EPA-650/2-75-061-Q
July 1975 Environmental Protection Technology Series
INFLUENCE OF AERODYNAMIC PHENOMENA
ON POLLUTANT FORMATION
IN COMBUSTION
VOLUME I. EXPERIMENTAL RESULTS
U.S. Environmental Protection Agency
Office of Research and Development
Washington, D.C. 20460
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EPA-650/2-75-061-a
INFLUENCE OF AERODYNAMIC PHENOMENA
ON POLLUTANT FORMATION
IN COMBUSTION
VOLUME I. EXPERIMENTAL RESULTS
by
Craig T. Bowman and Leonard S . Cohen
United Technologies Research Center
400 Main Street
East Hartford, Connecticut 06108
Contract No. 68-02-1092
ROAP No. 21BCC-014
Program Element No. 1AB014
EPA Project Officer: W. Steven Lanier
Control Systems Laboratory
National Environmental Research Center
Research Triangle Park, North Carolina 27711
Prepared for
U.S. ENVIRONMENTAL PROTECTION AGENCY
OFFICE OF RESEARCH AND DEVELOPMENT
WASHINGTON, D. C. 20460
July 1975
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EPA REVIEW NOTICE
This report has been reviewed by the National Environmental Research
Center - Research Triangle Park, Office of Research and Development,
EPA, and approved for publication. Approval does not signify that the
contents necessarily reflect the views and policies of the Environmental
Protection Agency, nor does mention of trade names or commercial
products constitute endorsement or recommendation for use.
RESEARCH REPORTING SERIES
Research reports of the Office of Research and Development, U.S. Environ-
mental Protection Agency, have been grouped into series. These broad
categories were established to facilitate further development and applica-
tion of environmental technology. Elimination of traditional grouping was
consciously planned to foster technology transfer and maximum interface
in related fields. These series are:
1. ENVIRONMENTAL HEALTH EFFECTS RESEARCH
2. ENVIRONMENTAL PROTECTION TECHNOLOGY
3. ECOLOGICAL RESEARCH
4. ENVIRONMENTAL MONITORING
5. SOCIOECONOMIC ENVIRONMENTAL STUDIES
6. SCIENTIFIC AND TECHNICAL ASSESSMENT REPORTS
9. MISCELLANEOUS
This report has been assigned to the ENVIRONMENTAL PROTECTION
TECHNOLOGY series. This series describes research performed to
develop and demonstrate instrumentation, equipment and methodology
to repair or prevent environmental degradation from point and non-
point sources of pollution. This work provides the new or improved
technology required for the control and treatment of pollution sources
to meet environmental quality standards.
This document is available to the public for sale through the National
Technical Information Service, Springfield, Virginia 22161.
Publication No. EPA-650/2-75-061-a
11
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TABLE OF CONTENTS
ABSTRACT iii
LIST OF FIGURES iv
LIST OF TABLES x
ACKNOWLEDGEMENTS xii
I. INTRODUCTION 1
II. EXPERIMENTAL INVESTIGATION 2
A. Experimental Configuration and Approach 2
B. Experimental Apparatus and Instrumentation 5
Combustor Facility 5
Proves 10
Sampling System 12
C. Experimental Results 15
Description of Experiments 15
Input-Output Experiments 17
Flow Field Mapping Experiments 36
III. CONCLUDING REMARKS - 8^
APPENDIX A - COMPUTATION OF THE SWIRL NUMBER
APPENDIX B - CCMBUSTOR PROBE DESIGNS, CALIBRATION PROCEDURES &J
AND REDUCTION OF PROBE DATA
APPENDIX C - GAS SAMPLING SYSTEM, CALIBRATION PROCEDURES
AND DATA REDUCTION TECHNIQUES 97
APPENDIX D - EXHAUST CONCENTRATION DATA 10U
APPENDIX E - SPECIES CONCENTRATION, TEMPERATURE AND
VELOCITY DISTRIBUTIONS 115
APPENDIX F - FUEL COMPOSITION 138
NOMENCLATURE
REFERENCES
iii
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ABSTRACT
Average concentration levels of the pollutants NO, N02, CO and unburned hydrocarbons
(THC) were measured at the exhaust of an axisyrametric combustor over a significant
range of operating conditions. In addition, detailed species concentration, tempera-
ture and velocity maps were obtained throughout the combustor for seven representative
operating conditions. Natural gas, a synthesized CH^/CO/H^ fuel or vaporized propane
issued through a central duct to mix and burn with an annular air stream within
the 1.8 m long cylindrical combustor. In a limited mumber of tests liquid propane
was utilized as the fuel. Major combustor input parameters were varied over the
following ranges: overall fuel/air equivalence ratio 0.5 - 1.35 air-fuel velocity
ratio 0.1 -40, inlet air swirl number 0 - 0.6, air flow rate 0.09 - 0.1^ kg/sec,
inlet air temperature 730 - 860°K and combustor pressure 1-7 atm. Water-cooled
probes were used to remove samples from the flow for on-line concentration analysis ,
and to measure temperature, velocity and flow direction.
Elevated pressure and introduction of swirl to the extent considered in the present
experiments, creates "unmixedness" in the combustor flow field, producing high local.
temperatures, which in turn results in enhanced NO formation and consumption of
hydrocarbons. Aerodynamic flame stabilization, achieved without benefit of swirl
or physical flameholders in systems having large air-fuel momentum flux ratios, pro-.
duces strong stirring which results in reduced temperatures and in relatively low
NO formation and hydrocarbon consumption rates.
iv
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LIST OF FIGURES
Figure No.
1 Recirculation Zone in Initial Region Downstream 3
of Coaxial Streams.
2 Schematic Diagram of Axisymmetric, Water-Cooled 6
Combustion System.
3 Injector and Swirl Vane Geometries. 8
^ Combustor Facility. n
5 Exhaust Sampling Rake. . ]_1
6 Combustor Probing Device. !2
7 On-line Gas Analysis System. lh
8 Average Species Concentrations in Combustor Exhaust
Gas. Combustion of Natural Gas with 0.132 kg/sec
of 7^0 Air, S=0, Va/V-j-22, at 1 atm. Species Con-
centrations Reported As Measured. 18
9 Exhaust Temperature Profiles for Non-Swirling (S=0) 19
Flow. Combustion of Natural Gas with 0.132 kg/sec
of 7^0°K Air, Va/Vf=22, at 1 atm.
10 Influence of Inlet Air Swirl and Combustor Pressure 20
on Average Exhaust NO Concentration for Combustion
of Natural Gas with 0.132 kg/sec of jJ|-00K Air,
Va/V±,=22.
11 Exhaust Temperature Profile for a Swirling (S=0.6) 22
Flow. Combustion of Natural Gas with 0.132 kg/sec
of 7^0°K Air, Va/Vf=22, at 3.5 atm.
12 Influence of Inlet Air Swirl on Average Exhaust CO 23
and THC Concentrations for Combustion of Natural
Gas with 0.132 kg/sec of 7^0°K Air, Va/Vf=22, at
1 atm.
13 Influence of Combustor Pressure on Average Exhaust 25
CO and THC Concentration for Combustion of Natural
Gas with 0.132 kg/sec of 7^0°K Air, VQ/Vf=22, 5=0.
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LIST OF FIGURES (CONT.)
Figure No.
11+ Influence of Air-Fuel Velocity Ratio on Average 26
Exhaust NO Concentration in a Swirling (3=0.6)
Flow. Combustion of Natural Gas with 0.132 kg/sec
of 740°K Air at 1 atm.
15 Influence of Air-Fuel Velocity Ratio on Average 27
Exhaust CO and THC Concentrations in a Swirling
(S=0.6) Flow. Combustion of Natural Gas with
0.132 kg/sec of 7^0 K°Air at 1 atm.
16 Influence of Air Flow Rate and Inlet Temperature 28
on Average Exhaust NO Concentrations for Combustion
of Natural Gas with Air, Va/Vf =22, S=0, at 1 atm.
17 Influence of Air Flow Rate on Average Exhaust CO 29
and THC Concentrations for Combustion of Natural
Gas with 7^0°K Air, Va/Vf=22, S=0, at 1 atm.
18 Influence of Inlet Air Temperature on Average Exhaust 30
CO and THC Concentrations for Combustion of Natural
Gas with 0.132 kg/sec of Air, Va/Vf=22, S=0, at 1 atm.
19 Average Exhaust NO Concentrations for Combustion of 32
Natural Gas, Synthesized Fuel and Vaporized Propane.
Combustion with 0.132 kg/sec of 7^0°K Air at 1 atm.
Va/Vf=22, for Natural Gas, Va/Vf=15 for the Synthesized
Fuel, Va/Vf =35 for Vaporized Propane.
20 Average Exhaust NO Concentrations for Combustion of 33
Vaporized and Liquid Propane with 0.136 kg/sec of
740 °K Air.
21 Average Exhaust THC Concentrations for Combustion of 34
Vaporized and Liquid Propane with 0.136 kg/sec of
740CK Air.
22 Average Exhaust CO Concentrations for Combustion of 35
Vaporized and Liquid Propane wiht 0.136 kg/sec of
7^0 °K Air.
23 Velocity Profiles for Baseline Case. Combustion of ho
Natural Gas (Tf=303°K) with 0.132 kg/sec of 752°K
Air, Va/Vf=22.9, S=0 at 1 atm.
vi
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LIST OF FIGURES (CONT.)
Figure No. Page
2^ Temperature Profiles for Baseline Case. l^
25 Oxygen Concentration Map for Baseline Case. Average k2
Exhaust Concentration = 0.02.
26 THC Concentration Map for Baseline Case. Average 1*3
Exhaust Concentration = 1300 ppm.
27 CO Concentration Map for Baseline Case. Average k$
Exhaust Concentration = 0.0l6.
28 C02 Concentration Map for Baseline Case. Average h6
Exhaust Concentration = 0.0881.
29 NO and N02 Concentration Maps for Baseline Case. hj
Average Exhaust Concentration: N0=156 ppm and
N02=30 ppm.
30 Comparison of the 02 Concentration Distributions at ^9
the Baseline Condition (4>-0.9) with Those at * ^ 0.7.
At the Lower Equivalence Ratio, Natural Gas (307°K)
is Burned with 0.136kg/sec of 75^° K Air, Va/Vf=29.2,
* = 0.68, S=0 at 1 atm.
31 Effect of Overall Equivalence Ratio on THC Concentration 50
Map. Operating Conditions Identical to Fig. 30.
32 Effect of Overall Equivalence Ratio on CO Concentra- 51
tion Map. Operating Conditions Identical to Fig. 30.
33 Effect of Overall Equivalence Ratio on C0? Concentra- 52
tion Map. Operating Conditions Identical~to Fig. 30.
3^ Velocity Profiles for Combustion of Synthesized Fuel 5^
(Tf=30y°K) with 0.132 kg/sec of 730°K Air, Va/Vf=lU.6,
S=0, 4> = 0.87, at 1 atm.
35 Temperature Profiles for Synthesized Fuel Combustion 55
Test.
36 C02 Concentration Maps for Synthesized Fuel and Base- 56
line (Natural Gas) Combustion Tests.
vii
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LIST OF FIGURES (CONT.)
Figure No. Page
37 NO Concentration Maps for Synthesized Fuel and 57
Baseline (Natural Gas) Combustion Tests.
38 Temperature Profiles in Swirling Flow Combustion 59
Test. Combustion of Natural Gas (Tf=360°K) with
0.1314- kg/sec of 750°K Air Having a Swirl Number
of 0.3, Va/Vf =23,*= 0.885, 1 atm.
39 Oxygen Concentration Maps for Swirling ( S=0.3) and 6l
Baseline (3=0) Combustion Tests.
40 Total Hydrocarbon Concentration Maps for Swirling 62
(3=0.3) and Baseline (S=0) Combustion Tests.
4l CO Concentration Maps for Swirling ( S=0.3) and 63
Baseline (3=0) Combustion Tests.
42 COp Concentration Maps for Swirling ( S=0.3) and 64
Baseline (3=0) Combustion Tests.
43 NO Concentration Maps for Swirling (3=0.3) and 65
Baseline (S=0) Combustion Tests.
kh Velocity ?Tofi~Les itv SwlTclixig, dcratoMia^OT YlON) 66
(3=0.3).
45 Direction of Flov in Conibustor with 3=0.3. 68
46 Temperature Profiles for Baseline (p=l atm) Combustion 69
Test and for High Pressure (P=3«6 atm) Combustion of
Natural Gas (Tf=305°K) with 0.137 kg/sec of 750°K Air,
Va/Vf=22.9, $=0.87, S=0.
4-7 Oxygen Concentration Maps for High Pressure (P=3.6 atm) 70
and Baseline (P=l atm) Combustion Tests.
48 Total Hydrocarbon Concentration Maps for High Pressure 71
(P=3.6 atm) and Baseline (P=l atm) Combustor Tests.
49 CO Concentration Maps for High Pressure (P=3.6 atm) 72
and Baseline (P=l atm) Combustion Tests.
viii
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LIST OF FIGURES (CONT.)
Figure No.
Page
50 C02 Concentration Maps for High Pressure (P=3.6 atm) 73
and Baseline (P=l atm) Combustion Tests.
51 NO Concentration Maps for High Pressure (P=3.6 atm) 7^
and Baseline (p=l atm) Combustion Tests.
52 Temperature Profiles for Combustion of Liquid Propane 76
with 0.137 kg/sec of 739°K Air, S=0.3, at 1 atm. Data
Obtained Using Calibrated Thermocouple Probe.
53 Temperature Profiles for Combustion of Liquid Propane 77
with 0.137 kg/sec of 739°K Air, S=0.3, at 1 atm.
Data Obtained Using Calibrated Thermocouple Probe.
51*- Oxygen Concentration Maps for Liquid Propane Combus- 78
tion Tests.
55 Total Hydrocarbon Maps for Liquid Propane Combustion 79
Tests.
56 CO Concentration Maps for Liquid Propane Combustion 80
Tests.
57 C02 Concentration Maps for Liquid Propane Combustion 8l
Tests.
58 NO Concentration Maps for Liquid Propane Combustion 83
Tests.
59 Exhaust Gas Sampling Probe. 88
60 Calibration Set-Up for Impact-Static Probe. 91
61 Impact-Static Probe Pitch Plane-Angle Calibration 92
Data.
62 Impact-Static Probe Yaw Plane-Angle Calibration 93
Data.
63 Correlation of Total Flow Angle. 95
6^ Double Sonic Orifice (DSO) Calibration . 96
ix
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LIST OF FIGURES (CONT.)
Figure No.
65 Exhaust Gas Analytical System. 99
55 Typical Mass Spectrum of Exhaust Gas Sample. 101
6 MBTH Formaldehyde Analysis Apparatus. 102
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LIST OF TABLES
Table
I Matrix of Nominal Test Conditions : Input-Output 16
Experiments
II Matrix of Nominal Test Conditions : Mapping 27
Til Legend for Mapping Data ; 3
Axial Probe Location
D-l Exhaust Concentration Data - Natural Gas Fuel, 10Ji
S=0.
D-2 Exhaust Concentration Data - Natural Gas Fuel, 10^
D-3 Exhaust Concentration Data - Natural Gas Fuel, 106
11=0.6.
D-'i' Exhaust Concentration Data - Synthesized Fuel, 10rf
S=0.
D-!; Exhaust Concentration Data - Synthesized Fuel, 108
f5=0.6.
D-6 Exhaust Concentration Data - Gaseous Propane . 109
D-Y Exhaust Concentration Data - Liquid Propane. 110
D-H 1'Lxhauot Tempc-rature J'rofiles for Natural Gus 111
Combustion at Various .
XI
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LIST OF TABLES (COM1.)
Table
E_i Baseline Case Concentration Distributions . 115
E_2 Species Concentration Distributions for Low 118
Equivalence Ratio Case.
E-3 Species Concentration Distributions for Synthesized 121
Gas Case .
E-U Species Concentration Distributions for Swirling 12^
Flow Case .
E-5 Species Concentration Distributions for High 127
Pressure Case.
E-6 Species Concentration Distributions for Liquid 129
Propane.
E-7 Baseline Case Velocity Profiles. 131
E-8 Baseline Case Temperature Profiles. 131
E-9 Velocity Profiles for Low Equivalence Ratio Case. 132
E-10 Temperature Profiles for Low Equivalence Ratio Case. 132
E-ll Synthesized Gas Case Velocity Profiles. 133
E-12 Synthesized Gas Case Temperature Profiles. 133
E-13 Velocity Profiles for Swirling Flow Case. 13^
E-lIf Flow Direction Data for Swirling Flow Case. 134
E-15 Temperature Profiles for Swirling Flow Case. 135
E-l6 Temperature Profiles for High Pressure Case. 136
E-17 Temperature Profiles for Liquid Propane Combustion. 137
p_l Natural Gas Composition. 138
F_2 Synthesized Fuel Composition. 138
F_3 Propane Fuel Composition. 139
xii
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ACKNOWLEDGMENTS
A number or individuals at UTEC made significant contributions to the experimental
investigation. Dr. ;i. F. Zabielski and Mr. r,. L. Dodge designed the sampling
system used in the investigation and developed the calibration and data reduction
procedures employed in the gas sampling portion of the experiments. Mr. T. rimun
wai3 responsible ^or fabrication of the sampling system and maintenance of the gas
analysis equipment. Mr. T. A. Murrin assisted throughout the experimental program,
and was responsible for operation of the combustor and for reduction of much of the
experimental data. Mr. R. Lohrach and Dr. M. II. Director contributed to the design
and calibration of the probes. Mrs. P. A. Rose, Mrs. P. Jankot and Mrs. ,T. Andrew
assisted in reduction and compilation of the experimental data and in the prepara-
tion of the final renort.
This research program was carried out under the sponsorship of the Environmental
Protection Agency, EPA Contract 68-02-1092, Durham, North Carolina, under the
direction of Mr. G. Blair Martin and Mr. W. S. Lanier, Project Officers.
xiii
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I. INTRODUCTION
Recent experimental investigations of factors affecting pollutant emissions from
various continuous combustion devices have shown that changes in operating conditions,
which alter the flow pattern in the combustion chamber, can have a substantial effect
on pollutant emissions. Heap and his co-workers (Ref. l) have shown that in Cas-
and pulverized coal-fired furnaces, changes in the velocity and swirl of the air stream
and in the fuel injection configuration produce significant changes in nitrogen
oxide emissions. ShoITstall and Larson (Ref. 2) observed substantial variations in NO
emissions from gas-fired industrial burners with changes in the gas injection method.
Furthermore, the work of Mellor, et al.(Ref. 3) and Jones and Grobman (Ref. k) provide
evidence that changes in the air flow distribution in the combustion chamber can signi-
ficantly affect the emissions of carbon monoxide and unburned hydrocarbons from gas turbines
All of these observations suggest that coupling between fluid dynamic and chemical
processes inside the combustion chamber is a major factor in governing pollutant
emissions from many combustion devices. Analytical studies of reacting flow fields
(Refs. 5-8) have indicated in a qualitative manner how mixing and turbulence can
influence pollutant formation. However, at the present time our understanding of the
nature of the coupling between fluid dynamic and chemical processes is insufficient
to permit extrapolation of results obtained from one combustion device to other
devices or to allow a priori determination of the effects of changes in operating
conditions on pollutant emissions.
The present report documents the results of an experimental and analytical investiga-
tion, sponsored by EPA Contract 68-02-1092, of the interaction between fluid dynamics
and chemistry in a combustor and the subsequent effects on pollutant formation and
destruction. The overall objectives of this investigation are:
(i) To examine experimentally the interaction between fluid dynamic and chemical
processes inside a model combustor as operating conditions are varied, and to
correlate changes in pollutant emissions with these variations.
(ii) To further develop an existing elliptic flow computational procedure for pre-
dicting fluid dynamic and chemical processes occurring in combustbrs.
(iii)To compare the experimental results with results from the combustor flow analy-
sis in order to evaluate and calibrate the theoretical model and to assist in
interpretation of the experimental results.
The report consists of two parts. Part 1 documents results from the experimental
investigation. Part 2 presents details of the combustor flow analysis together with
a preliminary comparison 'of analytical and experimental results for gaseous fuel
combust!on.
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II. EXPERIMENTAL INVESTIGATION
A. Experimental Configuration and Approach
The selection of the experimental configuration and approach utilized in the present
investigation was predicated on a number of requirements. For example, an experiments
geometry was desired which models tile essential features pf practical continuous flow
combustion devices. Furthermore, a versatile combustor design was sought to permit
independent variation of the various inlet parameters, e.g., air-fuel velocity ratio,
inlet air swirl, overall equivalence ratio, inlet air flow rate and temperature.and
pressure. The combustor design also must allow the accurate determination of inlet
conditions to facilitate meaningful comparisons between experimental and analytical
results. Finally, extensive combustor instrumentation was provided to permit measure-
ment of pollutant emissions and determination of the radial and axial variations of
temperature, velocity and species concentrations throughout'the flow field.
The experimental configuration selected for the present study is an axisymmetric
combustor in which a central fuel stream, gaseous or liquid, mixes with a coaxial
annular air stream. In this configuration, which resembles the primary combustion
zone in many practical combustion devices, the resulting flame will not be stabilized
in general, and blow-off (extinguishment) will occur. ' In order to stabilize a flame
in such a system, it is necessary to provide a continuous ignition source either by
means of a pilot flame or through creation of zones of hot recirculating combustion
gases. These stabilizing recirculation zones may be created by the introduction of
physical flameholders into the flow or by imparting a swirl component to the flow
(c.f. Ref. 9). Swirl stabilization techniques have been used extensively in continuous'
flow combustion devices such as furnaces and gas turbines. Hence, a significant por-
tion of the current effort focuses on swirl-stabilized flames.
At the outset of the program, it appeared that detailed characterization of inlet
conditions (a major program objective) using probe techniques would be difficult in
swirling flows because of the relatively large tangential velocity components
necessary to stabilize the flame. Heftce, alternative methods of flame stabilization
were examined. The use of physical flameholders (e. g., v-gutter and cylindrical-
shaped bodies or rearward-facing step*) was considered undesirable since the processes
which occur in the flame zone are influenced by the specific flameholder geometry
(Ref. 10), thereby complicating specification of the inlet conditions. Flame stabili-
zation by means of pilot flames was considered undesirable because of the primary
interest in flames stabilized on recirculation zones. A review of the recent litera-
ture on mixing of coaxial non-reacting jets indicated that flame stabilization might
be achieved in gaseous-fueled systems by judicious selection of the momentum flux
ratio of the fuel and air streams. This flameholding approach relies neither on a
physical flameholder nor on swirl. Rather, as shown in Fig. 1, it makes use of the
observation that a recirculation zone is formed in the initial region of coaxial
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FIG. 1
Recirculation Zone in Initial Region Downstream of Coaxial Streams.
'////////""""/////M^^^
N12-175-1
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streams when the outer, annular stream, i.e., the air stream, is moving very rapidly
relative to the central (fuel) jet. It was shown in Ref. 11 that recirculation zones
exist in a constant density, coaxial jet flow when the outer stream to inner stream
velocity ratio is 13 or larger. For higher velocity ratios, a circulating vortex is
established at the central Jet centerline which enhances mixing between the streams.
This "aerodynamic" phenomenon can be utilized to effect flameholding. In a practical
demonstration of aerodynamic flameholding as part of the present study, a fuel
injector was designed to produce an air-to-fuel velocity ratio of approximately 20:1
at stoichiometric operation using natural gas. The lip of the injector was made as
small as practicable, i.e.,~0.0l8 cm, to prevent flameholding on back-step-
associated recirculation zones. The particular injector design was based on the
requirement that the momentum flux ratio of the air stream relative to the fuel
stream exceeds 169, which evolves from a minimum air-to-fuel velocity ratio (Va/V.p)
of 13 for constant density flow. That is,
Vf2 > 169
Since the density ratio, Pa/Pf, for air at 800°K and nautral gas at 300°K is about
0.75, the design velocity ratio must exceed 15:1 to satisfy the criterion given by
Eq. (l). A margin of uncertainty was applied to arrive at the 20:1 injector actually
selected. In the course of the demonstration experiments, a natural gas-air flame
was stabilized over a wide range of inlet air temperatures and flow velocities and
overall fuel-air ratios. These data, presented later, constitute the foundation for
aerodynamic flame stabilization, an approach which relies on the formation of a recir-
culation zone from a momentum flux imbalance between concentric jets.
