Proceedings
of ASME
Turbo
Expo
2010:
Sea and Air
Proceedings
of ASME Turbo
Expo
2010:
Power
forPower
Land,for
SeaLand,
and Air
GT2010 GT 2010
July
14-18,
2010,
UK
June 14-18, 2010, Glasgow,Glasgow,
UK
GT2010-22300
GT2010-
METHANE OXY-COMBUSTION FOR LOW CO2 CYCLES: MEASUREMENTS AND
MODELING OF CO AND O2 EMISSIONS
Alberto Amato
School of Mechanical
Engineering
David Scarborough
School of Aerospace
Engineering
Robert Hudak
School of Aerospace
Engineering
Peter A. D’Carlo
School of Aerospace
Engineering
Jerry M. Seitzman
School of Aerospace
Engineering
David R. Noble
School of Aerospace
Engineering
Tim C. Lieuwen
School of Aerospace
Engineering
Georgia Institute of Technology
Atlanta, GA, 30332-0150
ABSTRACT
Concerns about green house gas emissions have
encouraged interest in hydrocarbon combustion techniques that
can accommodate carbon dioxide capture and sequestration.
Oxy-fuel combustion, where the fuel is combusted in oxygen
diluted with steam or CO2, is seen as one promising approach.
In this paper we focus on CO2 dilution effects and, in particular,
on CO and O2 emissions from these flames. The emissions
issue must be considered from a different perspective than
conventional power plants as the combustor effluents will be
sequestered, and, thus, their interactions with the terrestrial
atmosphere are not relevant. Equilibrium CO emissions from
CO2 diluted systems are much higher than conventional air
systems. In addition, for carbon sequestration application,
pipeline specifications impose limitations on CO and O2 levels
which also must then be controlled. Equilibrium and kinetic
modeling of CH4/O2/CO2 combustion systems was performed in
order to analyze CO2 dilution effects upon CO and O2 emissions
level. Companion experiments were also performed in an
atmospheric pressure, swirl stabilized combustor. These
numerical and experimental results quantify the strong
dependence of emissions on stoichiometry, CO2 dilution and
residence time.
hydrocarbon fuels into a high hydrogen mixture, or 2) postcombustion capture of CO2, which can be accomplished with
amine membranes (MEA), solvents or condensation processes.
This CO2 can then be sequestered in the ocean or deep saline
aquifiers, used for enhanced oil recovery (EOR), or for
enhanced coal bed methane recovery (ECBM) [5, 6].
The work described here is directed toward post flame
capture. In such cases, the combustion process typically
involves burning oxygen rather than air, in order to avoid the
large volumes of N2 which would otherwise need to be dealt
with. The flame temperature is controlled by diluting the
oxygen with either steam or CO2. This largely eliminates NOx
formation and enables the exhaust stream to be separated into
concentrated CO2 and water by a simple condensation process.
Several power cycles utilizing oxy-fuel combustion have been
proposed that either consist of slight modifications of existing
gas turbine combined cycles [7], as completely new concepts
(e.g., Graz cycle [8, 9], Matiant cycle [10]), as integration with
gasification processes for coal combustion [11-13], or even for
internal combustion engines for automotive applications [14].
Key challenges to the economical and technical feasibility of
oxy-fuel combustion power cycles concern oxygen production
by the air separation unit (ASU), the aerodynamic design of the
gas turbine which must be modified because of difference in
working fluid properties from that of air, and the combustion
process, which is the focus of this study.
Using CO2 as a diluent to burn natural gas impacts the
flame through changes in: 1) mixture specific heat and adiabatic
flame temperature, 2) transport properties (thermal
conductivity, mass diffusivity, viscosity) [15] 3) chemical
INTRODUCTION
Increasing concerns about climate change have promoted
interest in zero-CO2 emission hydrocarbon combustion
techniques. This can be done by removing carbon in essentially
two ways [1-4] 1) pre-combustion capture via reforming of
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kinetic rates [16-22], and 4) radiative heat transfer [23-26].
Recently, there has been a significant amount of work on
O2/H2O diluted CH4 flames by workers at NETL ([27, 28]),
who have reported results of calculations and large scale tests.
