Parameters for 

Properly Designed and Operated Flares 

Report for Flare Review Panel 

April 2012 

Prepared by 

U.S. EPA Office of Air Quality Planning and Standards (OAQPS) 

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information quality guidelines. It has not been formally disseminated by EPA. It does not represent 

and should not be construed to represent any Agency determination or policy.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

Acronyms Page i

ACRONYMS 

Acronym Definition 

AFTIR Active Fourier Transform Infrared 

API American Petroleum Institute 

ARI Aerodyne Research, Inc. 

BTU British Thermal Units 

CZ Combustion Zone Gas 

CCZ Fraction of Combustibles in the Combustion Zone Gas 

CFR Code of Federal Regulations 

DCS Distributed Control System 

EPA U.S. Environmental Protection Agency 

FHR Flint Hills Resources 

FHR AU Flint Hills Resources - Aromatics Unit 

FHR LOU Flint Hills Resources - Light Olefins Unit 

FLIR Forward Looking Infrared 

FTIR Fourier Transform Infrared Technology 

IMACC Industrial Monitor and Control Corporation 

INEOS INEOS ABS (USA) Corporation 

ISO International Standards Organization 

LFL Lower Flammability Limit 

LFLCZ Lower Flammability Limit of the Combustion Zone Gas 

LFLVG Lower Flammability Limit of the Flare Vent Gas 

LHV Lower Heating Value 

LFLVG, C Lower Flammability Limit of the Combustible Portion of the 

Flare Vent Gas 

MFR Momentum Flux Ratio 

MPC Marathon Petroleum Company, LP 

MPC Detroit Marathon Petroleum Company, LP Detroit Refinery 

MPC TX Marathon Petroleum Company, LP Texas City Refinery 

NESHAP National Emission Standards for Hazardous Air Pollutants 

NHV Net Heating Value 

NHVCZ Net Heating Value of the Combustion Zone Gas 

NHVLFL Net Heating Value of the Flare Vent Gas if Diluted to the Lower 

Flammability Limit 

NHVVG Net Heating Value of the Flare Vent Gas 

NHVVG-LFL Net Heating Value of the Flare Vent Gas if Diluted to the Lower 

Flammability Limit 

NSPS New Source Performance Standards 

OAQPS Office of Air Quality Planning and Standards 

PFTIR Passive Fourier Transform Infrared Technology 

SCF Standard Cubic Feet 

SDP Shell Deer Park Refinery 

SDP GF Shell Deer Park Refinery Ground Flare 

SDP EPF Shell Deer Park Refinery East Property Flare 

SR Stoichiometric Air Ratio 

TCEQ Texas Commission on Environmental Quality 
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Acronyms Page ii

Acronym Definition 

UFL Upper Flammability Limit 

Vmax Maximum Flare Tip Velocity Including, if Applicable, Center 

Steam at Which Flame Lift Off is Not Expected to Occur 
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Table of Contents Page i

TABLE OF CONTENTS 

1.0 INTRODUCTION ....................................................................................................... 1-1

2.0 AVAILABLE FLARE TEST DATA ............................................................................ 2-1

2.1 Flare Performance Studies and Test Reports ......................................... 2-1

2.2 Flare Vent Gas Constituents ................................................................. 2-3

2.3 Steam Injection Rates and Tip Design for Available Flare Test Data .... 2-5

2.4 Air Injection Rates and Tip Design for Available Flare Test Data ......... 2-6

2.5 Flare Test Methods ............................................................................... 2-7

2.6 Combining All Available Test Run Data ............................................... 2-9

2.7 Data Removed After Being Considered .............................................. 2-10

2.8 Determination of Combustion Efficiency Representing Good Flare 

Performance ....................................................................................... 2-11

3.0 STEAM AND FLARE PERFORMANCE .................................................................... 3-1

3.1 Lower Flammability Limit of Combustion Zone Gas for Steam-Assisted 

Flares ................................................................................................... 3-1

3.1.1 Flare Test Data and LFLCZ ........................................................ 3-5

3.1.2 The Le Chatelier Principle ........................................................ 3-7

3.1.3 Specific Test Data Not Fitting the Trend ................................. 3-11

3.1.4 Data Points with Good Combustion and High LFLCZ .............. 3-20

3.1.5 Excluding Pilot Gas ................................................................ 3-26

3.2 Combustible Gas Concentration in the Combustion Zone ................... 3-27

3.3 Heat Content Based Limit for Steam-Assisted Flares .......................... 3-28

3.4 Other Operating Parameters Considered for Steam-Assisted Flares..... 3-31

3.4.1 Net Heating Value .................................................................. 3-31

3.4.2 Steam Ratios ........................................................................... 3-34

4.0 AIR AND FLARE PERFORMANCE .......................................................................... 4-1

4.1 Stoichiometric Air Ratio ....................................................................... 4-1

4.2 TCEQ Test Data ................................................................................... 4-3

4.3 Other Test Data .................................................................................... 4-4

4.4 Analysis of Stoichiometric Air Ratio .................................................... 4-4

4.5 Considering LFLVG for Air-Assisted Flares .......................................... 4-7

5.0 WIND AND FLARE PERFORMANCE ...................................................................... 5-1

5.1 Introduction .......................................................................................... 5-2

5.2 Flare Flow Mixing Regimes ................................................................. 5-2

5.3 Efficiency Studies................................................................................. 5-4

5.4 Test Data Analysis .............................................................................. 5-10

6.0 FLARE FLAME LIFT OFF ......................................................................................... 6-1

6.1 Literature Review and Vmax Calculation ................................................ 6-1

6.2 Test Data Analysis ................................................................................ 6-2

6.3 Other Operating Parameters Considered for Flame Lift Off .................. 6-6
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Table of Contents Page ii

7.0 OTHER FLARE TYPE DESIGNS TO CONSIDER ..................................................... 7-1

7.1 Non-Assisted Flares.............................................................................. 7-1

7.2 Pressure-Assisted Flares and Other Flare Designs ................................. 7-2

8.0 MONITORING CONSIDERATIONS ......................................................................... 8-1

8.1 LFLCZ, LFLVG, and LFLVG,C ................................................................. 8-1

8.2 Ratio of NHVCZ to NHVVG-LFL ............................................................. 8-2

8.3 CCZ ....................................................................................................... 8-2

8.4 SR ........................................................................................................ 8-3

8.5 MFR ..................................................................................................... 8-3

8.6 Vmax ...................................................................................................... 8-4

9.0 REFERENCES ............................................................................................................ 9-1

TECHNICAL APPENDICES 

Appendix A. Brief Review Summary of Each Flare Performance Study and Test Report. 

Appendix B. Excel Workbook That Combines All Data Sets. 

Appendix C. Test Report Nomenclature Matrix. 

Appendix D. Detailed Calculation Methodologies For The Specific Parameters. 

Appendix E. Type and Amount of Components in Each Test Run by Test Report. 

Appendix F. Charts of Calculated and Measured LFL for Various Combustible Gases in 

Nitrogen and Carbon Dioxide. 

Appendix G. Details About Inerts and Further Explanation for Including an 

Equivalency Adjustment to Correct For Different Inert Behavior. 

Appendix H. Effect of Nitrogen and Carbon Dioxide on the LFL of Various 

Components: A Comparison of Le Chatelier Equation to Experimental 

LFL Values. 

Appendix I. Methodology for Calculating Unobstructed Cross Sectional Area of 

Several Flare Tip Designs. 

 
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Table of Contents Page iii

LIST OF TABLES 

Tables Page 

Table 2-1. Flare Performance Test Reports .............................................................................. 2-2

Table 2-2. Flare Vent Gas Constituents by Test Report ........................................................... 2-3

Table 2-3. Minimum, Maximum, and Average Volume Percents of Primary Constituents in Flare 

Vent Gas .......................................................................................................... 2-4

Table 2-4. Steam-Assisted Flare Tip Design Detail ................................................................. 2-5

Table 2-5. Air-Assisted Flare Tip Design Detail ...................................................................... 2-7

Table 2-6. Criteria To Exclude Data Points ............................................................................2-12

Table 3-1. Recommended Values of Coefficient of Nitrogen Equivalency for Water and Carbon 

Dioxide Relative to Nitrogen ...........................................................................3-10

Table 3-2. Test Run Detail for 11 Data Points with LFLCZ < 15.3% but Combustion Efficiency < 

96.5% .............................................................................................................3-12

Table 3-3. Olefin and Hydrogen Approximately Equal ...........................................................3-15

Table 3-4. High Hydrogen and Low Olefin ............................................................................3-16

Table 3-5. Higher Olefin and Low Hydrogen .........................................................................3-18

Table 3-6. Breakdown Of Steam Use For The 66 Test Runsa ..................................................3-21

Table 3-7. Potential LFLcz Thresholds based on LFLVG,C........................................................3-23

 
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Table of Contents Page iv

LIST OF FIGURES 

Figures Page 

Figure 3-1. Zabetakis Nose Plot For Methane And Inert In Air ................................................ 3-3

Figure 3-2. Time Sequence of Flare Vent Gas Volume Moving Through Flammability Region 

Source: (Adapted from Evans and Roesler, 2011) ............................................ 3-4

Figure 3-3. Combustion Efficiency vs. LFLCZ ......................................................................... 3-7

Figure 3-4. Combustion Efficiency vs. LFLCZ Adjusted for Nitrogen Equivalency .................3-11

Figure 3-5. Combustion Efficiency vs. LFLCZ for Category A Test Runs (see Table 3-7) ........3-23

Figure 3-6. Combustion Efficiency vs. LFLCZ for Category B Test Runs (see Table 3-7)........3-24

Figure 3-7. Combustion Efficiency vs. LFLCZ for Category C Test Runs (see Table 3-7)........3-24

Figure 3-8. Combustion Efficiency vs. LFLCZ for Category D Test Runs (see Table 3-7) ........3-25

Figure 3-9. Combustion Efficiency vs. CCZ ............................................................................3-28

Figure 3-10. Combustion Efficiency vs. Ratio of NHVCZ to NHVVG-LFL .................................3-29

Figure 3-11. LFLCZ vs. Ratio of NHVCZ to NHVVG-LFL ...........................................................3-30

Figure 3-12. Combustion Efficiency vs. NHVVG ....................................................................3-32

Figure 3-13. Combustion Efficiency vs. NHVCZ .....................................................................3-33

Figure 3-14. NHVVG vs. LFLVG ..............................................................................................3-34

Figure 3-15. Combustion Efficiency vs. S/VG by weight .......................................................3-35

Figure 3-16. Combustion Efficiency vs. S/VG by volume ......................................................3-36

Figure 3-17. Combustion Efficiency vs. S/HC by volume.......................................................3-37

Figure 4-1. Combustion Efficiency vs. SR (using TCEQ data) ................................................ 4-3

Figure 4-2. Combustion Efficiency vs. SR (using EPA-600/2-85-106 data) ............................. 4-4

Figure 4-3. Combustion Efficiency vs. SR (combined TCEQ and EPA-600/2-85-106 data) ..... 4-5

Figure 4-4. Combustion Efficiency vs. SR, zoomed (using TCEQ data) .................................. 4-6

Figure 5-1. Air Egression Into Flare Stack Source: (Smoot et al., 2009) .................................. 5-2

Figure 5-2. Images of Flow Mixing Regimes Source: (Seebold et al., 2004) ............................ 5-3
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Table of Contents Page v

Figure 5-3. Fuel Detection Downwind of Wake-Dominated Flare Source: (Johnston et. al, 2001)5-4

Figure 5-4. Flame Images Relating to Momentum Flux Ratio and Combustion Efficiency Source: 

(Johnson and Kostiuk, 2000) ............................................................................ 5-6

Figure 5-5. Combustion Efficiency vs. Momentum Flux Ratio, Seebold Data Source: (Seebold et 

al., 2004) .......................................................................................................... 5-7

Figure 5-6. Combustion Efficiency vs. Momentum Flux Ratio ...............................................5-12

Figure 5-7. Combustion Efficiency vs. Momentum Flux Ratio, zoomed (MFR < 3.0; wakedominated mixing regime) ..............................................................................5-13

Figure 5-8. Combustion Efficiency vs. Momentum Flux Ratio, further zoomed (MFR < 0.1) .5-14

Figure 5-9. Combustion Efficiency vs. Power Factor ..............................................................5-16

Figure 6-1. Conditions for Stable Flare Flame ......................................................................... 6-4
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Section 1.0: Introduction Page 1-1

1.0 INTRODUCTION 

Based on a series of flare performance studies conducted in the early 1980s, the EPA 

concluded that properly designed and operated flares achieve good combustion efficiency (e.g., 

greater than 98 percent conversion of organic compounds to carbon dioxide). It was observed, 

however, that flares operating outside “their stable flame envelope” produced flames that were 

not stable or would rapidly destabilize, causing a decrease in both combustion and destruction 

efficiency (Pohl and Soelberg, 1985). To define the stable flame envelope of operating 

conditions, the resulting regulations for flares (i.e., 40 CFR 60.18 and 40 CFR 63.11(b)), 

promulgated in their current form in 1998, included both minimum flare vent gas net heating 

value requirements and a limit on velocity as a function of net heating value. 