The aerodynamic flameholding technique, in principle, should result in a simplified
determination of inlet conditions. In addition, this approach provides a unique oppor-:
tunity to contrast swirling and non-swirling flow fields without the disruptive presence
of a physical flameholder. Swirl vanes have been installed in the annular air passage
for certain tests and omitted in others to provide swirl numbers, S, of 0, 0.3 and 0.6.:
The swirl number, defined as the ratio of the angular momentum flux to the product of
the axial momentum flux and an effective nozzle diameter, may be expressed in terms of
swirl vane angle and injector section geometry (see Appendix A),
The nature of the mixing-combusting flow field and, hence, pollutant formation rates
are known to be dependent on the fluid residence times within the recirculation zone
and the combustor and also on the temperature and pressure levels. Variations in
mean combustor residence times can be effected simply through changes in the initial
velocity of the air stream. In the present experimental study, the air flow rate was
varied from 0.088 to 0.137 kg/sec for a fixed injector configuration, i.e., a fixed
air/fuel area ratio, to produce significant changes in mean throughput velocity.
Fluid residence times within the recirculation zone are dependent on the mixing rates,
which are influenced by the air-fuel velocity ratio. In the present study,
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installation of injectors with different air/fuel area ratios provided air-fuel velo-
city ratios in the range 0.1 to 40. Variation in eombustor pressure was obtained
by the placement of a choke downstream from the combustor with allowed operation
at pressure levels of approximately k and 7 atm, and corresponding^ reduced flow
velocities. Operation at different temperature levels, for a fi*ed overall fuel-air
equivalence ratio, was accomplished by varying the temperature of the inlet air stream
from 730-860°K.
Recent analytical studies (e.g., Ref. 12) have shown that fuel composition can sig-
nificantly influence NO emissions from gas turbines operating on gaseous fuels.
This effect, which is believed to be due primarily to the variation in heating'value
with fuel composition, was examined in the present program. Most of the experiments
were conducted using natural gas, with a nominal higher heating value of 55 kj/g.
However, a number of tests v/ere conducted using a synthesized fuel (natural gas/'
hydrogen/carbon monoxide in approximate molar proportions of 1.0/0.39/0.3*0, with a
nominal heating value of ^2 kJ/g, and vaporized propane, with a nominal heating
value of 50 kJ/g.
Since many practical combustion devices operate on liquid fuels, a limited number of
experiments was carried out using liquid propane to provide a comparison with results
from the vaporized propane tests.
None of the fuels used in the present investigation contained organic-nitrogen
compounds, so that the only source of combustion-generated NO is atmospheric
(molecular) nitrogen.
B. Experimental Apparatus and Instrumentation
Combusjto£ .Facility
Tests were conducted in an instrumented, water-cooled combustion system (Fig. 2).
The facility design was based on the considerations outlined in the previous secticr^
with particular emphasis placed on acquisition of species concentration, temperature
and velocity distributions throughout the reacting flow field for comparison with
results obtained in the analytical study.
Air from a 30-atm supply, at flow rates up to 0.65 kg/sec, may be heated in an
electrical heater section to provide inlet air temperatures up to 1000°K. Within
the heater, the air flows through and around four, 6 m long stainless steel tubes
which may be supplied with as much as 720 kw of electrical power. The heated air
enters the circular annulus of a replaceable injector which is installed in the
12.23 cm diameter entry section. Fuel, introduced through three (air foil shaped)
struts into the center delivery duct, is brought into contact with the annular air
stream at the exit of the injector. Thereafter, mixing and chemical reaction
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Schematic Diagram of Axisyrmetric, 'Jater-Cooled Combustion Hysten,
ALL DIMENSIONS IN CM
.EXHAUST
SAMPLING
RAKE
FUEL
65432
-INSTRUMENTED 1=
COMBUSTOR
187.8-
20.9
REPLACEABLE
INJECTOR
m
FUEL
ENTRY-
SECTION
AIR
-AIR-
HEATER
WINDOW PORT LOCATIONS
z
I1
PORT
NO.
0
1
2
3
4
5
6
xCj(CM)
3.18
7.62
37.29
54.28
71.27
88.27
105.26
Xcj/ro
0.52
1.25
6.10
8.88
11.66
14.43
17.21
EXHAUST SAMPLING RAKE LOCATION AT 182.45 CM
P
M
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proceed at constant area in the remainder of the injection section and into the
instrumented combustor and extender sections. In order to impart swirl to the air
flow, straight swirl vanes are inserted into the annular passage of the injector.
The trailing edges of the swirl vanes are flush with the injector exit plane. The
three injectors utilized in the study are described in Fig. 3 in terms of the ratio
of the inner and outer diameters of the air annulus, Z = dh/d, and the nominal air-
to-fuel velocity ratio, m = Va/Vf, associated with air and natural gas coaxial jets
having an equivalence ratio, $, of 0.92. Swirl vane designs also are shown in Fig.
3, where the swirl number, S, has been computed from the injector geometry, Z, and
the angle of the swirl vanes, f] , according to the following expression (derived in
Appendix A ) :
"~3(,_Z2}I.5 an7? (2)
A swirl number of 0.3 connotes moderate swirl, while S = 0.6 results in a relatively
high swirl situation. A practical upper limit of S =0.8 exists for straight blades
from the standpoint of packaging the vanes.
Observation of the combusting flow may be made through the 6. it-cm diameter window
ports in the entry and combustor sections (Fig. 2). A pair of window ports 180 deg
apart are present at each location, with the exception of station no. 1. The loca-
tion of a port directly downstream of the injector exit plane allows an unhindered
view of the flame in the vicinity of the fuel delivery duct. A second port in the
entry section (no. 1) was located to permit acquisition of probe measurements very
close to the injector exit. These measurements should provide the inlet or initial
conditions needed in the analytical modeling effort. The combustor probing device
used to make these measurements is compatible with all ports and may replace a glass
window or water-cooled plug in any given port. Port locations are listed on Fig. 2.
The 12.23 cm diameter, 100.1 cm long instrumented combustor is divided into three
water-cooling zones of approximately equal length. Water flow can be set indepen-
dently in each zone, as needed, to keep wall temperature roughly constant along the
entire length of the combustor. Temperatures are set and monitored using thermo-
couples installed on the combustor surface, and at various depths in the combustor
wall and cooling passages. Two static pressure tap locations also are available
at opposite ends of the combustor. Flow exhausts from the combustor and extender
sections to the facility exhaust stack. Combustor extender (spool) pieces, 33. k cm
in length, are inserted when appropriate to fully contain the flame; the extender
consisted of two spool pieces for the entirety of the current experimental effort.
Chokes can be installed downstream of the extender to raise the pressure in the
combustor. To date, high pressure operation has been limited to about 7 atm. A
photograph of the comoustion facility is given in Pig. k.
-------
Injector and Swirl Vane Geometries
mini 111 ill
d0= 12.23 CM
FLOW
CD
i 11111111
.43 CM
INJECTOR STEP
VANE 0.163 CM
THICK 316 SS
INJECTOR
DESIGNATION
I
n
m
m, VELOCITY
RATIO*
20:1
8:1
0.3:1
df (CM)
6.314
4.727
0.757
Z = dh,d
0.677
0.508
0.084
fi(CM)
1.499
2.306
4.290
S
0.3
0.6
0.6
0.6
7? (DEC)
28
47
53
60
NO. OF
VANES
18
12
12
8
BASEDON NATURAL GAS AND * = 0.92
-------
Combustor Facility
it. juir
INSTRUMENTED «M|[ m ENTRY
COMBUSTOR MIS" SECTION
I si .,
AIR
HEATER
EXHAUST
SAMPLING 1
PROBE
// EXTENDER
T-O-F MASS
SPECTROMETER
-n
P
*.
-------
Probes
An exhaust probe rake is placed at the exit of the extender in certain tests to
sample flow which has had the entire test section length to mix and react. Compo-
sition information is determined on-line by aspirating flow through a five-probe rake
and analyzing the combined gas sample in a time-of-flight (TOF) mass spectrometer
and a Scott Model 119 Exhaust Analyzer. The identical, choked probes constituting
the rake (Fig. 5) are centered on equal areas to insure the acquisition of a flow
sample which is truly representative at the exhaust plane. Pressurized hot water
at UOO°K was used as the coolant for the probes to minimize wall-catalyzed reactions
and to prevent water condensation within the sample lines. Other details of the
rake design are discussed in Appendix B, which contains descriptions of all the
probes used in the investigation.
Temperature profiles at the exhaust plane and within the combustor were measured by
traversing a calibrated-heat-loss thermocouple probe across a combustor diameter.
Ordinarily, thermocouple materials limit application of these sensors to temperatures
below about 2000°K. It is recognized, however, that radiation and conduction heat
losses may reduce thermocouple temperature significantly below stream temperature.
If proper calibration information is acquired simultaneously with the required
temperature measurement, then the heat loss situation may be exploited to extend
the range of thermocouple utilization up to 2500°K (Appendix B). A thermocouple
probe of this type was applied without difficulty in the present combustion environ-
ment.
Temperature distributions in the combustor also were obtained with the double-sonic-
orifice (DSO) probe portion of a two-probe rake (Fig. 6) inserted through the window
ports. In the operation of this probe, ingested flow passes through two choked
orifices in series, resulting in an interrelationship between orifice stagnation
conditions:
(3)
/PTACDF\ /PTA*CDF\
\ vT7 /, \ v/T7
If orifice no. 1 is integrated with the probe tip, then Prp is the impact pressure
and T^is the total temperature to be determined. The parameter, F, which depends
on the ratio of specific heats and molecular weight, is essentially the same at both
orifices (i.e., FI » F2). Also, AJ and A| are readily established and PT and TT
are available from measurements at orifice no. 2. Thus, T may be calculated from
Eq. (3) if the ratio of flow coefficients, C^/Cj, , is known. Such information must
be generated in calibrations over a range of orifice Reynolds numbers at various
values of the ratio of wall temperature to stream temperature (see Appendix B).
Calibrations of DSO probes are laborious and often frustrating since the presence
of any foreign matter in the flow to deposit in or to scratch or pit throat regions,
10
-------
Exhaust Sampling Rake
PROBES CENTERED ON EQUAL AREAS
FLOW FROM ALL PROBES MIXED BEFORE
ON-LINE ANALYSIS
»COOLED WITH HOT WATER, WALLS
MAINTAINED AT ABOUT 400°K TO PREVENT:
(a) CATALYTIC REACTIONS
(b) CONDENSATION OF H2O
,
-
:
n
-------
Combustor Probing Device.
ALL DIMENSIONS IN CM
(a) IMPACT-STATIC PROBE
STATIC TAP
NOSE
STATIC
TAPI4)
OUTLET TUBE
MAT'L: 0.953 O.D. x
0.041 WALL-TYPE
316 S.S.
(b) DOUBLE-SONIC-ORIFICE (DSO) PROBE
0.396 IN-
FLOW-
\
FIRST
ORIFICE
OUTLET TUBE
MAT'L; 0.794 O.D. x
0.041 WALL-TYPE 316 S.S.
SECOND
ORIFICE
12
-------
can affect the accuracy of the calibration. This limitation generally precludes
application of the DSO technique to flows containing significant concentrations of
particulates (soot, liquid droplets or coal particles). In the current effort,
however, a probe was designed which minimizes flow coefficient corrections and which
performed satisfactorily in all of the gaseous fuel tests. In addition, measurements
obtained at identical test conditions with both the DSO probe and the thermocouple
probe generally were in agreement (see Appendix B), thereby establishing a high
confidence level in the temperature measurements.
The gas stream leaving the DSO probe is channeled to the sampling system (described
below) where species concentration levels are identified.
The companion probe on the combustor rake is an impact-static probe. Flow velocity
and direction can be determined from an interpretation of pressure differences between
various static pressure locations on the probe and the impact pressure. Pour static
pressures are sensed at taps located on a centerline circle hO deg from the probe
tip (Fig. 6). Additional taps located along the probe at 2.11 cm and U.02 cm from
the tip provide useful local axial pressure information. Calibration of the probe
in pitch and yaw and the use of this information to determine total flow angle is
discussed in Appendix B. The impact-static probe performed satisfactorily in all
regions of the combustor, except those regions with reverse flow. In these reverse
flow (recirculation) zones, meaningful measurements of gas velocity were not obtained.
Since the size, location and behavior of these recirculation zones significantly
influence the combustor flow field (see Section II-C), the lack of velocity data in
these zones is a major short-coming of the present experimental program.
_Sampling_Sv_s1;em
The gas samples withdrawn via the exhaust probe or combustor DSO probe are analyzed on-
line for the time-average concentration (mole fraction) of a number of species
using a time-of-flight mass spectrometer and a Scott Exhaust Analyzer. In the
schematic diagram of the sampling system, Fig. 7, the gas samples pass into a heated
(~1K>0°K), teflon-coated manifold. A portion of the gas sample is directed through
a heated, teflon-coated line to the TOF mass spectrometer, located close to the
probe assembly. In all experiments, the distance from the sample probe inlet to
the mass spectrometer inlet was less than 60 cm. The mass spectrometer was used to
measure the concentration of C02, 02 and NO in the gas sample. The remainder of
the gas sample passed successively through a condensate trap (~277°K) to remove
water vapor, and an unheated teflon-coated line to the Scott Analyzer. The Analyzer
located in the combustion facility control room approximately 10 m from the combustor
was used to measure the concentration of CO, C02, 02, NO, N02, and unburned hydro-
carbons (THC) in the gas sample. The redundant sampling instrumentation provided a
measure of the influence of the long line (and condensate trap) connecting the
sampling probe to the Analyzer on the gas sample received by the Analyzer. Output
signals from the TOF mass spectrometer and Analyzer were sent through a multiplexer
13
-------
On-Line Gas Analysis System.
ASPIRATED GAS
CALIBRATION
GASES
MBTH ANALYSIS
(FORMALDEHYDE)
ECO;
CONDENSATE
TRAP
BELLOWS
PUMP
SCOTT MODEL 119
EXHAUST
ANALYZER
(CO,C02, 02, NOX,THC)
TIME-OF-FLIGHT
MASS
SPECTROMETER
-------
and an analog-digital converter to be processed and stored in a signal averager.
Typically, at each test condition, sampling data were acquired for a period of 2-3
minutes. Upon completion of the tests, the stored data were processed immediately
using a direct link to a PDP-6 computer. Processed data received via teletype were
verified prior to initiation of subsequent tests. In a limited number of tests,
the formaldehyde concentration in the exhaust was measured using the MBTH wet
chemical method. Gas samples for these measurements were withdrawn from the heated
manifold and passed through the sampling device. Formaldehyde data were reduced
manually.
A detailed description of the TOF mass spectrometer, the Scott Exhaust Analyzer and
the MBTH wet chemical method is presented in Appendix C, together with a discussion
of calibration procedures, data reduction techniques and evaluation of experimental
errors.
C. Experimental Results
Dej3crip_t_ion £f_the_Expe_riments
The experimental program was comprised of two different types of tests: (l) input-
output tests, and (2) flow-field mapping tests. In the input-output tests, approximatel^-
fifty experiments were conducted with the objective of determining the influence of
various combustor input parameters on species concentration levels in the exhaust flow.
Generally, for a particular injector model and swirl vane insert, fuel-air equiva-
lence ratio was varied from fuel lean to rich conditions at two or more levels of
inlet air temperature, air flow rate and combustor pressure. Exhaust flow tempera-
ture profiles also were acquired for selected input-output test conditions. A matrix
of test conditions for natural gas, synthesized fuel and vaporized propane is given
in Table I. The ranges of parameters investigated encompass those found in practi-
cal combustion devices.
Results from the input-output tests were used to determine which combustor input
parameters had the greatest influence on pollutant emissions. From this informa-
tion, a matrix of test conditions was established for the flow field mapping ex-
periments. In each of these mapping experiments, the mole fractions of NO,
NOg, CO, COg, C>2 and total hydrocarbons, the total temperature and the axial flow
velocity and flow direction were measured throughout the combustor. Detailed com-
bustor maps developed from these data were studied to provide an understanding of the
changes in flow field characteristics and, ultimately, pollutant formation which
accompany variations in key input parameters. During each of the combustor mapping
tests, the two-probe rake was inserted and traversed along a diameter at every
window port location. Occasionally, both window ports at a particular port loca-
tion were entered in order to acquire data for assessing the symmetry of the flow
with respect to the combustor centerline. In addition, high-speed color motion
pictures (500 frames per sec) were taken at port location "0" for most test conditions
to provide visual records of the combusting flow in the vicinity of the injection plane.
15
-------
TABLE I. MATRIX OF NOMINAL TEST CONDITIONS: INPUT-OUTPUT EXPERIMENTS
Fuel*
NG
NG
NG
NG
NG
NG
NG
NG
NG
SYN
SYN
SYN
PG
PG
PL
PL
S
0
0
0
0
0.3
0.6
0.6
0.6
0.6
0
0.6
0.6
0
0.3
0.3
0.3
vaAf
22
22
22
22
22
22
22
9.*4-
0.2
15
15
15
35
35
__
P
(atra)
1
3.6
1
1
1
1
3.6
1.0
1.0
1.0
1.0
3.6
1.0
1.0
1.0
6.6
*
0.7-1.2
0.8-1.1
0.8-1.1
0.8-1.1
0.8-1.1
0.8-1.1
0.8-1.1
0.8-1.1
0.5-1-3
0.8-1.1
0.8-1.0
0.8-1.0
0.8-1.1
0.8-1.0
0.8-1.0
0.8-1.0
&
7*4-0
7kO
860
7*4-0
7*4-0
7*4-0
7*1-0
7*1-0
7*4-0
7*4-0
7*4-0
7**0
7*^0
7*4-0
7*4-0
7*4-0
(kg/sec)
0.132
0.132
0.132
0.088
0.132
0.132
0.132
0.132
0.132
0.132
0.132
0.132
0.132
0.132
0.132
0.132
* NG = natural gas
SYN = synthesized fuel
PG = gaseous propane
PL = liquid propane
16
-------
Inp_u1;-Output_Exp_eriments
Significant trends in the exhaust concentration and temperature data obtained in the
input-output experiments are illustrated by a series of graphs showing variations in
the exhaust emissions and temperature with changes in conibustor input parameters.
A complete compilation of the data is given in tabular form in Appendix D.
Average exhaust gas concentrations (dry basis) of CO, C02, 02, NO, N02 and THC for
a 1 atm, non-swirling natural gas-air flame in Fig. 8 typify the input-output
measurements. The natural gas used in the experiments is principally CH^ (>97%) with
small amounts of other gaseous hydrocarbons, C02 and N2. Detailed analyses of the
gaseous fuels are given in Appendix F. During acquisition of the data, the inlet
air flow rate was held constant, and a range of overall equivalence ratios was
obtained by varying the fuel flow rate. The concentration data shown in Fig. 8
are reported as measured and were obtained using the 22:1 injector; however, the
air-fuel velocity ratio varied between 16.6 and 27.4 in these tests as the fuel
flow was varied. As the overall equivalence ratio is increased, the exhaust 02
concentration decreases significantly while the CO and THC concentrations sub-
stantially increase. Concentrations of CO , NO and N02 in the exhaust gas are
only moderately dependent on overall equivalence ratio,$. Measured levels
of 02 and COp are approximately equal to the calculated equilibrium concen-
trations in the exhaust gas over the entire range of equivalence ratios investi-
gated. Equilibrium calculations assumed adiabatic combustion of the fuel-air
mixture for the measured inlet conditions. For 4>< 1, measured CO levels are
somewhat larger than the equilibrium values; however, as the equivalence ratio
increases, measured CO levels approach the equilibrium values. Measured THC concen-
trations are very large compared with calculated equilibrium values over the entire
range of equivalence ratios investigated. Measured NO concentrations are substan-
tially less than equilibrium values. Similar trends of CO, THC and NO emissions with
overall equivalence ratio have been reported for furnaces (Ref. 2) and gas
turbines (Ref. 3). The observed N02 concentrations are larger than equilibrium
levels in the exhaust gas. However, as noted in Appendix C, it is likely that some
N02 is produced by oxidation of NO and some N02 is lost during transfer of gas
samples to the Exhaust Analyzer, so that reported N02 levels must be considered as
uncertain.
Exhaust plane temperature profiles for several non-swirling flows are compared to
calculated adiabatic combustion temperatures in Fig. 9. Predictably, the exhaust
plane temperature is highest on the combustor centerline, and decreases as the cooled
combustor walls are approached. Exhaust plane temperature profiles are similar for
this combustor configuration for overall equivalence ratios between 0.8 and 1.2,
with small differences most notable near the combustor centerline. Measured center-
line temperatures are 10-15 percent lower than the calculated adiabatic combustion
temperatures.
The influence of inlet air swirl and pressure on the average NO concentration in the
exhaust of the combustor operating on natural gas-air is shown in Fig. 10. For
IT
-------
Average Species Concentration in Conibustor Exhaust Gas.
Combustion of Natural Gas with 0.132 kg/sec of 7^0°K Air,
g-0^ va/Vf=22, at 1 atm. Species concentrations reported
as measured.
500
oc
O
UJ
u
oc
LU
a
LU
_1
O
CM
O
CNI
O
U
O
O
0.6
0.8 1.0 1.2
OVERALL EQUIVALENCE RATIO,*
18
-------
FIG.9
Exhaust Temperature Profiles for Non-Swirling (S=0) Flow.
Combustion of Natural Gas with 0.132 kg/aec of 7^0°K Air,
Va/Vf=22, at 1 atm.
i.o K-
c/j
2
i
cr
ADIABATIC
COMBUSTION TEMPS.
TAD
i"
1500
2000
2500
EXHAUST TEMPERATURE - °K
19
-------
Influence of Inlet Air Svirl and Combustor Pressure on
Average Exhaust NO Concentration for Combustion of Natural
Gas with 0.132 kg/sec of 7^0°K Air, Va/V,,=22.
500
400
oc
Q
I
I
Z
o
DC
LL
UJ
_J
O
o
300
200
100 -
3.6 ATM
1 ATM
s
0
0.3
0.6
SYM
O
A
a
_L
0.6
0.8 1.0
OVERALL EQUIVALENCE RATIO,*
1.2
20
-------
the range of overall equivalence ratios investigated, the air-fuel velocity ratio
varied between 19.1* and 26.0. The data for no inlet air swirl (S=0) were taken
from Fig. 8. For a combustor pressure of 1 atm, increasing inlet air swirl raises
the NO concentration in the exhaust gas. This effect is more pronounced for the
lower overall equivalence ratios. Heap and his co-workers (Ref. l) have reported
similar behavior for NO emissions from nautral-gas-fired furnaces using axial fuel
injection. It should be noted, however, that other fuel injection geometries
(e.g., radial) can give variations in NO emissions with swirl which are significantly
different from that observed with the present experimental configuration (See, for
example, Ref. 1).
The double-lobed appearance of the exhaust plane temperature profile for a swirling
flow (Fig. 11) provides a partial explanation for the enhanced NO production. In
swirling flows, peak temperatures at the exhaust are significantly larger than in
the non-swirling flows and, in fact, are in excess of the adiabatic combustion tempera-
ture calculated for the overall fuel-air equivalence ratio. Thus, since the NO for-
mation rate is known to be strongly dependent on temperature, the higher temperatures
found in swirling flows result in higher NO levels than in comparable non-swirling
flows. For fixed inlet air swirl, increasing the combustor pressure generally
results in a significant increase in the exhaust NO concentration. A similar trend
of NO emissions with pressure has been reported by Mellor and his co-workers (Ref. 3)
for a gas turbine. Some portion of the increase in exhaust NO with increasing com-
bustor pressure may be attributed to increased gas residence time in the combustor.
For a pressure of 3.6 atm, residence time is roughly 0.075 sec versus about 0.02 sec
at 1 atm operation. At elevated combustor pressures, swirl has only a slight effect '
on the exhaust HO concentration. Reasons for this behavior are explored in a later
section of this report.
