of the furnace for both
analyzed cases (Figure 7.5). The peak temperature zone is further away from the burner
face at a distance of about 1.0 m.
The maximum gas temperature for the HiTAC mode is lower than that for the turbulent jet
flame mode although the combustion air for HiTAC firing mode is preheated up to 1223 K.
For example, for the design operation of burner, comparing HiTAC and conventional firing
modes, the maximum temperature difference is equal to 361K. This is the result of a very
intensive flue gas recirculation created by internal recirculation zones in HiTAC firing
mode.
Figure 7.5 Predicted temperature profile at a cross section through the fuel and one of the
air inlets at HiTAC mode [K]
The predicted gas temperature field uniformity is higher for the HiTAC mode, which is a
known and expected advantage of the HiTAC technology. However, one should be very
103
careful when using ‘more uniform temperature distribution’ to describe HiTAC
performance. This is demonstrated by the fact that the temperature near the burner zone is
very low, for example, 870K for one-burner regenerative burner system studied here. This
is very much lower than the furnace temperature, thus implying that temperature
distribution is not uniform.
Figures 7.6(a) shows the predicted flame volume by means of the above definition.
Figure7.6(b) gives the flame shape derived from the in-furnace gases measured value
according to the same definition. Table 3 shows that both the measured and predicted of the
flame geometric and physical characteristics.
The flame zone is almost limited to the volume of the hypothetical cylinder created by the
air jets because of their strong injection momentum. Consequently it leads to a rather long
flame. The furnace flame occupation coefficient was 3.174% for studied case, and this
value is 13.1 times for the same fuel capacity with a normal high velocity jet burner. The
large flame volume is consistent with the distribution of species.
The predicted and measured results are found to agree well with respect to the shape of the
flame zone. However the measured flame volumes are slightly larger than those predicted.
A possible reason for this is that only half a burner cycle was calculated in order to get the
steady-state condition in this regard. In fact, it is quite possible that the mixing of fuel and
air is in reality more intense than in the modelling due to the periodical switching of the
burner.
104
(a)
(b)
Figure 7.6 Predicted and measured flame shapes and volumes shown by the oxidation
mixture ratio RO=0.99
(a) Predicted (b) Estimated by experimental data
105
The predicted and measured values of the maximum normalized flame lengths are 12.95
and 12.74 respectively, and are calculated as the ratio of flame length to burner diameter as
shown in Figure 7.6 (b). The predicted flame length agrees well with the measured length
within a 4.3% measured error margin. Moreover, the predicted and measured maximum
normalized flame diameters are 2.4 and 2.25 respectively, defined as the ratio of the
maximum flame diameter to the burner diameter, as shown in Figure 3.5 (b). This implies
that the flame spread is in good agreement. However, while the predicted maximum flame
diameter occurs at the end of flame, the measured flame does not exhibit this feature. One
of possible reasons for this is that only half a burner cycle was calculated, and another
possible reasons for this could be the measurement was not taken at the rear of the furnace
(2.6 m from face of the burner), due to the test furnace construction limitation.
Figure 7.7 shows the typical difference from a conventional high velocity turbulent jet
burner. Comparing with the flame between the flame shapes and sizes for a HiTAC burner
(Figure 7.6(a) and a conventional high velocity turbulent jet burner (Figure 7.7). From the
figure it is very clear that the HiTAC flame spreads over a much larger volume than the
conventional flame. The furnace flame occupation coefficient in the case of HiTAC is 15.8
times larger than for the conventional flame mode. The predicted maximum normalized
flame length for the one-flame HRS is 12.95 at the design condition and the predicted
maximum normalized flame diameter is 2.4.
The heat release zone (chemistry reaction zone) for the HiTAC mode is significantly larger
than for the conventional flame mode. By implication, the firing rate in the HiTAC flame is
much smaller than that for the conventional flame mode. The maximum firing rate for the
flame mode is 1.87×106 kW/m3, which is 48.9 times higher than for the HiTAC mode
where the maximum firing rate is 3.82×104 kW/m3. The average flame heat release for the
flame mode is also much greater than that for HiTAC mode. This proportion is consistent
with the flame volume as shown above. It is a known fact that the flame volume is
significantly larger in HiTAC mode for the same type of fuel and the same firing rate.
Moreover, the value of firing rate also implies that the combustion noise in HiTAC mode is
much lower than in the flame mode since any approach that reduces combustion intensity
within a combustion reaction may be expected to reduce the sound power produced by a
flame. Further analysis can be found in references made in [118].
106
(b)
Figure 7.7 Predicted flame shape for Conventional burner
Figure 7.6b gives the measured flame shape according to the previous definition. The flame
zone is almost limited to the volume of the hypothetical cylinder created by the air jets
because of their strong injection. Consequently it leads to a rather long flame. The
measured flame volume was 0.27 m3, and the furnace flame occupation coefficient was
3.77% for studied case. This value is around 35 times larger than that for the same fuel
capacity with a normal high velocity jet burner. The large flame volume is consistence with
the distribution of species. The measured value of flame length was 2.8 m, and its
normalized flame length was 12.74, which was calculated as the ratio of flame length to
burner diameter as shown in Figure 3.5 (b). The flame almost spread the whole furnace
length. Moreover, the measured maximum flame diameter was 0.5 m and its normalized
flame diameter was 2.25, defined according to the ratio of the maximum flame diameter to
the burner diameter, as shown in Figure 3.5 (b). The measured large flame volume also
107
implies that HiTAC has a big chemical reaction zone, and consequently a uniform
temperature field. This is a result of very intensive flue gas recirculation created by internal
recirculation zones. Therefore, for the same total fuel thermal input in a furnace, the heat
release rate becomes more uniform than the normal combustion. Consequently, the
maximum firing rate is decreased. According our previous study numerically, the maximum
heat release rate for HiTAC is only 1/49 times for the same fuel capacity with a normal
high velocity jet burner. The low maximum firing rate also implies that the combustion
noise is possibly small since any approach that reduces combustion intensity within a
combustion reaction may be expected to reduce the sound power produced by a flame.
7.1.5 Heat Transfer Elevation
Heat transfer was evaluated in a test furnace equipped with the one-flame HRS has been
performed, which include a stationary sink and a moving sink. Figure 7.8 summarizes some
results.
Various heat flux densities were obtained depending on the type of charge used. The
highest values were obtained for the stationary heat sink. For the HiTAC mode with a
stationary sink the value of the heat flux density was on average 162.9 kW/m2. For the
flame mode it was in the range 91.4 kW/m2. This indicates that the heat flux density for
HiTAC mode is 1.78 times that for the conventional flame mode on the same type of sink.
The air cooling tubes characterize another distribution of the heat flux density. For the
HiTAC mode the value of the heat flux density was in the range 36.8 to 40.8 kW/m2. For
the flame mode it was in the range 21.0 to 27.2 kW/m2. A 59 % greater average heat flux
density for Case 0 was demonstrated.