Flares are often used at chemical plants and petroleum refineries as a control device for 

regulated vent streams as well as to handle non-routine emissions (e.g., leaks, purges, emergency 

releases); and since the development of the current flare regulations, industry has significantly 

reduced the amount of waste gas being routed to flares. Generally this reduction has affected the 

base load to flares and many are now receiving a small fraction of what the flare was originally 

designed to receive with only periodic releases of episodic or emergency waste gas that may use 

up to the full capacity of the flare. Many flare vent gas streams that are regulated by NESHAP 

and NSPS are often continuous streams that contribute to the base load of a flare; therefore, it is 

critical for flares to achieve good combustion efficiency at all levels of utilization. 

Available data suggest that there are numerous factors that should be considered in order 

to be confident that a flare is operated properly to achieve good combustion efficiency. Factors 

that can reduce the destruction efficiency capabilities of the flare include: 

Over Steaming. Using too much steam in a flare can reduce flare performance. Given 

that many steam-assisted flares are designed to have a minimum steam flow rate in 

order to protect the flare tip, over steaming has resulted, especially during base load 

conditions. In addition, operators acting cautiously to avoid non-compliance with the 

visible emissions standards for flares have liberally used steaming to control any 

potential visible emissions, also resulting in over steaming in some cases. 
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Section 1.0: Introduction Page 1-2

Excess Aeration. Using too much air in a flare can reduce flare performance. Airassisted flares operate similarly to steam-assisted flares; however, air is used as the 

assist-media instead of steam. 

High Winds. A high crosswind velocity can have a strong effect on the flare flame 

dimensions and shape, causing the flame to be wake-dominated (i.e., the flame is bent 

over on the downwind side of a flare and imbedded in the wake of the flare tip). This 

type of flame can reduce flare performance; and potentially damage the flare tip. 

Flame Lift Off. A condition in which a flame separates from the tip of the flare and 

there is space between the flare tip and the bottom of the flame due to excessive air 

induction as a result of the flare gas and center steam exit velocities. This type of 

flame can reduce flare performance; and can progress to a condition where the flame 

becomes completely extinguished. 

The observations presented in this report are a result of the analysis of several 

experimental flare efficiency studies and flare performance test reports. Section 2.0 summarizes 

these data and reports. In addition, scientific information from peer-reviewed studies and other 

technical assessments about flammability, wind, and flame lift off were used in this report. 

Sections 3.0 through 8.0 describe the development of our observations. Section 9.0 provides a 

list of documents referenced in this report. The primary observations are as follows: 

• To identify over steaming situations that may occur on steam-assisted flares, the data 

suggest that the lower flammability limit of combustion zone gas (LFLCZ) is the most 

appropriate operating parameter. Specifically, the data suggest that, in order to maintain 

good combustion efficiency, the LFLCZ must be 15.3 percent by volume or less for a 

steam-assisted flare. As an alternative to LFLCZ, the data suggest that the ratio of the net 

heating value of the combustion zone gas (NHVCZ) to the net heating value of the flare 

vent gas if diluted to the lower flammability limit (NHVLFL) must be greater than 6.54. 

Section 3.0 documents the analysis supporting these observations. 

• To identify excess aeration situations that may occur on air-assisted flares, the data 

suggest that the stoichiometric air ratio (SR) (the actual mass flow of assist air to the 

theoretical stoichiometric mass flow of air needed to combust the flare vent gas) is the 

most appropriate operating parameter. Specifically, the data suggest that, in order to 

maintain good combustion efficiency, the SR must be 7 or less for an air-assisted flare. 

Furthermore, the data suggest that the lower flammability limit of the flare vent gas 

(LFLVG) should be 15.3 percent by volume or less to ensure the flare vent gas being sent 

to the air-assisted flare is capable of adequately burning when introduced to enough air. 

Section 4.0 documents the analysis supporting these observations. 
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Section 1.0: Introduction Page 1-3

• The data suggest that flare performance is not significantly affected by crosswind 

velocities up to 22 miles per hour (mph). There are limited data for flares in winds greater 

than 22 mph. However, a wake-dominated flame in winds greater than 22 mph may affect 

flare performance. The data available indicate that the wake-dominated region begins at a 

momentum flux ratio (MFR) of 3 or greater. The MFR considers whether there is enough 

flare vent gas and center steam (if applicable) exit velocity (momentum) to offset 

crosswind velocity. Because wake-dominated flames can be identified visually, 

observations could be conducted to identify wake-dominated flames during crosswind 

velocities greater than 22 mph at the flare tip. Section 5.0 documents the analysis 

supporting these observations. 

• To avoid flame lift off, the data suggest that the actual flare tip velocity (i.e., actual flare 

vent gas velocity plus center steam velocity, if applicable) should be less than an 

established maximum allowable flare tip velocity calculated using an equation that is 

dependent on combustion zone gas composition, the flare tip diameter, density of the 

flare vent gas, and density of air. Section 6.0 documents the analysis supporting this 

observation. 

• LFLCZ could apply to non-assisted flares (i.e., the LFLCZ must be 15.3 percent by volume 

or less in order to maintain good combustion efficiency). Also, the same operating 

conditions that were observed to reduce poor flare performance associated with high 

crosswind velocity and flame lift off could apply to non-assisted flares. Finally, because 

of lack of performance test data on pressure-assisted flare designs and other flare design 

technologies, it seems likely that the parameters important for good flare performance for 

non-assisted, steam-assisted, and air-assisted flares cannot be applied to pressure-assisted, 

or other flare designs without further information. Section 7.0 documents the analysis 

supporting these observations. 

For purposes of this report, flare vent gas shall mean all gas found in the flare just prior to the 

gas reaching the flare tip. This gas includes all flare waste gas, flare sweep gas, flare purge gas, 

and flare supplemental gas, but does not include pilot gas, assist steam, or assist air. Also, 

combustion zone gas, a term only used for steam-assisted flares, shall mean all gases and vapors 

found just after a flare tip. Combustion zone gas includes all flare vent gas and total steam.
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Section 2.0: Available Flare Test Data Page 2-1

2.0 AVAILABLE FLARE TEST DATA 

This section identifies the data and reports that were used to support our investigation on 

the effects of flare performance with varying levels of steam (for steam-assisted flares); or air 

(for air-assisted flares); and high wind and flame lift off (for both types of flares). 

2.1 Flare Performance Studies and Test Reports 

Specific test run data were extracted from the experimental flare efficiency studies and 

flare performance test reports identified in Table 2-1. A brief summary of each study or report is 

provided in Appendix A. 

Data sets A through C in Table 2-1 are based on experimental data conducted on pilotscale test flares with tip sizes ranging from 3 to 12 inches (for steam-assisted flare designs); and 

1.5 inches (for the air-assisted flare design tested in data set C). Although data set A includes 

experimental data for an air-assisted flare, air flow rates and tip design were held confidential so 

it was not considered in our analysis (see Section 2.7); efforts to acquire this information from 

the authors were not successful. 

Data sets D through I in Table 2-1 are from steam-assisted flares located at various 

chemical and refinery facilities for which EPA Office of Enforcement and Compliance 

Assurance either requested studies pursuant to section 114 of the Clean Air Act, or required the 

study pursuant to a consent decree. With the exception of data set I (and the exception of data 

sets A through C), flare tip sizes (in terms of the effective diameter of the flare tip) for these data 

sets range from 16 to 54 inches (for steam-assisted flare designs). Data set I includes test data for 

a unique flare design and was not considered in our analysis (see Section 2.7). Data set J is based 

on experimental data from a 36-inch steam-assisted flare tip; and a 24-inch air-assisted flare tip. 

In general, the flare test runs were conducted at a high turndown ratio, which means the 

actual flare vent gas flow rate is much lower than what the flare is designed to handle. Data sets 

D through J focus completely on high turndown operating conditions. Data sets A through C 

offer some test data at low turndown ratios, while also offering test data at high turndown ratios. 
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Section 2.0: Available Flare Test Data Page 2-2

Table 2-1. Flare Performance Test Reports

Data 

Set

ID

Test Study 

(or Test Report) 

ID

Title Author Date 

Published

Test 

Method(s) 

Used

A EPA-600/2-83-052 Flare Efficiency Study. McDaniel, M. July 1983 Extractive

B EPA-600/2-84-095

Evaluation of the Efficiency 

of Industrial Flares: Test 

Results.

Pohl, J., et al. May 1984 Extractive

C EPA-600/2-85-106

Evaluation of the Efficiency 

of Industrial Flares: Flare 

Head Design and Gas 

Composition.

Pohl, J. and N. 

Soelberg.

September 

1985 Extractive

D MPC TX

Performance Test of a 

Steam-Assisted Elevated 

Flare With Passive FTIR. 

(Conducted in Texas City, 

TX)

Clean Air 

Engineering, 

Inc.

May 2010 PFTIR

E INEOS

Passive Fourier Transform 

Infrared Technology (FTIR) 

Evaluation of P001 Process 

Control Device at the 

INEOS ABS (USA) 

Corporation Addyston, Ohio 

Facility.

INEOS ABS 

(USA) 

Corporation

July 2010 PFTIR

F MPC Detroit

Performance Test of a 

Steam-Assisted Elevated 

Flare With Passive FTIR. 

(Conducted in Detroit, MI)

Clean Air 

Engineering, 

Inc. 

November 

2010 PFTIR

G

FHR AU

FHR LOU Flint Hills Resources

Clean Air 

Engineering, 

Inc. 

June 2011 PFTIR

H SDP EPF

Shell Deer Park Refining LP 

Deer Park Refinery East 

Property Flare Test Report.

Shell Global 

Solutions (US) 

Inc.

April 2011 PFTIR

I SDP GF

Shell Deer Park Site Deer 

Park Chemical Plant OP-3 

Ground Flare Performance 

Test Report.

Shell Global 

Solutions (US) 

Inc.

May 2011 AFTIR

J TCEQ TCEQ 2010 Flare Study 

Final Report.

Allen, David 

T. and Vincent 

M. Torres.

August 

2011

Extractive,

AFTIR, and 

PFTIR

 
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Section 2.0: Available Flare Test Data Page 2-3

2.2 Flare Vent Gas Constituents 

Table 2-2 identifies the components that were quantified in each experimental flare 

efficiency study and flare performance test report. With the exception of data set E, all test runs 

were based on flares burning propane, propylene, or a mixture of other refinery/petrochemical 

type gases (with some olefins and aromatics). In general, test runs containing only propane or 

propylene (with mixtures of inert) were from data sets A through C, and J; and test runs 

containing a mixture of combustible refinery gases and inerts were from data sets D, and F 

through I. Test runs associated with data set E were conducted while flaring 1,3-butadiene, in 

various mixtures of natural gas and nitrogen at a chemical plant. 