Average CO and THC concentrations in the exhaust gas for the three swirl numbers
investigated are given in Fig. 12 for a combustor operating at 1 atm. Imparting
swirl to the inlet air stream reduces levels of THC in the exhaust gas. A similar
variation in THC emissions with inlet air swirl was observed at a combustor pressure
of 3.6 atm. Increasing the inlet air swirl also results in a reduction in exhaust
CO concentrations for overall equivalence ratios less than one. However, for equiva-
lence ratios greater than one, inlet air swirl has little or no effect on exhaust CO
concentrations. The dashed line on Fig. 12 is the calculated equilibrium CO concen-
tration, assuming adiabatic combustion. For the two swirling flows investigated, the
measured exhaust CO concentrations are approximately equal to the equilibrium values
over the entire range of overall equivalence ratios. For the non-swirling flow,
the measured exhaust CO concentrations approach the equilibrium values forO>l. A
similar effect of inlet air swirl on CO emissions was observed at a combustor pressure
of 3.6 atm. Enhanced burnout of'THC and CO with swirl can be explained in terms of
associated changes which occur in temperature, velocity and 02 distributions within
the combustor (See results presented in the following section ), The influence
of an increase in pressure from 1 to 3-6 atm on average CO and THC exhaust con-
centrations is not large at any of the swirl numbers investigated. Data for
21
-------
Exhaust Temperature Profile for a Swirling (S»0.6) Flow.
Combustion of Natural Gea with 0.132 kg/sec of T^O°K Air,
Va/Vf=22, at 3-6 atm.
1500
2000
2500
EXHAUST TEMPERATURE-°K
22
-------
FIG. 12
Influence of Inlet Air Swirl on Average Exhaust CO and THC
Concentrations for Combustion of Natural Gas vith 0.132 kg/sec
of 7'K)0K Air, Va/Vf=22, at 1 atm.
10
,-2
10
-3
DC
Q
tr
LL
LU
O
H
10
,-5
10-6
0.6
0
0.3
0.6
SYM
CO
A
THC
o
A
a
EQUIL. CO
0.8 1.0
OVERALL EQUIVALENCE RATIO,*
1.2
10
-1
cc
a
<
10-2 £
uj
_i
O
o
o
10-3
-4
N12-133-2
-------
non-swirling flow.in Fig. 13 illustrate the extent of the pressure effect, i.e. a
modest decrease in THC and virtually no change in CO as pressure is increased.
As indicated earlier, the interaction between fluid dynamic and chemical processes
must be understood before it is possible to devise a model for pollutant formation
in combustors. One parameter which has been identified as key to the establishment
of the flow field within the combustor, the air-fuel velocity ratio, m, was variefl
during a series of input-output tests. Results from these tests are summarized in
Figs. Ik and 15. The swirl number was approximately 0.6 for these tests. For data
designated by nominal air-fuel velocity ratios of 22 and 9j the actual velocity
ratios varied between 19.k and 25.6 and between 7-6 and 10.3j respectively, as the
fuel flow rate was reduced at constant air flow rate. Data at a nominal velocity
ratio of 0.2 covered the range 0.13-0.26 as the fuel flow rate was decreased at corb-
stant air flow rate. A striking difference exists between NO exhaust concentrations
levels in swirl-stabilized flames for air-fuel velocities much greater than unity
and those for air-fuel velocities much less than unity. It is likely that this diff(
is due, in part, to a reduction in the volume of the hot zone in the vicinity of the
injection plane, resulting from the relatively small diameter fuel injector, Fig. 3
used for the Va/Vf =0.2 experiments.
Average exhaust CO and THC concentration measurements for high air-fuel velocity
ratios, i.e., 9-22, and low air-fuel velocity ratios of approximately 0.2 are shown
in Fig. 15. Equilibrium CO values are plotted as the dashed line. These data are
consistent with input-output data already presented, in that the set of initial
conditions which yields lowest levels of NO in the exhaust, generally also displays
the highest THC levels. Further, as observed for the exhaust NO concentration in
Fig. Ik, CO and THC concentrations in the exhaust gas are independent of air-fuel
velocity ratios in the range 9-22. Measured exhaust CO concentrations closely
respond to the calculated equilibrium values for these ratios. In fact, measured
calculated CO concentrations agree for all velocity ratios, m, investigated for o
all equivalence ratios exceeding 0.9. Leaner mixtures (*<0.9) yield CO levels for
m fa 0.2 which are somewhat larger than those predicted on the basis of equilibrium
considerations.
Variations in exhaust pollutant concentrations with inlet air velocity and pre-heat
level are shown in Figs. 16-18. For a fixed inlet air temperature, a reduction in
air flow rate (and hence velocity at constant PA) of $0% resulted in an increase of
about 30$ in NO in the exhaust (Fig. 16). This behavior, which can be attributed
primarily to changes in recirculation zone and combustor residence times, has been
noted by Heap,_et _al. (Ref. l) for gas-fired furnaces. The increased time for burn-
out also brought about reduced THC exhaust concentrations for all 0 investigated,
Fig. 17, However, no significant change in exhaust CO with air flow rate was obsena
At a fixed air flow rate, an increase in inlet air temperature from 7^2°K to 86o°K
resulted in an approximately doubling of the exhaust NO concentration. Other author
(Refs. 2 and 3) have noted similar trends of NO emissions with inlet air temperature
in natural gas-fired furnaces and gas turbine combustors. Exhaust CO concentrations
were unchanged for air inlet temperatures of 7*i20K and 860°K (Fig. 18), while exhaua
THC levels were generally lower at 860°K. By comparison, Mellor, et al. (Ref. 3)
2k
-------
FIG. 13
Influence of Conibustor Pressure on Average Exhaust
CO and THC Concentration for Combustion of Natural
Gas with 0.132 kg/sec of 7Uo°K Air, V,/Vf=22, 3=0.
1 fl-
<£ i
a 10~3
ee
U-
LU
_i
O
5
O
10-4
p
(ATM)
1
3.6
SYM
CO
THC
O
D
I
I
DC
Q
O
b
10-2 O
0.6
10
,-3
0.8 1.0
OVERALL EQUIVALENCE RATIO.*
1.2
N12-133-9
25
-------
Influence of Air-Fuel Velocity Ratio on Average Exhaust NO Concentration
in a Swirling (S=0.6) Flow. Combustion of Natural Gas with 0.132 kg/sec
of T^O°K Air at 1 atm.
40
Va/Vf
0.2
9
22
SYM
A
300
ir
Q
D.
a.
I
Z
O
LU
O
200
100
0.4
0.6 0.8
OVERALL EQUIVALENCE RATIO,'
1.0
1.2
26
-------
FIG.15
Influence of Air-Fuel Velocity Ratio on Average Exhaust
CO and THC Concentrations in a Swirling (3=0.6) Flow.
Combustion of Natural Gas with 0.132 kg/sec of
Air at 1 atm.
10-^
a
0.2
9
22
SYM
CO
A
THC
D
A
o
10-J
-------
Influence of Air Flow Rate and Inlet Temperature on Average
Exhaust WO Concentrations for Combustion of Natural Gas with
Air, Va/Vf=22, 3=0, at 1 atm.
400
1
1 1
w
(kg/sec)
0.088
0.132
SYM
742°K 860°K
A
0
300
cc.
Q
Q.
Q.
I
<
tr
UL
LU
_l
O
5
O
200
100
860°K
742°K
we
I
0.6
0.8 1.0
OVERALL EQUIVALENCE RATIO,*
1.2
28
-------
FIG. 17
Influence of Air Flow Rate on Average Exhaust CO and THC
Concentrations for Combustion of Natural Gas with 7^0°K Air,
Va/Vf=22, S=0, at 1 atm.
10~^
wa
(kg/sec)
0.088
0.132
SYM
CO
THC
D
O
10
,-3
cc
Q
Z
cc.
LL
QJ
_1
O
5
o
10-4
10
,-5
I
I
10
-1
cc
Q
cc
LL
LU
_l
O
8
10-'
10-3
0.6
0.8 1.0
OVERALL EQUIVALENCE RATIO,
1.2
N12-133-1
29
-------
Influence of Inlet Air Temperature on Average Exhaust
CO and THC Concentrations for Combustion of natural Gas
with 0.132 kg/sec of Air, Va/Vf=22, S-0, at 1 atm.
10
,-2
Ta( K)
742
860
SYM
CO
"
THC
O
D
10
-3
oc
0
2
O
O
<
en
LL
LLJ
_l
O
5
o
10-4
10-1
10-'
10
,-5
.0-3
0.6
0.8 1.0
OVERALL EQUIVALENCE RATIO,
1.2
30
-------
observed reductions in both CO and THC emissions from a gas turbine as the air inlet
temperature was increased from 1+56°K to 536°K.
A number of input-output tests were conducted to determine the effect of fuel composi-
tion (heating value) on exhaust NO concentrations. Two gaseous fuels with heating
values different from natural gas were employed in these tests a synthesized fuel
(CHti/Hg/CO), with a higher heating value of approximately k2 kJ/g-* and vaporized
propane, with a higher heating value of approximately 50 kJ/g. Measured exhaust NO
concentrations for these two fuels are compared with NO concentration data for natural
gas in Fig. 19. At zero swirl, exhaust NO levels for natural gas and the synthesized
fuel are essentially the same, while the NO levels for propane are somewhat larger.
Introduction of swirl increases exhaust NO concentrations for the synthesized
fuel, but to a lesser degree than observed with natural gas. For a given overall
fuel-air equivalence ratio, the firing rates (i.e., heat input per unit time)
for the three gaseous fueld are ordered as follows: synthesized gas > vapo-
rized propage > natural gas. From thermochemical considerations, one would expect
the trend in NO exhaust concentrations to follow that of the firing rates. The
observation that this is not the case, Fig. 19, suggests that there are significant
differences in the flow fields for the three gaseous fuels and that these differences
are influencing NO formation in the combustor. This point is examined in more
detail in the following section.
The final series of input-output tests was conducted using liquid propane as the
fuel. In these tests, the injector modification required for liquid fuel operation
is the only notable change in the combustor. The injector module for liquid propane
fuel resembled Injector I in that the air gap height, £, is 1.^99 cm. However, the
6.372 cm O.D. circular fuel delivery duct is capped off, in contrast to the configura-
tion used with gaseous fuels. A Delevan nozzle was installed at the center
of the cap, on the combustor centerline, to deliver a nominal 60 deg, solid cone
liquid spray. Thus, in the liquid fuel tests, a large base region was present.
Water cooling passages within the fuel delivery duct provided cooling of the base
region.
Significant trends in the exhaust sampling data are illustrated in Figs. 20-22.
Figure 20 shows the variation of the average NO concentration in the exhaust gas with
inlet air swirl, pressure and overall equivalence ratio for the combustor operating
on liquid propane and air. For purposes of comparison, exhaust NO concentrations for
*The fuel composition, and hence the heating value, was selected so as to provide air-
fuel velocity ratios which matched those employed in the natural gas tests. For the
composition selected, nominal air-fuel velocity ratios obtained using Injector I
approximate those obtained with natural gas using Injector II.
31
-------
Average Exhaust NO Concentrations for Combustion of Natural Gas,
Synthesized Fuel and Vaporized Propane. Combustion with 0.132 kg/sec
of T40°K Air at 1 atra. Va/Vf=22, for Natural Gas, Va/Vf=15 for the
Synthesized Fuel, Va'/Vf=35 for Vaporized Propane.
400
300
CC.
Q
a.
a.
I 200
z
o
cc
LL
UU
O
O
Z
100
FUEL
NAT.
GAS
SYN,
GAS
VAP.
OPANE
SYM
S = 0
D
A
O
S = 0.6
A
D-
Tnr
0.6
,0.8 1.0
OVERALL EQUIVALENCE RATIO.*
1.2
32
-------
FIG. 20
Average Exhaust NO Concentrations for Combustion of Vaporized
and Liquid Propane with 0.136 kg/sec of jhO°K Air.
cc
a
8:
i
g
o
DC
LL
111
O
500
400
300
o
z
200
100
VAPOR
A
LIQUID
A
D
s
0
0.3
0.3
P(ATM)
1
1
6.6
I
I
I
0.7
I
0.8 0.9 1.0 1.1
OVERALL EQUIVALENCE RATIO, $
1.2
R02-29-8
-------
Average Exhaust THC Concentrations for Combustion of Vaporized
and Liquid Propane with 0.136 kg/sec of TifO°K Air.
10-'
cc
Q
-------
FIG. 22
Average Exhaust CO Concentrations for Combustion of Vaporized
and Liquid Propane with 0.136 kg/sec of TUO°K Air.
0.05
0.04
EQUILIBRIUM CO
C3H8 GAS - 1 ATM -
0.03
cc.
Q
z.
o
cc
U-
LU
O
O
u
VAPOR
A
LIQUID
A
D
s
0
0.3
0.3
P (ATM)
1
1
6.6
0.02
0.01
0.7
A
I
0.8 0.9 1.0 1.1
OVERALL EQUIVALENCE RATIO, 4>
_
1.2
RO2-29-3
-------
vaporized propane also are plotted on Fig. 20. In all of the tests, the NO levels
in the exhaust are nearly independent of the overall fuel-air equivalence ratio.
For the swirling flow with liquid propane, a decrease in NO emissions was observed
when the combustor pressure was increased from one atmosphere to approximately 6.6 atm.
This trend is opposite to that observed using gaseous fuels and likely is due to
changes in the spray characteristics with increasing pressure (Ref. 13.). For the
swirling flow and a combustor pressure of one atmosphere, changing from prevaporized
to liquid propane produced a decrease in exhaust NO emissions.
Figure 21 shows the variation of the average THC concentration in the exhaust gas
with inlet air swirl, pressure and overall equivalence ratio for the combustor
operating on vaporized and liquid propane. In all of the tests, the THC levels in
the exhaust increased with increasing equivalence ratio. For a combustor pressure of
1 atm, the THC levels in the exhaust for vaporized propane decrease with increasing
inlet air swirl. For swirling flow and a combustor pressure of 1 atm, changing from
prevaporized to liquid propane produced very little change in the exhaust THC
emissions. For liquid propane, only a slight change in exhaust THC emissions was
observed when the combustor pressure was increased from one atmosphere to 6.6 atm.
Figure 22 shows the variation of the average CO concentration in the exhaust gas with
inlet air swirl, pressure and overall equivalence ratio for the combustor operating
on vaporized and liquid propane. For a combustor pressure of 1 atm, the CO levels
decrease slightly with increasing inlet air swirl. The measured exhaust CO levels
are very nearly equal to the calculated equilibrium CO levels over the entire
range of equivalence ratios investigated. For swirling flow and a combustor pressure
of one atmosphere, changing from vaporized to liquid propane produced very little
change in the exhaust CO emissions. For the swirling flow, with liquid propane, an
increase in combustor pressure from 1 atm to 6.6 atm produced a significant reduction
in the measured exhaust CO levels. At the elevated pressure, the measured exhaust
CO levels were less than calculated equilibrium values.
Fl£W_F^el1d_Ntep£ing_Expe_riments_
Examination of the results of the input-output tests, presented above, indicates
that pollutant emission levels are particularly sensitive to certain combustor input
parameters. NO and THC emissions are strongly dependent on swirl and pressure and
also are dependent on air-fuel velocity ratio. Changing fuel composition produces
some experimental trends which cannot be predicted on the basis of thermochemistry
alone, suggesting significant coupling between the fluid dynamic and chemical
processes in the combustor. Over most of the ranges of combustor input parameters
investigated, exhaust CO concentrations are the equilibrium levels; however, for
non-swirling flows and for swirling flows with low air-fuel velocity ratios, exhaust
CO levels significantly exceed equilibrium levels as the overall equivalence ratio
is decreased. Operation on liquid fuels produces some variations in emissions with
input parameters which are different from those observed using gaseous fuels. To
36
-------
investigate how changes in these significant input parameters influence the combustor
flow field and subsequent pollutant formation, detailed maps of the flow field
properties were obtained for the seven test conditions listed in Table II.
TABLE II. MATRIX OF NOMINAL TEST CONDITIONS: MAPPING EXPERIMENTS
*
Fuel
NG
NG
NG
NG
SYN
PL
PL
S
0
0
0.3
0
0
0.3
0.3
v.A,
22
22
22
22
15
--
--
P(atm)
1
1
1
3.6
1
1
1
,
0.9
0.7
0.9
0.9
0.9
0.9
1.0
Ta(°K)
7^0
7^0
7^0
7^0
7^0
7^0
71*0
wa (kg/sec)
0.132
0.132
0.132
0.132
0.132
0.132
0.132
* NG = natural gas
SIN - synthesized fuel
PL = liquid propane
In these mapping experiments, the radial distributions of axial flow velocity,
flow direction, temperature and species concentrations were acquired at all six
combustor port locations. Variations in swirl number, pressure, overall equivalence
ratio and fuel heating values are represented by the test conditions selected.
During this effort, the two-probe rake was traversed in a "stop-go" mode at each
axial location, i.e., the probe was moved to a desired radial position, data were
taken and the probe was moved to a new radial position, and so on. A typical
radial traverse consisted of 8-12 measurements. Approximately U-5 minutes were
required at each radial position to conduct on-line species concentration determina-
tions.
Dynamic pressure, 0.5 0v , and flow direction were computed from impact pressures
and differential static pressures (i.e., impact pressure minus static pressure)
following the method given in Appendix B. DSO impact pressure and second orifice
total temperature and pressure provided the necessary input for the determination
of stream temperature. Ultimately, velocity was extracted from the temperature and
dynamic pressure information.
An examination of the data generated during the mapping experiments reveals that
there are several common features for all five of the gaseous fuel test conditions.
It is worthwhile, therefore, to consider the data for one of the test conditions in
37
-------
some detail and to point out similarities and differences with respect to the re-
maining four test conditions. The test condition chosen for illustrating the sig-
nificant characteristics of the confined, mixing/reacting flow, designated the
baseline condition, involves the combustor operating on natural gas-air at an
overall equivalence ratio of about 0.9» a pressure of 1 atm, with no swirl on
the inlet air.
A preliminary description of the flow field was obtained from high-speed (500 frame/sec,
motion pictures of the flow in the vicinity of the injector. The motion pictures
for "baseline" operation disclosed a light blue flame whose initial ignition loca-
tion varied with time from axial positions downstream of port location "0" to
positions inside the fuel delivery duct. The frequency of movement of the ignition
location was estimated to be 40-100 Hz. On the basis of these observations, it
can be concluded that the aerodynamically-induced recirculation zone is very mobile
and, therefore, that its influence is extended over a larger portion of the flow
than would be surmised on the basis of its size. At any given moment it might be
far downstream of the injector exit so that the central fuel and annular air
streams are unaffected by its presence, or it might be partially swallowed by the
fuel delivery duct in which case hot air is recirculated to mix and burn with fuel
upstream of the injector exit plane. Therefore, actual flow conditions at the in-
jector exit will reflect recirculation behavior to some degree, thereby complicating
the specification of initial conditions. The observed instability is primarily a fluid
dynamic instability and is not the result of a coupling of the combustion process
with the acoustic properties of the combustor or mechanical properties of the
injector.* Similar recirculation zone instabilities have been observed in non-
reacting flow fields for geometries nearly identical to the present combustor
geometry (Ref. ih).
Baseline case velocity and temperature distributions at several axial locations
are given in Figs. 23 and 24, respectively. In these figures and in all others to
follow, the symbols in Table III will be used to identify the axial location of
the probe tip. In addition, the extent of the fuel delivery port and air annulus
will be indicated on the figures by vertical dashed lines normal to the abscissa.
* It should also be mentioned, for completeness, that a pressure instability was
noted for combustor operation at 1 atm. Visicorder records of the signals from
a pressure transducer located at the downstream end of the combustor section re-
vealed that the magnitude of the fluctuations increased with overall equivalence
ratio while the frequency of the temporal component remained unchanged at 130 Hz
for both swirling and non-swirling flows. Amplitudes of 0.092 atm, 0.136 atm and
0.1^3 atm were recorded for 4> = 0.8, 1.0 and 1.2, respectively, during tests at
zero swirl.
38
-------
TABLE III. LEGEND FOR MAPPING DATA
Port
No.
1
2
3
k
5
6
Axial Probe
Tip Location*
x(cm)
1.93
31.60
48.59
65.58
82.58
99-57
x/ro
0.32
5.17
7.95
10.73
13.50
16.28
Symbol**
V
A
0
D
D
o
Injector exhaust plane is at
x=0
** Solid symbols indicate either:
(l) entry through opposing port or
(2) non-basline data in comparative
plots.
Generally, when data from ports on both sides of the combustor are presented, the
ordinate label will be placed to read horizontally in the upper right hand corner
of the plot.
The velocity profile in the vicinity of the injector exit plane for the baseline
aerodynamically-stabilized flame, Fig. 23, is highly distorted relative to the
theoretical, inviscid situation of a step profile with a 22.9:1 velocity ratio,
as the result of the recirculation zone being imbedded in the fuel injection port
and "boundary layer effects. In addition, the chaotic movements of the recirculation
zone, discussed above, effectively stir the fuel and air streams to reduce gradients
(except those associated with the wall boundary layer) prior to the x=31«6 cm axial
location. Downstream temperature distributions, Fig. 24, also are relatively flat
and unchanging with longitudinal distance, suggesting that only a small portion of
the total available chemical heat release occurs beyond the initial stirred zone
and that such heat production is approximately matched by heat loss to the cold
combustor wall. The behavior exemplified by the data in Figs. 23 and 24 is that of
a stirred reactor, of length lees than 31.6 cm, followed by a plug flow region in
which the (inviscid) velocity is 82± 8 m/sec and the temperature is approximately
1900°K. Such a system appears to be chemistry-limited in that mixing is rapid
relative to characteristic chemical reaction times.
By virtue of the rapid transport associated with the fluid mechanics in the base-
line case, relatively high oxygen concentrations are available throughout the fuel
stream (See Figs. 25 and 26). Depletion of oxygen is rapid upstream of the -x=31.6-cm
39
-------
Velocity Profiles for Baseline Case. Combustion of Natural Gas
(Tf=303cK) with 0.132 kg/sec of T52°K Air, VQ/Vf =22.9, S=0, at
1 atm.
RADIAL POSITION, r/rc
-------
FIG. 24
Temperature Profiles for Baseline Case.
2000-
V^
1.0
0.5
0.5
RADIAL POSITION, r/ro
-------
Oxygen Concentration Map for Baseline Case.
Exhaust Concentration = 0.02.
Average
0.20-
RADIAL POSITION , r/rc
-------
FIG. 26
THC Concentration Map for Baseline Case.
Exhaust Concentration = 1300 ppm.
Average
0.025-
THC MOLE FRACTION
A
RAD IAL POSITION, r/rn
N12-139-5
-------
location, i.e. in the highly stirred region of the combustor. In this region
there is significant oxidation of the hydrocarbon fuel to form CO, Fig. 27, which
in turn reacts forming C02, Fig. 28, and there is significant nitrogen oxide forma-
tion, Fig. 29. Downstream from the x=31.6 cm location, oxygen is consumed at a
slower rate as the hydrocarbons continue to burn out. Concurrent with the hydrocarbon
burnout, the CO increases and then burns out and C02 and nitrogen oxides form.
Referring to Fig. 25, oxygen concentration at the most upstream measurement loca-
tion is seen to decrease rapidly from the air annulus position, as through a shear
layer, and achieve a constant concentration level from r/ro « 0,k to the centerline
(r/r0=0). Examination.of the CO and C02 concentration profiles, Figs. 27 and 28,
at the same location also shows relatively constant concentration levels in the
vicinity of the centerline. This central well- mixed portion of the flow,
covering about 6$% of the area of the fuel delivery duct, reflects the time-
average extent of the recirculation region.
The regions of NO and N02 formation in the combustor are indicated in Fig. 29. At
the axial location nearest the injection plane, the NO concentrations are very small;
however, in the vicinity of the fuel injector lip, N02 concentrations in excess of
the NO concentrations were measured. In several recent investigations of nitrogen
oxide formation in gaseous turbulent diffusion flames (Refs. 2, 16 and 17), relative.!?
large (NO )/(NO) ratios were measured near the combustion zone. It should be noted,
however, that there is considerable difficulty in making quantitative measurements
of N02 in the combustor (See Appendix C). It was not possible to measure N02 con-
centrations in fuel-rich regions in the combustor because of problems with the
stainless-steel thermal converter in the chemiluminescent analyzer. Similar
limitations of stainless-steel converters have been reported by other investigators
(Ref. 15). Furthermore, in regions where N02 measurements were obtained there is
considerable uncertainty in the observed levels because of potential sources and
sinks for NO during transfer of the gas sample to the Exhaust Analyzer. Signi-
ficant levels of NO and NO are achieved within 5 combustor radii downstream
of the injector exit. Thereafter, N02 concentration remains relatively constant
with axial distance, while NO continues to form at an approximately steady rate;
the centerline value of NO increases at about 1.36 ppm/cm.