In addition, recent demonstration tests have been performed by
Clean Energy Systems, under support of the US DOE [29],
while extensive experimental and numerical studies of oxynatural gas fired furnaces have been conducted at the IFRF in
the framework of the OXYFLAM project [30, 31] and at the
Swedish Department of Energy and Environment [32]. Our
objective in this study is to focus on the O2/CO2 problem.
Burning methane in an O2/CO2 diluent raises new combustion
opportunities and challenges for both emissions and operability.
As operability issues are discussed in a companion paper [33],
we focus attention here on emissions.
The emissions issue must be considered from a different
perspective than conventional power plants as the combustor
effluents will be presumably sequestered, and, thus, their
interactions with the terrestrial atmosphere are not relevant.
However, the system’s emissions are still of interest. For
example, it is known that CO emissions are augmented due to
higher equilibrium levels as well as slower CO oxidation rates
[34]. CO2 competes with O2 for atomic hydrogen, leading to
of
CO,
mainly through the
reaction
formation
CO2+H↔CO+OH [35]. High levels of CO are, in essence, a
loss in efficiency, and results in increased fuel costs for the
same power output; besides high local CO levels may have
corrosion implications. For similar reasons, exhaust O2
emissions are also of interest because of the cost and power
consumption associated with generating O2 and possible
corrosion in pipeline for CO2 transportation [36].
Diluted oxy-combustion systems have a different degree of
freedom than conventional air systems, where variations in
stoichiometry and fuel/air ratio are coupled. In these systems,
stoichiometry and fuel/air ratio are decoupled. As such, it
seems clear that such systems will be operated at or near
stoichiometric to minimize both fuel and O2 usage, with an
O2/CO2 ratio set by temperature limitations. We see at least
three combustion issues which must be addressed for
CH4/O2/CO2 systems, summarized in Figure 1, which plots the
two key degrees of freedom of this problem: stoichiometry and
flame temperature. High stoichiometries and flame temperatures
are associated with excessive CO emissions, while low
stoichiometries and high flame temperatures with excess O2
emissions. As such, lower temperature considerably broadens
the fuel/air ratio sensitivity of these emissions. However, low
temperature operation is limited by blowoff. The rest of this
paper focuses on measurements and calculations of these CO
and O2 emissions.
Figure 1. Operational space of a premixed, CH4/O2/CO2 flame in the
equivalence ratio – flame temperature space.
EQUILIBRIUM TRENDS
Equilibrium calculations provide a useful perspective on
emissions trends, given a system with a sufficiently long
residence time. This section presents example calculations of
these equilibrium tendencies, obtained used the GasEq solver
[37], at pressure p=1atm or 15atm and at initial temperature
Tin=500F(~533K).
Figure 2. Dependence of adiabatic flame temperature upon the mole
fraction of oxygen in the reactants (Tin = 533K, p = 1atm).
First, we consider the dependence of adiabatic flame
temperature Tad upon oxygen levels in the reactant stream,
shown in Figure 2. As expected, increasing O2 levels in the
oxidizer mixture increases adiabatic flame temperature.
Figure 3 plots the dependence of equilibrium CO concentration
upon flame temperature for CH4/O2/CO2 flames at different
equivalence ratios (colored lines). Also shown for reference is a
methane/air flame (black lines), where flame temperature is
varied with equivalence ratio (bottom loop is lean conditions,
top is rich).
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These results illustrate that adiabatic flame temperature and
stoichiometry exert powerful, but generally opposing influences
upon CO and O2 levels. In order to facilitate observing these
trends simultaneously, Figure 5 plots iso-contours of O2 and CO
levels at p=1atm. This figure illustrates the regions of high and
low sensitivity and the opposing trends of O2 and CO in most
situations.
Figure 3. Dependence of CO equilibrium concentration in wet
combustion products on adiabatic flame temperature, pressure and
equivalence ratio (Tin = 533K, p = 1atm and 15atm).
Figure 5. Dependence of CO and O2 equilibrium concentration (in ppm)
in dry combustion products on adiabatic flame temperature and
equivalence ratio: solid lines refer to CO while dashed lines to O2 (Tin =
533K, p = 1atm).