Table 2-2. Flare Vent Gas Constituents by Test Report

Flare Vent Gas Constituent

A: 

EPA 2-83-052 

B: 

EPA 2-84-095 

C: 

EPA 2-85-106 

D: 

MPC TX 

E: 

INEOS 

F: 

MPC Detroit 

G1: 

FHR AU 

G2: 

FHR LOU 

H: 

SDP EPF 

I: 

SDP GF 

J: 

TCEQ 

-----Combustibles-----

1-Butene Υ Υ Υ Υ Υ

1,3-Butadiene Υ Υ Υ Υ Υ Υ

Acetylene Υ Υ Υ Υ

Benzene Υ Υ Υ

Carbon Monoxide Υ Υ Υ Υ Υ

Cis-2-Butene Υ Υ Υ Υ Υ Υ

Ethane Υ Υ Υ Υ Υ Υ

Ethyl Benzene Υ Υ Υ

Ethylene Υ Υ Υ Υ Υ Υ

Hydrogen Υ Υ Υ Υ Υ Υ

Hydrogen Sulfide Υ Υ

Iso-Butane Υ Υ Υ Υ Υ Υ

Iso-Butylene Υ Υ

Methane Υ Υ Υ Υ Υ Υ Υ Υ

Methyl Acetylene Υ

n-Butane Υ Υ Υ Υ Υ Υ

Pentane and Heavier Alkanes Υ Υ Υ Υ Υ Υ

Propane Υ Υ Υ Υ Υ Υ Υ Υ Υ

Propylene Υ Υ Υ Υ Υ Υ Υ Υ

Toluene Υ Υ

Trans-2-Butene Υ Υ Υ Υ Υ

Xylenes Υ

Total Combustibles In Flare Vent Gas 1 1 1 14 2 15 17 16 16 20 3

(1)
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Section 2.0: Available Flare Test Data Page 2-4

Table 2-2. Flare Vent Gas Constituents by Test Report (Continued) 

Flare Vent Gas Constituent

A: 

EPA 2-83-052 

B: 

EPA 2-84-095 

C: 

EPA 2-85-106 

D: 

MPC TX 

E: 

INEOS 

F: 

MPC Detroit 

G1: 

FHR AU 

G2: 

FHR LOU 

H: 

SDP EPF 

I: 

SDP GF 

J: 

TCEQ 

-----Other-----

Nitrogen Υ Υ Υ Υ Υ Υ Υ Υ Υ Υ Υ

Oxygen Υ Υ Υ Υ Υ

Carbon Dioxide Υ Υ Υ Υ Υ Υ

Water Y Y Y 

Total Other Constituents

In Flare Vent Gas 1 1 1 4 1 3 2 3 4 4 1

1 – For data set J, tests were not performed with all three combustibles; tests were either performed with propane and methane, or 

propylene and methane. 

Data sets D, and F through I, used flare vent gas with methane and hydrogen as the 

primary combustibles, and data sets D and F also had significant amounts of olefins in the flare 

vent gas. Table 2-3 shows the range and average of methane, hydrogen, olefins, and nitrogen in 

the flare vent gas for each data set. More specific details regarding flare vent gas constituents are 

discussed in Appendix A. Also, chemical composition for each test run (by test report) used in 

the steam data analysis is discussed in section 3.0 of this report. 

Table 2-3. Minimum, Maximum, and Average Volume Percents of 

Primary Constituents in Flare Vent Gas

Test Report 

% 

Hydrogen 

(Average) 

% 

Methane 

(Average) 

% Total 

Olefins 

(Average) 

% Total Other 

Combustibles 

(Average) 

% 

Nitrogen 

(Average) 

EPA-2-83-052 8.8–100

(46) 

0–9

(54) 

EPA-2-85-106 12–18

(15) 

82–88

(85) 

MPC TX 3.1–24

(14) 

3.8–41

(28) 

11–44

(19) 

7.6–43

(17) 

8.2–35

(21) 

INEOS 0–61

(20) 

2.4–33

(18) 

36–78

(61) 

MPC Detroit 7.0–55

(23) 

16–46

(30) 

4.5–65

(20) 

4.2–24

(13) 

5.7–70

(16) 

FHR AU 13–47

(30) 

29–75

(55) 

0.018–0.47

(0.13) 

3.7–12

(6.5) 

3.5–16

(7.2) 

FHR LOU 20–30

(27) 

55–69

(63) 

1.2–7.7

(3.1) 

3.2–4.2

(3.7) 

0.90–9.0

(2.9) 

SDP EPF 37–62

(52) 

8.5–31

(18) 

0.010–0.49

(0.14) 

11–20

(13.8) 

10–27

(16) 

TCEQ 0–6.9

(4.0) 

0–100

(24) 

0–14.8

(2.2) 

0–83

(70) 
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Section 2.0: Available Flare Test Data Page 2-5

2.3 Steam Injection Rates and Tip Design for Available Flare Test Data 

A steam-assisted flare uses steam at the flare stack or flare tip for purposes including, but 

not limited to, protecting the design of the flare tip, promoting turbulence for mixing or inducing 

air into the flame. Test data are available for nine different steam-assisted flares when 

considering the data sets listed in Table 2-1. 

There are several different ways steam can be injected into the flare waste stream. The 

location of steam injection on each of nine steam-assisted flares varied between the data sets. 

The steam-assisted flares had steam injected through either: nozzles located above the main flare 

tip opening (upper steam), nozzles on an external ring around the top of the flare tip (ring steam), 

a single nozzle located inside the flare prior to the flare tip (center steam), or internal tubes 

interspersed throughout the flare tip (lower steam). The location of steam injection can change 

the nominal flare tip diameter. An effective diameter of the flare tip considers the location of 

steam injection by subtracting the obstructed exit area of the flare tip (i.e., area of any stability 

tabs, stability rings, and steam tubes) from the total exit area of the flare tip. 

Table 2-4 summarizes the design detail (including steam injection location and effective 

flare tip diameter) of each steam-assisted flare tip used for each data set. For most performance 

tests, not only were the steam injection locations different, but also the steam rate varied between 

test runs. For certain data sets and steam injection locations, the steam rate was held constant 

over all test runs. Owners and operators are limited regarding how much they can reduce steam 

flow to the flare tip because steam-assisted flares often have a manufacturer’s minimum steam 

requirement in order to protect the integrity and life of the flare tip. 

Table 2-4. Steam-Assisted Flare Tip Design Detail 

Data 

Set ID

Flare Tip

Manufacturer and

Model Number

Effective 

Diameter1

(inch)

Tip Design and Steam Injection Test Rates

At or Above Tip Inside Tip 

Upper 

(lb/hr) 

Ring 

(lb/hr) 

Lower 

(lb/hr) 

Center 

(lb/hr) 

A

John Zink

STF-S-8 5.86 Varied None None None

 
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Section 2.0: Available Flare Test Data Page 2-6

Table 2-4. Steam-Assisted Flare Tip Design Detail (Continued) 

Data 

Set ID

Flare Tip

Manufacturer and

Model Number

Effective 

Diameter1

(inch)

Tip Design and Steam Injection Test Rates

At or Above Tip Inside Tip 

Upper 

(lb/hr) 

Ring 

(lb/hr) 

Lower 

(lb/hr) 

Center 

(lb/hr) 

B

Energy and Environmental 

Research Corporation; and 

other manufacturer designs2

3, 6, and 12 Varied None None None

C

Unknown Commercial Coanda 

(tulip) Flare Tip 12 140 None None None

D

Callidus Technologies

BTZ-IS3/US-24-C 23.25 None3 None Varied 500

E

John Zink 

EEF-QS-16 16 Varied None None None

F

NAO Inc.

20” NFF-RC 16 None Varied None 300

G

Callidus Technologies

BTZ-US-16/20-C 20 None Varied None 500

G

Callidus Technologies

BTZ-1S3-54C 54 None None Varied 2,890

H

John Zink

EEF-QA-36-C 36 Varied None None Varied

J

John Zink 

EE-QSC-36” 36 Varied None None Varied

1 – The effective diameter of each flare tip was either directly extracted from the test report or calculated from effective area 

reported in the test report. The effective diameter (or area) considers the portion of the area that is occupied by obstructions and 

not available for flare vent gas to flow through; it is determined by subtracting the obstructed exit area of the flare tip (i.e., area of 

any stability tabs, stability rings, and steam tubes) from the total exit area of the flare tip. 

2 – Three simple pipe flare heads were designed and built by Energy and Environmental Research Corporation for testing. A

retention ring was used on the 3-inch flare head during some testing, so the effective diameter would be less than 3 inches during 

those specific tests. In addition to these simple pipe flare heads, three commercial 12-inch diameter pipe flares were also tested; 

these flares were supplied by various flare manufacturers, but the specific design of the flare heads was held confidential. 

3 – The flare tip is equipped with upper steam; however, it was not used during any test runs. (Dickens 2011) 

2.4 Air Injection Rates and Tip Design for Available Flare Test Data 

An air-assisted flare uses assist air at the flare tip for purposes including, but not limited 

to, protecting the design of the flare tip, promoting turbulence for mixing, and inducing air into 

the flame. Test data are available for three different air-assisted flares when considering the data 

sets listed in Table 2-1. However, the experimental data for the air-assisted flare associated with 

data set A were not considered in our analysis (see Section 2.7) because air flow rates and tip 

design were held confidential. Table 2-5 summarizes the design detail of each air-assisted flare 

tip used for each data set. 
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Section 2.0: Available Flare Test Data Page 2-7

Air is injected into the flare waste stream through nozzles located above the main flare tip 

opening. Air injection rates were varied during each test run; ranges of injection rates for each 

air-assisted flare tip tested are provided in Table 2-5. 

Table 2-5. Air-Assisted Flare Tip Design Detail 

Data 

Set ID

Flare Tip

Manufacturer and

Model Number

Effective 

Diameter1

(inch)

Range of Tested Air Injection Rates

(lb/hr)

A

John Zink

STF-LH-457-5 Unknown Unknown

C Unknown 1.5 8,100 to 150,000

J

John Zink 

LHTS-24/60 24 250 to 4,700

1 – The effective diameter of each flare tip was either directly extracted from the test report or calculated from effective area 

reported in the test report. The effective diameter (or area) considers the portion of the area that is occupied by obstructions and 

not available for flare vent gas to flow through; it is determined by subtracting the obstructed exit area of the flare tip (i.e., area of 

any stability tabs, stability rings, and steam tubes) from the total exit area of the flare tip. 

2.5 Flare Test Methods 

Measuring emissions from a flare can be difficult and dangerous because flares lack an 

enclosed combustion chamber, may be elevated, and come in many different designs and sizes. 

With combustion taking place at and above the tip of the flare, the combusted gases are released 

into the atmosphere in any direction given the meteorological conditions and flare vent gas 

velocity that exist at that moment. Although extractive techniques have been used to measure 

emissions from flares, they require placement of a hood-like structure, sampling rake with 

multiple sample ports, or other scheme to ensure representative collection of the flare plume. 

This renders the use of extraction methods for testing industrial flares impractical and relegated 

to research studies, usually on smaller flares. 

Recent technological advances have produced remote sensing instruments capable of 

indicating the presence of combustion products (e.g., carbon dioxide, carbon monoxide, and 

select hydrocarbons) without the safety hazards introduced by physically extracting a sample of a 

flare plume. The remote sensing techniques that have been used on flares discussed in this report 

include: active Fourier transform infrared (AFTIR) and passive Fourier transform infrared 

(PFTIR). The main difference between AFTIR and PFTIR is that AFTIR requires the remote 
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Section 2.0: Available Flare Test Data Page 2-8

sensor be aligned to an artificial light source; whereas PFTIR simply detects infrared radiation 

emitted as heat (i.e., PFTIR uses thermal imaging). Table 2-1 identifies whether each test report 

used extractive, AFTIR, or PFTIR test methods to determine the combustion efficiency of a 

flare. The majority of these reports used PFTIR, which involves using a spectrometer positioned 

on the ground to view hot gases from the flare which radiate spectra that are unique to each 

compound. The PFTIR tests were performed and analyzed by one company, and we are unaware 

of other companies currently using this technique on flares. 

Although AFTIR and PFTIR remote sensing offers an attractive alternative to 

characterize emissions from flares, AFTIR and PFTIR are relatively expensive, new tools that 

currently have no approved methods for universal use on flares. Furthermore, for these remote 

sensing techniques, accurate fitting of measurement and reference spectra for chemical species of 

interest at representative flare temperatures are pivotal in accurately characterizing industrial 

flares. Currently, high temperature spectra are not available for all chemical species that may be 

found in flare vent gas. 