The total radiation heat flux density for a stationary heat sink depends very much on the
combustion mode. For the HiTAC mode with the stationary heat sink, the value of the
radiation heat flux density was in the range 191.0 to 205.4 kW/m2. For the conventional
mode it was in the range 108.7 to 117.6 kW/m2. The average difference of total radiation
for these two firing modes is the same proportion as the net heat flux as shown above.
108
Total radiation heat flux density along the furnace wall depends also very much on the
combustion mode. For the HiTAC mode the value of the heat flux density was around 200
kW/m2. This value is similar to the total radiation heat flux on the top of the stationary heat
sink. For the conventional mode it was in the range 138 to 151 kW/m2.
Net heat flux for sink, kW/m2
200
160
120
80
40
0
0
500
1000
1500
2000
2500
3000
Distance from burner face, mm
Case 0
Case 2
Case 4
Case 1
Case 3
Figure 7.8 Predications of heat flux absorbed by the charge along central line on the
surface of sink
Case 0: Test furnace with a one-burner HRS without any charge or heat sink.
Case 1: Test furnace with conventional turbulent jet flame without any charge or heat sink.
Case 2: Test furnace with a one-burner HRS and with a stationary heat sink whose surface
temperature is equal to 20OC and constant,
Case 3: Test furnace with conventional turbulent jet flame with a stationary heat sink whose
surface temperature is equal to 20OC and constant.
Case 4: Test furnace with a one-burner HRS with a moving steel slab which initial surface
temperature is equal to 20OC.
109
7.1.6 Simulation of a Moving Slab
In preparation to calculate a real industrial furnace with a moving heat sink, a moving slab
is assumed in HiTAC test furnace equipped with a one-burner HRS (High-cycleRegenerative burner System). The slabs’ initial surface temperature was equal to 20OC.
The moving slab was assumed to be made of a low carbon steel in the form of a moving
plate. The moving slab was treated as a charge and its heating was calculated for one side
facing the in-furnace processes. The total heat transfer surface of the moving slab was equal
to 2.945 m2 and the heating capacity of the furnace was 1.25t/h. The steel slab moves with
velocity equal to 0.000833 m/s along the furnace length beginning from the furnace “inlet”
located below the burner.
The distribution of the surface temperature of the moving slab heated under the one-flame
HRS are shown in Figure 7.9. The slab’s surface temperature increases gradually and is
fairly uniform across the furnace width. It should be noticed that the slab end temperature is
608.5 K. This is due to the limited length of the test furnace and due to the initial surface
temperature, which is equal to 20 oC. The test furnace can be treated as a short section of a
real heating furnace.
110
Figure 7.9 Predicted temperature distribution of moving slab with one-flame HRS
7.2 Study of the HiTAC furnace with a twin-flame HiTAC burner
7.2.1 Experimental and verification of the modelling prediction.
Experiments measurements and numerical simulations on this semi-industrial-furnace
equipped with two sets of two-flame HRS have been performed. Further more, the
numerical modelling results were validated by comparing with experimental results
including in-furnace species and temperature. In this work, only the profiles of the
compositions of CO and O2 in furnace are compared between predicted and measured. The
predicted were carried out when burner B and D were fired, and the measured values are
average data after several cycles of burner’s firing.
Figure 7.10 shows the comparisons for the concentrations of CO by depth through the
furnace chamber at x=0 for z=1.7m along y direction, at x=−0.475 m for z=1.4m (near
burner B) along y direction, and at x=0 for z=0.75m along y direction, respectively.
Generally, the overall agreement may be accepted considering the complexity of
111
fluctuations of turbulent combustion process in the furnace, the assumptions made in the
numerical model and the uncertainty of the experimental data. The measured shows larger
zone of CO existing than that of the predicted, the most possibility reason is only half cycle
of firing was simulated.
Figure 7.11 presents the measured and predicted concentrations of O2 by depth through the
furnace chamber at x=−0.475 for z=1.4 m (near burner B) along y direction respectively.
The predicted concentrations of O2 follow similar trend to measured values, exhibiting the
same peak concentration location in Figure 4 (b), which is very close to the burner B.
112
5000
Predicted-CO
Measured-CO
3000
2000
1000
0
0
100
200
300
400
500
600
700
Distance from furnace horizontal central plane, mm
800
(a)
5000
CO (ppm,dry)
4000
Predicted-CO
Measured-CO
3000
2000
1000
0
0
100
200
300
400
500
600
700
Distance from furnace horizontal central plane, mm
800
(b)
400
CO (ppm,dry)
CO (ppm,dry)
4000
Predicted-CO
300
Measured-CO
200
100
0
0
200
400
600
Distance from furnace horizontal central plane, mm
800
(c)
Figure 7.10 Predicted and measured CO profiles in the upper part of the furnace
(a) x=0, z=1.7m, (b) x=-0.475 m, z=1.4m, (c) x=0., z=0.75m
113
Concentration of O2,/%,dry)
12
Predited-O2
9
Mearsured
6
3
0
0
100
200
300
400
500
600
700
800
Distance from burner central surface
Figure 7.11 Predicted and measured O2 profiles in the upper part of the furnace at x=-0.475
m, z=1.4m
7.2.2 Effect of flame configurations
Figure 7.12 depicts the temperature distributions on the horizontal plane, including the
central plane, of injection ports of combustion air for different flame configurations which
used in this study are Counter mode, fired by combining of burners B and D, Single-side
mode, fired by combining of burners C and D and Stagger mode, fired by combining of
burners A and D.
T, K
(a)
(b)
(c)
Figure 7.12 Temperature profiles (K) at various firing locations
(a) Counter mode (b) Single-side mode, (c) Stagger mode
114
The peak temperatures of the gases are 1856, 1821 and 1900 K for combustion modes of
counter, Single-side and stagger, respectively. The maximum flame temperature occurs at
the case of stagger mode. Results from experimental study show the same conclusion. This
is the fact that the hot gases recirculation formed by burners for stagger mode reduces the
heat loss from flame.
These maximum temperatures in the furnace at various modes are lower though the
combustion air was preheated to 1311 K. This is a result of very intensive flue gas
recirculation created by the internal recirculation zones.
The highest temperature zones are found to occur in the middle of the combustion chamber
at the burner level, farther away from the burner face. The counter mode gives the longest
distance from burner face where the maximum temperature appears, while the stagger mode
gives the shortest.
Furthermore, because the flame from this burner is less visible (sometimes invisible) than
the combustion flame from a conventional burner because of the oxidizer diluted by the
internal recirculation, it is therefore generally accepted that flame length is not suitable
parameter for characterizing flame size. Instead, it is necessary to demonstrate flame shape
and size using a comprehensive numerical simulation.
Figure 7.13 shows the influences of flame configurations on flame volumes. Because of
difference of burner’s configurations, the flame shape and size are difference.