CHEMICAL KINETICS CO AND O2 CALCULATIONS
Equilibrium calculations are useful for providing insight
into limiting emissions levels that are approached given a
sufficiently long residence time. In reality, the actual combustor
residence time is finite and so emissions levels do not actually
reach their equilibrium levels. As such, kinetic calculations are
needed to determine the characteristic times associated with CO
and O2 evolution to determine their values as a function of the
actual residence time.
Calculations were performed using the PREMIX module in
CHEMKIN 4.1, using the GRI 3.0 mechanism. The t = 0 point
was defined as the point where the heat release rate drops to
10% of its maximum value, as shown in Figure 6. For
reference, a typical value of combustor residence time (defined
by ratio of combustor volume to burned gas volume flow rate)
for an F-class gas turbine is about 40ms.
Figure 4. Dependence of O2 equilibrium concentration in dry
combustion products on adiabatic flame temperature, pressure and
equivalence ratio (Tin = 533K, p = 1atm and 15atm).
Note that CO levels are much higher for the CO2 diluted system
than the fuel lean air system at a given temperature. They are
also lowest at lean conditions and increase monotonically with
equivalence ratio.
Increasing pressure and decreasing
temperature reduce equilibrium levels for lean mixtures, as
might be expected from Le Chatelier’s principle. These pressure
and temperature sensitivities are much lower for rich mixtures,
reflecting the shift in key mechanism controlling equilibrium
levels at these high stoichiometries (water gas shift reaction).
Similarly, Figure 4 shows that the concentration of O2 increases
with flame temperature and decreases with equivalence ratio
and pressure. The pressure and temperature sensitivity to
fuel/air ratio is inverted from the CO case, reflecting the fact
that O2 is a major and trace equilibrium product under lean and
rich conditions, respectively, opposite that of CO.
CO emissions
We first consider CO concentrations. There is a very
important difference between the combustion process
considered here and lean, premixed combustion that should be
emphasized. This can be understood by considering the ratio
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between the maximum value of CO and its equilibrium value, as
shown in Figure 7.
Figure 7. Dependence of CO overshoot upon flame temperature and
stoichiometry. Overshoot is defined as the maximum concentration
divided by equilibrium concentration. (Tin = 533K, p = 1atm and 15atm).
Figure 6. Procedure for defining the end of the flame region (t = 0) for
a typical heat release profile from premixed flame calculations.
Note the large “CO overshoot” in the flame relative to
equilibrium for the fuel lean condition (φ=0.9) at atmospheric
pressure. At 1800 K, for example, peak CO levels in the flame
exceed equilibrium levels by a factor of nearly 30. Even larger
overshoots are present at leaner equivalence ratio. Increasing
pressure increases CO overshoot values at lean equivalence
ratios further. This overshoot is a major design challenge for
lean, premixed systems which must achieve both acceptable CO
levels at low power and NOx at higher power. Significant
attention is given to designing systems to achieve sufficient
residence time for CO burnout, and minimal CO quenching by
cooling holes. Moreover, insufficient residence time for CO
burnout limits turndown for industrial gas turbines, as low
power operation is not possible without CO levels that exceed
permitting restrictions.
For the oxyfuel system running slightly rich, this sensitivity is
not present because of the small value of the “CO overshoot”.
While CO levels are substantially higher than lean systems,
there is almost no overshoot – typical values are 2 or 3. This
implies substantially less sensitivity of exhaust CO levels to
combustor design details. In other words, similar CO levels
will be experienced regardless of quenching by walls or dilution
jets. Another way of looking at these data is to consider the
actual value of maximum CO concentrations in the flame, as
plotted in Figure 8. This figure shows that the peak CO levels
in the flame are much less sensitive to stoichiometry than
equilibrium values.
Figure 8. Maximum value of CO concentration reached in premixed
flame. (Tin = 533K, p = 1atm and 15atm).
We next consider post-flame relaxation of CO emissions
toward equilibrium. Figure 9 plots the dependence of CO
concentration on adiabatic flame temperature, under both
equilibrium and fixed residence time conditions, where
τres=40ms. Notice the strong influence of flame temperature
upon combustion relaxation time. At high temperatures (T >~
2000K), the fixed residence time and equilibrium values are the
same. At lower temperatures, the fixed residence time and
equilibrium CO concentrations diverge progressively as the
temperature decreases. Note that in all cases, the fixed
residence time result is less sensitive to temperature than the
equilibrium result.