The test report for data set J evaluated the performance of remote sensing technologies 

against extractive techniques. The test report for data set J concluded that the mean difference 

and standard deviation of the reported AFTIR and PFTIR combustion efficiency values increase 

as the reported extractive combustion efficiency values decreases; however, both the AFTIR and 

PFTIR methods actually compare very well to the extractive test results for combustion 

efficiencies reported as 90 percent or greater. For combustion efficiencies reported as 90 percent 

or greater, the test report for data set J states that the mean difference of combustion efficiency 

values averaged 2.5 percentage points different between extractive and AFTIR, and 

2.2 percentage points different between extractive and PFTIR. Based on these conclusions, the 

data collected from all the reports in Table 2-1 were combined and used to support our 

investigation on the effects of flare performance. 
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Section 2.0: Available Flare Test Data Page 2-9

2.6 Combining All Available Test Run Data 

All data sets identified in Table 2-1 were combined into an Excel workbook (see 

Appendix B) and the data were separated by flare type (i.e., steam-assisted versus air-assisted). 

The Excel workbook identifies each specific test run by the exact test condition and run 

identification used in each individual report. For data sets A through C, test run data from tables 

provided within the reports had to be extracted and manually entered into the Excel workbook. 

Raw test data in the form of Excel worksheets were available for data sets D through J, which 

eliminated the need to manually enter data into the Excel workbook for these sets. Each 

individual test run is identified in the “All Run Data” tab of the Excel workbook. 

The amount of detail provided per test run varies between each data set. Also, the 

nomenclature used to describe a variable is different between each data set. For example, data 

set A uses the term “Lower Heating Value (Btu/scf)” when identifying the net heating value of 

the flare vent gas, and data set D uses the term “Vent Gas HV” to describe the same variable. 

Appendix C shows the nomenclature that each data set uses and how it is mapped to one 

common term used in the Excel workbook. 

In some cases, a data set did not explicitly provide a variable, but it could be calculated 

using details from the test reports. For example, for some data sets, in order to calculate a 

volumetric flow rate of the flare vent gas for a specific test run, known values for the mass flow 

rate, molecular weight of the flare vent gas, and a conversion factor for molar volume of an ideal 

gas (379.48 scf/mol) were used. These cases are identified with the words “Calc Eq. D.##” in 

Appendix C; where “##” is the specific calculation methodology number. Each calculation 

methodology is described in Appendix D. 
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Section 2.0: Available Flare Test Data Page 2-10

2.7 Data Removed After Being Considered 

A total of 582 steam-assisted test runs (118 of these runs came from tests performed on a 

steam-assisted flare, but no steam was used during the test) and 111 air-assisted test runs were 

considered in our analysis. However, 270 of the steam-assisted test runs (no steam was used 

during the test for 109 of these runs) and 67 of the air-assisted test runs were removed prior to 

any final analysis. 

Data sets B and I were not used in any of our analysis. Data set B does not provide 

enough data to determine a flare vent gas flow rate, which is critical to calculating the various 

operating limits and parameters we examined. Data set I provides performance testing data for a 

unique flare design that did not operate in the same way as the other flares and the test data did 

not appear consistent. The design is a multistage steam-assisted enclosed ground flare with three 

different stages, which become active at successively higher flows. The flare has 92 horizontallymounted burners (basically a refractory lined steel shell into which 92 raw flare vent gas burners 

discharge). Because the flare tested in data set I is so different from the flare designs in other 

data sets, it is not appropriate to combine and compare its results with the others. 

In addition to excluding data sets B and I in their entirety, Table 2-6 identifies various 

reasons why an individual test run was removed prior to any final analyses described in this 

Report. Each individual test run removed from the analysis is identified in the “Removed Data” 

tab of the Excel workbook (see Appendix B). Each individual air-assisted or steam-assisted test 

run remaining (after removing data due to the reasons described in Table 2-6) is identified in 

either the “Air Data Used All Analysis” or “Steam Data Used All Analysis” tabs of the Excel 

workbook depending on flare tip type. 
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Section 2.0: Available Flare Test Data Page 2-11

2.8 Determination of Combustion Efficiency Representing Good Flare Performance 

The PFTIR testing measures carbon dioxide, carbon monoxide, and hydrocarbons in the 

plume of the flare in order to calculate combustion efficiency. Several current regulations, 

including NSPS and NESHAP, require non-flare control devices to be installed and operated to 

achieve 98 percent destruction efficiency. Therefore, it seemed reasonable to assume that a 98 

percent destruction efficiency represents good performance for flares as well. However, most of 

the flare data was reported in terms of combustion efficiency, making it necessary to estimate a 

combustion efficiency equivalent to 98 percent destruction efficiency as a means for determining 

which test runs (in reviewing the flare test data) demonstrated good performance. 

According to the John Zink Combustion Handbook (Baukal, 2001), destruction efficiency 

is a measure of how much of the hydrocarbon is destroyed; and combustion efficiency is a 

measure of how much the hydrocarbon burns completely to yield carbon dioxide and water 

vapor. Baukal states that combustion efficiency will always be less than or equal to the 

destruction efficiency; and a flare operating with a combustion efficiency of 98 percent can 

achieve a destruction efficiency in excess of 99.5 percent. The relationship between destruction 

and combustion efficiency is not constant and changes with different compounds; however, we 

believe Baukal’s estimation of 1.5% difference is a reasonable assumption. Extrapolating this to 

98 percent destruction efficiency, and also considering the variability in results from the different 

test methods used in this analysis (e.g., PFTIR vs. AFTIR, vs. extractive sampling methods), it 

was determined that a combustion efficiency of 96.5 percent in the flare test data demonstrates 

good flare performance. 
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Section 2.0: Available Flare Test Data Page 2-12

Table 2-6. Criteria To Exclude Data Points 

Criteria Explanation

Test report did not record combustion efficiency for a specific test run. 

It was determined there was not enough information to be able to use 

Test report recorded combustion efficiency as 0% for a specific test run. 

The flare flame was completely snuffed out and the data point was notrends. However, the data point was reviewed to determine conditions combustion (all data points in this category had a LFLCZ greater than 

this report for an explanation on LFLCZ).

Test report recorded that the extraction probe positioning for a specific test run was located in the flame. 

The specific test run was considered invalid because the extraction tecgood sample of the flare plume. 

Test report recorded that the extraction probe positioning for a specific test run was uncertain. 

The specific test run was considered invalid because the extraction tecobtained a good sample of the flare plume. 

Test report recorded a specific test run time as less than 5 minutes. 

The specific test run has too much uncertainty and variability in the rethere were four data points in data set D (i.e., runs 6-1, 8-1, 10-1, and that were reported as having a run time greater than 5 minutes; howevincluded in this removal category because several of the minutes in theither showed zero entries for the PFTIR data. These 4 runs had less tthat were not zero or not affected by wind. 

Test report recorded single test runs and an average of the specific single test runs; the single test runs were removed, but the average was kept. 

The single test runs were considered duplicative because each run wasame conditions. 

Test report recorded a specific test run as smoking. The specific test run was considered out of compliance because visibl

violation with the current regulation, and determined that test runs thacompliance should not be used to establish operating parameters for g

Test report did not record enough information to determine the flare vent gas flow for a specific test run. 

It was determined there was not enough information to be able to use 

Test report recorded that the flare vent gas flow rate of a specific test run was less than 10 pounds per hour. 

The specific test results for these runs are based on an extractive test mshowed very different CE values than other similar runs except that thgas flow rates less than 10 pounds per hour. The extractive test methodetected the waste gas compositions because flow was too low. 

Specific to only data set H, the test report concluded that the "GE Panametric flow readings must be in error when nitrogen concentrations in the SDP EPF line were greater than 30v%". 

The flare vent gas flow rates (above 30% N2) for data set H are reportobservation is limited to ten specific test runs. 

Specific to only data set C, two specific test runs were reported as achieving greater than 99% combustion efficiency, yet the fraction of combustible in the stream was less than 2%. 

The specific test results for these two test runs are based on an extractmay not have correctly detected the waste gas compositions. Given thstream, it is not possible for these two test runs to have achieved greatefficiency (the combustibility of the stream is too low). 
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Section 3.0: Steam and Flare Performance Page 3-1

3.0 STEAM AND FLARE PERFORMANCE 

Steam is used in some flares as a design feature to protect the flare tip. Steam injection 

also promotes smokeless burning in a flare. A key factor to smokeless burning is having enough 

waste gas momentum as it exits the flare burner so that sufficient amounts of air can mix with the 

waste gas and achieve complete combustion. Steam injection is the most common technique for 

adding momentum to low-pressure gases. In addition to adding momentum, steam also provides 

smoke suppression benefits of gas dilution and participates in the chemistry of the combustion 

process (Baukal, 2001). Steam will react with hot carbon particles in soot, removing the carbon 

before it can cool and form smoke. Steam will also react with intermediate combustion products 

to form compounds that readily burn at lower temperatures (Castiñeira, 2006). Using too much 

steam in a flare (over steaming) can result in a flare operating outside its stable flame envelope, 

reducing the destruction efficiency capabilities of the flare. Moreover, the cooling effect from 

use of excessive steam may actually inhibit dispersion of flared gases. In extreme cases, over 

steaming can actually snuff out a flame and allow waste gases to go into the atmosphere 

unburned (Peterson, 2007). 

To identify over steaming situations that may occur on steam-assisted flares, the data 

suggest that the lower flammability limit of combustion zone gas (LFLCZ) is the most appropriate 

operating parameter. Specifically, the data suggest that, in order to maintain good combustion 

efficiency, the LFLCZ must be 15.3 percent by volume or less for a steam-assisted flare. As an 

alternative to LFLCZ, the data suggest that the ratio of the net heating value of the combustion 

zone gas (NHVCZ) to the net heating value of the flare vent gas if diluted to the lower 

flammability limit (NHVLFL) must be greater than 6.54. Section 3.1 documents the analysis 

supporting this observation. Sections 3.2 through 3.4 explain other operating conditions that we 

investigated for good combustion efficiency for steam-assisted flares. 

3.1 Lower Flammability Limit of Combustion Zone Gas for Steam-Assisted Flares 

The lower flammability limit (LFL) is an important chemical property when considering 

combustibility of a gas mixture. The LFL of any mixture is the lowest concentration of that 

mixture in air at which the mixture will burn. Mixtures with a relatively high LFL are less 
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Section 3.0: Steam and Flare Performance Page 3-2

flammable when released to the air than mixtures with a relatively low LFL. A gas mixture with 

a relatively high LFL requires a larger volume of the mixture to burn in a specific volume of air, 

than would a mixture of gases with a relatively low LFL being combusted in that same volume 

of air. The LFL of a mixture is therefore influenced by both the type and amount of chemical 

components (including inerts) present in the gas being burned and is a significant parameter 

when assessing whether a mixture being combusted with an open flame will adequately combust. 

The combustion zone of a steam-assisted flare includes the gas mixture that is created by 

the flare vent gas and the steam that is supplied to the flare. The flare vent gas includes all waste 

gas, sweep gas, purge gas, and supplemental gas, but does not include pilot gas, or assist media. 

Therefore, the combustion zone gas includes all the gases injected into the combustion zone of 

the flare except the pilot gas. See Section 3.1.5 for a discussion of why pilot gas is not included 

in the combustion zone gas. The chemical components and their relative amounts in the 

combustion zone for each test run used in the data analysis for this section can be seen in 

Appendix E by test report. The LFLCZ is the resulting LFL of the mixture that is created by 

combining both the flare vent gas and steam. This parameter was considered as a means to take 

into account the effect of steam on the capability of the flare vent gas to burn. 