The temperature and concentration distributions typically are not symmetric about
the combustor centerline. The apparatus itself was eliminated as the source of
this asymmetry following extensive inspections and careful alignment of all com-
ponents at the test site and recalibrations of traversing gear and analytical
instrumentation. Thus, the asymmetry must be related to the erratic motion of
the recirculation zone, and it may be anticipated that profile three-dimensionaltiy
will be aggrevated (or improved) with changes in air-fuel.velocity ratio and swirl
number. This preliminary observation is explored in more detail later in the
report.
The general flow field characteristics discussed above are typical of all of the
experimental conditions investigated in the flow field mapping experiments. How-
ever, as combustor operating conditions are varied from the "baseline" conditions,
-------
CO Concentration Map for Baseline Case,
Exhaust Concentration = 0.016.
0.040-
FIG. 27
Average
RADIAL POSITION .
N12-139-2
-------
C02 Concentration Map for Baseline Case.
Exhaust Concentration = 0.0881.
Average
0.09-
1.0
RADIAL POSITION. r/rr
-------
NO and N02 Concentration Maps for Baseline Case.
Average Exhaust Concentration: N0=156 ppm and
N02=30 ppm.
FIG. 29
200-
NO,N02 MOLE FRACTION (PPM)
RADIAL POSITION, rhr
N12-139-4
-------
there are noticeable changes in the velocity, temperature and species concentration
distributions in the ccmbustor. In the following sections, these differences'will
"be discussed in detail.
i. Effect of Overall Equivalence Ratio
Decreasing the overall fuel/air equivalence ratio from the baseline value of 0.9
(~ 11 percent excess air) to 0.7 (~^0 percent excess air), with all other inlet
conditions remaining the same,produces only minor changes in the velocity and tempera-
ture distributions shown in Figs. 23 and 2U. Observed temperature levels decrease
by 100-1509K and the maximum flow velocity decreases to ?0±5 m/sec, primarily
because of density increases associated vith the decreased temperatures.
The 02 concentration distributions for4>=0.7, Fig. 30, are qualitatively similar to
those found for the baseline conditions, although the Og concentrations generally
are larger for the <|>=0.7 operating condition. The time-mean radial extent of the
recirculation zone for the $=0.7 operating condition, as indicated by the region
of relatively constant 02 concentrations, is approximately the same as for the base-
line case.
Oxidation of the hydrocarbon fuel to CO is more rapid at the lower overall equiva-
lence ratio, resulting in hydrocarbon levels which are lower, Fig0 31, and initial
CO levels which are higher, Fig. 32, than for baseline operation. Although initial
CO levels are high for the 4>=0.7 operating condition, CO burnout rates are rapid
and exhaust concentration levels are only about 36$ of the levels measured for the
baseline case.
At the axial station nearest the injection plane, COo concentrations for the flow
with ^0 percent excess air are somewhat larger than were measured for the baseline
operating condition, Fig. 33« Near the combustor wall, a region of locally high
C02 concentration is observed, suggesting a region of recirculating flow just
downstream from the back-step at the injection plane. Similar evidence of recircula-
ting flow in this region was obtained in the swirling flow experiment. At axial
stations further downstream from the injection plane, C02 concentrations in the kO
percent excess air flow are generally Icrwerthan in the baseline flow, consistent
with the equivalence ratio trend of the exhaust C02 concentration reported
previously Fig. 8.
Except for the axial station near the injection plane, the NO concentration profiles
are not significantly different for the two equivalence ratios investigated. Near
the injection plane, NO concentrations near the combustor centerline are somwehat
larger for$= 0.7 than for= 0.9. These differences appear to correlate with
the temperature measurements, in that the measured temperatures near the centerline
are larger f or =0.9. The above observations, together with the fact
that the exhaust NO concentrations for the two overall equivalence ratios are
approximately the same, suggest that there are no significant differences in the
NO formation rate. This conclusion is consistent with the observation that there
are only very slight differences in the temperature and 02 concentration profiles
in the regions of maximum NO formation in the ccmbustor.
U8
-------
FIG. 30
Comparison of the Og Concentration Distributions at the Baseline Condition
(*=*0.9) with Those at«t>^0.7. At the Lower Equivalence Ratio, Natural Gas
(307°K) is Burned with 0.136 kg/sec of T5^°K Air, Va/Vf=29.2, 4>=0.68,
3=0 at 1 atm.
!
i.o
$=0.9
IXE = 0.020)
POS. 1
POS. 6
It
0.5
0.25-
02 MOLE FRACTION
0.20 4>=0.7
(KE = 0.058)
I
II
0.5
1.0
RADIAL POSITION. r/rr
1*9
N12-153-13
-------
Effect of Overall Equivalence Ratio on THC Concentration Map.
Operating Conditions Identical to Fig. 30-
0.020-
RADIAL POSITION, r/rf
-------
FIG. 32
Effect of Overall Equivalence Ratio on CO Concentration Map.
Operating Conditions Identical to Fig. 30.
-0.04-
CO MOLE FRACTION
RADIAL POSITION . r/r
N12-153-16
51
-------
Effect of Overall Equivalence Ratio on C02 Concentration Map,
Operating Conditions Identical to Fig. 30.
CO2 MOLE FRACTION
RAD IAL POSITION. r/rr
-------
ii. Effect of Fuel Composition
Substitution of the synthesized fuel (CHi/CO/H2) for natural gas in the baseline
case lowers the heating value by about 2k% and, more significantly, reduces the
air-fuel velocity ratio from 22.1 to 15.1*. This last change brings the air-fuel
momentum flux ratio uncomfortably close to the minimum value necessary to sustain
an aerodynamic flameholding capability. The recirculation region for these condi-
tions undergoes large scale displacements which leads to a flow field with major
three-dimensional features. In particular, the centrally-located, highly mobile
recirculation region may merge with other recirculation regions, e.g., the annular
region associated with the injector back-step. Detailed maps for the synthesized
fuel-burning combustor support these remarks. Large distortions of the velocity
distribution at x=31.6 cm, Fig. 3k, testify to the unsettled flow situation which
exists. The peculiar temperature distribution, Fig. 35, at that same station creates
the impression that the recirculation region is located off-center and that the fuel
stream has been channeled toward the combustor wall. These observations are con-
sistent with the central recirculation region merging with or establishing inter-
mittant contact with a portion of the recirculation region asosciated with the
injector step. In this event, the recirculation zone may occupy a large, irregular
region encompassing the centerline and extending to cover a portion of the combustor
wall. The fuel stream is forced away from the central portion of the combustor to
move around the recirculation region. All symmetry is destroyed in the initial
portion of the combustor and three-dimensional effects persist for a significant
distance downstream. As indicated in Figs. 3k and 35, three-dimensional effects
are extreme at 5 combustor radii, but only minor at 10.7 radii.
In spite of the quite different composition of the synthesized fuel versus natural
gas (Appendix F), and the major nonuniformities described earlier, measured species
concentration levels and profile shapes for the two fuels are remarkably similar.
Pigs. 36 and 37 show comparisons of the concentration distributions of C02 and
NO, respectively, for the two fuels. For the synthesized fuel, the C02 concentra-
tions in the conbustor are somewhat larger than for natural gas, reflecting the
increased effective carbon/hydrogen ratio of the synthesized fuel. Immediately
donwstream from the injector, relatively large C02 concentrations near the wall
and depleted concentrations in the center of the combustor are consistent with
the three-dimensional flow behavior discussed above. NO also appears in larger
concentrations adjacent to the wall than in the center of the combustor at the
first axial location. At axial locations further downstream, the NO concentration
profiles are nearly identical for the synthesized fuel and natural gas, and the
exhaust concentrations are approximately the same for these two fuels. These
observations suggest that there is no significant difference between the NO forma-
tion rates for these two fuels for the set of operating conditions investigated.
This conclusion is consistent with the temperature and 02 concentration profiles
measured in the combustor for the two fuels. One additional observation (not
shown) is the presence of N02 in the vicinity of the air-fuel shear layer at
x=1.93 cm at levels exceeding the local NO concentrations; a similar situation
was noted earlier for the baseline case.
53
-------
Velocity Profiles for Combustion of Synthesized Fuel (Tf=309°K)
with 0.132 kg/sec of T30°K Air, Va/Vf=1^.6, S=0, * =0.87, at 1 atm.
--75 \
I I,
50
25
POS. 2 /
I I
1.0
0.5 0 0.5
RADIAL POSITION, r/ro
1.0
-------
FIG. 35
Temperature Profiles for Synthesized Fuel Combustion Test.
2000-
RADIAL POSITION, r/r.
R1-60-1
-------
C02 Concentration Maps for Synthesized Fuel and Baseline
(Natural Gas) Combustion Tests.
0.09-
CO2 MOLE FRACTION
RADIAL POSITION . r/rr
-------
FIG. 37
NO Concentration Maps for Synthesized Fuel and Baseline
(Natural Gas) Combustion Tests.
200-
s.
s
NO MOLE FRACTION,(PPM)
NATURAL GAS
(XE = 156PPM)
IZ
POS. 1
SYNTHESIZED GAS
165PPM)
X
1.0
1.0
0.5
0.5
RADIAL POSITION, r/rr
57
N12-153-14
-------
iii. Effect of Swirl
In mapping experiments discussed so far, the flame is stabilized on a recirculation
region produced as the result of a momentum flux deficit "between the fuel and air
streams, without imparting swirl to either flow. The great mobility of the recircu-
lation zone results in rapid transport of heat, mass and momentum and the formation of
an essentially completely mixed flow. The mixed flow downstream of the recirculation
zone contains reactive species which continue to 'undergo chemical changes. Distri-
butions of temperature, velocity and species concentrations in such flows are char-
acteristically flat in the center portion of the combustor and the maximum (or
minimum) is at the centerline. These chemical- reaction-limited systems display
relatively low NO levels and high THC levels.
When a swirl component is introduced on the inlet air flow, the appearance of
the combustion zone in the vicinity of the injector changes dramatically. High-
speed motion pictures reveal that the extent of the axial movement of the ignition Ipca
tion is reduced and that the ignition location rotates around the injector lip with a
frequency of approximately 150 Hz. Syred and Beer (Ref. 18) have reported similar
periodic instabilities in swirl combustors, resulting from a precession of vortices
about the axis of symmetry. In the flow fields investigated in Ref. 18', the pre-
cessing vortex core (PVC) was located between the boundary of the recirculation
zone and the zero streamline, corresponding to a region near the fuel injector lip
in the present experimental configuration. Syred and Beer have correlated the FVC
frequency, f , in terms of a non-dimensional frequency parameter, fd3/Q^ and the
Reynolds number of the inlet air stream for a range of fuel injection geometries
and overall fuel/air equivalence ratios. The frequencies of the rotational instabili-
ties observed in the present experiments can be expressed in terms of the Syred and
Beer correlation, suggesting that the observed instability is a PVC. Associated
with the PVC, Syred and Beer report a radial-axial instability resulting from eddy
shedding from the FVC. A similar axial instability is observed in the present
experiments. Hence, in the vicinity of the injection plane, the flow is not
axisymmetric and steady, but instead is a three-dimensional, time -dependent turbulent
flow. The axial extent of the FVC in the present experiments is not known, however
Syred and Beer report that the FVC dissipates rapidly in the axial direction
Although there are substantial time-dependent, three dimensional effects, at the
S=0.3 condition investigated, the time-mean flow appears to resemble a classical
axisymmetric diffusion flame. High temperatures develop off centerline, Fig.. 38, by
virtue of a separation of the air and fuel streams, resulting in a mix ing -limited
situation. In the extreme mix ing -limited case of a "flame sheet", oxygen moving
toward the centerline and fuel moving away from the centerline react exclusively
within a very narrow, high temperature annular region, thereby sustaining the local
temperature. Reaction products display sharp concentration peaks in the combustion
region, while reactant species concentrations fall of f abruptly at the boundary of
58
-------
FIG. 38
Temperature Profiles in Swirling Flow Combustion Test. Com-
bustion of Natural Gas (Tf=360°K) with 0.13^ kg/sec of 750°K
Air Having a Swirl Number of 0.3, Va/Vf=23, 4>=0.885, 1 atm.
2500-
RADIAL POSITION, r/rr
R1-60-3
59
-------
the flame. In this manner, the flame sheet acts as a sink for reactants and a
source of products.
Mixing-limited behavior is clearly manifested in the species concentration data
acquired for the swirling flow. Figs. 39 a"d 40 contrast the separation of the
fuel and air streams at initial axial locations (x/rQ «s 5.1?) for the swirling flow
with the non-swirling "baseline case. Similar radial distributions of fuel and air
in turbulent diffusions flames have been reported by Bilger (Ref. 19). The increase
in hydrocarbons near the combustor wall at x/r = 5-17 is probably due to entrain-
ment of fuel in the recirculating flow in the vicinity of the back-step at the
injection plane. Fig. 4l compares the CO concentration distributions for the swirling
and non-swirling flow. The initial rate of CO formation in the swirling flow is
substantially larger than for the non-swirling flow. This observation, together
with the rapid disappearance of hydrocarbons and oxygen in the swirling flow, Figs.
39 and 40, suggests that swirl significantly increases the rate of oxidation of the
fuel to CO. Since the exhaust CO levels are lower for the swirling flow than for
the non-swirling flow, it appears that swirl also results in an increased rate of,
CO burnout.
The radial concentration distributions of C02 and NO, Figs. 42 and 43, for the swirlim
flow are significantly different from those in the non-swirling baseline case.
Imparting swirl to the inlet air stream results in C02 and NO concentration profiles
which exhibit peaks off the combustor centerline, at approximately the same radial
position as the maximum temperature. The C02, NO and temperature peaks (and, hence,
unmixedness) persist for long distances in the combustor, and spread at an angle
of approximately 5.5 deg. This angle is consistent with spreading rate data obtained
for unconfined turbulent diffusion flames (Ref. 20).
Combustion in the swirling flows resembles the classical turbulent diffusion flame.
These flows are characterized by substantial unmixedness, higher combustion rates
and locally higher temperatures in the flow field, resulting in increased rates
of NO formation. In swirling flows, the region of maximum energy release occurs
off the combustor centerline, as evidenced by the temperature and concentration
profiles. Evidently, the unmixedness increases with an increase in the swirl number
from 0.3 to 0.6, since results from the input-output experiments show that exhaust
levels of NO rise and THC levels decrease as the swirl number is increased.
The persistance of swirl-induced effects is evident from the concentration and
temperature distributions and also from the axial velocity and flow direction
distributions. Axial velocity profiles for the S=0.3 flow Fig. 44 are found to
be symmetric and to vary in a nearly linear fashion with radial distance in the
central portion of the combustor, i.e., (V-V(?)/r= constant. The acceleration rate
of the flow in the center of the combustor decreases with axial distance, and a
substantial velocity deficit still exists at the centerline at 16.3 combustor radii
(x=99-57 cm).
60
-------
FIG. 39
Oxygen Concentration Maps for Swirling (S=0.3) and Baseline
(S=0) Combustion Tests.
02 MOLE FRACTION
RADIAL POSITION, r/rfl
61
-153-2
-------
FIQ.40
Total Hydrocarbon Concentration Maps for Swirling (S=0.3) and
Baseline (S=0) Combustion Tests.
0.03-
s = o
(YE = 1300 PPM)
THC MOLE FRACTION
1.0
0.5 0 0.5
RADIAL POSITION . r/r0
1.0
62
-------
FIG. 41
CO Concentration Maps for Swirling (S=0.3) and Baseline (S=0)
Combustion Tests.
0.08-
CO MOLE FRACTION
RADIAL POSITION , r/r.
N12-153-10
63
-------
C02 Concentration Maps for Swirling (S=0.3) and Baseline (S=0)
Combustion Tests.
0.10-
C02MOLE FRACTION
0 0.5
RADIAL POSITION , r/ro
61*
-------
FIG. 43
NO Concentration Maps for Swirling (S=0.3) and Baseline (S=0)
Combustion Tests.
NO MOLE FRACTION ( PPM
°-5 0 0.5
RADIAL POSITION , r/rn
65
N12-153-6
-------
\
Velocity Profiles in Swirling Combustor Flow (S=0.3).
- -100
II
150-
VELOCITY, M/SEC
POS6
--50
I I
I I
I I
\
1.0
0.5
0.5
1.0
RADIAL POSITION, r/r0
66
-------
Measured flow angles are given in Fig. 45. With the impact-static probe moving
toward the combustor centerline in the horizontal plane, pitch angles are measured
in the horizontal plane, and yaw angles are measured in the vertical plane which is
parallel to the main flow direction. Pitch angles, oi} are positive (negative) for
flow moving towards the probe from the wall (centerline). In the S=0.3 flow, pitch
angles are small, a £ 10 deg. The flow in the pitch plane is directed toward the
wall for 0.5 ^ r/r0 ^ 1«° and toward the centerline for radial locations less than
r/r0 «0.5. Yaw angles, g, are small near the centerline of the comb us tor and
increase as the wall is approached. For typical yaw angles of 25-30 deg, the tan-
gential velocity is approximately one-half of the axial velocity. The tangential
current moves down one side of the combustor and up the other in a spiral fashion
along the entire combustor length. Total flow angles, 6 in Fig. 4l, where e^cos"1
[(cos or) (cos 0)1, (approximately equal to the yaw angle in the present case), provide
further verification of the persistance of swirl-induced effects. Thus, for example,
maximum total flow angle is only reduced 8 deg between the x=31.6 cm and the x=82.6
cm axial locations.
iv. Effect of Combustor Pressure
Increased pressure leads to enhanced reaction rates and a shift from chemistry-
limited to mixing-limited conditions. It may be anticipated, therefore, that tem-
perature and concentration profiles will develop off-axis peaks as the pressure
level is raised, and that NO production will increase relative to the baseline
case, and THC concentrations will decrease. This diffusion flame behavior is
indeed displayed by the temperature map of Fig. 46 and the various concentration
maps presented in Figs. 47-51, for combustor operation at about 3.6 atm with all
other fuel and air input .conditions being the same as the baseline case.
Oxygen and THC concentration profiles, Figs, 47 and 48, for 3.6 atm reflect sig-
nificant separation of the air and fuels stream, and resemble profiles encountered
in classical diffusion flames. As in the swirling flow, combustion is concentrated
in an off-axis annular region, characterized by maxima in temperature and in the
concentrations of CO, Fig. 49, C02, Fig. 50, and NO, Fig. 51. At the elevated
pressure the rate of oxidation of hydrocarbons to CO is rapid due to the combina-
tion of high peak temperatures and longer residence times. Centerline CO concen-
tration levels at 3.6 atm are very high relative to the 1 atm baseline case in
the first meter of combustor length, reflecting the rapid initial hydrocarbon oxida-
tion and the limited availability of oxygen in the central portion of the combustor
due to unmixedness. Ultimately, far downstream in the extender section, oxygen is
more thoroughly diffused throughout the flow and CO oxidation proceeds unhindered.
At the exhaust sampling station, therefore, average CO and C02 concentrations are"
comparable for 1.0 and 3^6 atm.
Increased temperatures and longer residence times associated with the high pressure
operating conditions result in an appreciable increase in the NO formation rates
67
-------
Direction of Flow in Combustor with S»0.3.
40
30
ui 20
o
10
§ 3-^-°-
-10
A V A
0.5
RADIAL POSITION, r/ro
1.0
68
-------
FIG. 46
Temperature Profiles for Baseline (P=l atm) Combustion
Test and for High Pressure (P=3«6 atm) Combustion of
Natural Gas (Tf=305°K) with 0.137 kg/sec of 750°K Air,
Va/Vf=22.9, $=0.87, S=0.
RADIAL POSITION, r/r0
69
-------
Oxygen Concentration Maps for High Pressure (P=3.6 atm) and
Baseline (P=l atm) Combustion Tests.
P= 1.0 ATM
{ VE = 0.020)
-0.25-
O2 MOLE FRACTION
P = 3.6 ATM
(VE = 0.0276)
- - 0.20
POS. 1
1.0
0.5
0.5
1.0
RADIAL POSITION . r/rr
TO
-------
FIG.48
Total Hydrocarbon Concentration Maps for High Pressure
(P=3.6 atm) and Baseline (P=l atm) Combustion Tests.
THC MOLE FRACTION
P- 1.0 ATM
(YE =1300 PPM)
0 0.5
RADIAL POSITION , r/r
N12-153-15
71
-------
CO Concentration Maps for High Pressure (P=3.6 atm) and
Baseline (P=l atm) Combustion Tests.
0.12
CO MOLE FRACTION
P = 3.6 ATM
(VE =0.0207)
0 0.5
RADIAL POSITION , r/ro
72
-------
FIG.50
C02 Concentration Maps for High Pressure (P=3-6 atm) and
Baseline (P=l atm) Combustion Tests.
P = 1.0 ATM
(YE =0.0881)
POS. 1
0.09-
C02 MOLE FRACTION
P = 3.6 ATM
(VE =0.0875)
--0.03
RADIAL POSITION. r/rr
73
N12-153-9
-------
NO Concentration Maps for High Pressure (P-3.6 atm) and
Baseline (P=l atm) Combustion Tests.
FIG. 61
T500-
.. 400
P = 1.0 ATM
(VE= 156 PPM)
NO MOLE FRACTION (PPM)
1.0
0.5 0
RADIAL POSITION . r/ro
0.5
71*
-------
relative to baseline operation (Pig. 51). Introduction of swirl has only a minor
effect on exhaust NO levels (see results from input-output experiments) at the
elevated pressure, suggesting that swirl produces only minor changes in the flow
field.
N02 levels measured at 3.6 atm were much smaller than those at 1 atm. No N02 was
found in the shear region near the lip of the fuel injector, in contrast to the
results of the 1 atm tests.
v. Liquid Propane Tests
Results from the input-output experiments indicate that operation of the combustor
on liquid propane produces variations in pollutant emissions levels with input
parameters which differ from those found with gaseous fuels. To gain some pre-
liminary understanding of the flow field when burning liquid fuels and to evaluate
experimental problems encountered when probing two-phase reacting flows, two
mapping tests were carried out with the combustor operating on liquid propane.
Nominal test conditions for these experiments are listed in Table II.
High-speed motion pictures of the reacting flow field in the vicinity of the
injection plane reveal that the fuel-spray is surrounded by a light blue-orange
flame (at one atmosphere pressure) rotating at a frequency which correlates
approximately with the tangential velocity of the swirling air stream. In addi-
tion, a significant axial movement of the ignition location is observed, and the
fuel spray pattern downstream from the injector appears to be non-steady.
Although time-dependent, three-dimensional effects are evident in the swirling
flow liquid propane tests, the time-mean flow resembles a classical axisymmetric
diffusion flame (except, perhaps, in the vicinity of the spray) and is qualitatively
similar to that observed in the swirling natural gas flame. The temperature
profiles for the liquid propane tests, Figs. 52 and 53, exhibit peaks off-center-
line. Peak temperature levels in the liquid propane tests were approximately the
same as found in the swirling natural gas flame, although lower centerline
temperatures were found in the propane tests. The 02 and THC concentration maps
for liquid propane combustion, Figs. 5^ and 55, also resemble those found in
the swirling natural gas flame, with oxygen being confined, more or less, to
an. annular region near the combustor wall and fuel being confined near the
combustor centerline. The CO and C02 concentration maps for liquid propane
combustion, Figs. 56 and 57, are qualitatively similar to those found in the swirl-
ing natural gas flame, with CO being confined to a region near the centerline and
co exhibiting concentration peaks off-centerline. in the liquid propane tests,
hy§rocarbon and CO oxidation rates appear to be somewhat slower than in the swirling
natural gas flame. These reduced oxidation rates are reflected in the results of
the input-output tests which show generally higher THC and CO emissions for the
liquid propane tests. Examination of the concentration maps suggests that the
-------
FIG. 52
Temperature Profiles for Combustion of Liquid Propane with 0.137 kg/sec
of 739 K°Air, S=0.3, at 1 atm. Overall Fuel-Air Equivalence Eatio = 0.92.