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Figure 9. Dependence of CO concentration in wet combustion products
at τres=40ms upon adiabatic flame temperature (Tin = 533K, p = 1atm).
Figure 11. Dependence of CO concentration in wet combustion
products upon adiabatic flame temperature (Tin = 533K, p=15atm).
To illustrate the effect of pressure, Figure 10 plots the CO
concentrations at fixed residence time and different pressures.
As it can be observed, CO concentrations at 15atm are
considerably lower than those at 1atm especially in lean
conditions, where CO concentrations are lower. This is due to
the combination of two effects, (i) the lower CO equilibrium
concentration (as described in the previous section) and (ii) the
faster CO relaxation kinetics at higher pressures. The latter
effect can be seen from Figure 11, which compares the fixed
residence time and equilibrium results at 15atm. Notice that the
two curves diverge from each other at temperatures much lower
than in the 1atm case, meaning that equilibrium values are
reached faster at high pressure.
Figure 12. Dependence of O2 concentration in wet combustion products
at τres=40ms upon adiabatic flame temperature. (Tin = 533K, p = 1atm).
O2 emissions
We next consider the trends for O2 concentrations. We start
in Figure 12 showing the dependence of O2 concentration on
adiabatic flame temperature both at equilibrium and at τres =
40ms. Similar conclusions as discussed for CO (see Figure 9)
can be seen here as well. An important point that should be reemphasized is the substantially reduced temperature sensitivity
of the fixed residence time results relative to the equilibrium,
even more the case here than for CO. To illustrate, consider
φ=1.05. Equilibrium considerations suggest that O2 emissions
can be substantially reduced by decreasing flame temperature.
For example, there is a two order of magnitude difference
between levels at T = 1800 and 2200 K. The important
dependence of relaxation times upon temperature substantially
changes the fixed residence time result. Note that the difference
between O2 emissions at fixed residence time for these two
temperatures changes from a 100x reduction to a 2x reduction.
Figure 10. Dependence of CO concentration at τres =40ms in dry
combustion products upon adiabatic flame temperature at two
pressures. (Tin = 533K, p = 1atm and 15atm).
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EXPERIMENTAL FACILITY
Parallel measurements were obtained to study emissions
trends in a swirling, atmospheric combustor shown in Figure 15.
This system is a duplicated from an experimental rig at Sandia
National Laboratories [38]. The facility consists of a
swirler/nozzle, combustor, and exhaust sections. Premixed gas,
consisting of CH4 and O2/CO2 mixture flows through a
swirler/nozzle section. The nozzle is an annular tube with inner
diameter of 28mm. The center body has an outer diameter of
20 mm. The overall flow area remains constant at 3.0 cm2
inside the nozzle. Tests were performed with a six-vane, 45o
swirler, which is located in the annulus between the centerbody
and nozzle wall. The theoretical swirl number is 0.85. The fuel
is injected 150 cm upstream of the combustor to achieve a
premixed condition. The combustor consists of a 508 mm (20
inches) long quartz tube, with a 115 and 120 mm inner and
outer diameter, respectively. It rests in a circular groove in a
base plate. The O2, CO2, and fuel flow rates are measured with
mass flow meters. Reactant mixtures were preheated to 500F (~
533K).
Particular attention was paid in order to eliminate air
leakages, which is an important “contaminant” in the low O2
exhaust. The combustor was insulated with ceramic cloth and
sealed at the bottom and top edges of the quartz tube using an
8” steel pipe. The 8” pipe features flanges on either end for
attachment of the combustor base plate and an exhaust plate to
allow for sealing of the quartz from the atmosphere. Without
these sealing provisions, we found that atmospheric air is pulled
into the quartz tube, leading to inaccurate O2 measurements.
To illustrate pressure effects, Figure 13 compares the
temperature dependence of O2 concentration at fixed residence
time at two pressures. Again, similar results as in the CO case
occur (see Figure 10): there are significant differences between
the two pressures due to faster relaxation times and lower
equilibrium values at high pressure. The latter effect is
illustrated in Figure 14 (as done in Figure 11), where the fixed
residence time and equilibrium results at 15atm are shown as a
function of adiabatic flame temperature. Notice that the two
curves diverge from each other at temperatures much lower than
in the 1atm case, meaning that equilibrium values are reached
faster at high pressure.