115
(a)
(b)
(c)
Figure 7.13 Flame shapes shown by the oxidation mixture ratio RO=0.99 for difference
firing modes
(a) Counter mode (b) Single-side mode (c) Stagger mode
116
The flame volume and the furnace flame occupation coefficient are summarized in Table 1.
The Counter flame occupies the smallest fraction of the furnace volume and the Single-side
flame occupies the biggest fraction of the furnace volume. Consistent with this conclusion,
the Counter mode has maximum local peak firing rate and maximum flame combustion
intensity. This implies that the combustion noise in Counter mode is larger than in the other
two modes since any approach that reduces combustion intensity within a combustion
reaction may be expected to reduce the sound power produced by a flame.
Figure 7.14 shows the influences of flame configurations on NO emissions. Generally, the
NO emission is at low level for all studied cases. The differences in NO emissions can be
clearly seen among these cases. The stagger mode has highest emissions of NO and the
Single-side mode leads to lest emissions. The NO emission corresponds to the flame
NO emission, mg/MJ
maximum temperature. A higher the flame temperature leads to a higher the NO emission.
60
40
20
0
Counter
Parallel
Stagger
Firing Configureation
Figure 7.14 Effect of firing configurations on NO emissions
7.2.3 Effect of excess air ratio
The internal recirculation of the flue gas is key issue to successful contemporary
regenerator burner system. Therefore the momentum of the combustion air or fuel defining
117
this internal recirculation plays a very important role in this type of burner. For a given fuel
flux, the excess air ratio plays a central role. The effects of the excess air ratio on NOx level
have been investigated numerically in this study for excess air ratios from 1.04 to 1.45. In
all of these simulations, the thermal input of the fuel to the burner has been kept constant.
Other boundary conditions have been also kept constant during the simulation comparison
in the test case used above. The increase of excess air ratio increases the velocity of
combustion air.
Figure 7.15 depicts the measured and predicted NO emission effected excess air ratio for
the counter mode. The measured NO emission generally agreed with the predicted values,
NO emission, mg/MJ
and exhibited similar changes in ratios of excess air.
60
Predicted
Experied
45
30
15
0
1,0
1,1
1,2
1,3
1,4
Excess Air Ratio for counter mode
1,5
Figure 7.15 NOx emission vs excess air ration during the combustion of counter mode
Generally, NO emission increases with increasing excess air ratio and this trend of increase
become weak after excess air ratio is higher than of 1.3. This is caused by the associated
increase in oxygen availability and flame temperature. On one hand, the velocity of injected
air increases with increase of excess air ratio, more combustion products are entrained into
the root of the flame, thus the temperature of flame decreases. However, the velocity of
increased injected air leads to better mixing, and maximum temperature of flame increases.
On the other hand, combustion reaction rates are depressed because the fuel and air mixture
is diluted by gases entrained from the furnace atmosphere. Therefore, chemical reaction
118
zone is not wholly confined by a burner tunnel or quarl. Consequence, the concentration of
oxygen is quite sensitive to NO formation.
This result also indicates that the production rate of NO is also controlled both by turbulentmixing and by chemical kinetics. When excess air ratio increases a slight from
stoichiometric ratio, both the chemical kinetic (oxygen concentration) and turbulent-mixing
contributes the NO product rate. If the excess air ratio increases further, NO product rate
would be controlled major by turbulent-mixing.
7.2.4 Effect of Fuel/Air injection momentum ratio NO emission
Previous we argued that the actual NOx emission is obscured by the ratio of the fuel and air
injection momentum. In Figure 7.16, NOx emissions for various fuel/air injection
momentum ratio were shown.
It was found that the fuel/air momentum ratio has a strong influence on the NO emission in
the furnace. Less the ratio of fuel/air momentum, larger the NO production, especially
when the ratio of fuel and air momentum is low.
The explanations for this trend have been shown in previous. Summary, the emissions of
NO at this technology depend on the oxygen availability and flame temperature. A higher
ratio of fuel/air injection momentum, for example, a higher momentum of fuel, or a lower
momentum of air, decreases the mixing of fuel and air within the primary combustion zone.
This leads to a lower maximum temperature and more uniform distribution of reactants
inside the flame and larger chemical reaction region. If the changing of momentum of ratio
is caused also by excess air ratio, availability of oxygen plays also very important role for
the emissions of NO.
119
NO emission, mg/MJ
100
80
Predicted-Stagger
Measured-Counter
Predicted-Counter
60
40
y = 0,6204x -1,7991
20
0
0,05
0,15
0,25
0,35
Fuel/Air Injection Momentum Ratio
Figure 7.16 NO variations with the ratio of the fuel/air injection momentum
7.3. OPTIMAL DESIGN OF A HiTAC FURNACE
7.3.1 Flame Entrainment Ratio
Figure 7.17 presents the distribution of Rfe along flame length (excess air ratio equal 1.09).
Here, the flame length is calculated as the distance between the end of fuel nozzle and the
axial location of the oxidation mixture ratio equal to 0.99.
When the normalized distance—defined by the ratio of the flame length to the burner
diameter—equals 1, the entrainment is very small. Above 1, the entrainment increases with
increasing of the flame length as show in Figure 7.17 with 1.09 of excess air ratio. The
highest flame entrainment ratio 4.1 occurs at the distance of 1.41 m corresponding to half of
the flame length approximately. At this cross section, the diameter of the flame is around
0.51m (the normalized flame diameter 2.32). After this section, the flame entrainment ratio
almost keeps constant.
120
6,0
ALFA=1.09
5,0
ALFA=1.15
Entrainment Ratio
ALFA=1.26
4,0
3,0
2,0
1,0
0,0
0
2
4
6
8
10
12
14
z/D
Figure 7.17 Flame entrainment ratio vs. excess air ratio (ALFA)
The internal recirculation of the flue gas is the key issue to a successful realizing the
HiTAC technology. Therefore the momentum of the combustion air and the fuel that
determine this internal recirculation plays a very important role. For a given fuel flow rate,
the excess air ratio plays a central role. The effects of the excess air ratio on NOx level and
flame volume have been investigated numerically, for the excess air ratios from 1.09 to
1.26. In all of these simulations, the fuel input to the burner has been kept constant as
shown in Table 7.1. Other boundary conditions, for example the distribution of the cooling
tube wall temperature, have been also kept constant. The influences of the excess air ratio
on the flame volume, the flame entrainment ratio, Rfe, and NOx emissions have all been
investigated.
Figure 7.18 shows the furnace flame occupation coefficient, RFOC for various excess air
ratios. The RFOC decreases with the increasing of the excess air ratio. For example, when
the ratio of excess air is 1.09, a much larger flame volume appears; around 2.8 times larger
than when λ =1.26. The explanation for this phenomenon is that the lower excess air ratio
limits the mixing of the fuel with the combustion air in the primary combustion region,
leading to a lower combustion rate. Thus, a larger combustion reaction zone is required.