Data Obtained Using Calibrated Thermocouple Probe.
2500-
RADIAL POSITION, r/r.
76
-------
FIG. 53
Temperature Profiles for Combustion of Liquid Propane with 0.137 kg/sec of
739°K Air, 3=0.3, at 1 atm. Overall Fuel-Air Equivalence Ratio - 1.04.
Data Obtained Using Calibrated Thermocouple Probe.
2500-
RADIAL POSITION, r/ro
R02-29-2
T7
-------
Oxygen Concentration Maps for Liquid Propane Combustion Tests.
FIC
0.12-
RADIAL POSITION, r/ro
78
R02-29-10
-------
Total Hydrocarbon Maps for Liquid Propane Combustion Tests.
0.10-
4> = 1.04
-------
FIG. 66
CO Concentration Maps for Liquid Propane Combustion Tests.
0.14-
0.5
1.0
RADIAL POSITION, r/rr
80
-------
FIG. 57
CO Concentration Maps for Liquid Propane Combustion Tests.
0.09-
RADIAL POSITION, r/r.
81
R02-29-5
-------
slower oxidation rates primarily are the consequence of a greater separation of the
fuel and air streams (i.e. decreased availability of 02 near the centerline) in the
liquid fuel tests. The NO concentration profiles, Fig. 58, in the liquid propane
tests also exhibit off-centerline peaks. The radial locations of these peaks
correspond approximately with the maximum temperatures. NO concentrations in the
liquid propane tests generally are lower than measured in the swirling natural gas
flame. The lower NO levels in the liquid fuel tests probably result from a reduction
in the high-temperature,, 02-rich volume in the combustor.
The temperature and concentration distributions presented above document the diffusion-
flame character of the flow in the liquid propane combustion tests. In these tests,
unmixedness, which persists for long distances in the combustor, has a significant
influence on hydrocarbon and CO oxidation rates and on NO formation. It can be
anticipated that spray characteristics (in particular, spray cone angle) will
significantly influence radial temperature and concentration gradients in the
flow. As combustor operating pressure increases, the spray cone angle decreases
(Ref. 13), and the region of maximum temperature will be somewhat closer to the
combustor centerline, encompassing a somewhat smaller volume. Hence, it is reasonable
to expect lower NO emissions than found in the atmospheric pressure tests. This
observation is in agreement with the trend established in the input-output tests.
Certain difficulties were encountered in probing the two-phase regions of the flow
field. The relatively cold exterior probe surfaces served to quench chemical reac-
tions and, in the vicinity of the fuel spray, substantial deposits of soot were formed
on the probe. These deposits frequently blocked the probe inlets, hence it was not
Possible to obtain flow field data at the x=1.92 cm location.
82
-------
FIG. 58
NO Concentration Maps for Liquid Propane Combustion Tests.
200-
IMOMOLE FRACTION (PPM)
1.04
E = 260 PPM)
I
1.0
$=0.92
(VE = 255 PPM)
I
0.5
0.5
RADIAL POSITION, r/r.
1.0
83
R02-29-11
-------
III. CONCLUDING REMARKS
Prom the results of the flow-field mapping experiments, several general observations
can be made concerning the interaction between the flow field and the pollutant
formation and destruction processes. For the range of conditions investigated the
primary hydrocarbon combustion zone, where most of the hydrocarbon fuel reacts to
form CO, is located relatively close to the injector, i.e., within an axial length
of about 5 combustor radii. In this zone which contains the stabilizing ^circula-
tion zone, a significant fraction of the total available chemical energy is released
and high levels of CO and NO are established. Peak CO levels in the primary zone '
are significantly larger than equilibrium levels,calculated based on inlet conditions.
Measured N02 concentrations in the primary combustion zone are relatively large.
Near the injection plane, N02 concentrations in excess of the NO concentrations
were measured in the shear region separating the inner and outer flows. Downstream
from the primary combustion zone, the remaining hydrocarbon fuel and the CO oxidize.
There are no significant increases in temperature in this post-combustion zone;
however, because of the relatively high temperatures, significant additional NO
formation occurs. In the post-combustion zone, the N02/N0 ratio decreases, primarily
because of the increasing NO concentration.
In the present experiments, the reacting flows can be roughly characterized as
chemistry-limited or mixing-limited, depending on the relative rates of chemical
reaction and mixing. In the non-swirling, atmospheric pressure flames, the large
scale fluctuations of the stabilizing recirculation zone enhance mixing of the fuel
and air streams and result in a primary combustion zone which resembles a stirred '
reactor. Downstream from the primary zone, the flow resembles plug flow. These
flows appear to be chemistry-limited in that mixing rates are rapid relative to
chemical reaction rates. As a result of the well-mixed nature of the flow, tempera-
tures are moderated so that hydrocarbon and CO oxidation rates and NO formation
rates are relatively slow. Hence, chemistry-limited flows are characterized by
relatively high THC and CO emissions and relatively low NO emissions.
increasing combustor pressure or imparting swirl to the inlet air streams results in
a shift from chemistry-limited to mixing-limited behavior. In mixing-limited flows
the time-mean flow field exhibits the characteristics of a diffusion flame, with '
separation of the fuel and air streams and with chemical reaction occuring principally
in an annular region surrounding the combustor centerline. Maximum temperatures ln
the region are considerably larger than found in the chemistry-limited flows,
resulting in increased hydrocarbon and CO oxidation rates and NO formation rates.
«ndC^Bi^~lllBlted fl°WS are characterized by relatively low THC and CO emissions
ana relatively high NO emissions.
High-speed motion pictures of the reacting flow field in the vicinity of the injection
-------
plane reveal significant fluctuations in the flame structure, and several different
types of flow instabilities have "been identified. Although there are substantial
time-dependent, three-dimensional effects, for most of the experimental conditions
investigated the time-mean flow is nearly axisymmetric.
present experimental results indicate that much of the interesting fluid dynamic and
chemical phenomena occur in a region very near the injection plane, containing the
stabilizing recirculation zone. More detailed information on the velocity, tempera-
ture and species concentration distributions are required in the near-injector region
if the significant features of the flow field and the interaction with the pollutant
formation process are to "be understood. Of particular importance are velocity and
flow direction measurements in the recirculation zone. An effort is required to
characterize the unsteadiness in the flow field and the subsequent effect on
pollutant formation. In the present experiments, penetration of the recirculation
zone in the gaseous fuel injector was observed for a range of experimental condi-
tions. This penetration made identification of the inlet conditions difficult,
and, hence,introduced an additional uncertainty in the comparison of the experimental
data with results from the combustor flow analysis. Some effort should be devoted
to the characterization of the injector flow upstream from the penetration region
of the recirculation zone.
85
-------
APPENDIX A
COMPUTATION OF THE SWIRL NUMBER
The derivation to follow is taken from Ref.21 where a swirl number, S, which is
used to characterize swirling flows, is defined in terras of the nozzle torque, T,
the jet thrust, £, and an effective nozzle diameter, de = d (l-Z2)^, i.e.,
S=t-/f-de (A-l)
For the swirl generator in Fig. 3, vanes of constant chord and angle are mounted on
a hub of diameter, d-^, and extend to the wall of the nozzle tube, of diameter d.
In passing over the vanes the air at velocity, V, is accelerated to V/cos r\, where
n is the blade angle. The axial and tangential components of the air velocity is-
suing from the blades are V and V tan n, respectively. If the air-fuel momentum
flux ratio is large with respect to unity and/or the hub diameter is small with
respect to the outer diameter, the components of the momentum associated with the.
fuel stream can be ignored. This is the case for the present experiments.
Considering an elemental annulus the torque is
dr = 27r/JV(Vtan7?) r2dr , ,
or
I -I /I -7O \ I
(A-3)
Then
S = 1" "(i-z2)3/2 tan7? (A'U)
verifying that the swirl number is dependent exclusively on nozzle and blade geo-
metry.
86
-------
APPENDIX B
COMBUSTOR PROBE DESIGNS AND CALIBRATION PROCEDURES AND REDUCTION OF PROBE DATA
The probing devices employed in the current investigation included: (l) a five-
probe rake for exhaust flov sampling, (2) a calibrated conduction heat-loss thermo-
couple probe and (3) a multi-purpose rake consisting of a double-sonic orifice probe
and a flow field measuring probe. All of the probes vere designed specifically for
operation within the severe environments that are characteristic of the EPA combus-
tion rig. Since these designs contain several unique features and since the ap-
plication of the various probes to combustion measurements is somewhat complicated,
each probe type is described in detail and calibration and data reduction proce-
dures are discussed.
Exhaust Gas Sampling Rake
The five-probe, exhaust gas sampling rake is placed just downstream of the second
spool piece, approximately 1.8 m from the plane at which fuel and air are brought
into initial contact, to aspirate'the combustion gases for subsequent analysis.
Placement of the five probes was determined by dividing one-half of the combustor
circular cross-section into five equal concentric area segments and noting the
radial distances to the mid-point of each area segment. A probe was then located
at each mid-point along a diameter, or its corresponding position 180 deg away as
dictated by packaging constraints. With this arrangement, each probe aspirates
sample with a composition which is approximately an average for the segment in
which it is located. An identical choke is placed in each probe to insure that
equal mass flows are aspirated for every segment. Gas from all five probes enters
a common chamber where mixing occurs to produce a fluid with the same area-averaged
composition as that at the combustor exhaust. It is this fluid which is analyzed
using a time-of-flight mass spectrometer or Scott Analyzer.
Quenching of the gas sample and cooling of the stainless steel probes is accom-
plished using high pressure, hot water which flows in jackets surrounding the
sample flow passage (Fig.59)« Operation of the rake at gas temperatures to 2500 °K and
pressures from 1 to 7 atm is permissible for 1.8 kg/sec of cooling water at k atm
and an inlet temperature of about 1*00°K. The internal design of the probes was
achiev6^ through a one-dimensional analysis of the governing geometric and flow
rameters for ^9 most severe environmental conditions expected. Area variations,
waH friction and heat transfer were accounted for in the analysis through the use
of influence coefficients (Ref. 22). It was desired in the design of the probes to:
(1) inhibit chemical reaction of the material captured, by the combined effects of
aerodynamic expansion and convective cooling, (2) minimize wall catalytic activity
t,y keeping wall temperature below about 520°K, (3) prevent condensation of water by
maintaining a minimum wall temperature of ii000K throughout, and (U) delay shocking
down of the flow downstream of the choke until the flow total temperature was less
than the auto-ignition limit, i.e., about 1000°K. The resulting probe design,
Fig.59» has a length, of 15.88 cm and an inlet hole diameter of O.l6 cm. A minimum
area section located U.6l cm from the probe entry chokes the flow. The question of
.chemical reaction of material within the probe was assessed by following the growth
of an ignition parameter in the probe. This parameter, defined as the ratio of gas
87
-------
Exhaust Gas Sampling Probe.
ALL DIMENSIONS IN CM
NOT TO SCALE
00
OO
-15.88-
\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\
A
Ol
p
s
-------
residence time to chemical ignition delay time (Ref.20), signifies ignition when its
value is equal to or greater than unity. For gas samples initially at conditions
which are representative of the most severe thermal environments of interest, e.g.,
T=2500 K, P=7 atm, no ignition occurs in the probe according to the one-dimensional
analysis. At the aft end of the probe the flow which was accelerated to supersonic
speeds in the h deg divergent section, must shock down and rise to the stagnation
temperature. Chemical reaction will not occur, however, since the stagnation temp-
erature is reduced significantly below 1000°K in this probe design. The total
volumetric flow for the five probe rake is in excess of 0.65 std m3/hr at 1 atm,
which is the minimum flow required by the Scott Analyzer.
Calibrated Heat-Loss Conduction Probe
This water-cooled thermocouple probe is designed to measure temperatures in excess
of the platinum-rhodium and iridium-rhodium melting points which usually limit
the use of thermocouples. The basic design philosophy is "to deliberately apply a
known error (heat loss) to the sensing thermocouple and to measure its subsequent
temperature. The amount of heat loss used is sufficient to cool the thermocouple
below the melting point.
The probe consists of three thermocouples including an iridium-60$ rhodium/iridium
thermocouple which protrudes from a copper base into the stream to record a tempera-
ture, TA, and two platinum-10$ rhodium/platinum thermocouples installed on the base
to record base temperature, TB. In order to obtain the true local stream tempera-
ture, Trj^the measured stream thermocouple temperature must be increased by the con-
duction and radiation heat-loss corrections, T_ and T_, respectively:
L R
TT = TA + Tc + TR (B-I)
= 0.556 (tA -Mc + tR)
where the small t denotes the measured temperature. Manipulation of corrections
(supplied by the manufacturer) yields
TT =0.556 [(I + A) tA- AtB + BtA375] (B_2)
where for combustor weight flow,w,
Assech (0.59w)
-I? / 05 (B-3a,b)
B=33.6 X 10 I2/W0'5
In typical situations, conduction and radiation heat loss corrections account for
about a 25$ reduction in the stream temperature measurement.
-------
Multi-Purpose Rake
I. Impact-Static Probe
The impact-static probe, Fig. 6, consists of a cylinder with a hemispherical tip
having a impact tap at its center and four symmetrically-located static pressure
taps on a circle HO deg back from the center. A static tap also is located 2.12 cm
from the probe stagnation point, on the 0.95 cm diameter cylindrical portion. The
pressure measured at this location was subtracted from the impact pressure for the
calculation of dynamic pressure and, ultimately, axial flow velocity. Water at
about ljOO°K and h atm was used to cool the 5 cm long probe.
Local flow direction in the combustor was determined from the four differential
pressures associated with the four static pressure taps on the hemispherical tip
and the impact pressure, using suitable calibration information. Calibration data
were acquired in the UARL U-in.x h-in. continuous (cold) flow wind tunnel for free
stream dynamic pressures of 2-5 in. H20 corresponding to Reynolds numbers based on
the tunnel width of 0.5-1.3x10^. These cold flow dynamic pressures approximate the
levels found within the combustor. The Reynolds numbers are 3 to 6 times greater;
however, no important variations in the calibration data were discovered over the
range of Reynolds numbers tested.
Prior to installation of the probe, the uniformity of the tunnel flow field was
verified by means of a traversing impact rake available with the facility. Subse-
quently, the probe was inserted into a rotatable disk, Fig. 60, in the tunnel side-
wall. A range of probe yaw angles (-20 degsfrs+UO deg) relative to the known flow
direction was set by rotating the disk. In addition, the probe mounting disk was
fitted with angled brackets such that the probe could be pitched (-kO deg-iai+Uo deg)
relative to the main flow direction. In this fashion flow angles comprised of both
pitch and yaw components could be studied.
Differential pressures were measured between the probe impact tap and each of the
static taps for several combinations of pitch and yaw angles. Read-out capability
was provided with Magnehelic- pressure gages (accuracy-better than 2% of full scale
for 2 in. 1^0 and 10 in. H^O gages). It was found convenient for data reduction
purposes to display this cold flow calibration information as shown in Figs. 6l and
62. The ordinate in each case is the difference between the differential pressures
measured for static taps 180 deg apart, normalized by the dynamic pressure, q, e.g.,
L(Pip-P1)-(pfp,-P2)]/q. If the impact-static probe is visualized as moving toward
the combustor centerline in a horizontal plane, then the pitch plane is the hori-
zontal plane and the yaw plane is vertical. Static taps nos. 3 and 1 are in the
pitch plane, with tap no. 3 closest to the centerline for most of the traverse
during the course of a typical test. Pitch angles, a, are positive (negative) for
flow moving towards the probe from the wall (centerline). Static taps nos. k and
2 are in the yaw plane, with tap no. U over tap no. 2. Yaw angles, 6, are positive
(negative) for flow moving towards the probe from above (below).
Examination of Fig. 62 reveals that the yaw plane pressure factor is only a weak
function of pitch angle for -20^a<+20. Within this range, therefore, a good esti-
90
-------
Calibration Set-up for Impact-Static Probe.
FIG. 60
NOT TO SCALE
SIDE VIEW
FLOW DIRECTION
\
YAW ANGLE, 0
(NEGATIVE YAW
SHOWN)
TOP VIEW
FLOW DIRECTION
PITCH ANGLE, a
(NEGATIVE PITCH SHOWN)
TUNNEL SIDE WALL
_FLOW FIELD
SENSING .
PROBE
////////////////////////A
TYPICAL PITCH
BRACKET
P1-27-3
91
-------
FIG. 61
Impact-Static Probe Pitch Plane-Angle Calibration Data.
n
o.
tr
O
oc
D
CO
CO
UJ
oc
0.
UJ
<
a.
I
o
-0.4 -
-0.8 -
-1.2
-40
-20
0 20
PITCH ANGLE, a - DEG
92
60
PI-27-2
-------
FIG. 62
Impact-Static Probe Yaw Plane-Angle Calibration Data.
-0.4
40 60
YAW ANGLE, 0- DEC
80
93
P1-27-1
-------
mate of yaw angle can be determined for entry into Fig.6l to establish a pitch angle.
A minimal amount of iteration between Figs.61 and 62 is then required to refine the
values of a and g. Finally, a total flow angle, 9, can be computed from,
9 = cos'1 [(cosa) (cos/3)] (s-U)
This parameter is found to correlate with the geometric average pressure,
AP.-P| V
-[(
L\
q
as shown in Fig. 63.
II. Double-Sonic-Orifice (DSO) Probe
The DSO probe, Fig. 6, contains a nominal 0.127 cm diameter choked orifice inte-
grated with the tip and a second orifice, 0. 168 cm in diameter, at some distance
downstream of the first orifice. The orifices are sized sc that choking occurs at
both orifices when gas is aspirated through the probe.
In principle, by measuring the stagnation pressure and temperature upstream of the
second orifice during aspiration, Prp2 and TT2' respectively, and the stagnation
pressure of the first orifice, i.e., the impact pressure obtained by "dead-ending"
the probe, PT,, the stagnation temperature of the stream can be determined. That
is, for a constant mass flow, and assuming that no significant molecular weight or
ratio of specific heat changes occur in the flow between the two orifices, the de-
sired local stream temperature, TT , is given by
, ., (CD**) PT, I2 (B-6)
TI
In order to use Eq.. (B-6) for stream temperature determination, suitable calibra-
tion information for orifice flow coefficients must be available. These coef-
ficients, which reflect boundary layer effects, depend on throat Reynolds number- and
the ratio of wall temperature to flow stagnation temperature. As Reynolds numbers
are increased and/or wall temperature ratios are decreased, "boundary layer displace-
ment thickness decreases and the flow coefficient approaches unity.
DSO calibration data were obtained by placing the probe in uniform streams of high
temperature air and natural gas-air combustion products. Local stream temperatures
were measured using calibrated thermocouples and the calibrated heat-loss conduc-
tion probe. A summary of the calibration data is provided in Fig. 6k. Also shown
are analytical results for various first orifice wall temperature ratios, (Ty/Tip)^
computed using an integral boundary layer code (Ref.22). It should be noted that
the flow coefficient correction for Reynolds numbers of interest is approximately
10$. At high Reynolds numbers, where Cv^C^ 21 1> "the actual throat area ratio,
A* /A*, is found to be about 0.662 versus a nominal value of (0. 127/0. l68)2 = 0.572.,
-------
Correlation of Total Flow Angle.
FIG. 63
SOLID SYMBOLS DENOTE NEGATIVE YAW ANGLE
"BEST FIT"
LINE THROUGH
DATA
20 30 40
TOTAL FLOW ANGLE, 6 - DEG
R-35-1
-------
Double Sonic Orifice (BSD) Calibration.
0.75
MD
ON
M
r
M
CN
<
CM
Q
O
Q
O
oc
UJ
H
UJ
Q_
I-
z
UJ
O
u.
LL.
Ol
O
O
1
u.
0.73
0.71
0.69
0.67
0.65
103
I
I
I
5 104 2
SECOND ORIFICE THROAT REYNOLDS NUMBER
105
P
8
-------
APPENDIX C
GAS SAMPLING SYSTEM, CALIBRATION PROCEDURES AND DATA REDUCTION TECHNIQUES
The gas samples withdrawn through the five probe exhaust rake or the combustor DSO
probe are analyzed on-line for the time-averaged concentration of a number of
species. The sampling system, the various pieces of analytical instrumentation, the
calibration procedures and the data reduction techniques used to obtain species
concentration data are discussed herein.
Sampling System
A schematic diagram of the sampling system used in the present investigation is
shown in Fig. 7. The gas samples withdrawn by the hot-water-cooled sampling probe
pass into an electrically-heated HiOOoK), teflon-coated, stainless-steel manifold.
The sample pressure in this manifold is maintained at approximately 300 torr by
means of a mechanical vacuum pump (not shown in Fig. 7). A portion of the gas
sample in the manifold is directed through an electrically-heated, teflon-coated,
stainless-steel line to the inlet of a time-of-flight mass spectrometer. The
sample pressure at the inlet of the mass spectrometer is maintained at approximately
10 torr by means of a remotely-actuated micrometer valve. The TOF mass spectro-
meter was located close to the probe assembly. In all of the experiments, the
distance from the sample probe inlet to the mass spectrometer inlet was less than
60 cm. The mass spectrometer was used to measure the concentration (mole fraction)
of C02, 02 and NO, prior to removal of water vapor from the sample. The remainder
of the gas sample passed through a short electrically-heated, teflon-coated,
aluminum line to a condensate trap (~277°K), where most of the water vapor in the
sample was removed. The gas sample then passed through an unheated, teflon-coated,
aluminum line and a stainless-steel bellows pump to a Scott Model 119 Exhaust
Analyzer. The stainless steel bellows pump increased the sample pressure from sub-
atmospheric levels to 1 atm as required by the Exhaust Analyzer. The Analyzer was
used to measure the concentration (mole fraction) of CO, C02, 02,NO, N02 and un-
burned hydrocarbons (THC) in the "dry" gas sample. The Exhaust Analyzer was lo-
cated in the combustion facility control room, approximately 10 m from the combus-
tor. The redundant sampling instrumentation provided a measure of the influence
of the long line (and condensate trap) connecting the sampling probe to the Exhaust
Analyzer on the sample received by the Analyzer. In a limited number of the input-
output tests, the formaldehyde concentration in the exhaust gas was measured using
the MBTH wet chemical method. Gas samples for these measurements were withdrawn
from the heated manifold via an unheated pyrex line and passed through the reagent
solution.
Time-of-Flight Mass Spectrometer
The time-of-flight (TOF) mass spectrometer used in the present investigation was
developed at United Aircraft Research Laboratories under the direction of Dr. M
F. Zabielski. The instrument is portable and rugged, and was specifically designed
for the harsh environments encountered in combustion facility test cells. The mass
97
-------
spectrometer consists of two separate packages: (l) a portable vacuum station
housing the spectrometer, to which the electronics associated with the ion source
and ion detector are attached, and (2) a small control unit. This configuration
not only permits the spectrometer to be coupled to the sampling device by an ex-
tremely short gas inlet, but also allows control by an operator in a remote loca-
tion.
The vacuum station consists of a 30U stainless steel tee-type chamber connected via
a liquid-nitrogen cooled baffle to a 6-inch oil diffusion pump, backed by a mechan-
ical pump. The vacuum chamber can be pumped to a pressure in the low 10~8 torr
range, and has a typical operating pressure of 1 x 1Q~° torr. The inlet system,
consists of a stainless steel tube with the flow in the inlet line controlled by
a remotely-actuated micrometer valve.
In the present investigation, the TOF mass spectrometer was used to scan the m/e
range of 12-lA. The output signal from the TOF mass spectrometer was sent to on-
line data storage and processing equipment. A description of the data storage
and processing equipment, the data processing techniques and the procedures used in
calibration of the mass spectrometer are discussed later.
Scott Exhaust Analyzer
The Scott Model 119 Exhaust Analyzer used in the present investigation provides for
the simultaneous analysis of CO, COg, NO or NOg, 02 and total hydrocarbons (THC).