Figure 3-1 shows the boundaries of flammability for several different inerts in methane 

and air mixtures (Zabetakis, 1965). The plot is referred to as a Zabetakis plot, or “nose plot”, 

because of its shape and represents the concentrations of fuel (methane in this case), inert and air, 

and the conditions in which combustion will occur. Note that the air concentration is determined 

by subtracting the methane and inert concentrations from 100 percent. The line hitting the y-axis 

near the bottom of the figure is the LFL with no inert added (5% for methane). The upper 

flammability limit (UFL) is the line hitting the y-axis at a higher level (about 15%). The x-axis 

shows the quantity of inert added to the methane and air mixture. The curves show that the UFL 

falls rapidly for mixtures with increasing amounts of inert and the lowest UFL value occurs at 

the maximum amount of inert at which combustion can still be supported. At this amount of 

inert, the UFL has been reduced to be equal to the LFL and combustion can only occur at this 

concentration. An amount of inert above this maximum would render the mixture nonflammable, because there would not be enough air to sustain combustion. 
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Section 3.0: Steam and Flare Performance Page 3-3

Figure 3-1. Zabetakis Nose Plot For Methane And Inert In Air 

At the UFL, at any inert concentration, the combustible-inert-air mixture is oxygen 

limited; there is enough fuel to burn but just enough oxygen to sustain combustion. Any less 

oxygen, the mixture is not flammable. At the LFL, at any inert concentration, there is plenty of 

oxygen with respect to the amount of fuel available. At any combustible concentration less than 

the LFL, the fuel is too lean and the mixture is not flammable. 

Adapted from Evans and Roesler (2011), Figure 3-2 uses a Zabetakis plot for methane in 

air with water as the inert in order to illustrate how a ‘pocket’ of flare vent gas and steam mixture 

interacts with the atmosphere as it travels from the flare tip through the combustion zone. For 

simplicity, Figure 3-2 assumes that the flare vent gas is methane and center steam has been 

added, so the flare vent gas and steam mixture is 25 volume percent methane and 75 volume 

percent water. Just prior to the flare vent gas and steam mixture being released from the flare tip 

into the combustion zone, it contains no oxygen and is therefore above the UFL (this point is not 
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Section 3.0: Steam and Flare Performance Page 3-4

represented on Figure 3-2 because this condition would exist off the graph). However, from the 

point of release at the flare tip, the flare vent gas and steam mixture is quickly diluted by the 

ambient air. At point A on Figure 3-2, the flare vent gas and steam mixture has mixed with air to 

about 15 percent methane, 45 percent water, and 40 percent air; and the gas mixture is still too 

rich to burn. As the gas mixture continues mixing with air it approaches the UFL (point B on 

Figure 3-2). For this example, the UFL of the gas mixture is 8.3 percent methane, 25 percent 

water, and 66.7 percent air. Just beyond point B, the gas mixture is within the UFL and LFL and 

will burn. While the gas mixture is burning it also continues to be diluted until the concentration 

of the gas mixture is just at the LFL (point C on Figure 3-2) of 5.8 percent methane, 17.5 percent 

water, and 76.7 percent air. Using this thought process for other mixtures, it becomes clearer that 

if the flare vent gas and steam mixture are already near the LFL when the gas mixture is released 

into air, the concentration could quickly fall below the mixture’s LFL (point D on Figure 3-2) 

before having a chance to ignite and burn completely; and therefore, it is possible to dilute the 

flammable gas mixture to a mixture with no combustible characteristics. 

Figure 3-2. Time Sequence of Flare Vent Gas Volume 

Moving Through Flammability Region 

Source: (Adapted from Evans and Roesler, 2011) 

Blue: No combustion, above UFL (fuel rich)

Red: Combustion, within flammability region

Yellow: No combustion, below LFL (fuel lean)

Flare

Tip

A

B

D

C

55

Methane, Percent

15

14

13

5

6

7

0 5

Water, Percent

50

8

9

10

11

12

10 15 20 25 30 35 40 45

Just After

Flare Tip (A)

UFL (B)

LFL (C)

Post Combustion (D)

Zabetakis Plot:

Methane in air with water as the inert.

Water
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Section 3.0: Steam and Flare Performance Page 3-5

In addition, the inerts (e.g., nitrogen, carbon dioxide, steam, etc.) of the combustion zone 

gas, which are either included in the flare vent gas or are added as assist steam, compete for 

space with the air and combustible components. The inert gases reduce the flammable range of 

the gas-air mixtures (increase the LFL and decrease the UFL). Some of the heat from the 

combustion reaction is absorbed by the inert gases (Molnarne et. al., 2005), which cools the 

flame and slows the propagation of combustion. This is especially a concern for mixtures with 

higher LFL values. The higher the inert concentration present in the combustion zone gas, the 

higher the LFLCZ. When the higher LFL combustion zone gas mixes with air, it takes less air to 

dilute the gas mixture below the LFL and therefore this diluting effect can happen much more 

quickly. Also, because there are fewer combustible gas molecules in the combustion zone gas, 

more mixing is required to get the combustible components near oxygen molecules. Therefore, 

the higher the LFL of the combustion zone gas, the more difficult it is to achieve and maintain 

good combustion. 

Section 3.1.1 discusses the flare test data and how these data suggest that 15.3 percent 

LFLCZ may be an appropriate threshold (an operating condition representing good combustion). 

Section 3.1.2 describes the method for calculating LFLCZ for mixtures using Le Chatelier’s 

principle and the level of accuracy expected given different flare vent gas compositions. Section 

3.1.3 analyzes data points that had a LFLCZ less than 15.3 percent but a combustion efficiency 

less than 96.5 percent. Section 3.1.4 analyzes the 66 data points with good combustion efficiency 

and a LFLCZ greater than 15.3 percent, and provides possible further categorizing of the LFLCZ to 

provide more chemical component specificity to the LFLCZ limit. Section 3.1.5 provides an 

explanation for excluding pilot gas in these analyses.

3.1.1 Flare Test Data and LFLCZ 

The LFLCZ was examined for 312 steam-assisted flare test runs to determine if there is a 

LFLCZ threshold that would indicate whether good combustion would occur. Figure 3-3 shows 

that as the LFLCZ increases, the combustion efficiency of a flare deteriorates. The vertical dotted 

line in Figure 3-3 marks the threshold where all test runs (47) at a LFLCZ of less than or equal to 

10.0 percent achieved a combustion efficiency of 96.5 percent or greater. However, there are 

several test runs with gas mixtures that have a LFLCZ greater than 10.0 percent that had good 
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Section 3.0: Steam and Flare Performance Page 3-6

combustion efficiencies. Looking at the extent of test points both above and below 96.5 percent 

combustion efficiency reveals that test runs with a LFLCZ of 15.3 percent or less resulted in good 

combustion efficiency for all runs except eleven. At a LFLCZ greater than 15.3 percent, there are 

several test runs with good combustion efficiency, but for every run with good combustion, there 

is at least one run with bad combustion. For example, using test runs with a LFLCZ of 20.0 

percent or less, there are 30 test runs with good combustion and 33 test runs with combustion 

efficiency less than 96.5 percent. 

The vertical solid line in Figure 3-3 shows a threshold where most test runs at a LFLCZ of 

less than or equal to 15.3 percent achieved a combustion efficiency of 96.5 percent or greater. 

This level maximizes the number of test runs (105) with a combustion efficiency of 96.5 percent 

or greater that are below 15.3 percent LFLCZ while minimizing the number of test runs (11) that 

have a combustion efficiency less than 96.5 percent. There are 11 test runs (highlighted in red on 

Figure 3-3) with a LFLCZ between 10.0 and 15.3 percent that did not achieve a combustion 

efficiency of 96.5 percent or greater; further detail about the 11 test runs is provided in Section 

3.1.3. Also evident from Figure 3-3 are several test runs that achieved good combustion but all 

had a LFLCZ greater than 15.3 percent. There are 197 data points with a LFLCZ greater than 15.3 

percent and 61 of these points have a combustion efficiency of greater than 96.5 percent. These 

data points are discussed further in Section 3.1.4. 
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Section 3.0: Steam and Flare Performance Page 3-7

Figure 3-3. Combustion Efficiency vs. LFLCZ

3.1.2 The Le Chatelier Principle 

Le Chatelier’s principle was used to estimate lower flammability limits of the flare 

combustion zone gas for each test run. Le Chatelier’s principle for determining flammability 

limits for mixtures uses the reciprocal of the volume weighted average over the LFL of the 

individual compounds for estimating the gas mixture’s LFL (Equation 3-1). 

∑=

















=

n

j j

a

LFL

LFL

j

1

100

χ

 (Eq. 3-1)

Where: 

LFLa = Lower flammability limit of any gas mixture ‘a’, volume %. 

a = Placeholder for any gas mixture. 

n = Number of components in gas mixture ‘a’. 

j = Individual pure component in gas mixture ‘a’. 

40%

50%

60%

70%

80%

90%

100%

5% 10% 15% 20% 25% 30% 35% 40% 45% 50%

Combustion Efficiency (%)

LFLCZ (%)

10.0%

15.3%
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Section 3.0: Steam and Flare Performance Page 3-8

xj = Concentration of individual pure component j in gas mixture ‘a’, volume %. 

LFLj = Lower flammability limit of individual pure component j, volume %. 

Le Chatelier’s equation was originally limited to binary mixtures of combustible gases 

(Coward et al., 1919). Coward et al. (1919) generalized the equation for mixtures containing 

more than two combustible gases. Coward and Jones (1952) presented a method for using Le 

Chatelier’s equation when inert gases are included in the mixture. This method entails assigning 

amounts of inert gas to the combustible gases, identifying a LFL for each combustible–inert 

combination from previous experimental data, and then using Le Chatelier’s equation to 

calculate the LFL for the full mixture. For example, a mixture of 50:20:30 percent 

nitrogen/hydrogen/methane could be broken up into two mixtures, one with 60:40 percent 

nitrogen/hydrogen (30% nitrogen and 20% hydrogen from original mixture, or total of 50% of 

the original mixture) and one with 40:60 nitrogen/methane (20% nitrogen and 30% methane 

from original mixture, or total of 50% of the original mixture). Experimental data of nitrogen and 

hydrogen, and nitrogen and methane mixtures, such as the data graphed in Appendix F, can then 

be used to determine each partial mixture’s LFL. Looking at Appendix F, the LFL of a 

60:40 percent nitrogen/hydrogen mixture is 6.5 percent; and 12.5 percent for a 40:60 percent 

nitrogen/methane mixture. Then using Le Chatelier’s equation, a LFL of 8.6 percent is 

determined for the full mixture: 

8.6%

12.5

50

6.5

50

100

=

+

 (Eq. 3-2)

This method is not practical for determining the LFL of flare vent gas and steam mixtures 

because flare vent gases can change at any moment and consist of a wide variety of chemicals. 

The LFL data for mixtures with nitrogen (or carbon dioxide) and a combustible is limited; and 

this type of information is essentially unavailable for steam. The results of the method also vary 

significantly depending on how the combustibles and inerts are paired together. 

Karim et al. (1985) describes another way to address mixtures with inerts, which is to 

include the inert in Le Chatelier’s equation and assume a LFL for the inert gas of infinity. This 

zeros out the term for the inert in the equation since one divided by infinity is zero. The LFLCZ in 

Figure 3-3 assumes inerts have an infinite LFL (or a reciprocal of the LFL equal to zero). 
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Section 3.0: Steam and Flare Performance Page 3-9

However, assuming infinity for all inerts as Karim et al. (1985) suggests does not address 

the difference in behavior for the different inerts. Molnarne et al. (2005) describes a method for 

taking into account non-nitrogen inerts. This method is used in the International Standards 

Organization (ISO) Standard 10156 and was originally presented by Besnard (1996). In this 

method, a nitrogen equivalency factor is used that varies depending on both the type of inert(s) 

and combustible compound(s) in the mixture. Molnarne, et al. describes the possible values for 

nitrogen equivalency for the specific combustion gas and inert combinations. Equation 3-3 

incorporates this method into Le Chatelier’s equation. More details about inerts and further 

explanation for including an equivalency adjustment to correct for different inert behavior are 

provided in Appendix G. Appendix G also provides a more detailed discussion about LFL and 

the accuracy of calculated LFL values with respect to varying amounts of inerts and specific 

combustible gases that can be contained in a mixture. Appendix G does not specifically discuss 

the flare test data, but includes important concepts and considerations helpful in understanding 

the discussions contained in Sections 3.1.3 and 3.1.4. 