121
2000
3,5
1900
RFOC, %
3,0
1800
2,5
1700
2,0
Temperature, K
4,0
1600
1,5
1,0
1500
1,09
1,15
Air Excess Ration
1,26
Flame-Occupation-Coefficient
The maximum temperature
Figure 7.18 Excess air ratio vs. flame occupation coefficient and peak temperature
It is also apparent that the maximum temperature in the furnace is strongly affected by the
excess air ratio as well. A larger excess air ratio leads to a larger maximum temperature
(Figure 7.12). One of the possible reasons is that the oxygen concentration in reaction zone
increase as excess air ratio increase, which resulting in a higher flame temperature. Another
reason is a smaller flame entrainment at a larger excess air ratio leads a weaker dilution of
the flue gas, thus a higher the flue gas’s peak temperature.
Figure 7.17 presents the distribution of the flame entrainment ratio for different excess air
ratio. Generally, when the excess air ratio increases, the flame entrainment decreases
because of the smaller flame volume as interpreted previous. This also implies that a larger
gross injection momentum does not mean a larger flame entrainment ratio. This is different
from the conventional high velocity jet burner. Furthermore, the profile of the flame
entrainment ratio is also affected by the excess air ratio. When the excess air ratio increases
above 1.09, the entrainment ratio has a maximum value at the half of the flame length.
Figure 7.19 shows the flame entrainment ratio as a function of the normalized distance,
which is defined as the ratio of flame length to burner diameter, for HiTAC test furnace
equipped with two sets of two-flame HRS fired with staggered and counter modes.
122
Flame entrainment Ratio
4
3
2
1
0
0
2
4
6
Normalize distance
8
10
Stag. Tf =293 K
Stag. Tf =573K
Stag. Tf =873
Counter Tf =293K
Figure 7.19 Effect of fuel temperature and flame locations on flame entrainment ratio
The effect of fuel temperature is obvious. When the fuel temperature is 293K, the location
of the maximum flame entrainment is away from the flue gas outlet. In this case the CO
emission is zero. When the fuel temperature is increased, the fuel velocity increases, the
location of the highest flame entrainment is moved more closely to the outlet, and the
combustion efficiency is decreased. According to the simulation, the emissions of CO are
118, 183 and 250 ppm (mass) for the fuel temperatures 573, 873 and 1273 K, respectively.
Based on these results, an optimum furnace width can be designed.
The counter mode offers a different profile of the flame entrainment ratio since the flows of
two firing burners ru n into each other. The profile of the flame entrainment ratio is
symmetrical (Fig.7.18). The value of entrainment increases with flame length from both
burner faces. There are two peaks for the entrainment ratio and they occurr at about 4
burner diameters from both burner faces. The flame entrainment is very small at the middle
of the two firing burners. This is because the flow direction of the flue gases in the flame is
changed from a vertical burner face to single-side burner face as shown in Figure 7.12.
123
Generally, the highest flame entrainment ratio is found at the same normalized distance
(around 3) although they occurred at different locations since the total momentum of the
burners for the two modes studied was kept constant. This could be used to estimate an
optimum furnace design.
7.3.2 Optimal Design of HiTAC Furnace
The purpose of the design of a high performace industrial furnace is to provide a high and
uniform temperature in the furnace and a low NOx emission without sacrifying the
combustion efficiency. HiTAC has the characteristics of slow combustion diluted by
internal flue gas recirculation. In order to maintain enough internal flue gas recirculation for
a specific burner configuration, a minimum flow area in the furnace cross section is
requried. Furthermore, the width of furnace (distance from burner face to exhaust exit)
should also be maintained otherwise the problem of ejection of unburned fuel gases through
the discharge burner arises. In this part of the work these two parameters are discussed.
It is clearly seen from the above results and discussion that a high entrainment ratio of the
flue gas should be maintained. Therefore, according the flame shape and the flame
entrainment ratio, the miminun flowing area of recirculation zone and furnace length can be
determined. The minimum cross section of the flue gas recirculation flow could be founded
at the section where the maximum flame diameter exists.
Figure 7.20 shows the scheme of the flame zone and the flue gas recirculation zone for a
HiTAC furance equipped with any type of HRS. In this study, it is assumed that all the
positive velocities across the cross section of the furnace are included in the flame section.
This is a good assumption in the part of furnace from the burner face to the maximum flame
cross-section.
124
mr
mout
mout-2
mf
mo
Figure 7.20 The scheme of structure of the flame zone and recirculation in an enclosed
combustion chamber
The mass balance for this part of the chamber extending from the burner face to any flame
section is as follows:
m0 + mr = m f + mout
(7.1)
where, mo is the mass flow rate including the fuel and the combustion air, mf is the mass
flow rate through the flame cross-section, and mr is the mass flow rate of the flue gas
recirculation and mout is the mass flow rate of the gases leaving the chamber.
When Equ. 7.1 is divided by m0, one obtains:
mr m f m0 − mout
=
−
m 0 m0
m0
(7.2)
The flow rate, m, flows any cross section can be calculated as:
m = ρv A
(7.3)
Here, ρ , v and A are the flue gases average density, the velocity and area, respectively
According to the definition of the flame entrainment ratio in Eq.3.29, and Equ.7.2, the area
of flue gas recirculation can be calculated from Equ.7.4, expressed as:
125
m0 − mout
)
m0
ρ r vr
mo ( R fe −
Ar =
(7.4)
The minimum cross section of the flue gas recirculation flow could exist at the section
where the maximum flame entrainment ratio, Rfe. The critical area of the flue gas
recirculation, Ar,cr can be determined by the maximum flame entrainment ratio Rfe.
mo − mout
)
mo
ρ r vr
mo ( R fe ,max −
Ar ,cr =
(7.5)
It can be seen that the criteria minimum area of combustion chamber is function of:
•
the maximum flame entrainment rate of the burner, Rfe,max,,
•
the initial total mass flow rates of the burner, mo,
•
density of in-furnace flue gases, thus furnace temperature,
•
velocity of in-furnace flue gases which are effected by furnace temperature and the
length of the furnace along the vertical direction of the burner’s face.
•
the type of the HiTAC system. i.e., one-flame or two-flame HiTAC system.
According to the design of the studied burner system in this paper,
mout = 0.8m0
(7.6)
One obtains:
Ar ,cr =
mo ( R fe ,max − 0.2)
ρ r vr
(7.7)
Here, Ar,cr is the minimum cross area needed for flue gas recirculation, mo is initial total
mass flow rate of the burner. ρ and v are the density and the velocity respectively, r denotes
the recirculated flue gas in the combustion chamber.
126
From the Eq.7.7, it is easy to calculate a minimum flowing area of the recirculation flue
gas. For example, if this one-flame HiTAC system is installed on a cylindrical furnace, the
maximum flame entrainment ratio is 4.1 at the half of the flame length and this flame
diameter is 0.51 m. The temperature of the furnace is around 1373 K, and the average
density is around 0.28. Additionally the average velocity of the recirculating flue gases is
approach 1 m/s in the cross section of Rfe,max at this furnace temperature. Therefore, the
minimum diameter of the chamber is 0.7 m. This is approximately 3.2 times of the burner
diameter.