The Exhaust Analyzer, Fig.65s is an integrated system, with flow controls for sample
zero and calibration gases conveniently located on the control panel. The incoming
gas sample passes through a refrigeration condenser ( 275°K), to remove residual
water vapor. As the sample passes from the condenser, it is filtered to remove
particulate matter. The Exhaust Analyzer is comprised of five different pieces
of analytical instrumentation. Beckman Model 315B Non-Dispersive Infrared (NDIR)
Analyzers were used to measure the CO and C02 concentations (mole fractions) in the
gas sample. Concentration ranges available on the CO analyzer were from 0-200 ppm
to 0-15 % on several scales. Concentration ranges available on the CO
analyzer were Q-h% and 0-16$. The accuracy of the NDIR analyzers is nominally ±ijj
of full scale. A Scott Model 125 Chemiluminescence Analyzer was used to measure
the NO and N02 concentrations in the gas sample. Concentration ranges available -
with this instrument were from 0-1 ppm to 0-10,000 ppm on several scales, with a
nominal ±1% of full scale accuracy. The thermal converter used in the chemilumine-
scent analyzer was stainless steel, and was operated at a temperature of approxi-
mately 1000°K. A Sectt Model 150 Paramagnetic Analyzer was used to measure the Oo
concentration in the gas sample. Concentration ranges available with this instru-
ment were from 0-1$ to 0-25$ on several scales, with a nominal accuracy of ±lj£ of
full scale. A Scott Model Il6 Total Hydrocarbon Analyzer was used to measure the
hydrocarbon concentration in the gas sample. This analyzer utilizes an unheated
flame ionization detection system to provide for measurement of hydrocarbons (as,
carbon) in concentration ranges from 0-1 ppm to 0-10$, with a nominal accuracy of
±1$ of full scale. Output signals from the various analyzers are displayed on
chart recorders and also are sent to the on-line data storage and processing equip-
ment.
98
-------
Exhaust Gas Analytical System.
CO2 NDIR
ANALYZER
CO NDIR
ANALYZER
TOP
MASS
SPECTRO
METER
CONTROL
AD
CONVERTER
AND SIGNAL
PDF 6
DATA
LINK
CHEMI LUMINESCENT
ANALYZER
PARAMAGNETIC
ANALYZER
FID
HYDROCARBON
ANALYZER
'
-------
Data Storage and Processing Equipment
Output signals from the TOP mass spectrometer and from the various components of
the Exhaust Analyzer were recorded on chart recorders and also sent through a multi
plexer to a Northern NS-575 analog-digital converter and signal averager, Fig. 7.
Averaging of the output signals of the various analytical instruments vas required
"because of low-amplitude, low-frequency fluctuations in the sample line pressure.
These fluctuations were the result of temperature and density fluctuations in the
reacting flow field. Typically, at each test condition, sampling data were ac-
quired for a period of 2-3 min. Upon completion of the test, the stored data were
dumped directly into a PDP-6 computer for processing. Data from the Scott Exhaust
Analyzer were processed immediately by the computer, and the processed data vere
transmitted to the test operators via teletype for verification prior to initia-
tion of the next test. Output from the TOF mass spectrometer was stored in the
computer, for later analysis via several data reduction computer codes. A typical
mass spectrum of an exhaust gas sample as stored in the computer is shown in Fig. gg
MBTH Formaldehyde Analysis
In a limited number of the input-output experiments, the formaldehyde concentration
in the exhaust gas was measured using the MBTH (3- methyl-2-benzothiazolone hydra-
zine hydrochloride) wet chemical method. This method is discussed in detail in
Ref.2.1* and only a brief summary of the technique will be given here. A schematic
diagram of the apparatus is shown in Fig. 67. A gas sample is continuously with-
drawn from the heated manifold and bubbled through a 0.05$ aqueous solution of
MBTH. The volume of gas sample passed through the reagent solution depended on the
formaldehyde concentration in the sample. Typically, gas samples of 1-2 liters
(STP) were used, requiring a collection time of 5-10 min. Water soluble
aliphatic aldehydes in the gas sample react with MBTH to form an azine. After
collecting a gas sample, the reagent solution was taken to the UARL Analytical
Chemistry Laboratory where the resulting azine is oxidized by a ferric chloride-
sulfuric acid solution to form a blue cationic dye in acid media. The concentra-
tion of this dye, which is related to the aldehyde concentration, is determined by
measuring the absorptivity of the solution at 628 nra. There are no significant
interferences of other gaseous, water-soluble compounds with the MBTH method.
Calibration Procedures
The TOF mass spectrometer and the components of the Exhaust Analyzer were cali-
brated using various mixtures of methane, carbon dioxide, carbon monoxide, oxygen,,!
nitric oxide, nitrogen dioxide in nitrogen or air. The calibration gases were sup-
plied by Scott Research Laboratories, Plumsteadville, Pa. The specified tolerances
of the calibration gas composition were ±2.% of the cylinder analysis as supplied by
the manufacturer. Spot checks of the gas composition were made at UARL, and the,
compositions were found to be within the specified tolerances.
The TOF mass spectrometer was calibrated by measuring the mass spectra of various
calibration gas mixtures. From the known mixture composition and the observed mass
spectra, mass discrimination effects in the inlet and the fragmentation patterns
100
-------
FIG. 66
Typical Mass Spectrum of Exhaust Gas Sample.
A
A
m/e 12 14 16 17 18
28 29 32
40
44
Ml 2-139-1
101
-------
MBTH Formaldehyde Analysis Apparatus.
FIG. 67
PRESSURE GAGE
FLOW
METER
VACUUM
PUMP
MBTH SOLUTION
FROM
SAMPLE
MANIFOLD
102
N12-161-a
-------
for various species of interest were determined. Mass discrimination effects and
fragmentation patterns are required if a unique determination of gas sample composi-
tion is to be obtained from the measured mass spectra. The Scott Exhaust Analyzer
was calibrated prior to each test by flowing a zero gas (hydrocarbon-free N«) and
a calibration gas through each of the instruments in the Analyzer. As part of the
calibration procedure of the Exhaust Analyzer, the converter efficiency was deter-
mined as a function of converter temperature by flowing known mixtures of N02/air
mixtures through the converter. A converter efficiency, defined in terms of per-
cent N02 dissociation, greater than 9%% was obtained for the converter operating at
a temperature of 1000°K.
Sampling System Effects
To evaluate the effect of the sample line, condensate traps and particulate filters
on the composition of the gas sample, calibration gas mixtures vere flowed through
the entire sampling system and the composition of the gas was measured using both
the TOF mass spectrometer and the Exhaust Analyzer. These experiments served two
purposes: (l) to determine if the sampling system served as a source or sink for
any of the gaseous species of interest and, (2) to establish the time required to
attain a steady-state measurement of the composition. When various N2-diluted
mixtures of CH^, CO, C02, 02 and NO were flowed through the sampling system, both
the TOF mass spectrometer and the Exhaust Analyzer gave essentially the gas analysis
indicated on the gas cylinders. A steady-state measurement of concentration was
attained at the TOF mass spectrometer in less than 30 sec and at the Scott Exhaust
Analyzer in less than 60 sec. When various N02-air mixtures were flowed through
the sampling system, approximately 120 sec were required to establish a steady-
state measurement at the Exhaust Analyzer approximating the cylinder analysis.
Sawyer, et. al. (Ref.l6) and Mellor, et al.(Ref.15) have reported N02 losses in'sample
line condensate traps and on particulate filters with soot deposits. In the pre-
sent study, noticeable losses in N02 were observed in flowing N02-air mixtures
through dirty particulate filters, in experimental studies of nitrogen oxide
formation in prenuxed laminar flames (Ref. 25 and 26), significant levels of NOo
vere found in the flame zone. One investigator (Ref. 26) has attributed these
observations to the conversion of NO to N02 in the sampling probe, in a reclnt
analytical study of nitrogen oxide reaction, in probes (Ref^T). «*! T^f
,
that there could be significant interconversion s^i i
si i
steel and quartz probes, particularly in regions of the flow'with relative^
large 0-atom concentrations (i.e. in hot, 0 -rich regions). Considering the
potential sources (oxidation of NO in a sanjle containing 02) and sinks oj L
in the probe and sampling system and the difficulty in using a chemi luminescent
analyzer to measure N02 in gas samples containing appreciable concentrations of
hydrocarbons ^reported levels of N02 in combustion gas samples must be viewed
as having a hxgh degree of uncertainty. This uncertainty, together with the
possibility of the conversion of some of the NO in the gas sample to NO
suggests that reported levels of NO also are uncertain. However, it is believed
that the reported nitrogen oxide levels, vhich are uncertain, do provide a valid
qualitative measure of the variations of nitrogen oxide concentrations with
changes in combustor operating conditions.
103
-------
TABLE D-l. EXHAUST CONCENTRATION DATA - NATURAL GAS FUEL, S = 0
4>
0.719
0.797
0.823
0.892
0.915
0.967*
1.058
1.164
0.752
0.875
0.969
1.068
0.778
0.871
0.980
1.068
0.803
0.886
0.987
1.047
P
(atm)
1.01
1.00
0.99
1.00
1.00
1.00
1.01
1.00
3.58
3.65
3.79
3.76
1.00
1.00
0.98
1.00
1.00
1.00
1.00
0.99
Ta
(°K)
750
744
738
738
738
742
740
735
749
745
739
750
860
859
860
863
74o
733
742
739
Vvf
27.4
24.9
23-9
22.1
21.5
20.3
18.4
16.6
26.6
22.9
20.6
19.0
29.4
26.6
23.4
21.6
24.2
21.8
19.8
18.6
wa
(kg/sec)
0.132
0.133
0.131
0.133
0.132
0.132
0.132
0.131
0.134
0.134
0.134
0.134
0.132
0.132
0.133
0.133
0.087
0.088
0.088
0.088
*f
(kg/sec)
0.00567
0.00637
0.00645
0.00708
0.00725
0.00763
0.00835
0.00916
0.00604
0.00703
0.00777
0.00855
0.00616
0.00689
0.00778
0.00851
0.00420
0.00468
0.00523
0.00554
NO
(ppm)
163
151
163
156
155
166
154
157
500
470
423
402
302
284
281
274
196
190
221
200
N02
(ppm)
20
37
27
30
35
30
-
-
31
10
-
-
45
50
22
-
37
20
_
THC
(ppm)
333
537
782
1310
1670
1904
3130
4770
60
477
1010
i4oo
128
424
1380
3410
357
711
1620
2400
CO 02
(%) Mass Spec.
(*)
0.6l
1.14
1.35
1.69
1.91
2.19
2.96
3-91
0.77
2.07
.3.76
4.88
0.78
1.37
2.39
3.71
0.912
1.352
2.374
3.292
5.08
4.15
3-95
2.99
2.02
2.00
1.31
0.55
4.03
2.76
1.68
0.89
3.94
2.54
1.69
1.14
3.80
2.64
1.28
0.80
C02
Ex. Anal. Mass Spec.
(*)
8.05
8.59
8.50
8.81
8.58
8.4l
8.69
8.74
8.72
8.75
8.36
8.43
8.56
8.46
8.39
8.20
9-35
8.94
8.79
8.93
8.24
8.85
8.65
9.06
8.70
8.59
8.52
8.37
9.12
9.24
8.85
8.86
9.01
9-03
8.87
8.67
9.40
9.40
9.12
9.07
Note: All species concentrations (mole fractions) are reported on a dry "basis .
* HoCO =0.5 ppm
-------
TABLE D-2. EXHAUST CONCENTRATION DATA - NATURAL GAS FUEL, S = 0.3
4>
0.785
0.876
0.964*
1.079
P
(atm)
0.99
0.99
0.99
0.99
Ta Va/Vf
(°K)
751 26.1
743 23. 4
744 21.2
742 19.4
>
(kg/sec )
0.134
0.134
0.135
0.134
wf
(kg/sec )
0.00631
0.00704
0.00778
0.00867
NO
(ppm)
260
252
226
195
N02 THC CO 02 C02
(ppm) (ppm) (%} Ex. Anal. Mass Spec. Ex. Anal. Mass Spec.
(*) W
19 3
16 7
18
- 74
0.21 3.55
0.60 2.56
1.50 1.23
3.52 0.6l
3.88 9.12 9.46
2.73 9.77 9.91
1.13 9.89 10.1
0.35 8.11 8.33
o
VI
Note: All species concentrations (mole fractions) are reported on a dry basis.
= 0.6 ppm
-------
TABLE D-3. EXHAUST CONCENTRATION DATA - NATURAL GAS FUEL, S = 0.6
4>
0.795
0.876
0.983
1.078
0.771*
0.878
0.980
1.080
0.778
0.821+
0.862
0.927
0.955*
1.02
1.056
1.11
0.511
0.616
0.711*
0.817
0.917
1.02
1.11
1.22
1.33
p
(atm)
0.99
0.99
0.99
1.00
3.**2
3.60
3.59
3.55
1.02
1.02
1.01
1.02
1.01
1.02
1.01
1.02
1.01
1.01
1.01
1.01
1.01
1.01
1.01
1.01
1.01
Ta
(°K)
739
737
731*
736
7**7
7l*2
736
761
71*1*
71*6
75l*
7l*l
756
753
71*8
7**3
71*0
7**0
739
7UO
71*2
772
770
769
770
Va/Vf
25.6
23.1
20.7
19.1*
26.0
23.2
20.7
19-5
10.8
10.3
9-9
9-1
9.0
8.3
8.0
7.6
0.30
0.25
0.21
0.19
0.17
0.16
O.lU
0.13
0.12
w
(kg /sec)
0.132
0.133
0.131*
0.131*
0.133
0.135
0.131*
0.131*
0.131
0.131
0.132
0.131
0.132
0.132
0.132
0.132
0.137
0.137
0.137
0.137
0.137
0.137
0.137
0.136
0.136
wf
(kg/sec)
0.00627
0.00698
0.007.88
0.00866
0.00618
0.00710
0.0078U
0.00861*
0.00609
0.006*15
0.00682
0.00727
0.00755
0.00805
0.00832
0.00879
0.001*25
0.00507
0.00589
0.00671
0.00751*
0.00836
0.00918
0.00993
0.0109
NO
(ppm)
299
280
269
2l*6
518
1*38
1*13
351
276
288
303
273
320
273
266
233
117
108
117
131
11*5
191*
203
169
131
N02 THC
(ppm) (ppm)
-
-
-
-
-
-
-
21
20
-
22
-
_
_
_
18
22
33
32
32
-
-
-
U
1*
10
1*1*
8
5
11
65
3
3
k
3
5
1*
19
17
20
95
2l*0
1*35
525
780
1250
5100
7850
CO
(*)
0.1U2
0.611*
1.79
3.98
0.100
0.371
1.59
i*.oo
0.092
0.162
0.257
0.1*67
1.1*1
1.70
U.19
3.7l*
0.67
l.Ol*
1.2
1.07
1.11
1.75
2.61*
5.83
7.67
02
Ex. Mass
Anal. Spec.
{%}
3.1*3
1.66
0.58
0.06
1.76
0.15
0.10
3.73
3.52
2.26
1.6l
0.61
0.68
0.09
0.1*U
9.60
9.1*0
8.02
6.02
i*.30
1.86
0.63
0.28
0.05
3.79
1.85
0.73
0.12
3.76
1.82
0.25
^
_
3.1*7
2.21
t-
_
Im
w
^
mm
_
M
-
-
"
C02
Ex. Mass
Anal. Spec.
("to)
9.18
9.70
9-73
8.91
9.55
9-78
9.52
8.86
9.96
9-79
9.82
9-95
9-!*9
8.76
8.1*9
5-92
6.30
6.53
7.61*
8.52
9-5**
9.70
7.98
6.91*
9.1*6
10.3
10.0
9.UU
9.63
10. 1*
10.3
8.92
9.38
10.1
_
_
_
_
_
_
..
-
"
All species concentrations (mole fractions) are reported on a dry basis.
= 0.5 ppm
-------
TABLE D-k. EXHAUST CONCENTRATION DATA~ SYNTHESIZED FUEL, S = 0
p
* (atm)
0.779 1-02
0.87^ 1.02
0.972 1.02
1.063 1-02
rib
728
7^0
732
735
Va/Vf (kg /sec)
16.9 0.137
15. U 0.137
13.8 0.137
12.7 0.137
(kg/sec)
0.00867
0.00976
0.0109
0.0119
NO
(ppm)
165
165
165
115
N02 THC CO
(ppm) (ppm) (%}
25
15
28
20
170 0.90
U25 1.1+0
1125 2.15
2500 3.00
°2
1+. 62
3.62
2.75
2.12
C02
(to
9.50
10.0
10.0
9.40
Note: All species concentrations (mole fractions) are reported on a dry basis.
-------
TABLE D-5. EXHAUST CONCENTRATION DATA - SYNTHESIZED FUEL, S = 0.6
..
0.791
0.889
0.991
0.793
0.892
1.000
p
(atm)
1.00
1.00
1.00
3.59
3.75
3.82
(°K) Va/Vf
71*5 18.0
71*! 16.0
71*! ll*.5
752 17.7
7^9 15.8
71*6 ll*.l
(kg/sec)
0.137
0.137
0.136
0.137
0.137
0.137
(kg/sec)
0.00881
0.00989
0.0110
0.00879
0.00990
0.0111
NO
(ppm)
2^41
2^0
221
561
567
N02
(ppm)
28
26
1*1*
18
27
THC
(ppm)
3
3
3
I*
1*
1*
CO
0.162
0.666
1.72
0.089
0.557
1.85
Ex. Anal
U.16
2.61*
1.29
3.62
1.76
Mass Spec.
(*)
l*.08
2.27
1.08
3-51
1.77
0.25
COg
Ex. Anal.
(4}
9.76
10.1*
10.5
10.2
11.1
11.0
Mass Spec.
9.6
10.8
10.6
10.1
11.3
11.1*
NOTE: All species concentrations (mole fractions) are reported on a dry basis.
-------
TABLE D-6. EXHAUST CONCENTRATION DATA. - GASEOUS PROPANE
*
0.804
0.907
0.973
1.115
o.8i4
0.927
1.033
1.128
p
(atm)
1.01
1.00
1.00
1.00
1.01
1.01
1.01
1.01
Ta
730
724
738 t
738
733
744
744
747
Va/Vf
4l.3
36.6
34.8
30.6
39-1
37.5
34.3
30.3
(kg/sec )
0.137
0.137
0.136
0.137
0.136
0.136
0.136
0.136
s
0
0
0
0
0.3
0.3
0.3
0.3
NO
(ppm)
260
250
270
255
350
350
355
325
N02
(ppm)
15
18
17
25
30
25
18
-
THC
(ppm)
4oo
1425
1900
4500
50
350
600
1550
CO
(mole %)
0.94
1.63
2.23
4.10
0.59
1.24
2.51
3.75
°2
(mole %)
-
-
-
-
4.45
3.05
1.65
1.05
C02
(mole %)
9-30
10.1
10.1
10.1
10.3
10.2
10.5
10.1
Note: All species concentrations (mole fractions) are reported on a dry basis.
-------
TABLE D-7. EXHAUST CONCENTRATION DATA - LIQUID PROPANE
*
0.824
0.935
1.037
1.153
0.731
0.835
0.947
1.046
. p
(atm)
1.00
1.00
1.00
1.00
6.27
6.95
6.99
6.44
°TK
728
729
726
730
747
756
754
751
(kg/sec)
0.135
0.135
0.135
0.135
0.136
0.136
0.136
0.136
s
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
NO
(ppm)
250
255
260
260
170
150
165
185
N02
(ppm)
30
35
20
-
15
5
10
-
THC
(ppm)
40
165
600
1850
50
75
175
700
CO
(mole $)
0.79
1.77
2.63
4.10
0.18
0.18
0.4l
1.20
02 C02
(mole %} (mole %)
9-80
10.6
10.6
10.5
7.00 10.8
4.10 10.6
1.70 11.8
0.25 11.8
Note: All species concentrations (mole fractions) are reported on a dry basis.
-------
TABLE D-8. EXHAUST TEMPERATURE PROFILES FOR NATURAL GAS
COMBUSTION AT VARIOUS $*
^
1.0 (wall)
0.95
0.89
0.79
0.68
0.58
0.47
0.35
0.17
0.06
0. ((j^
0.10
0.20
0.33
0.45
0.56
0.68
0.77
0.88
TEMPERATURE, °K
§- 0.80
1066
1175
1332
1509
1631
1716
1765
1827
1855
1883
1885
1890
1897
1881
1881
1848
1749
1635
1478
§>= 1.0
1088
1170
1320
1525
1669
1755
1828
1894
1919
1945
1958
1953
1972
1982
1935
1920
1841
1738
1620
$ = 1.2
1105
1177
1340
1505
1670
1754
1846
1928
1994
1996
1996
2029
2028
2003
1966
1935
1829
1702
1583
*P=1 atm, S=0, Ta=740°K, w_=0.132 kg/sec
111
-------
TABLE D-9. EXHAUST TEMPERATURE PROFILES FOR NATURAL GAS
COMBUSTION AT VARIOUS INLET AIR CONDITIONS*
r/'o
1.0 (wall)
0.95
0.89
0.79
0.68
0.58
0.47
0.35
0.17
0.06
0. (£)
0.10
0.20
0.33
0.45
0.56
0.68
0.77
0.88
TEMPERATURE, °K
Ta = 742°K
wa= 0.089 kg /sec
1093
1097
1256
1422
1572
1737
1760
1833
1893
1893
1902
1897
1891
1900
1825
1816
1762
--
1532
Ta = 860°K
wa= 0.132 kg/sec
1183
1277
1499
1650
1779
1883
1927
2001
2027
2044
2026
2026
2028
2039
2017
1999
1919
1800
1667
*P = 1 atm, S = 0, $ S 1.0
112
-------
TABLE D-10. EXHAUST TEMPERATURE PROFILES FOR NATURAL GAS
COMBUSTION - S = 0.6
*".
1.0 (wall)
0.95
0.89
0.79
0.68
0.58
0.47
0.35
0.17
0.06
0. ((£)
0.10
0.20
0.33
0.45
0.56
0.68
0.77
0.88
TEMPERATURE, UK
$ = 0.8
P = 3.4 atm
1429
1587
1766
1900
1957
2060
2177
2246
2167
2083
2011
1998
2026
2014
2244
2259
2204
2126
2014
$ = 1.0
P = 3.4 atm
1595
1821
1989
2238
2390
2472
2539
2550
2525
_ _
2391
2386
__
2390
2486
2490
2490
--
2303
$= 0.8
P = 1 atm
1195
1464
1633
1704
1803
1873
1784
1639
1601
1808
1901
1845
1618
*Ta = 740°K,
s 0.133 kg/sec
113
-------
TABLE D-H- EXHAUST TEMPERATURE PROFILES FOR SYNTHESIZED
GAS COMBUSTION AT VARIOUS $*
'/'
1.0 (wall)
0.96
0.89
0.80
0.69
0.59
0.48
0.37
0.27
0.17
£ °-06
^ 0.04
0.14
0.25
0.35
0.46
0.56
0.67
0.78
0.87
TEMPERATURE, °K
$= 0.8
1210
1283
1410
1564
1717
1855
2040
2168
2182
2102
1934
1835
1808
1902
2098
2203
2157
2019
1895
1807
$= 0.9
1334
1413
1490
1637
1821
2012
2130
2190
2169
2060
1927
2170
1789
1790
2053
2213
2213
2122
2020
1916
$- 1.0
1416
1445
1569
1690
1887
2035
2096
2110
2062
1984
1859
1760
1750
1838
2017
2132
2146
2126
2064
1989
*P = 1 atm, S = 0.6, TJ, = 740°K, w = 0.137 kg/sec
111*
-------
TABLE E-l. BASELINE CASE CONCENTRATION DISTRIBUTIONS
Natural Gas-Air
Overall Equivalence Ratio =0.9
Pressure = 1 atm
Inlet Air Swirl Number = 0
Inlet Air Temperature = 74o°K
Air-Fuel Velocity Ratio =22.1
x/r0 - 0.32
r/ro
Species
CO (molefo)
C02 (mole$>)
02 (mole10
0.495
1.23
1.04
11.5
0.7
-
>10
0.062
2.27
1.9
7.6
1.2
>10
0.254
2.22
1.9
7.3
1.6
>10
0.328
2.04
1.71
7.3
1.5
>10
0.516
1.1
.94
13.0
0.7
13.0
>10
0.54l
1.02
.84
14.2
0.3
11.0
>10
0.561
0.85
.66
15.1
0.4
9-0
0.4i6
1.66
1.51
7.3
1.0
>10
0.599
0.71
.60
16.6
6.0
0.462
1.4
1.21
9.05
1.0
>10
0.653
0.56
.44
17.4
5.0
0.478
1.34
1.14
10.0
1.0
>10
x/ro = 5 « 17
r/ro
Species
CO (mole%)
C02 (mole^i)
Op (mole$)
NO (ppm)
NOX (ppm)
THC (molefo)
Ollf.)