( )( ) ( )( )

2 2 2 2

1 1

100

, ,

1

e H O H O e CO CO

n

j j

a

N x N x

LFL

LFL

j

− − − −

















=

∑=

χ

 (Eq. 3-3)

Where: 

LFLa = Lower flammability limit of any gas mixture ‘a’, volume %. 

a = Placeholder for any gas mixture (e.g., flare combustion zone gas, ‘cz’; flare vent 

gas with center steam, ‘vgcs’; or flare vent gas ‘vg’). 

n = Number of components in gas mixture ‘a’. 

j = Individual pure component in gas mixture ‘a’. 

xj = Concentration of individual pure component j in gas mixture ‘a’, volume %. 

LFLj = Lower flammability limit of individual pure component j, volume %. The LFL of 

individual compounds are based on experimental data and are provided in a 

Bureau of Mines Bulletin entitled: “Flammability Characteristics of Combustible 

Gases and Vapors” (Zabetakis 1965). All inerts, including nitrogen, are assumed 

to have an infinite lower flammability limit (e.g., LFLN2 = ∞).

Ne,H2O = Coefficient of nitrogen equivalency for water relative to nitrogen, unitless. See 

Table 3-1. 
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Section 3.0: Steam and Flare Performance Page 3-10

xH2O = Concentration of water in gas mixture ‘a’, volume fraction. 

Ne,CO2 = Coefficient of nitrogen equivalency for carbon dioxide relative to nitrogen, 

unitless. See Table 3-1. 

xCO2 = Concentration of carbon dioxide in gas mixture ‘a’, volume fraction. 

Table 3-1. Recommended Values of Coefficient of Nitrogen Equivalency for Water and 

Carbon Dioxide Relative to Nitrogen 

Combustible Component i in Flare Vent Gas Ne,H2O Ne,CO2

Methane 1.87 2.23 

Ethane 1.40 1.87 

Propane 1.51 1.93 

Ethylene 1.68 1.84 

Propylene 1.36 1.92 

Hydrogen 1.35 1.51 

All Other Combustibles 1.50 1.87 

Source: Molnarne et al. (2005) 

Figure 3-4 is the same plot as Figure 3-3, except the LFLCZ has been adjusted for nitrogen 

equivalency. All data points have essentially shifted to the right; and the amount each test run 

shifted is dependent on the amount of water and carbon dioxide in the combustion zone gas. By 

applying this adjustment, one of the 11 test runs highlighted (in red) in Figure 3-3, is no longer 

between the 10.0 and 15.3 percent LFLCZ region. (See Section 3.1.3 about these test runs.) The

differences between Figures 3-3 and 3-4 do not appear significant; however, from the perspective 

of assuring that a LFLCZ of 15.3 percent or less specifies a flare with good combustion, it is 

important as will become more evident throughout this section and Appendix G. 
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Figure 3-4. Combustion Efficiency vs. LFLCZ Adjusted for Nitrogen Equivalency 

3.1.3 Specific Test Data Not Fitting the Trend 

The test runs highlighted red in Figures 3-3 through 3-8 (Figures 3-5 through 3-8 are 

shown in Section 3.1.4) do not seem to fit the trend found for the other test runs that good 

combustion efficiency is achieved with a LFLCZ less than 15.3 percent depicted by the vertical 

blue line in each figure. These test runs have a LFLCZ of less than 15.3 percent but they did not 

have a combustion efficiency of greater than 96.5 percent. The specific test reports associated 

with these test runs were examined. These test runs could be in error or have a higher variability 

than other test runs, the estimate of the LFL could be incorrect, or the LFL may not be an 

appropriate indicator of combustion efficiency. Given the previous discussion in Sections 3.1.1 

and 3.1.2, the LFL appears to be a good indicator of combustion efficiency (assuming wind and 

velocity of the flare vent gas are within reasonable limits as discussed in Sections 5 and 6 of this 

report). Therefore, this section investigates the possibility of inaccuracies in testing or calculating 

the LFL (the appropriateness of using Le Chatelier’s equation for the mixture). Note that four of 

40%

50%

60%

70%

80%

90%

100%

5% 10% 15% 20% 25% 30% 35% 40% 45% 50%

Combustion Efficiency (%)

LFLCZ (%)

10.0%

15.3%
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Section 3.0: Steam and Flare Performance Page 3-12

the test runs are within the statistical uncertainty for the test method. As determined by Clean Air 

Engineering (2010a), the measured combustion is ±1.5% at a confidence level of 99%. 

These data points (Table 3-2) were scrutinized against all of the other test runs to see if 

they were unique results that indicated an issue with the accuracy in the test results, or whether 

they are indicating a previously unseen trend. Figure 3-3 presents 11 data points in red; however, 

Figures 3-4 though 3-8 only indicate 10 data points that do not fit the trend. The eleventh data 

point (Flint Hills Resources LOU-C) shown in Figure 3-3 has a LFLCZ greater than 15.3 percent 

after the nitrogen equivalency adjustment was made. Therefore, the data point follows the trend 

of the majority of the data and is not discussed further in this section. Section 3.1.3.1 describes 

any test run specific information revealed by the review of the test reports. Section 3.1.3.2 

includes observations from examining these test runs and comparing them to other test runs. 

Section 3.1.3.3 provides conclusions regarding these test runs. 

 Table 3-2. Test Run Detail for 11 Data Points with 

LFLCZ < 15.3% but Combustion Efficiency < 96.5%

Test Site Condition 

ID Run ID 

Combustion 

Efficiency 

(%) 

%Combustible

In CZ 

(%) 

%Inert

In CZ 

(%) 

%Olefin

In CZ 

(%)a

%H2

In CZ 

(%) 

MPC TX B 8-2

b

95.8 28.0 72.0 5.06 5.72

MPC TX B 9-2 89.1 27.0 73.0 4.89 5.56

MPC TX D 7-2

b

96.0 20.5 79.5 8.89 1.36

MPC Detroit B 8-1 94.6 28.0 72.0 4.20 10.5

MPC Detroit C 4-1 92.7 21.9 78.1 14.8 1.67

MPC Detroit C 4-2 93.7 22.3 77.7 14.5 1.83

MPC Detroit C 5-1 93.1 19.8 80.2 13.2 1.74

MPC Detroit C 5-2 91.6 19.4 80.6 12.3 1.70

MPC Detroit D 8-1

b

96.2 24.8 75.2 2.75 11.6

FHR LOU LOU-A 2.0(1)b95.9 35.0 65.0 2.83 10.6

FHR LOU LOU-C 3.0(2)

b, c

95.8 30.8 69.2 0.50 9.03

a Olefins include: 1-Butene, 1,2-Butadiene, 1,3-Butadiene, Cis-2-Butene, Ethylene, Iso-Butylene, Propadiene, 

Propylene, and Trans-2-Butene. 

b

 These test runs are within the testing uncertainty of 96.5%; they each had combustion efficiencies less than 1.5% 

below 96.5%. 

c After making the nitrogen equivalency adjustment for this run, the LFLCZ is greater than 15.3% and therefore 

follows the trend for the majority of the data points (see Figures 3-3 and 3-4). 
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Section 3.0: Steam and Flare Performance Page 3-13

3.1.3.1 Review of Test Reports 

The MPC TX test report (Clean Air Engineering, Inc., 2010a) discusses a potential 

problem with some reported data. The report says that when the wind came from a certain 

direction steam may have interfered with the pilot due to a “shaping steam ring failure.” The 

report states that “the shaping steam ring failure caused the shaping pressure to be unbalanced, 

with one side of the concentric ring starved for steam. As the wind blew towards the area of 

unbalance, the flame sheared off of the flare tip and was extinguished.” For all of the test runs 

where this was noted in the raw data, the efficiency was more variable on a minute-by-minute 

basis during the testing than for the other runs. This could mean that the efficiencies for test runs 

affected by the flare steam ring failure and wind could be biased low. All of the MPC TX data 

points included in Table 3-2 (i.e., B 8-2, B 9-2, and D 7-2) were affected by this issue. The raw 

data showed that for test run B 8-2, there were 9 out of 12 minutes tested affected by this issue; 

for test run B 9-2, there were 10 out of 10 minutes affected; and for test run D 7-2, there were 4 

out of 10 minutes affected. 

The MPC Detroit test report (Clean Air Engineering, Inc., 2010b) stated that if the wind 

was “blowing directly away from the PFTIR at more than 5 mph,” the PFTIR view of the flare 

plume could be obscured causing inaccuracies in the testing. The wind was blowing in the 

“wrong” direction (for the PFTIR analysis) for all of the MPC Detroit Condition C runs and for 

the B 8-1 test run. However, the wind speeds were not excessive during these runs (average wind 

speed for these test runs was less than 3.6 miles per hour). The test run data revealed no issues 

with the MPC Detroit test run D 8-1 and the Flint Hills Resources (LOU) test run LOU-A 2.0(1). 

3.1.3.2 Comparison of Data Runs 

The data points listed in Table 3-2 can be divided, generally, into three groups when 

considering the combustion zone gas: one with olefin and hydrogen both approximately equal to 

5 percent; one with high hydrogen content (9 to 11.6%) and low olefin content (0.5 to 2.8%); and 

the third group with relatively high olefin content (8.9 to 14.8%) and low hydrogen content (less 

than 2%). All of the available test data were reviewed to determine if there were any similar test 

runs to those in Table 3-2. The information found for each group is discussed below. 
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Section 3.0: Steam and Flare Performance Page 3-14

Olefin and Hydrogen Approximately Equal 

The test runs at MPC TX numbered B 8-2 and B 9-2 did not fit the trend, and both had 

olefin and hydrogen content at about 5 percent. Looking at the full set of test data, there were 

several other test runs with a similar composition yet had combustion efficiencies greater than 

96.5 percent and a LFLCZ greater than 15.3 percent. Table 3-3 shows all the test runs where 

olefin and hydrogen concentrations were approximately equal; data with an olefin to hydrogen 

ratio from 0.81 to 1.23 were reviewed. As evident by Table 3-3, there were several data points 

with very similar concentrations of olefin, hydrogen, other combustibles and inerts. “Other 

combustibles” include primarily methane and other alkanes. 

High Hydrogen and Low Olefin 

The test runs MPC Detroit B 8-1 and D 8-1, and Flint Hills test LOU-A 2.0(1) had a 

higher concentration of hydrogen than olefin by about 2.5 to 4 times. Table 3-4 provides the 

information for all test runs with relatively high hydrogen and low olefin concentrations (ratio of 

olefin to hydrogen of 0.12 to 0.55). All the test runs in Table 3-4 have less olefin than hydrogen 

by a factor of between 2 and 8 and there is inert content of between 49 and 88 percent. Test runs 

MPC Detroit D 8-1 and Flint Hills LOU-A 2.0(1) have measured combustion efficiencies that 

are within the statistical uncertainty of 96.5%. Also, Table 3-4 shows several other test runs that 

have similar conditions and all had combustion efficiencies of 96.5% or greater. These test runs 

do not appear to indicate a unique trend. 