Figure 13. Dependence of O2 concentration at τres=40ms in dry
combustion products upon adiabatic flame temperature for different
pressures (Tin = 533K, p = 1atm and 15atm).
Figure 15. Combustor configuration used for emissions sampling.
Assembly on left and cut-away view of model on right (dimensions in
mm).
Figure 14. Dependence of O2 concentration at τres =40ms in dry
combustion products upon adiabatic flame temperature for different
pressures (Tin = 533K, p = 15atm).
O2 emissions were obtained with an Advanced Instruments
Inc. GPR-1200 Portable ppm Oxygen Analyzer. This O2
analyzer has the ability to auto range to four different scales, 010 ppm, 0-100 ppm, 0-1000 ppm, and 0-25%, allowing for the
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full spectrum of O2 emissions of interest to be measured with
the same device. The sensor was zeroed with pure nitrogen and
calibrated at 800 ppm oxygen in a nitrogen bath calibration gas.
CO emissions were obtained with a LAND Lancom Series II
Combustion Products Analyzer. This analyzer was zeroed with
air and calibrated with 3000 ppm and 30,000 ppm CO in
nitrogen bath calibration gases.
The combustion gases were sampled with a water quenched
probe illustrated in Figure 16. The probe was designed using
three concentric tubes with 1/8”, 1/4”, and 3/8” outer diameters.
Assembly of the probe was performed such that the incoming
cooling water would flow between the 3/8” and 1/4” tubes and
then stagnate at the probe tip and return the heated water
between the 1/4” and 1/8” tubes. This design allows the
emissions sample to cool very rapidly in order to freeze the O2
and CO reactions. The sample was then filtered through a
coalescing filter to extract any water prior to the vacuum pump
used to draw out the emissions sample. After the vacuum pump,
the sample was then run to either the O2 sensor or CO sensor.
Care was taken to use all stainless steel tubing and to minimize
leaks or losses in order to acquire quality O2 measurements.
Calculations using Henry’s law [39] show that, assuming an
average temperature of 350K inside the probe, up to 0.3% of
the oxygen can dissolve in the condensing water which would
not then be measured by the O2 sensor.
2000, 2050, and 2100 K). No measurable sensitivity to
stoichiometry at a fixed adiabatic flame temperature was
observed. Note the slightly higher temperature loss relative to
the CH4/air case, probably reflecting increased radiation
emissions from carbon dioxide.
In addition, axial traverses were performed for a CH4/CO2/O2
flame at φ = 1.10 and Tad = 1900K, as reported in Figure 18.
The axial temperature gradient revealed an expected trend, with
temperature decreasing with downstream distance.
Figure 17. Difference ∆T between measured flame temperature T and
adiabatic flame temperature Tad for CH4/Air and CH4/CO2/O2 flames.
Figure 16. Internal assembly of water cooled sample probe (a) and
sample probe assembly with water inlet and outlet on right (b).
Flame temperature measurements were obtained using a
grounded, K-type thermocouple from Omega Engineering,
model number KMQXL-125G-24. The sheath material of this
thermocouple is Nickel-Chrome based, providing good
performance in heavily oxidizing environments. Radiation
losses were minimized by inserting the 0.125” diameter
thermocouple into a 0.145” diameter ceramic sleeve.
Figure 17 plots measured temperatures for CH4/Air and
CH4/CO2/O2 flames, where ∆T represents the difference
between the measured flame temperature T and the theoretical
adiabatic flame temperature Tad. This difference ranges
between 300 and 400K. The two curves “with ceramic” and
“without ceramic” illustrate the effect of radiation losses from
the thermocouple tip. The same measurements (with ceramic)
were repeated for CH4/CO2/O2 flames, as reported in Figure 17:
two fuel-air ratios (φ = 0.90 and 1.10) were tested for different
theoretical adiabatic flame temperatures (Tad = 1900, 1950,
Figure 18. Temperature measurements for CH4/CO2/O2 flame at φ =
1.10 and Tad = 1900K. Axial position is measured starting from the
dump plane.