According to this, it is easy to give the optimal design of HiTAC furnace. For example, for
the stagger firing mode of the burner studied and for the design condition, the Rfe,max is 2.6
according to this calculation. The temperature of the furnace is around 1373 K, and the
average flue gas density is around 0.28. Additionally the average velocity of the
recirculating flue gases is approaching 1 m/s in the cross section of Rfe,max at this furnace
temperature. The area of flame cross section at the Rfe,max is 0.155168 m2, which occurs at
1.2m from burner face.
If this burner is placed in a furnace with a cylindrical combustion chamber, the critical area
of flue gas recirculation is 0.04m2, therefore, the critical diameter is 0.5m. This is
approximately 2.7 times of the burner diameter.
Another parameter for optimal design is the width of the furnace. Figure 7.21 depicts the
distribution of temperature, through the central point of the burner B and D, and flame
entrainment ratio along the direction of the furnace width for HiTAC the test furnace with
two-flame HRS burners in the staggered firing mode. The optimal design of the width of
furnace depends on the optimal location of the peak flame entrainment ratio.
127
3
1800
2
1400
1
Flame temperature
1000
Flame entrainment
600
Flame entrainment ratio
Temperature, K
2200
0
0
2
4
6
8
Normalize Distance
10
Figure 7.21 Distributions of Temperature and Flame entrainment ratio along the direction
of the furnace width
The following three aspects are considered to determine the optimum location of peak
flame entrainment ratio. According to the characteristics of HiTAC, a large entrainment
ratio leads to low NOx emissions. It therefore follows that if the peak temperature occurs at
the location of the peak flame entrainment ratio, the lowest NOx emission will be realized
for a specific type of burner. Therefore, the location of the peak flame entrainment ratio
should be designed to exist as close to the firing burner as possible. Secondly, the maximal
uniform temperature profile in the furnace can be achieved when the peak flame
entrainment ratio is located at the middle of the furnace width. Finally, the combustion
efficiency in the flame increases with the increase in the flame width. In order to achieve
complete combustion, a sufficiently large furnace width is necessary. To summarize, the
maximum flame cross section should be at least half of the furnace length.
Consequently, for a HiTAC furnace fitted with two-flame HRS, the optimum furnace
length is in the range of 1.2−2.4m for the burner studied in a staggered firing configuration.
This is approximately 6.5−13 times the burner diameter.
Similarly, the optimum design of the HiTAC furnace equipped with a one-flame HRS can
also can be determined. It is assumed that the one-flame HRS is installed on a cylinder
combustion chamber, and the temperature of the furnace is 1373 K, the average flue gas
128
density is around 0.28 and the average velocity of recirculation flue gas is approaching 1
m/s at the cross section of Rfe,max at this furnace temperature. For the above design, the
maximum flame diameter is 0.5m and exists at the half of the flame length and the
entrainment ratio is equal to 4.0. Therefore, the minimum diameter of the chamber is 0.7 m.
This is approximately 3.2 times the burner diameter.
It can be seen for the same fuel capacity, for example, 200 kW, that the optimum width of
combustion chamber equipped with two-flame HRS is 1.0 m, this value for one-flame HRS
is 0.7 m. Therefore it can be concluded that the optimum width of a combustion chamber
equipped with a two-flame HRSis larger than that for a one-flame HRS.
7. 4 CONCLUSIONS
A HiTAC test furnace equipped with HRS (both one-flame and two-flame) was studied
experimentally and numerically and the main conclusions are as follows:
1. For the combustion model, the Eddy-Dissipation-Concept with multi-step chemical
reactions is a more suitable numerical model for HiTAC, especially when modelling is
applied to large scale industrial furnaces.
2. The concepts, including oxidation mixture ratio, furnace-gas-temperature-uniformityratio, Furnace Flame Occupation Coefficient and Flame entrainment ratio, were defined
to describe the characteristics of HiTAC, and are used to optimize the design of a
HiTAC furnace and burner.
3. The benefits of HRS are quantitatively demonstrated by mathematical models. They are:
lower peak temperature, larger flame volume, a more uniform thermal field, lower local
firing rate, higher heat transfer, higher energy utilizing efficiency and lower combustion
noise.
4. Operation parameters, including oxygen concentration, combustion air temperature, fuel
temperature, fuel flow rate, excess air ratio and flame locations have strong influences
129
on the combustion and NO emissions in the HiTAC furnace. The optimum combination
of these parameters should be considered
5. The criteria for the choice of the diameter and length of a furnace fitted with HRS
burners are proposed in order to achieve an optimum design for HiTAC operation.
130
8 NO formation and destruction Mechanism study
For HiTAC, the large quantities of recirculation of combustion products are entrained into
the fresh reactants before combustion thus higher peak temperature (T<1850 K) is lack. As
a result, thermal NOx is suppressed and much of NO maybe form mainly by mechanisms
that are insignificant in most conventional combustors. One of most possibility routines is
the NO formation via nitrous oxide mechanism (N2O).
In the present work, N2O-intermediate NO model is developed. This model is accompanied
with exiting thermal, prompt, and reburning NO models for prediction of the NO formation
and emission in a semi-industrial furnace equipped with a HiTAC burner. The sensitivity of
furnace temperature and oxygen availability on NO generation rate has been also
investigated. The predicted results were compared and analysis with experimental values.
The numerical simulations on this semi-industrial furnace equipped with one-flame HiTAC
system have been validated in our previous study by comparing with experimental results
including energy balance, in-furnace species and temperature. The agreements of predicted
and measured are very encouraging, especially the numerical simulation shows a very good
prediction of the HiTAC flame’s shape and size. In the present study, only NO models were
concentrated.
8.1 Experimental and verification of NO emissions models
Figure 8.1 shows the results from experimental measurement and numerical prediction with
and without N2O route at different excess air ratio. The measured values were obtained at
the average of burner outlet and chimney of the furnace. The estimated uncertainties of the
measured NOx values were 95 percent confidence level. The experimental data shown were
the values of NOx (NO+NO2) and the numerical calculations are only for NO. However,
the gradual oxidation of a small amount of NO to NO2 in the immediate post flame gases
does not alter the total NOx emission significantly. Additionally, in order to eliminate the
dilution effect by excess air, the NO concentrations are corrected to mg/MJ fuel.
131
NO, mg/MJ
80
Predicted without N2O route
Mesured
Predicted with N2O route
60
40
20
0
1,0
1,1
1,2
1,3
Air Excess Ratio
Figure 8.1 Effects of excess air ratio on NO emission
Generally, it can be seen that NO emissions are lower than the limit of emission standards
for industrial furnace (100mg/MJ) although combustion air was preheated 1207K.