The MPC Detroit test run B 8-1 is similar to several test runs in Table 3-4 but it does 

have a relatively high inert concentration (72%) for the group. The MPC Detroit test runs A 4-2 

and A 3-1 both have similar amounts of inert at 74 and 72 percent respectively. However, there 

is more hydrogen and olefin in the B 8-1 run than these runs. 
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Section 3.0: Steam and Flare Performance 

Table 3-3. Olefin and Hydrogen Approximately Equal 

Test Site Condition Run CE 

%OlefinCZ

(%) 

%H2-

CZ

(%) 

Ratio 

%Olefin-CZ / 

%H2-CZ

Other 

Combustibles 

in CZ 

(%) 

Inert 

in CZ 

(%) 

Fraction 

Combustible 

in CZ 

Steam 

Fraction 

of CZ 

MPC TX A11 4-1 95.5% 2.84 3.53 0.81 9.78 83.86 0.16 0.77 MPC TX C 3-2 99.2% 6.34 7.82 0.81 24.33 61.51 0.38 0.55 MPC TX A11 12-1 61.6% 1.38 1.69 0.82 4.35 92.57 0.07 0.89 MPC TX LTS 2-5 99.4% 6.91 8.41 0.82 23.94 60.75 0.39 0.53 MPC TX LTS 2-6 97.2% 7.40 8.89 0.83 19.20 64.51 0.35 0.54 MPC TX A11 6-1 92.5% 2.26 2.71 0.83 8.64 86.39 0.14 0.81 MPC TX B 4-2 99.3% 6.89 8.09 0.85 23.59 61.44 0.39 0.54 MPC TX LTS 2-2 99.3% 6.88 8.07 0.85 23.56 61.49 0.39 0.54 MPC TX A11 2-1 95.6% 3.62 4.25 0.85 11.82 80.31 0.20 0.70 MPC TX LTS 2-3 99.0% 6.96 8.01 0.87 23.60 61.42 0.39 0.54 MPC TX B 5-2 99.0% 6.16 7.09 0.87 20.88 65.86 0.34 0.59 MPC TX B 10-2 96.7% 4.69 5.39 0.87 15.94 73.99 0.26 0.69 MPC TX B 6-1 98.4% 5.54 6.31 0.88 18.54 69.61 0.30 0.64 MPC TX B 6-2 98.5% 5.53 6.29 0.88 18.63 69.56 0.30 0.64 

MPC TX B 9-2 89.1% 4.89 5.56 0.88 16.58 72.97 0.27 0.68 

MPC TX B 7-2 97.9% 5.27 5.98 0.88 17.87 70.88 0.29 0.65 MPC TX B 7-1 97.8% 5.32 6.03 0.88 17.77 70.88 0.29 0.65 

MPC TX B 8-2 95.8% 5.06 5.72 0.88 17.18 72.05 0.28 0.67 

MPC TX B 9-1 97.6% 4.89 5.53 0.89 16.39 73.20 0.27 0.68 MPC TX B 8-1 97.6% 5.14 5.80 0.89 17.15 71.92 0.28 0.66 MPC TX B 5-1 98.7% 6.21 7.01 0.89 20.86 65.92 0.34 0.59 

MPC Detroit LTS 8 98.7% 7.12 6.78 1.05 25.83 60.28 0.40 0.56 MPC Detroit LTS 7 99.3% 10.77 10.20 1.06 21.65 57.38 0.43 0.53 MPC Detroit LTS 5 99.1% 8.76 7.92 1.11 23.56 59.76 0.40 0.55 MPC Detroit LTS 4 98.9% 8.03 7.24 1.11 22.86 61.87 0.38 0.57 MPC Detroit LTS 1 98.9% 8.70 7.39 1.18 23.64 60.27 0.40 0.55 

MPC TX C 3-1 98.9% 3.94 3.20 1.23 20.84 72.02 0.28 0.66 

Pink = CE less than 96.5%, and LFLCZ less than 15.3% (Test Runs from Table 3-2). 

Gray = LFLCZ greater than 15.3%. 
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Section 3.0: Steam and Flare Performance 

Table 3-4. High Hydrogen and Low Olefin 

Test Site Condition Run CE 

%OlefinCZ

(%) 

%H2-CZ

(%) 

Ratio 

%Olefin-CZ / 

%H2-CZ

Other 

Combustibles 

in CZ 

(%) 

Inert 

in CZ

(%) 

Fraction 

Combustible 

in CZ 

Stea

Fract

of C

FHR LOU LOU-B 1.0 (2) 99.4% 1.49 12.38 0.12 36.84 49.29 0.51 0.4

MPC Detroit D 3-1 97.5% 2.04 16.49 0.12 11.07 70.39 0.30 0.6

FHR LOU LOU-A MIN (3) 98.8% 1.59 11.47 0.14 22.98 63.96 0.36 0.6FHR LOU LOU-A 8.5 (1) 78.8% 0.62 3.86 0.16 7.89 87.63 0.12 0.8FHR LOU LOU-A 3.0 (2) 99.1% 1.30 8.01 0.16 18.18 72.50 0.27 0.7FHR LOU LOU-A 4.0 (2) 98.6% 1.06 6.41 0.17 14.57 77.95 0.22 0.7FHR LOU LOU-A 5.0 (2) 96.5% 0.89 5.30 0.17 11.92 81.90 0.18 0.8

MPC Detroit D 5-1 91.6% 1.85 11.05 0.17 7.49 79.61 0.20 0.7

FHR LOU LOU-A 6.0 (1) 88.6% 0.86 5.11 0.17 10.41 83.61 0.16 0.8

MPC Detroit D 6-1 98.4% 3.43 20.30 0.17 13.17 63.10 0.37 0.6

FHR LOU LOU-A 6.0 (2) 95.1% 0.79 4.67 0.17 10.29 84.25 0.16 0.8FHR LOU LOU-A 5.0 (1) 95.5% 1.01 5.93 0.17 12.13 80.94 0.19 0.8FHR LOU LOU-A 2.0 (3) 98.6% 1.66 9.53 0.17 22.45 66.36 0.34 0.6

MPC Detroit D 4-1 88.6% 1.84 10.31 0.18 7.25 80.60 0.19 0.7

FHR LOU LOU-A 4.0 (1) 96.5% 1.25 6.91 0.18 14.19 77.64 0.22 0.7FHR LOU LOU-A 3.0 (1) 97.2% 1.59 8.09 0.20 16.89 73.43 0.27 0.7FHR LOU LOU-A MIN (2) 99.2% 2.10 10.05 0.21 26.28 61.57 0.38 0.5

MPC Detroit D 9-1 98.6% 3.84 16.98 0.23 14.80 64.37 0.36 0.6

FHR LOU LOU-A 2.0 (2) 98.6% 1.65 7.29 0.23 25.65 65.41 0.35 0.6

MPC Detroit D 8-1 96.2% 2.75 11.61 0.24 10.43 75.22 0.25 0.7

FHR LOU LOU-A MIN (1) 97.5% 3.01 12.04 0.25 24.27 60.67 0.39 0.5

FHR LOU LOU-A 2.0 (1) 95.9% 2.83 10.57 0.27 21.56 65.04 0.35 0.6

MPC Detroit D 7-1 83.5% 2.19 8.06 0.27 8.28 81.47 0.19 0.8MPC Detroit A 3-2 98.6% 2.69 8.77 0.31 18.26 70.28 0.30 0.6MPC Detroit A 2-2 98.5% 3.23 9.73 0.33 19.84 67.20 0.33 0.6MPC Detroit D 10-1 81.8% 2.12 5.40 0.39 7.27 85.21 0.15 0.8

MPC Detroit B 8-1 94.6% 4.20 10.46 0.40 13.30 72.04 0.28 0.6

MPC TX LTS 2-4 98.7% 5.10 11.02 0.46 16.24 67.64 0.32 0.5

MPC Detroit A 1-1 99.0% 4.51 9.00 0.50 23.26 63.23 0.37 0.5MPC Detroit A 7-1 95.2% 2.84 5.62 0.50 13.87 77.67 0.22 0.7MPC Detroit A 8-1 92.0% 2.31 4.53 0.51 12.09 81.07 0.19 0.7MPC Detroit A 4-2 97.7% 3.18 6.18 0.51 16.46 74.18 0.26 0.6MPC Detroit B 1-1 98.9% 9.07 17.61 0.52 27.10 46.21 0.54 0.4MPC Detroit A 3-1 98.0% 3.48 6.65 0.52 18.13 71.74 0.28 0.6MPC Detroit A 9-1 81.3% 1.93 3.66 0.53 9.90 84.51 0.15 0.8MPC Detroit A 6-2 96.7% 2.88 5.38 0.54 14.74 76.99 0.23 0.7MPC Detroit A 2-1 97.7% 3.98 7.25 0.55 19.91 68.86 0.31 0.6

Pink = CE less than 96.5%, and LFLCZ less than 15.3% (Test Runs from Table 3-2). Gray = LFLCZ greater than 15.3%.
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Section 3.0: Steam and Flare Performance Page 3-17

Higher Olefin and Low Hydrogen 

There are five test runs from Table 3-2 that are in this category and they are shown in 

Table 3-5 in the pink rows along with all other test runs where the olefin concentration was 

greater than the hydrogen concentration (except for those runs where the olefin concentration 

was just above hydrogen which are shown in Table 3-4). Only one of the five test runs is within 

the statistical uncertainty of 96.5% combustion efficiency (MPC TX D 7-2). Table 3-5 shows 

several test runs with good combustion efficiencies that match the conditions of the five test 

runs. However, none of the data points had the same or higher amounts of inerts than these five 

test runs had. These five runs have the highest amounts of inerts (between 78 to 81%) of any of 

the test runs with similar compositions in Table 3-5. 

3.1.3.3 Conclusions on Ten Data Points 

The data points in Table 3-2 were expected to achieve good combustion efficiency 

because their LFLCZ were determined to be less than 15.3 percent using Le Chatelier’s equation 

and adjusting for nitrogen equivalency. However, these test runs achieved combustion 

efficiencies between 89.1 and 96.2 percent. These data points were reviewed to determine if 

these anomalies were an indication that the 15.3 percent LFLCZ was incorrect and should be 

adjusted to better fit the data or if the points were just a few inaccurate test results in a group of 

over 300 test points. By comparing the compositions and results of other test runs with the tests 

in Table 3-2 it can be speculated whether the tests are inaccurate or a clue to something else. 

For the two test runs where the olefin and hydrogen concentration are approximately 

equal (MPC TX B 9-2 and B 8-2), there are several examples of test runs of almost identical 

conditions where a combustion efficiency of greater than 96.5 percent was achieved. Also, both 

of these test runs had the majority of the testing time affected by the steam ring failure. It appears 

likely that these two test runs had inaccurate results and that their results should not impact the 

15.3 percent LFLCZ value. 
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Section 3.0: Steam and Flare Performance Page 3-18

Table 3-5. Higher Olefin and Low Hydrogen 

Test Site Condition Run CE 

%OlefinCZ

(%) 

%H2-CZ

(%) 

Ratio 

%Olefin-CZ / 

%H2-CZ

Other 

Combustibles 

in CZ 

(%) 

Inert 

in CZ

(%) 

Fraction 

Combustible 

in CZ 

St

Fra

of 

MPC TX D 4-1 98.4% 11.03 1.83 6.03 12.63 74.51 0.25 0

MPC TX D 7-2 96.0% 8.89 1.36 6.54 10.23 79.51 0.20 0

MPC TX D 4-2 98.2% 11.01 1.68 6.56 12.20 75.11 0.25 0

MPC Detroit C 3-1 96.8% 16.62 2.50 6.64 6.95 73.93 0.26 0MPC Detroit C 2-1 98.4% 18.59 2.61 7.12 8.72 70.07 0.30 0MPC Detroit C 2-2 99.1% 20.97 2.93 7.16 9.46 66.63 0.33 0

MPC Detroit C 5-2 91.6% 12.33 1.70 7.26 5.37 80.60 0.19 0

MPC Detroit C 1-2 99.4% 26.15 3.47 7.54 12.70 57.67 0.42 0

MPC Detroit C 5-1 93.1% 13.18 1.74 7.58 4.93 80.15 0.20 0MPC Detroit C 4-2 93.7% 14.54 1.83 7.93 5.89 77.74 0.22 0

MPC Detroit C 3-2 97.4% 17.53 2.15 8.17 7.11 73.22 0.27 0

MPC Detroit C 4-1 92.7% 14.78 1.67 8.83 5.48 78.06 0.22 0

MPC TX E 3-3 98.3% 12.32 1.10 11.20 13.69 72.89 0.27 0MPC TX E 5-2 97.5% 10.68 0.94 11.33 11.91 76.46 0.24 0MPC TX E 4-3 97.7% 11.52 1.01 11.38 12.86 74.60 0.25 0MPC TX E 5-1 96.8% 11.44 0.86 13.37 11.50 76.21 0.24 0MPC TX E 3-1 98.6% 12.84 0.93 13.74 12.76 73.47 0.27 0MPC TX E 4-1 98.1% 12.12 0.85 14.31 11.80 75.23 0.25 0

Pink = CE less than 96.5%, and LFLCZ less than 15.3% (Test Runs from Table 3-2). 