3.14
6.83
5.0
83
127
1.9
0.125
3.08
6.75
5.1
85
122
.182
0.254
2.8
6.74
5.65
88
130
1.6o
0.395
2.34
6.63
6.5
70 .
110
1.32
0.495
1.95
6.25
7.55
70
108
1.20
0.711
1.4
5.65
9.05
52
88
0.9
0.852
1.05
5.28
10.6
50
76
0.79
0.944
0.8
5.13
11.6
43
68
0.685
115
-------
TABLE E-l ( CONT.)
x/r^ = 5.17*
r/ro
Species
CO (mole$)
C02 (raolefo)
02 (mole
-------
TABLE E-l (CONT. )
x/ro = 16.28
r/ro
Species
CO (mole0/,)
C02 (mole%)
02 (molefo)
NO (ppm)
NO, (ppm)
THC (mole%)
~1
0.121
2.4
7.4
4.85
183
215
0.35
0.249
2.27
7.34
5.0
186
219
0.32
0.395
2.03
7.43
5.11
180
210
0.27
0.545
1.62
7.63
5.3
162
190
0.215
0.665
1.2
7-74
5.52
147
173
0.145
0.790
0.81
8.03
5.55
137
160
0.09
0.869
0.6l
8.15
5.66
132
154
0.070
0.940
0.49
8.12
5.73
128
148
0.070
117
-------
TABLE E-2. SPECIES CONCENTRATION DISTRIBUTIONS FOR LOW EQUIVALENCE RATIO CASE
Natural Gas-Air
Overall Equivalence Ratio =0.7
Pressure - 1 atm
Inlet Air Swirl Number = 0
Inlet Air Temperature =7UO°K
Air-Fuel Velocity Ratio = 27
x/r0 =0.32
r/ro
Species
CO (mole$)
C02 (molei)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (mole$)
0 (?.)
3.81
2. U3
5.60
16.2
33.2
>10
0.079
3.81
2.68
6.30
17.5
-
>10
0.191
3.8l
2.62
6.00
17.5'
.-
>10
0.2k5
3.81
2.62
5.65
18.5
-
>10
0.32U
3.23
2.20
6.20
15.7
-
>10
O.U07
2.68
1.82
6.30
1U.5
-
>10
O.U91
1.78
1.26
9.55
8.70
-
>10
r/ro
Species
CO (mole$)
C02 (mole$)
02 (mole10
0.836
0.28
O.U7
18.8
2.3
6.50
3.60
0.9UU
0.26
1.02
19.0
3.3
6.3
1.75
x/ro =5-17
r/ro
Species
CO (mole
-------
TABLE E-2. (CONT.)
x/r0 = 5.17*
r/ro
Species
CO (molefo)
COp (mole$>)
02 (molefo)
NO (ppm)
NOX (ppm)
THC (mole'/o)
0 (? )
3.21+
6.75
5.^5
92
lUU
1.37
0.1U6
3AO
6.70
5.10
100
lU3
1.57
0.31+9
3.1+1+
6.75
1+.80
100
-
1.55
0.582
3.00
6.35
6.05
78
130 .
1.65
0.719
2.65
6.07
7.05
68
113
l.6o
0.857
2.20
5.75
8.05
53
91
1.6o
0.91+8
1.82
5.61
8.60
1+2
80
1.68
* Traverse made in combustor on opposite side of centerline from previous tab
le
x/ro =7.95
' r/ro
Species
CO (mole$)
C02 (mole'/o)
02 (mole^)
NO (ppm)
NOX (ppm)
THC (mole/0)
0 T?,^
3.07
6.53
5.U5
117
185
1.05
0.287
2.50
6.30
6.90
117
169
0.70
O.U86
1.70
5.85
8.75
97
0.^5
0.753
1.00
5-32
11.2
72
102
- 0.22
0.9U8
0.36
U.81
13.0
5U
76
0.15
x/ro = 10.73
r/ro
Species
CO (mole'/o)
COp (mole$)
02 (mole^)
NO (ppm)
NOX (ppm)
THC (molefo)
0 (?,)
2.60
6.80
6.05
lU?
190
0.52
0.262
1.90
6.56
7.^0
132
152
o.Uo
0.528
i.Uo
6.25
8.1+7
120
157
0.23
0.786
0.55
5.76
10.5
88
113
0.07
0.9UU
0.21
5.32
11.6
75
93
0.0k
119
-------
TABLE E-2 (CONT.)
x/r0 = 13.50
r/ro
Species
CO (mole$)
C02 (mole$)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (molefo)
0 (?,)
2.26
6.70
6.50
160
198
0.35
0.195
2.13
6.6l
6.90
150
185
0.31
0.382
1.70
6.50
7.55
1U3
177
0.22
0.6ll
0.99
6.h2
8.80
120
150
0.10
?
0.699
o.Ui
6.2U
10.1
100
122
o.oUU
0.9UU
o.Uo
6.15
10.7
87
10U
0.01
x/rQ = 16.28
r/ro
Species
CO (molefo)
C02 (molefo)
02 (molefo)
NO (ppm)
NOX (ppm)
THC (mole#)
0 (?. )
1.57
6.75
7.15
170
188
0.21
0.520
0.96
6.70
8.25
lUo
162
0.778
0.21
6.57
9.^5
120
135
0.092 0.037
'
0.9UO
0.11
6.U9
9.87
105
125
0.03
120
-------
TABU2 E-3. SPECIES CONCENTRATION DISTRIBUTIONS FOR SYNTHESIZED GAS CASE
Synthesized Fuel-Air
Overall Equivalence Ratio
Pressure = 1 atm
0.9
Inlet Air Swirl Number = 0
Inlet Air Temperature = 74o°K
Air-Fuel Velocity Ratio = 15.8
X/r0 =0.32
r/ro
Species
CO (mole*)
C02 (mole'/o)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (mole$)
x/r0 = 5.17
r/ro
Species
CO (mole$)
C02 (mole/0)
NO (ppm)
NOX (ppm)
THC (molefo)
0 ((£ )
13.7
2.20
5.64
0.8
o(£>
4.70
7.30
5.00
77
113
1.88
0.129
13.7
1.60
6.25
0.6
>10
0.262
14.8
5*64
0.6
>10
0.407
15.0
1.10
5-37
0.5
>10
0.536
8.20
0.70
11.8
0.4
>10
0.649
4.30
0.50
15-9
0.3
8.85
0.778
2.10
0.90
17-5
0.5
7.5
4.10
0.852
l.4o
2.15
16.3
3.2
10.0
2.35
0.121
4.45
7-30
5.12
77
118
1.75
0.245
4.00
7.30
6.12
75
115
1.60
0.395
3.15
6.90
7-38
67
110
1.08
- ' i
0.511
2.65
6.70
8.13
57
107
0.87
^ mm H »«*
1
0.711
1-95
6.10
9.90
57
90
0.60
* i ii-n
0.848
1.55
5-90
10.8
50
75
0.50
0.944
1.20
5.60
11.5
67
0.37
~
1.25
2.4o
16.3
2.5
5.0
1.75
~^
x/r0 =5.17*
r/ro
Species
CO (mole$)
COp (mole^)
02 (mole^)
NO (PP"i)
NOX (PPM)
THC (mole*)
0«fc)
4.10
6.80
6.45
65
112
1.92
O.l4l
4.4o
6.95
5.80
77
120
2.15
0.
4
7
5
83
125
2
0.282
4.50
7.00
5.50
83
125
2.08
^^ ^- ^-«
0.399
- ' i
4.4o
7.20
5.38
80
118
2.02
-' .1
0.524
^ «^-»^_
4.50
6.90
5-70
78
118
2.15
0.736
^"^- i^^ i
4.o6
6.90
6.20
72
107
2.17
0.877
3.60
6.85
6.52
65
104
2.02
0.956
3.4o
6.90
6.66
60
105
1.90
* Traverse made in combustor on opposite side of centerline from previous table.
121
-------
TABLE E-3 (CONT.)
x/r0 =7-95
r/ro
Species
CO (mole%)
C02 (mole$)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (molefo)
0 (<£)
5.1*5
7.00
if. 87
115
11*7
1.92
0.129
5-25
7.00
5.00
107
150
1.75
0.279
U.65
7.00
5.62
105
153
1.1*1*
0.1*11
1*.10
6.90
6.25
100
152
1.19
0.532
3.^5
6.90
7.00
96
lUl
0.9k
0.661
2.90
6.90
7.65
93
132
0.75
0.790
2.10
7.00
8.37
83
115
0.1*7
0.857
1.1*5
7.00
8.88
77
99
0.38
0.956
1.15
6.90
9.38
65
8?
0.29
x/r0 = 10.73
r/ro
Species
CO (raolefo)
C02 (mole$)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (mole$)
0 (<£)
-
7.05
I*. 83
1^5
190
1.53
0.125
_
7.05
>K96
Ikl
188
1.39
0.2J+5
.
7.00
5.U6
132
171
1.20
0.383
_
6.95
6.21
123
160
0.95
0.516
_
7.10
6.62
12k
163
0.62
0.578
_
7.16
7.18
116
11*5
0.56
0.769
_
7.27
7.66
97
130
0.33
0.857
_
7.38
8.00
86
107
0.26
0.9kk
7.27
8.U2
83
97
0.21
x/r0 =10.73
r/ro
Species
CO (mole$)
C02 (mole'JO
02 (mole$)
NO (ppm)
NOX (ppm)
THC (mole^)
0 (£ )
0.125
0.378
i 1
-
7.00
^.39
li*2
~
7.00
1*.15
1^5
i
i
;
1.1*0
1.1*0
-
7-35
3.82
152
-
1.57
0.578
-
7.^5
3.68
126
-
0.769
-
8.27
3-08
105
-
1.50 1.32
0.857
_
8.80
3.00
99
132
1.25
0.91*1*
_
8.65
3.25
88
12k
1.20
Traverse made in combustor on opposite side of centerline from previous table.
122
-------
TABLE E-3 (CONT.)
x/r0 =16.28
r/ro
Species
CO (mole$)
C02(mQle$)
Oo (mole$)
NO (ppm)
NOX (ppm)
IRC (mole
-------
TABLE E-4. SPECIES CONCENTRATION DISTRIBUTIONS FOR SWIRLING FLOW CASE
Natural Gas - Air Inlet Air Swirl Number = 0.3
Overall Equivalence Ratio =0.9 Inlet Air Temperature = 74Q °K
Pressure - 1 atm Air-Fuel Velocity Ratio = 22
x/r0 =0.32
r/ro
Species
CO (molefo)
C02 (mole$)
02 (moleio.o
0.099
4.31
2.74
1.27
17.0
-
>io.o
0.187
4.75
2.97
1.24
21.0
-
XLO.O
0.266
5.01
3.25
1.57
22.0
-
XLO.O
r/r0 0.495
Species
CO (mole*)
COp (mole$)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (molefo)
x/r0 - 5.17
r/ro
Species
CO (molefo)
C02 (molefo) !
02 (mole10.0
I
7.4o
' 6.68
80
145
-
2.64
0.536
2.97
2.47
7.44
10.0
-
>io.o
0.125 .0.270
7.13 6.37
6.76 6.76
0.77 1.03
148 166
-
2.55 1.57
0.599
1.99
l.8o
12.0
8.0
26.0
>io.o
0.391
5.33
7.3^
1.65
162
-
0.90
0.702
0.97
1.12
16.7
4.0
15.0
5.1
0.524
4.13
7.25
3.18
151
-
0.42
0.328
5.00
3.28
2.32
22.0
-
XLO.O
0.807
0.43
0.76
18.7
3.0
10.0
2.2
0.412
4.42
3.13
3-70
19.0
-
XLO.O
0.923
0.21
0.85
19-1
3.0
9.0
0.7
0.644
3.44
7.39
3.65
l4o
-
0.36
0.786
2.90
7.23
4.66
118
-
0.46
0.869 0.941
2.58 2.12
6.83 6.68
5.50 6.87
97 75-
125 98-
0.84 1.54
121*
-------
TABLE E-4 (COHT.)
Species
CO (mole$)
!02 (mole%)
WO (ppm)
>x (ppm)
THC (molefo)
5.63
7.98
0.7
247
0.097
5.33
8.1
0.74
255
0.082
U.68
8.33
0.99
253
0.060
2
D
0.383
3-57
8.57
1.88
250
0.054
2.8
235
x/ro =16.28
CO
co2
2
NO
THC
516
42
55
3
D49
0.657
1.77
8.59
3.7^
216
216
o.o44
0.782
i.Uo
8.51
^.35
197
198
0.0^7
0.852
1.26
8.U8
4.6l
190
189
o.o4o
L7
L7
<
0.93
8.45
4.98
'6
7
O.oUl
4.62
8.33
0.85
285
0.032
4.66
8.42
0.87
289
0.029
4.27
8.54
0.80
289
0.0245
3.47
8.85
1.20
0.019
2.4l
9-05
1.92
263
0.012
1.74
8.87
2.78
242
0.007
1.43
8.64
3.64
225
220
0.006
1.36
8.59
215
206
O.OC
0.96
8.73
4.30
205
200
0.005
125
-------
TABLE E-U (CONT.)
x/r0 = 5.17*
r/ro
Species
CO (molefo)
C02 (molef0)
02 (mole$)
NO (ppm)
NOX (ppm)
THC (mole
-------
TABLE E-5. SPECIES CONCENTRATION DISTRIBUTIONS FOR HIGH PRESSURE CASE
Natural Gas-Air
Overall Equivalence Ratio =0.9
Pressure =3.6 atm
Inlet Air Swirl Number = 0
Inlet Air Temperature = 740 °K
Air-Fuel Velocity Ratio = 22.0
x/ro =0.32
r/ro
Species
CO (molefo)
C02 (molefo)
Og (molefo)
NO (ppm)
N0x (ppm)
THC (molefo)
0 (?. )
4.95
1.98
2.15
9.0
>10
0
4.88
1.93
2.15
9.2
>10
0.079
5.07
1.95
9.2
>10
0.183
5.50
2.09
1.44
9.3
>10
0.254
4.93
1.98
0.80
9.7
>10
0.324
3.86
1.74
0.50
9.0
>10
0.408
2.86
1.42
0.90
5.7
>10
r/ro
Species
CO (molef,)
C02 (molefo)
02 (molefo)
NO (ppm)
NOX (ppm)
THC (mole%)
0.453
2.34
1.22
1.87
5.0
>Io
0.491
1.7
0.93
4.06
3.4
>io
0.536
0.96
0.58
9-3
2.5
>io
0.578
0.25
0.23
16.7
0.8
3.4
0.619
0.05
0.06
20.5
0.2
1.45
0 . 707
0.04
21.1
0.1
0.85
0.802
0.04
21.2
0.2
0.59
0 940
o.o4
21.2
0.1
0.43
x/ro =5.17
r/ro
Species
CO (molefo)
COo (molefo)
0,^~ (molefo)
NO (ppm)
NOX (ppm)
THC (mole%)
0 (^
1.1.7
4.19
0.70
95
-
4.10
0.145
11.1
4.83
0.58
143
-
3.3
0.266
9.18
5.92
0.90
2CA
-
2.3
i
0.412
5.83
7.07
2.45
272
.
1.0
0.532
3.12
7.07
5.1
307
H
0.46
0.632
1.59
6.54
8.15
295
Ml
0.21
0.731 I
0.593
4.97
12.4
225
218
0.065
0.869
0.15
3.71
15.4
148
150
0.052
1 0.940
0.08
2.75
16.9
93
100
0.052
127
-------
TABLE E-5 ( CONT.)
x/ro = 5.17* 1
r/ro
Species
CO (molefo)
GOp (molefo)
02 (molefo)
NO (ppm)
NOX (ppm)
TUG (mole%)
o (5) : 0.129
10.8
5.33
0.15
220
-
1.35
9.38
5.82
0.1*0
263
-
1.07
0.274
7.03
6.93
1.18
350
-
0.65
0.407
4.43
7.38
3.20
390
-
0.33
0.532
2.50
7.24
5.60
375
-
0.17
0.6^9
1.15
6.53
8.55
345
-
0.0725
0.769
0.34
5.03
12.U
235
230
0.033
0.857
0.052
3-77
14.6
140
148
0.024
0.940
0.02
3.32
15.7
92
120
0.028
* Traverse made in combustor on opposite side of centerline from previous table
x/ro = 10.73 1
r/ro
Species
CC (mole$)
C02 (mole^)
Op (mule$)
NO (ppm)
NOX (ppm)
THC (mole$)
0 (t)
9.80
6.05
0.23
300
0.78
0.133
8.58
6.60
0.60
370
0.53
0.262
6.1*2
7.20
1.30
406
0.35
0.399
V37
7.43
2.90
1*50
0.205
0.520
2/65
7-30
5.05
1*1*5
0.12
0.678
0.78
6.30
9.20
3U5
0.01*5
0.778 i 0.865
0.17
5.27
11.9
255
21*3
0.019
0.32
l*.i*0
13.7
177
180
0.015
0.931
0.02
4.1Q
13.9
155
170
0.017
x/ro = 13.50 1
r/ro
Species
CO (mole%)
C02 (mole%)
Og (mole"/o)
NO (ppm)
NO (ppm)
Tllfi (mole^o)
o (<£ ); o.sooj
7.!*7
6.72
0.5
1*30
.
0.23^
5.80
7.25
1.19
472
_
0.18
0.395
3.63
7.68
2.85
511
0.082
0.607
1.01
7.07
7.15
1*20
0.016
0.765 ! 0.936
0.175
5.81*
10.6
310
0.0025
0.02
5.16
11.9
240
240
0.001
x/ro - 16.28
r/ro
Species
CO (mole/0)
CO.^ (mole%)
02 (mole^)
NO (ppm)
NOX (ppm)
THC (mole%)
o (V.)
7.2k
7.15
0.7
508
-
0.13
0.187
5.68
7.63
l.ll*
577
-
0.087
0.353
3.8l
8.07
2.38
603
-
0.01*8
0.520 i 0.657
1.54
8.08
5.07
650
-
0.017
0.042
7.20
7.96
1*60
-
0.001*14
i
0.786
0.1
6.1*8
9.60
370
-
0.0008
0.936
0
6.14
10.7
344
-
0.00036
128
-------
TABLE E-6. SPECIES CONCENTRATION DISTRIBUTIONS FOR LIQUID PROPANE
Liquid-Propane Air
Overall Equivalence Ratio
Pressure = 1 atm
0.9
x/rn =5.17
r/ro
Species
CO (mole $)
C02 (mole %)
02 (mole %)
NO (ppm)
NOX (ppm)
THC (mole %)
0 (?.)
10.4
6.65
0.60
128
5.40
0.104
10. k
6.80
0.57
128
4.Uo
x/r0 = 10.73
r/ro
Species
CO (mole $)
C02 (mole $)
02 (mole $)
NO (ppm)
NOX (ppm)
THC (mole %}
0 (<£ )
12.3
5.25
0.19
160
»
3.45
0.125
13.6
5.80
0.19
155
2.60
Inlet Air Swirl Mumoer = 0.3
Inlet Air Temperature = 74-1 °K
Vr - °'
0.399
8.70
7.35
1.4o
128
2.55
0.391
8.70
7.70
0.77
190
0.90
"-^ ^ ^«
136 kg/sec
0.528
6.80
7-70
2.25
125
1.65
__
0.516
-^-^ ^ ^ « ^_
6.65
8.30
1.7
180
0.82
i
0.632
5-25
7-50
3.75
81
1.3
0.636 1
^ ^^-_
4.70
8.30
3-3
165
0.55
- ^ ^-^
0.802
2.95
6.35
8.0
57
97
1.3
0.923
1.90
5.25
11.6
47
79
1.0
0.778
2.57
7.85
6.1
123
170
0.35
O.Q?^
1.28
6.60
9.2
85
105
0.26
129
-------
TABLE E -- 6. (CONT.)
Conditions identical to previous page except:
Overall Equivalence Ratio = 1.04
x/r0 =5-17
r/ro
Species
CO (mole $)
COg (mole $>)
02 (mole $)
NO (ppm)
NOX (ppm)
THC (mole %)
0 (£)
9.75
6.70
0.75
95
-
9.80
0.129
9.70
6.70
0.75
98
-
9.60
0.245
9.35
6.90
0.95
97
-
8.90
0.383
8.60
7.10
1.4
9^
-
7.25
0.520
7.15
7.35
2.5
85
-
5-50
0.640
5.50
7.30
4.2
76
135
3.55
0.802
3.35
6.50
7.7
60
99
1.90
0.923J
2.10
5.60
11.5
46
V*
1-35
x/rn = 10.73
r/ro
Species
CO (mole %)
C02 (mole %)
02 (mole $)
NO (ppm)
NOX (ppm)
THC (mole $)
0 (? )
12.3
6.35
0.17
145
.
4.45
0.125
12.3
6.40
0.17
147
_
4.25
0.270
11.4
6.70
0.30
165
_
3-25
0.370
10.0
7.40
0.55
182
_
2.20
0.520
7.60.
8.00
1.3
188
,-.
1.25
0.653
5.55
8.30
2.6
165
200
0.88
0.790
3.4o
8.15
5-3
125
Itfi
0.62
0.915
1-95
7.4o
8.0
90
105
0.45
130
-------
TABLE E-7. BASELINE CASE VELOCITY PROFILES*
x/r0=0.32
r/r0 V
(tn/sec)
0.76 92
0.59 93
0.5^ 91
0.52 8?
0.51 82
0.1+6 U5
0.33 11
0.23 22
0 . 12 21+
0 26
x/rQ=5.17
r/r0 V
(m/sec)
i
0.76 86
0.61+ 76
0.51 75
0.37 75
0.25 81+
0.06 71+
0.08'S 7!+
0.17 79
x/rQ=5.l7+
r/r0 V
(m/sec)
0.76 82
0.61+ 85
o.5l
0.38 82
0.21+ 83
0.06 80
0.08'6 80
0.17 80
x/r0=7.95
r/r0 V
(m/sec)
0.76 71+
0.63 87
0.51 80
0.36 77
0.22 81
0.11 81+
0.01-6 82
0.18 71
'
x/rQ=10.73
r/r0 V
(m/sec)
0.76 76
0.63 81
0.1+9 75
0.39 89
0.21+
0.13 82
0.02-S 80
0.17 83
" :
x/rQ=13.50
r/r0 V
(m/sec)
0.76 67
0.63 77
0.52 83
0.38 8U
0.27 75
0.12 90
0.01-6 81
0.09 90
0.17 88
x/r0=l6.28
(m/sec)
«__
0.76 80
0.65 78
0.52 7!i
0.38 80
0.23 82
0.11 66
0.02-E 85
0.09 S3
0.16 90
Velocity profile, except for x/ro . 5.17+ taken on opposite aide of a from +he
majority o± the concentration and temperature profiles
TABLE E-8. BASELINE CASE TEMPERATURE PROFILES
x/rQ=0.32
r/r0 T
0 . 0 81+0
0 . 12 820
0.25 850
0.37 790
0 . 51 760
0.59 760
0.6'+ 780
x/r0=5.i7
v / v T1
o.o 1880
0.12 1910
0.25 1870
0.39 181+0
0.51 1800
0.71 1620
0.81+ 11+80
0.93 lt+20
x/r0=5.17*
r/r0 T
V '")
o.o 1900
0.12 i860
0.25 1870
0.39 181+0
0.53 1850
0.71 1800
0.81+ 171+0
o . 93 1600
"/r.,.7.95
r/r0 T
( K)
__ «_^_
o.o 191+0
0.11+ 1920
0.25 1900
0.1+1 1090
0.5^ 181+0
0.65 1750
0.78 1650
0.93 iM+0
x/r0=10.73
r/ro T
( K)
o.o 1960
0.13 191+0
0.27 1920
0.38 1880
0.52 1880
0.63 1830
0.78 1630
0.93 11+00
x/r0=13.5o
T IT T
/ O
(^ v \
1\ )
. .