The numerical calculated NO emission without N2O-route model is lower than that
obtained by experimental measurement of NOx emissions. This difference is clearer at the
cases of lower excess air ratios. For example, the NO model without N2O route could
almost not predict NO emissions in the case of excess air ratio equal to 1.04 because of the
predicted value is almost zero. Although only NO emission was predicted, these differences
are much larger than errors led by ignoring of reading of NO2, which is only in order of
10% during experiments.
When N2O route was added to predict NO emission, the calculated NO emission showed a
significant improvement. For instance, the NO emission obtained based on the models with
N2O route is around 2.7 times larger than that predicted without N2O route in the case of
excess air ratio equal to 1.09. At 1.15 of excess air ratio, this relation is 2.4. The
approximate percentages of NOx production by nitrous and other three mechanisms
(thermal, prompt and reburning) are 99:1, 73:27 and 70:30 at excess air ratios equal to 1.04,
1.09 and 1.15, respectively. This result implies that NO emission via N2O intermediate is
significant and for lower excess air ratio this model will gain more importance. This
conclusion is consistent with the model results in the work [33] in the case of methane
combustion in a stirred reactor.
132
Comparing with experimental data, the results of the NO emission predicted with N2O
route is very encouraging when the excess air ratio is equal or less than 1.15. However,
when excess air ratio is equal to 1.25, predicted NO emission with N2O-route was 52
mg/MJ, which is almost 2 times higher than measurement. In this case, proposed model of
NO emissions via N2O-mechanism fails to predict NO emission, existing NO models
(thermal, prompt and reburning NO) are enough exactly to predict NO emissions as shown
in Figure 8.1.
In order to further understand this result, the variation of the maximum furnace temperature
is given in Figure 8.2 as a function of excess air ratio. It can be seen that the peak
temperature in the furnace rises as excess air ratio increases. This is the fact that the oxygen
concentration in reaction zone increase as excess air ratio increase, which resulting in a
higher flame temperature. When excess air ratio is lower than 1.15, the peak temperature is
lower than 1850 K. When excess air ratio is raised to 1.25, the peak temperature is over
1900 K.
Max. T, K
1950
1850
1750
1650
1,0
1,1
1,2
1,3
Excess Air Ratio
Figure 8.2 Effect of excess air ratio on maximum temperature
Additionally, Figure 8.2 also clearly shows that the NO concentration is very sensitive to
the excess air ratio and increases linearly with excess. This is caused by the associated
133
increase in oxygen availability and flame temperature. On one hand, a larger excess air
ratio leads to increase of oxygen availability, consequently, increase formation of NO via
N2O+O. On the other hand, the velocity of increased injected air leads to better mixing, and
maximum temperature of flame increases, thus NO formation rate increase since NO
formation rate depends on temperature. As reported by Sarofim and Flagan [119] that the
rate of thermal-NOx formation is significant only at temperatures greater than 1823 K and
doubles for every increase in temperature of about 40 K. Therefore, when flame
temperature is lower than 1850K, NO emission rate via N2O-route plays a very important
role. When flame temperature is higher than 1850K, NO formation rate via thermal-NO is
of importance. Present thermal-NO model woks well because of its self-compensating (big
errors in measurements and derived reaction rates [120]).
Therefore, developed NO model via N2O-machanis have to be further improved in order to
obtain high accuracy prediction of NO emissions. This will be further argued in the next
part.
8.2 Effects of NO model on the NO formations
In order to further exam effects of the NO model on NO formation and destruction in the
test furnace, it is very useful to investigate the temperature profiles as shown in Figure 8.5,
which at the cross-section through the fuel and one air nozzle in the test furnace at different
excess air ratio. Generally, there is a very large chemical reaction zone in all cases. This is
also verified by experimental studies. The highest temperature zones are found to occur in
the middle of the combustion chamber at the burner level, farther away from the burner
face. HiTAC burner used in this study limit the mixing of the fuel with the combustion air
at the initial stage of the combustion. Large combustion products are entrained into the root
of the flame because of the high injection momentum of it’s nozzles, this reduce the oxygen
availability in the primary combustion zone. The fuel that is not burned in this region
gradually mixes with air to complete the combustion. This results in a more uniform
temperature profile, which corresponds to a lower NO production rate.
134
Predicted NO contours at the cross-section through the fuel and one air nozzle in the test
furnace at different excess air ratio (λ) with and without N2O-route were shown in Figure
8.3 and Figure 8.4. It can be seen that the predicted NO formations with and without N2Ointermediate mechanism are quite difference. Firstly, the locations of maximum NO
formation are different. For example, when N2O-intermediate mechanism has not been
included in the cases of excess air ratios equal to 1.09 and 1.15, these locations are far from
the burner zone, which is consistent with where the peak temperature appears. This implies
that temperature plays the major role for the formation of NO. However, when N2Ointermediate mechanism was considered, the location of the maximum NO formation is
closer to the burner zone. In particular, these locations are occurred near the nozzles of fuel
and air inlet when excess air ratios equal to 1.09, 1.15 and 1.25. This implies that NO
formation and destruction are strong affected by the radical species as well as flame’s
temperature.
135
(a) without N2O-route
(b) with N2O-route at
at λ= 1.04
λ= 1.04
(c) without N2O-route
(d) with N2O-route at
at λ= 1.09
λ= 1.09
Figure 8.3 Predicted NO contours (mass%) at the cross-section through fuel and one air
nozzle in the test furnace at excess air ratio (λ) equal to 1.04 and 1.09 with and without
N2O-route
136
(e) without N2O-route
(f) with N2O-route at λ=
at λ= 1.15
1.15
(g) without N2O-route
(h) with N2O-route at
at λ= 1.25
λ= 1.25
Figure 8.4 Predicted NO contours (mass%) at the cross-section through fuel and one air
nozzle in the test furnace at excess air ratio (λ) equal to 1.15 and 1.25with and without
N2O-route
137
(a) λ=1.04
(b) λ=1.09
(c) λ=1.15
(d) λ=1.25
Figure 8.5 Temperature distribution at the cross-section through fuel and one air nozzle in
the test furnace at different excess air ratio (λ)
Secondly, the distributions of NO formation with N2O-intermediate mechanism is limited
in the major chemical reaction zone as shown in our previous works, which the flame
volume was figured out by experimental and numerical calculating studies.
138
Basing on these analyses, it could be concluded that the prediction of NO formation and
destruction with N2O-intermediate mechanism is more reasonable.
It is stipulated that for the case of excess air ratio equal to 1.25, the prediction of NO
formation and destruction with N2O-intermediate mechanism is also reasonable. This
implies that developed N2O-intermediate mechanism is able to predict the profiles of NO
formation and destruction in high temperature air combustion, but it gives an over
prediction of NO emission. Therefore, one of improvement solution is considering the limit
of temperature. Therefore, the equation (3.13) can be modified as following:
[
R N 2O _ NO _ i
⎧− min REBU _ NO _ i , RKIN
⎪⎪
=⎨
⎪0
⎪⎩
]
T < 1850 K
T ≥ 1850 K
kg/m3s
(8.1)
This improvement was used to predict NO emission at the case of excess air ratio equal to
1.25. The predicted NO emissions were 35 mg/MJ, and the NO profile in the furnace was
similar as before. This improvement is effective.