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COMPARISON OF EXPERIMENTAL AND NUMERICAL
RESULTS
Comparison of the 1-D model results described earlier and
experimental data for this swirling burner is not straightforward,
primarily because of uncertainties associated with residence
time (because of flow recirculation in the vortex breakdown
bubble). In addition, heat losses cause the actual flame
temperature to be significantly lower than the theoretical
adiabatic flame temperature, as already shown by the data
reported in Figure 17 and Figure 18. This effect was partially
compensated for by using the actual measured temperature at
the probe locations. Finally, CO and O2 levels are both quite
sensitive to variations in stoichiometry and temperature in
parameter space regions where their concentrations are low,
implying that small uncertainties in temperature and flow rate
can lead to significant variations in predicted O2 and CO levels.
All the data reported here were measured at a fixed probe
location, 46 cm from the base plate, in the combustor center.
The flow rates were adjusted at each flame temperature in order
to give the same average burned combustor velocity of 10m/s,
based upon a calculated adiabatic flame temperature. As such,
this probe location was estimated to correspond approximately
to a bulk residence time of ~46ms.
visualization of various trends, Figure 20 replots these data as a
function of equivalence ratio.
Figure 20. Comparison between calculated and experimental data for
CO concentration in dry products: CO levels are represented as
function of equivalence ratio at different flame temperatures (Tin = 533K,
p = 1atm).
Figure 21. Comparison between calculated and experimental data for
O2 concentration in dry products: O2 levels are represented as function
of equivalence ratio at different flame temperatures (Tin = 533K, p =
1atm).
Figure 19. Comparison between calculated and experimental data for
CO concentration in dry products: CO levels are represented as
function of flame temperature at different equivalence ratio (Tin = 533K,
p = 1atm).
Figure 21 plots the dependence of the measured O2 emissions
on adiabatic flame temperatures at different equivalence ratios.
Similar to CO emissions, the measured O2 emissions disagree
with the numerical model primarily in the regions of high
sensitivity (where values are kinetically controlled, rather than
by equilibrium) described previously. However, the same
qualitative trends are observed both in numerical and in
experimental results; i.e., the O2 emissions monotonically rise
with flame temperature and decrease with stoichiometry. To
facilitate the visualization of these trends, Figure 22 plots these
same experimental data as a function of equivalence ratio at
different flame temperatures.
Figure 19 plots the dependence of the measured CO emissions
on temperature for different equivalence ratios. It can be
observed that the measured CO emissions are in good
agreement with the model at high equivalence ratios. Near
stoichiometric conditions, where the sensitivity to experimental
uncertainties is greater, the difference between numerical and
experimental values is larger. However, the same qualitative
trends are observed both for numerical and experimental
measurements; i.e., the CO emissions monotonically rise with
flame temperature and stoichiometry. To facilitate the
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Copyright © 2010 by ASME
4.
5.
6.
7.
Figure 22. Comparison between calculated and experimental data for
O2 concentration in dry products: O2 levels are represented as function
of equivalence ratio at different flame temperatures (Tin = 533K, p =
1atm).
8.
CONCLUDING REMARKS
Emissions concerns associated with high CO and O2 levels
place limitations on the flame temperature and feasible
stoichiometries of operation of premixed methane oxy-fueled
systems. Equilibrium CO emissions are higher than in air due to
higher CO2 levels. In addition, at a fixed residence time, CO
emissions are also higher because of slower burnout of the
intermediate CO formed in the flame. Equilibrium CO
emissions are an exponentially increasing function of flame
temperature. However, this temperature sensitivity is
substantially reduced at a fixed residence time. Pressure lowers
CO emissions because of faster chemical kinetics and decreased
equilibrium levels. CO emissions, thus, are minimized by
operating at low stoichoimetry, flame temperature and high
pressure. Equilibrium O2 emissions are also an exponentially
increasing function of flame temperature. Similar to CO, this
sensitivity is substantially reduced at a fixed residence time. O2
emissions, thus, are minimized by operating at high
stoichiometry, low flame temperature and high pressure.
9.
10.
11.
12.
13.
14.
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Copyright © 2010 by ASME
Copyright © 2010 by ASME