0.0 1930
0.12 1920
0.21+ 1890
0.38 1890
0.51 1850
0.61+ 1770
0.78 1600
0.86 H+8o
x/rQ=l6.28
r/ro
( K)
c.o 1890
0.12 1890
0.21+ 1830
0.38 1830
0.53 1800
0.66 1730
0.78 1580
0.86 11+60
0.9^ li+OO j 0.93 1370
* Temperature profile taken on opposite side of
data.
from majority of the temperature
131
-------
TABLE S-9. VELOCITY PROFILES FOR L0₯ EQUIVALENCE RATIO CASE
x/r0=0.32
r/r0 V
(w.
sec)
0.76 80
0.68 82
0.57 95
0.52 76
o.M+ -
- -
x/r0=5.17
r/r0 V
(V,
sec)
0.76 68
O.k7 66
0.28 65
0.06 6l
x/r0=5.17+
r/rn V
/ o / /
("A
sec)
0.76 73
0.62 73
O.U2 68
0.19 66
0.06 66
0.08-E 66
6k
x/r0=7.95
r/r° t V/
(m/\
sec)
0.76 65
O.kQ 72
0.28 82
0.02 81
0.17~£ 72
x/r0=10.73
r/rn V
/ O / /
(V,
sec)
0.76 63
0.50 7^
0.2k 7k
o.oi 76
0.17~fi 73
x/r0=13.50
r/r0 V
(m/
sec)
0.76 63
0.57 80
0.39 72
0.16 66
O.Ol"^ 70
0.17 7^
x/r0=l6.28
r/r0 V
(V,
sec)
0.52' 76
0.25 6k
0.0(|) 7^
0.16 7U
__
TABLE E-10. TEMPERATURE PROFILES FOR LOW EQUIVALENCE RATIO CASE
r/r° TO
o.o 1050
0.08 960
0.19 890
0.25 890
0.33 850
O.itl 800
O.U8 720
0.57 710
0 . 78 710
0.93 730
x/rn=5.17
r/r0 T
(°V\
"/
o.o 1900
0.29 1830
O.U8 1720
0.75 1510
x/r0=5.17+
r/r0 T
o.o i8Uo
o.ik i£>ko
0.3^ i860
0.72 1750
o.Qk 1670
0.93 1520
x/r0=7.95
r/r0 T
o.o 1900
0.28 i860
O.kQ 1760
0.75 1550
0.9^ 1730
x/r0=10.73
r/r T
o.o 1890
0.26 1830
0.52 1750
0.78 1^50
0.93 Il8o
x/r0=13.50
r/r0 T
o.o 1830
0.18 18UO
0.37 1800
0.61 1680
0.78 lU8o
0.93 12^0
x/r 0==16. 23
T* / T* T1
/ O -tp \
0.0
0.2k 1800
O.k6 1730
0.76 1U50
0.92 1220
* ALL CONDITIONS SAME AS BASELINE EXCEPT *=- 0.7; Velocity profiles except for
x/r0 = 5.17-*- taken on opposite side of ^from the majority of the concentration
and temperature profiles. Temperature data at x/rQ = 5.17 + also taken on
opposite side of C from most concentration data.
132
-------
TABLE E-ll. SYNTHESIZED GAS CASE VELOCITY PROFILES*
x/ro=
r/r0
0.76
0.63
0.50
0.36
=0.32
( V/
sec)
71
35
37
Ik
x/r0=5.17
y
r/r0 (m/
sec)
0.76 30
o.6k 31
0.52 22
>5
)
x/r0=10.73
1
V,
r/r0 (m/
sec)
0.76 7!+
0.6k 80
0.52 82
0.38 75
0.25 81+
0.19 87
o.o(g) 86
0.08 85
0.17 78
x/r0=lo.73 +
/ (V/
r/r0 (W
sec)
0.76 79
0.6k 8k
0.39 86
0.19 89
O.o(g) 89
0.08 86
0.17 85
x/r0=l6.28
V .
r/r0 ("V,
sec)
NO DATA
TABLE E-12. SYNTHESIZED GAS CASE TEMPERATURE PROFILES*
x/rQ=0.32
r/ro (°K)
0.0 860
0.13 800
0.26 770
OA2 680
0.53 700
0.63 730
0.77 780
0.83 850
0.93 880
x/r0=5.i7
/ T
r/ro (°K)
o.o 1970
0.13 1930
0.25 1920
O.IK> 1875
0.50 1800
0.70 1670
0.83 1580
0.93 ikko
x/i-0=5.17+
/ T
r/r0 (oK)
NO DATA
x/r0=7.95
/ T
r/ro (°K)
o.o 1820
0.13 i860
0.26 1920
O.IK) 1800
0.52 1820
0.63 1800
0.77 1670
0.8k 1530
0.86 11+50
0.93 1370
x/r0=ao.73
/ T
r/ro (OK)
o.o 1850
0.13 1850
0.2U 1820
0.38 1820
0.51 1780
0.55 1800
0.77 1590
0.85 li+70
0.93 1350
'
"
--
x/r0=lo.73+
/ T
r/ro (°K)
0.58 1880
0.76 1760
0.85 1570
0.93 1330
1
^ ^ _^
x/r0=l6.28
..._._.. .
r/ro (oK)
o.o 1750
0.13 1760
0.26 1750
0.38 17^0
0.50 1720
0.63 l6Uo
.,
^Velocity profiles except for those at x/rn=5.i7 + and x/r -lo 7^ 4- v
side of C for the mainri-Hr n-p -t-v L ^ x/r -10.73 + taken on opposite
siae 01 ^ ior -one majority of the concentration and temperature ds-hn v I
data at x/rn-5.17 + and x/r -in 7-5 + i ^- , temperature data. Temperature
aa^d. ao */io ? ±, ana x/ro-10.73 + also taken on opposite side of e from mo^t
concentration data. ^ irom most
133
-------
TABLE E-13. VELOCITY PROFILES FOR SWIRLING FLOW CASE*
x/r0=0.32
r/r0 V
(m/sec)
0.76 56
0.65 67
0.58 77
0.56 58
0.54 48
0.50 35
0.44 ...
0.28 26
0.24 22
0.17 25
0.07 26
0.03"£ 16
0.15 27
» -
x/r0=5.17
r/ro v
(m/sec)
0.76 120
0.64 108
0.50 107
0.38 80
0.25 70
0.13 49
O.O(g) 33
0.09 39
0.17 28
x/r0=5.17+
r/ro V
(m/sec)
0.76 110
0.64 116
0.50 97
0.38 70
0.25 60
0.13 44
0.0(£) 36
0.08 23
0.15 30
x/r0=7.95
r/r0 V
(m/sec)
0.76 124
0.64 112
0.50 103
0.38 81
0.25 6l
0.13 49
O.O(fi) 45
0.08
0.17 37
.
x/ro=10.73
r/r0 V
(m/sec)
0.76 113
0.63 105
0.51 103
0.38 97
0.24 91
0.12 82
O.O(g) 74
0.09 62
0.17 65
x/r0=13.50
r/r0 V
(m/sec)
0.76 111
0.62 114
0.50 108
0.38 101
0.25 85
0.11 91
0.01-6 76
x/ro=l6.28
r/r V
(m/sec)
0.76 111
0.64 114
0.51 100
0.39 98
0.25 93
0.13 89
o.o(e) 81
0.08 71 \ 0.08 74
0.17 45 ! 0.17 78
* Velocity profiles except for x/rQ -- 5-17+taken on opposite side of (£ from the majority
of the concentration «nri temperature profiles. All conditions the same as baseline
case except S = 0.3.
TABTE E-14. FLOW DIRECTION DATA FOR SWIRLING FLOW CAGE*
I i 1
x/ro = 0.32
r/r0 Q- 0 6
i (deg)
0.58 -14 29 32
0.56 -5 38 38
0.54 8 --
0.50 9 10 13
0.40 11 2 11
x/r0 =5.17
r/ro a $ B
(deg)
0.76 -10 30 32
0.64 -15 39 41
0.50 -15 32 35
0.38 -
0.25 -7
0.13 6
0.0 -6 15 16
x/r0 =5-17+
r/ro » 3 9
(deg)
0.76 -9 35 36
0.63 -10 29 30
0.51 7 47 47
0.38 10
0.25
0.13
0.0 7 25 26
i 1
x/r0 =10.73
r/ro <* P e
(deg)
0.76 -10 27 29
0.63 -5 31 31
0.51 -10 31 32
0.38 -5 31 31
0.24 8 27 28
0.12 7 27 28
0.0 8 9 12
x/r0 =13.50
r/ro a 3 6
(deg)
0.76 -12 25 28
0.62 -10 23 26
0.50 -6 24 25
0.38 4 26 26
0.25 -6 25 26
0.01 13 14 19
0.08 11 9 i4
0.17 17 5 18
* Profiler, except for x/rQ =5.17+taken on opposite sides of (£ from the majority
of the concentration and temperature profiles.
134
-------
TABU: E-15. TEMPERATURE PROFILES FOR SWIRLING FLOW CASE
r/ro
0.90
0.88
0.80
0.78
0.68
0.59
0.1+9
0.39
0.28
0.26
0.18
0.10
0.06
*o.o3~k
X0.12
*0.ll+
-X0.23
O.25
*0.32
0.3!+
*0. 1+3
-xO.53
-x-0.59
*o.6s
*0.75
O.81+
*0.87
TEMPERATURE (°K)
Vr0 = 5.17
...
1923
2062
211+6
2215
2230
2230
2212
2192
2111
2022
2005
. . .
2073
2150
2208
2321+
2310
233^
221+6
* t
_
x/r0 - 10.73
1879
2003
2090
211+1+
2181+
2221+
2221+
*
221+3
2199
213*+
2176
2128
*
2137
2170
2193
2183
211+1+
2056
i
* Most local concentration data taken on this side of
combust or g
135
-------
TABLE E-16. TEMPERATURE PROFILES FOR HIGH PRESSURE CASE*
x/r0=0.32
T
r/ro (oK)
o.o 900
0.0? 1080
0.18 1190
0.26 980
0.33 920
o.ho 820
OA5 770
0.^9 780
0.53 760
o . 58 800
0.62 830
0.70 850
0.80 830
0.93 770
x/r0=5.17
T
r/ro (oK)
o.o 1980
0.13 1980
0.27 1980
o.Ui 1980
0.53 1920
0.63 1820
0.76- 1720
0.85 1580
0.93 1^30
x/r0=7-95
T
r/ro (°K)
0.0 2020
0.12 2070
0.27 2130
o.ho 2100
0 . 53 20^0
0.65 1970
0.76 1730
0.85 15^0
0.93 lUio
x/r0=10.73
T
r/ro (°K)
o.o 19^0
0.26 2020
o.ho 2070
0 . 51 2020
0.68 1910
0.76 1700
0.86 1550
0.92' 1^90
x/ro=13.50
T
r/r° ' («K)
o.o 2130
0.20 2080
0.39 2110
0.60 2100
0.76 1800
o ..93 1600
x/r0=l6.28
T
r/r° (°K)
o.o 2150
0.19 2190
0.35 2130
0.51 2110
0.65 2010
0.78 i860
0.93 1580
Conditions same as baseline case except P = 3.6 atm.
136
-------
TABLE E-1T. TEMPERATURE PROFILES FOR LIQUID PROPANE COMBUSTION
_ _
r/ro
0.90
0.80
0.78
0.68
0.58
0.56
0.49
0.47
0.38
0 ^6
0.29
0.27
0.18
0.17
0.16
* 0.03
0.05
0.13
0.14
0.15
0.24
0.25
0.35
0.44
0.55
0.57
0.66
0.75
0.77
0.87
0.88
TEMPERATURE (°K)
* = 0.92 ± o.oi
x/ro = 5.17
1631
1795
1984
2091
2112
2121
2063
*"
1975
1871
1791
1789
-
1817
1905
2003
2075
2099
2076
1882
-
x/rQ = 10.73
1617
1881
2052
2134
2201
2211
;
2148
2038
1921
1881
"
1883
1898
2027
2142
2187
2173
2104 -
7864
-____
* = i.o4 ± o.oi
x/rQ =5.17
1611
1859
1952
2041
2059
_
204l
2004
-
1913
1823
1745
1736
1780
1876
1946
2039
2061
2035
1852
- _ _
. _
x/rQ = 10.73
1816
2012
2106
2183
2215
2206
-
2151
-
2058
1982
1911
-
1863
~
-
1893
1987
2089
2150
2163
-
1917
__ _______
137
-------
APPENDIX F
FUEL COMPOSITION
Natural Gas
TABLE F-l. NATURAL GAS COMPOSITION
Species
City
C2H6
co2
C3H6,C3H8
02+Ar
N2
i-CUHio
n-C,H10
Mole Percent
Trailer No. 1
96.26
1.89
0.70
0.2k
0.10
0.73
o.oU
o.oU
No. 2
96.71
1.87
0.55
0.19
O.OU
0.55
0.05
0.0k
No. 3
96.11
2.17
0.59
0.28
0.08
0.66
0.06
0.05
Average
96.21
1.98
0.6l
0.2k
0.07
0.65
0.05
o.ok
Synthesized Fuel (CH^/H2/CO)
TABLE F-2. SYNTHESIZED FUEL COMPOSITION*
Species
H
C2H6
1 u
3H8
N
C0
Mole Percent
27.18
69.83
0.85
0.83
0.27
*COwas mixed with fuel mixture prior
to injection into burner. The nomi-
nal CHj/Hg/CO ratio was 1/0.39/0.3!*.
138
-------
APPENDIX F (CONT.)
Propane (Natviral Grade)
TABLE F-3. PROPANE FUEL COMPOSITION
Species
C3H8
C^H10 and
lighter fractions
Mole Percent
96
k
139
-------
NOMENCLATURE
A Defined by Eq. (B-3a)
A* Orifice Throat Diameter, cm2
B Defined by Eg.. (B-3b)
CD Nozzle flow coefficient
d Outer diameter of air annulus, cm
de Effective diameter for swirl number = dd-Z2)0'^ cm
df Fuel delivery duct inner diameter, cm
dn Inner diameter of air annulus, cm
d0 Combustor duct diameter, cm
f Frequency, Hz
F Function of ratio of specific heats and molecular weight
£ Annular air gap height, cm
m Air-fuel velocity ratio
P Combustor pressure level, atm
P1'P3 Pitch plane static pressure, in HpO
P2'pi| Yaw plane static pressures, in HpO
PTJ_ Impact pressure , torr
PT2 Downstream orifice total pressure, torr
Ap Geometric average pressure, Eg. (B-5), torr
Q Volumetric flow rate,cm3/sec
1 Dynamic pressure, torr
r Radial dimension, cm
140
-------
NOMENCLATURE (COST.)
Combustor radius, cm
Q Swirl number, Eq_. (A-l)
t Temperature, °R
T Temperature, °K
Uncorrected free stream temperature, °K
Conduction correction, °K
Radiation correction, °K
,TT Free stream stagnation temperature, °K
Downstream orifice temperature, °K
Axial velocity, m/sec
Mass flow, kg/sec
Axial (longitudinal) dimension, cm
Average exhaust concentration
Ratio of inner and outer diameters of the air annulus
Pitch angle, deg.
Yaw angle, deg.
Swirl vane angle, deg.
Total flow angle, deg.
Jet thrust, kg-m/sec2
Density, kg/m3
Nozzle torque, kg-m2/sec2
Overall fuel/air equivalence
stoichiometric
-------
NOMENCLATURE (COM?. )
Subscripts
a Air stream
f Fuel stream
w Wall
1 Orifice no. 1 (upstream orifice) of DSO probe
2 Orifice no. 2 (downstream orifice) of DSO probe
-------
REFERENCES
1. ' a) M. P. Heap, T. M. Lowes and R. Walmsley: The Effect of Burner Parameters on
Nitric Oxide Formation in Natural Gas and Pulverized Fuel Flames, Paper
presented at First American Flame Days Meeting, Chicago, Illinois, September
X;7 \ £- «
b) M. P. Heap, T. M. Lowes and R. Walmsley: Emission of Nitric Oxide From Large
Turbulent Diffusion Flames, Fourteenth Symposium (international) on Com-
bustion (The Combustion Institute, Pittsburgh, 1973), p. 883.
2> D. R. Shoffstall and D. H. Larson: Aerodynamic Control of Nitrogen Oxides
and Other Pollutants from Fossil Fuel Combustion, EPA Report 650/2-73-033a
October 1973. '
3. a) A. M. Mellor, R. D. Anderson, R. A. Altenkirch and J. H. Tuttle: Emissions
from and within an Allison J-33 Combustor, Combust. Sci, and Tech. 6, 169
(1972).
b) J. H. Tuttle, R. A. Altenkirch and A. M. Mellor: Emissions from and within
an Allison J-33 Combustor II: The Effect of Inlet Air Temperature, Combust.
Sci. and Tech. ]_, 125 (1973).
**' R' E' Jones and J- Grobman: Design and Evaluation of Coinbustors for Reducing
Aircraft Engine Pollution, AGARD Document CP-125, April 1973.
5 F. C. Gouldin: Role of Turbulent Fluctuations in NO Formations, Combust
Sci. and Tech. 9_, 17 (197*0.
6. R. F. Anasoulis and H. McDonald: A Study of Combustor Flow Computations
and Comparison with Experiment, EPA Report 650/2-73-0^5, December 1973.
T« L< S' Caretto: Modeling Pollutant Formation in Combustion Processes,
Fourteenth Symposium (international) on Combustion (The Combustion
Institute, Pittsburgh, 1973), p. 803.
8. D. B. Spalding: Mathematical Models of Continuous Combustion, Proceedings
of the Symposium on Emissions from Continuous Combustion Systems (Plenum
Press, New York, 1972), p. 3.~~
9. J. M. Bee'r and N. A. Chigier: Combustion Aerodynamics (Halsted Press
Division, John Wiley & Sons, Inc., New York,1972).
-------
REFERENCES (CQKT.)
10. C. T. Bowman, L. S. Cohen and M. N. Director: Nitric Oxide Formation in
Combustion Processes with Strong Recirculation, EPA Report EPA-R2-73-
291, July 1973.
11. T. Rozenman and H. Weinstein: Recirculation Patterns in the Initial Region
of Coaxial Jets, NASA CR-1595 , May 1970.
12. A. J. Giramonti: Advanced COGAS Power Systems for Low Pollution Emissions,
Paper presented at the ACS Symposium on Novel Combined Power Cycles
April 1973. '
13. J. B. Kennedy and J. B. McVey: An Experimental Study of Fuel Spray
Vaporization Rates, United Aircraft Research Laboratories Report UAR-N139,
September
1^. F. K. Owen: Laser Velocimeter Measurements in Free and Confined Coaxial
Jets with Recirculation, AIAA Paper 75-120, presented at 13th AIAA Aero-
space Sciences Meeting, January 1975,
15. J. H. Tuttle, R. A. Shisler and A. M. Mellor: Nitrogen Dioxide Formation
in Gas Turbine Engines - Measurements and Measurement Methods , Report
PURDU-CL-73-06 , December 1973.
16. N. P. Cernansky: Formation of NO and N©2 in a Turbulent Propane/Air
Diffusion Flame, Ph.D. Thesis, University of California, Berkeley,
California, I9lk.
17. R. W. Schefer, R. D. Matthews, N. P. Cernansky and R. F. Sawyer: Measure-
ment of NO and N02 in Combustion Systems, Paper 73-31, Fall Meeting,
Western Section/The Combustion Institute, 1973.
18. N. Syred and J. M. Bee': Combustion in Swirling Flows: A Review, Combust.
and Flame 23_, 1^3 (197^).
19, R. W. Bilger: Turbulent Diffusion Flame Structure and Implications for
Suppression of Pollutant Formation, Paper presented at the Spring Meeting,
Western and Central Sections /The Combustion Institute, 1975.
20. L. S. Cohen and R. N. Guile: Measurements in Freejet Mixing/Combustion
Flows. AIAA J. 8, 1053
-------
REFERENCES (COOT.)
21. N. M. Kerr and D. Fraser: Swirl. Part I: Effect on Axisymmetrical
Turbulent Jets, J. Inst. Fuel 38, 519 (1965).
22. A. H. Shapiro: The Dynamics and Thermodynamics of Compressible Fluid
Floy. Vol. I. (Ronald Press, New York, 1953).
23. R. J. Flaherty: Arbitrary Pressure Gradient Integral Technique for
Predicting Boundary Layer and Thermal Parameters. J. Aircraft 11,
293 (197*0. ~~~
2U. Methods of Air Sampling and Analysis (American Public Health Association,
Washington, D. C., 1972), p. 199.
25. E. L. Merryman and A. Levy: Nitrogen Oxide Formation in Flames: The
Roles of N02 and Fuel Nitrogen, Fifteenth Symposium (international) on
Combustion (to be published).
26. J. D. Allen: Probe Sampling of Oxides of Nitrogen from Flames, Combust.
and Flame 2jt_, 133 (1975).
27. H. Amin: Analytical Study of Probe Effects on Sampling of Nitrogen
Oxides from Flames (to be published).
-------
TECHNICAL REPORT DATA
(Please read liiainiciions i»n tlic reverse before completing)
1. REPORT NO.
A. TITLE AND SUBTITLE
In fluence of Aerodynamic Phenomena on Pollutant
Formation in Combustion
Volume I. Experimental Results
.REPORT DATE
July 1975
7. AUTHOR(S)
Craig T. Bowman and Leonard S. Cohen
6. PERFORMING ORGANIZATION CODE
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORGANIZATION NAME AND ADDRESS
United Technologies Research Center
400 Main Street
East Hartford, CT 06108
12, SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
Industrial Environmental Research Laboratory
Research Triangle Park, NC 27711
3. RECIPIENT'S ACCESSION" NO.
10. PROGRAM ELEMENT NO.
1AB014:
11. CONTRACT/GRANT NO.
68-02-1092
13. TYPE OF REPORT AND PERIOD COVERED
Final; 7/73 - 12/Tj
14. SPONSORING AGENCY CODE
15. SUPPLEMENTARY NOTES
report gives results of the measurement of average concentration levels
of NO, NO2, CO, and unburned hydrocarbons (THC) at the exhaust of an axisym-
metric combustor over a significant range of operating conditions. In addition, it g
detailed species concentration, temperature, and velocity maps throughout the com-
bustor for seven representative operating conditions. In the combustor, natural ga >
a synthesized CH4/CO/H2 fuel, or vaporized propane issued through a central due
mix and burn with an annular air stream in a 1. 8 m long cylindrical duct. In a few
tests, liquid propane was the fuel. Major combustor input parameters varied over
following ranges: overall fuel/air equivalence ratio 0. 5-1. 3, air/fuel velocity r*t10.
0.1-40, inlet air swirl number zero-0. 6, air flow rate 0. 088-0.137 kg/sec, inlet ai
temperature 730-860K, and combustor pressure 1-7 atm. Water-cooled probes wer
used to remove samples from the flow for on-line concentration analysis and to me
sure temperature, velocity, and flow direction. Elevated pressure and introduction
of swirl, to the extent considered in the present experiments, create 'unmixednes
in the combustor flow field which in turn results in enhanced NO formation and co -
sumption of hydrocarbons. Aerodynamic flame stabilization produces strong stirri
which results in relatively low NO formation and hydrocarbon consumption rates
17.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFlERS/OPEN ENDED TERMS
Air Pollution
Combustion
Aerodynamics
Nitrogen Oxide (NO)
Nitrogen Dioxide
Methane
Propane
Flow Distribution
13. DISTRIBUTION STATtMbNf
Unlimited
Air Pollution Control
Stationary Sources
Chemical Heat Release
Flame Stabilization
Flow Field Mapping
Low-Btu Gas
19. SECURITY CLASS (This Report)
Unclassified
20. SECURITY CLASS (This page)
Unclassified
EPA Form 2220-1 (9-73)
------- |