8.3 NO emissions From Coflow Gas Jet Combustion Study
In this study here, the following NO formation and destruction models were used:
• Thermal NO formation
• Prompt NOx route
• N2O route
• NO-reburning
The numerical results show that the N2O-intermediate mechanism is very importance
during HiTAC operation at low preheated temperatures. For example, Figure 8.6 presents
139
the influences of oxygen concentration on NO emissions at the case of the co-flow gas jet
in the combustion test case. It shows that when the oxygen concentration is larger than
10%, thermal and prompt NO emissions play a dominating role. However, when the
oxygen concentration is less than 10%, the NO emissions formed from the nitrous oxide
mechanism (N2O) have a substantial influence. For LPG combustion with a 1173K air
preheat temperature, the approximate percentage of NO production by the nitrous oxide to
Zeldovich and prompt mechanism vary from 5:95 at 10% oxygen concentration to 95:5 at
5% oxygen concentration.
Generally, NO emissions decrease with reduction in oxygen concentration because of the
lack of a peak temperature in the flame.
140
2,0
3000
1,5
2000
1,0
1000
0,5
0
0,0
0
5
10
15
20
ratio of NO via N2O and
total
NO emission, mg/MJ
4000
25
Oxygen concentration in combsution air,mass%
Thermal+Prompt-NO
N2O-path-NO
N2O-path-NO/NO
Figure 8.6 Effect of oxygen concentration on NO emissions at an air temperature of 1173K
and fuel temperature of 299K for gas jet combustion in the co-flow test
Furthermore, the increase in the preheat temperature of air leads to an increase of the NO
production. However, this increasing trend is weak at low preheat temperatures. Figure 8.7
gives the variation of the NO emission with the preheat air temperature. A lower maximum
temperature in the flame and NO formation from the N2O-intermediate mechanism can
explain this very well.
141
NO emission, mg/MJ
3000
2500
2000
1500
1000
500
0
1000
1200
1400
1600
Preheated air temperature, K
Figure 8.7 Effect of preheated air temperature on NO emissions at a fuel temperature of
299K for the case of gas jet combustion in the co-flow test
The influence of fuel flow rate on the NO emissions are also considered. Generally, an
increase in the fuel flowrate increases the NO emissions as shown in Figure 8.8. Figure 8.8
also indicates that this increase is much smaller at a lower oxygen concentration than at a
higher oxygen concentration.
NO emission,mg/MJ
2500
2000
1500
1000
500
0
0,E+00
5,E-06
1,E-05
2,E-05
2,E-05
Fuel flow rate QF,kg/s
Oxygen=15,4%(vol.)
Oxygen=9,4%,(vol)
Figure 8.8 Effect of fuel flowrate on NO emission at an air temperature of 1173K and a
fuel temperature of 299K for gas jet combustion in the co-flow test
142
8.4 Conclusions
N2O-route is developed to predict NO formation and emission during high temperature air
combustion because of its lack of peak temperature. This model was used to simulate NO
formation and emission for a semi-industrial furnace equipped with HiTAC burner. The
predicted results were compared with experimental values. The results show that NO
emission formed by N2O-intermediate mechanism is of outstanding importance during low
peak temperature. It can give more reasonable predicted profile of NO formation and
destruction. Developed model provides a basis for further studies. Furthermore, increasing
excess air ratio leads to increasing of NO emission.
143
9 Thermodynamics investigation of the ‘Flameless’
combustion
The available energy of fuel Eav is the difference between Gibb’s energy of reactant at
ambient condition and that of burned produce as the same condition:
E av ≡ Ex I = G I − G E
(9.1)
Exergy, Ex and Gibb’s free energy, G, are expressed by using temperature, T, enthalpy, H,,
and entropy, S:
Ex = H − H 0 − T0 ( S − S 0 )
(9.2)
G = H − TS
(9.3)
Where, subscript 0 denotes the ambient condition.
The thermodynamic analysis is usefulness to describe quantitatively the regime of the
flameless combustion. Here, only the H-T diagram is used.
In Figure 9.1, the HR and Hp curves are the relation of enthalpy of unburned mixture and its
preheat temperature and the relation of enthalpy of chemically equilibrated burned product
and its temperatures, respectively. It can be seen that the HR is increased linearly with the
increasing of the preheated temperature. The difference of the enthalpy of unburned
mixture at different oxygen concentrations in the reactants is not so larger since it
dependence strong on the temperature. However, the Hp-curve increases exponentially in
the high temperature range due to the increase of apparent specific heat caused by the
thermal dissociation in product. The effect of the oxygen concentration on the enthalpy
difference is significant, and the pure oxygen combustion has the biggest enthalpy
difference.
144
The maximum temperature rise in the combustion process of specific condition is an
isenthalpic combustion. This temperature rise is decreased obviously with the oxygen
concentration in the mixture, and the increasing of the reactant temperature. It becomes
zero at the cross point of the two curves, which corresponds to the adiabatic limit
temperature.
Ideally, the combustion without temperature rise can be realised by an isothermal
combustion. The enthalpy difference at the isothermal combustion implies the maximum
available energy. It can be seen that the oxygen-combustion has the maximum available
energy.
12000
Taut
Enthalpy, KJ/kg
8000
Tmax-NOx
4000
0
-4000
-8000
-12000
0
1000
2000
Temperature, K
3000
4000
HR-CH4/O2
Hp-CH4/O2
HR-CH4/21%O2
Hp-CH4/21%O2
HR-CH4/5%O2
Hp-CH4/5%O2
Figure 9.1 Enthalpy-temperature diagram for CH4 mixture with different oxygen level at
the stoichiometric
In order to clear demonstrate the effect of the oxygen concentration and the reactants’
temperature on the temperature rise, Figure 9.2 shows the relationship of the excess
enthalpy
and
inlet
temperature
at
different
oxygen
concentration
for
a
methane/oxygen/nitrogen mixture. Here the excess enthalpy, ∆h, is calculated as:
145
Δh =
Tad − Tin
Tin − Tref
(9.4)
Here, Tad is the adiabatic temperature (K) of the reactant mixture with respect to inlet
temperature Tin (K) of the reactant mixture, K. Tref is reference temperature, taken 273K, for
example in this work.
Figure 9.2 shows that the lower excess enthalpy during the combustion process can be
achieved at a higher reactant temperature for the combustion with all the oxygen
concentrations range.
Excess Enthaply in log( (Tad-Tin)/(Tin-Tref))
100,0
Nonisothermal
Combustion
10,0
100 % oxygen
1,0
Quasiisothermal
Combustion
21% oxygen
5% oxygen
0,1
0
1000
2000
3000
Temperature of the Reactant, K
Figure 9.2 Excess enthalpy in log versus inlet temperature at different oxygen
concentration for a methane/oxygen/nitrogen mixture
Therefore, it is easy to find the ways to get a low temperature rise in the combustion
process, for example, the followings may be taken:
• preheating unburned mixture,
• decreasing oxygen concentration (or lean combustion) in the unburned mixture, or
• increasing the heat loss from the flame, or flame cooling
146
Preheating unburned mixture is very effective way to increase the energy utilization for the
industrial furnace, where a larger enthalpy of the flue gas exists. This technology has been
well studied. In fact, this is also effective solution to improve temperature uniformity.
If the second way is used, the combustion stability should be considered. A preheated
temperature is higher than the fuel’s autoignition temperature is preferred, which is widely
used in the application of the furnace. We can also expect that any other ways to increase
fuel flammability limitation technology might be used as well, for example, the catalytic
combustion.
The flame cooling is another technology to improve the combustion temperature
uniformity. There are many different techniques which could be classified as flame cooling.
A flame cooling tube might be also used to increase flame heat loss in order to get a low
temperature rise during the combustion processor. The porous combustion is another
example, where the heat released in the oxidative region is transferred to the unreacted
mixture passing through a conduction channel inside the solid material or a radiative one
inside the pores. Direct Flame Impingement (DFI) can be also classified this catalogue.
Nitric oxide could be ranked to the most relevant pollutant when the combustion process is
employed in the industries. Thermal NOx is favour when the temperature is higher than
1800K. It is well knowledge that the most effective way to depress NOx production is low
the flame temperature for the high-temperature industrial application of fossil fuel
combustion technologies.
Obviously, the HiTAC, or MILD, or Flameless combustion concept is a combination of the
first two items (and the third). The major characteristics of this available knowledge are:
•
The reactants must exceed self-ignition temperature,
•
This process evolves in a rather narrow temperature range especially when NOx
emission is considered, and
•
The available energy is smaller, which leads to a low combustion intensity.
It is generally agreed that the major characteristics of the HiTAC (or MILD) combustion is
led by the low temperature increase in the combustion process. This is the fact that a low
temperature increase makes a combustion chamber more similar as a well-stirred-reactor.
147
Again, the effect of heat release on the combustion characteristic is less at the case of a low
temperature rise. For example, a larger flame entrainment is found during the HiTAC
condition in or previous work.
It is possible to liberate the above three limitations if only the mild temperature rise during
the combustion process is considered. For example, the use of oxyfuel combustion can
increase the combustion intensity and maintain a mild temperature rise as studied in this
work.
Basing on the above discussion, this new combustion phenomenon might be named as a
Quasi-Isothermal Combustion, or QIC.
It is obviously that the maximum allowable excess enthalpy of the QIC is decided by the
fuel and oxidizer’s physical properties, and the combustion process, such as inlet,
operations parameters. Additionally, the combustion stability should be maintained. The
temperature of the reactants is higher that the fuel auto ignition is preferred.
Generally, for the QIC, the maximum allowable excess enthalpy of the reactant mixture
with respect to inlet temperature during combustion may be decided by the maximum
excess enthalpy of the oxyfuel combustion (100% oxygen concentration) with respect to its
self-ignition temperature. For methane, this value is 2.93 as shown in Figure 9.2.
The studied burner in this paper is also one example of the QIC. It is realised by decreasing
oxygen concentration before the combustion occur, and by flame cooling through the fluegas internal recirculation.
148
10 Conclusion
The investigation has been performed using a single fuel jet flame facility with cross and
co-flowing, and a semi-industrial furnace equipped with HiTAC burners. Experimental,
numerical and theoretical analyzing investigations are adopted.
In this work, a ‘chemical’ flame volume and ‘chemical’ flame length were used to describe
this ‘invisible flame’. Results from single jet flame study show that:
•
Flame length increases with either the decrease of oxygen content, or increase of
oxidizer temperature, or decrease of fuel temperature. Furthermore, the flame length is
independent of the fuel flow rate and the diameter of the fuel nozzle for the studied
cases.
•
Flame volume increases either with the decrease of oxygen content and increase of
oxidizer temperature, or with the reduction of fuel temperature, or with the increasing in
fuel firing rate. Flame volume depends very much on the oxygen concentration in the
oxidizer.
•
Influences of high temperature and low oxygen concentration in the oxidizer on the
flame Froude number, Frf were examined. Regimes of momentum- or buoyancy-control,
were determined on the assumption that oxidizer temperature and oxygen concentration
are changeable. A simple correlation of the ‘flame’ length and volume with flow
parameters has been derived in terms of a flame Froude number for momentumbuoyancy transition jet flame under the HiTAC condition. The criteria constants of the
dimensionless flame volume V* and the dimensionless flame length L* to assess
momentum– or buoyancy–control flame are given.
Additionally, the entrainments of this ‘invisible’ flame have been numerically and
theoretical studied. Conclusions are:
•
The uniformity of the heat release in reacting jets has strong effect on the flame
entrainment. More uniform the heat release, larger the entrainment. The effect of heat
release reduces the entrainment in the near field of the reacting jets with the same
factor of the characteristic ratio (Tf/To)0.5.
149
•
The entrainment increases as the oxygen concentration is decreased. Furthermore, the
entrainment is independent of the fuel flow rate and the preheated temperature of the
oxidizer for the investigated temperature range (1073-1573K).
•
The effect of the oxygen concentration and preheated temperature of the oxidizer on
buoyancy was examined. A correction Richardson coordinate, where the effect of the
oxygen concentration (stoichiometric ratio) is included, was derived to describe the
local influence of buoyancy force along the chemical flame length under the high
temperature and oxygen deficient oxidizer condition. It can be concluded that the
buoyancy force increases with the reduction of the oxygen concentration in the
oxidizer.
•
The global behaviour of the entrainment was revealed. The entrainment of jet flames
can be identified as two regimes: (a) the near field where entrainment coefficient is
positive; and (b) the far field where entrainment coefficient is negative. Corrections of
entrainment rates were derived in terms of a Frf number for momentum-buoyancy
transition jet flame under the high temperature and low oxygen concentration oxidizer
condition. Furthermore, the maximum entrainments along the flame length are
estimated
Further on, the benefits of HiTAC technology are quantitatively demonstrated by
mathematical models. These benefits are: lower peak temperature, larger flame volume,
more uniform thermal field, lower local firing rate, higher heat transfer, higher energy
utilizing efficiency and lower combustion noise.
NOx formation and destruction during this new combustion phenomenon has been studied
numerically. It was found that the NO formation via N2O mechanism may be important.
The approximate percentage of NO production by the nitrous oxide according to the
Zeldovich and prompt mechanism varies from 5:95 at 10% oxygen concentration to 95:5 at
5% oxygen concentration.
Finally, a thermodynamic analysis of combustion process shows that oxyfuel combustion is
able to increases the available energy of the flameless combustion, thus higher combustion
intensity. Additionally, the flameless oxyfuel burner does not need preheating the oxidizer,
this extend the concept of the HiTAC.
150
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