Gray = LFLCZ greater than 15.3%.
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Section 3.0: Steam and Flare Performance Page 3-19

The three test runs that had relatively high hydrogen content, all about 10.5 percent, and 

lower olefin content [FHR LOU-A 2.0(1), MPC Detroit B 8-1 and D 8-1] were compared with 

other test runs. The FHR LOU-A 2.0(1) and MPC Detroit D 8-1 were both very similar test runs, 

and they both had combustion efficiencies within the statistical uncertainty of the method. There 

were also several test runs that were very close matches to these runs that all achieved good 

combustion. These test runs do not appear to signify a separate trend or raise doubt regarding the 

15.3 percent LFLCZ threshold. The MPC Detroit test run B 8-1 can be nearly matched to a few 

test runs, including some with more inert that still achieved greater than 96.5 percent combustion 

efficiency; indicating that these test runs are likely due to inaccurate test results. 

The remaining test runs in Table 3-3 (MPC TX D 7-7 and MPC Detroit C 4-1, 4-2, 5-1, 

and 5-2) are all very similar to each other, except that the MPC TX run has a much lower olefin 

content and a higher amount of other combustibles (and it is the only run in this group with a 

combustion efficiency within the statistical uncertainty of 96.5%). There are several similar runs 

with good combustion efficiency shown in Table 3-5, but they all have lower amounts of inert 

than those runs shown in Table 3-2 — 78 to 81 percent inert for the test runs with a combustion 

efficiency less than 96.5 percent compared to 58 to 76 percent inert for runs achieving greater 

than 96.5 percent combustion efficiency. Because there are five test runs all similar in 

composition and results, it seems that there could be issues with one or more components acting 

to inhibit the combustion to some degree; thereby interfering with the accuracy of the Le 

Chatelier’s equation. Azatyan et al. (2005a, 2005b, and 2007) showed that small amounts of 

propylene can react with hydrogen and form chain terminating products that inhibit some of the 

combustion reactions (see Appendix G for further discussion). However, these test runs have 

larger amounts of olefins and a small amount of hydrogen; it is unknown whether inhibition is 

possible with this situation. Also, the test runs contain other combustibles, such as methane and 

propane. Therefore, it is difficult to make any conclusion. The flare steam analysis was 

performed on 312 data points; all but 17 had some amount of olefin in the gas mixture, 203 test 

runs with at least 1 percent olefin, and 198 test runs with both hydrogen and olefin. If there are 

chain terminating reactions from olefins or another compound, the inhibition affected very few 

data points with respect to a LFLCZ of 15.3 percent. 
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Section 3.0: Steam and Flare Performance Page 3-20

3.1.4 Data Points with Good Combustion and High LFLCZ

Out of the 312 data points available for this analysis, 164 test runs achieved a combustion 

efficiency of 96.5 percent or greater; Figure 3-4 shows 66 of these runs (or 40%) have a LFLCZ

greater than 15.3 percent and achieve greater than 96.5 percent combustion efficiency. This 

means that for approximately 40 percent of the test runs that achieved a combustion efficiency of 

96.5 percent or greater, and if LFLCZ were being used to monitor flare performance, owners 

would need to reduce the amount of inert in the combustion zone (i.e., reduce the amount of 

assist steam being used), or add additional combustible gas in order to lower the LFLCZ below 

the threshold of less than 15.3 percent. The test data were scrutinized to see if there were any 

commonalities or trends that could explain why the 66 test points still achieved good combustion 

efficiency with a higher than expected LFLCZ. Approximately half (36) of the 66 test points are 

from INEOS, EPA, and TCEQ data sets and have only one or two flammable gases. Gas 

compositions for these tests included nitrogen, propane, propylene, 1,3-butadiene, and natural 

gas. The other 30 test points are from Shell, Marathon, and Flint Hill Resources data sets and 

have 10 or more combustible constituents in the flare vent gas. The primary gas components for 

these tests were hydrogen and methane. 

Similar to the data points discussed in Section 3.1.3, the 66 data points do not fit the 

overall trend because the calculation of the LFLCZ is inaccurate, the combustion efficiency 

measurements are inaccurate, or the flammability is not an appropriate parameter to represent 

flare performance or it does not completely explain the performance of a flare. This section 

explores two explanations for the 66 data points: (1) The potential that some of the steam may 

not enter the combustion zone due to wind effects, and (2) The LFLCZ representing good 

combustion may vary with the composition of the flare vent gas. 

3.1.4.1 Partial Steam Contribution 

One factor that may contribute to overestimated LFLCZ values is that the calculation 

methodology for LFLCZ that was used in this analysis assumes all assist steam is incorporated 

into the combustion zone gas. However, it is possible that some of the steam delivered from the 

upper and/or lower ring steam nozzles may actually not participate in mixing with the 
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Section 3.0: Steam and Flare Performance Page 3-21

combustion zone because wind can blow the flame away from a portion of the nozzles causing 

steam to be shot directly into the atmosphere without ever hitting the combustion zone. With this 

in mind, the LFLCZ values could be overestimated for some test runs depending on the steam 

injection configuration and the wind direction and velocity. 

The location of steam injection on each of the steam-assisted flares varied between test 

data sets (see Table 2-4 in Section 2.0). Table 3-6 shows the breakdown of steam use for the 

66 test runs that have a LFLCZ greater than 15.3 percent and achieve greater than 96.5 percent

combustion efficiency when the LFLCZ is adjusted for nitrogen equivalency. Sixty-one out of the 

66 test runs that have a LFLCZ greater than 15.3 percent (and achieve greater than 96.5% 

combustion efficiency), were using steam that potentially did not mix completely with the other 

combustion zone gases. This means that for these test runs, the LFLCZ that was calculated in this 

analysis may actually be too high, which would explain why these test runs achieved greater than 

96.5 percent combustion efficiency but the calculation shows a LFLCZ greater than 15.3 percent. 

Table 3-6. Breakdown Of Steam Use For The 66 Test Runsa 

Steam Configuration Test Run Count Test Reports 

Upper 18 EPA-600/2-83-052; EPA-600/2-85-106; 

TCEQ; and INEOS 

Upper and Center 26 TCEQ and SDP EPF 

Lower and Center 12 FHR LOU and MPC TX 

Ringb and Center 5 MPC Detroit and FHR AU 

No Steam 5 EPA-600/2-83-052 

a Those test runs that have a LFLCZ greater than 15.3% and achieve greater than 96.5% combustion efficiency when 

the LFLCZ is adjusted for nitrogen equivalency. 

b The specific test reports do not clarify whether the ring steam is located in an upper or lower position at the flare 

tip. 

In a recent report, Marathon Petroleum Company reviewed their performance testing 

results against wind conditions and flare plume position to characterize how wind speed affects 

the amount of steam mixing with the combustion zone gas (Cade et al., 2010). The report derived 

a “Steam Contribution Factor” calculation methodology to predict conditions when non-mixing 

steam is occurring; however, the methodology is not practical to apply to all flares because it 

relies heavily on flare design, specific test data and flame images. 
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Section 3.0: Steam and Flare Performance Page 3-22

The amount of steam that mixes in the combustion zone at a given moment would depend 

on the wind speed and direction and would vary moment-by-moment. Because of the difficulties 

in determining how much upper, lower, and ring steam mixes with the combustion zone, and 

since the upper, lower, and ring steam is designed to be injected into the combustion zone gas; it 

is logical to assume that in the majority of situations complete mixing occurs and includes all the 

steam, regardless of where it is injected, when calculating the LFLCZ. However, to estimate the 

possible effect of this assumption, a worst case analysis was conducted assuming no upper, 

lower, and ring steam participated in mixing with the combustion zone (i.e., only center steam 

was used in the LFLCZ calculation). The LFLCZ for 37 of the 66 test runs, with only center steam 

considered, computed to less than 15.3 percent. Therefore, even if it were known that the wind 

was great enough on the day of testing to remove all upper, lower, and ring steam from the 

combustion zone, this would only resolve just over half (56%) of the test runs with good 

combustion efficiency but high LFLCZ (greater than 15.3%). 

3.1.4.2 More Specific LFLCZ Thresholds 

Another detail that could be considered when establishing a LFLCZ threshold on the entire 

312 test runs is the commonalities in composition of flare vent gas. Figures 3-5 through 3-8 are 

plots of the same 312 test runs shown in Figures 3-3 and 3-4; however, the test runs are divided 

into one of the four figures depending on the test run composition. The lower flammability limit 

of the combustible portion of the flare vent gas (LFLVG,C) was used to define the composition 

categories. This scheme was developed based on the observation that the smaller the LFL of a 

pure component, the greater the amount of inert that chemical can withstand and still propagate a 

flame (Kondo, 2006a). Therefore, to allow the greatest amount of inert (or highest LFLCZ), the 

data were divided into categories associated with their LFL without inert. 

Four categories were developed and are shown in Table 3-7. For example, if the 

combustible portion of the flare vent gas for a test run consisted of 30 percent propane (and the 

other 70% was nitrogen), the test run would fall into category A because the LFLVG,C is equal to 

that of propane, or 2.1 percent. Table 3-7 also shows the potential LFLCZ threshold for each 

category based on Figures 3-5 through 3-8. The test runs that do not fit the trends in these 

figures, as having not achieved a combustion efficiency of 96.5 percent or greater (but are 
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Section 3.0: Steam and Flare Performance Page 3-23

meeting the potential LFLCZ threshold), are highlighted in red and have been discussed in 

Section 3.1.3. 

Table 3-7. Potential LFLcz Thresholds based on LFLVG,C

Category ID LFLVG,C

(%) 

Potential Threshold 

for LFLCZ 

(%) 

A ≤ 2.4 ≤ 24.0 

B >2.4 and ≤ 2.8 ≤ 18.5 

C >2.8 and ≤ 3.8 ≤ 16.0 

D >3.8 ≤ 14.5 

Figure 3-5. Combustion Efficiency vs. LFLCZ for Category A Test Runs (see Table 3-7) 

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40% 50% 60% 70%

Combustion Efficiency (%)

LFLcz (%)

0.24
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Section 3.0: Steam and Flare Performance Page 3-24

Figure 3-6. Combustion Efficiency vs. LFLCZ for Category B Test Runs (see Table 3-7) 

Figure 3-7. Combustion Efficiency vs. LFLCZ for Category C Test Runs (see Table 3-7) 

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40% 50% 60% 70%

Combustion Efficiency (%)

LFLcz (%)

0.185

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40% 50% 60% 70%

Combustion Efficiency (%)

LFLcz (%)

0.16
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Section 3.0: Steam and Flare Performance Page 3-25

Figure 3-8. Combustion Efficiency vs. LFLCZ for Category D Test Runs (see Table 3-7) 

By categorizing the data into these four categories, one test run (in Category D) no longer 

meets the LFLCZ threshold (≤14.5%) and has a combustion efficiency of 96.5 percent or better. 

However, 22 test runs now meet the LFLCZ threshold (in Table 3-7), and have an efficiency of 

96.5 percent or greater. This reduces the number of test runs that have good combustion 

efficiency, yet do not meet the LFLCZ threshold, from 66 to 45 test runs. All but two of the 22 

test runs that meet the thresholds in Table 3-7 and have good combustion efficiency are in 

Category A. Generally, Category A includes test runs with only 1 or 2 combustible components 

and Categories C and D have several constituents and most include hydrogen and methane. 

Categorizing did not improve the results for Categories C and D. The test runs that make up a 

given category generally are from specific test reports. For example, SDP EPF test runs are all in 

Category C, while Category D is almost entirely FHR AU or FHR LOU test runs. The SDP EPF, 

FHR AU and FHR LOU test runs are for mixtures with higher amounts of hydrogen. 

Combustible mixtures that have components that enhance the combustion of other constituents 

would also be a reason that a mixture with a higher calculated LFLCZ could have a combustion 

efficiency of greater than 96.5 percent. As discussed in Appendix G, hydrogen can enhance the 

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40% 50% 60% 70%

Combustion Efficiency (%)

LFLcz (%)

0.145
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Section 3.0: Steam and Flare Performance Page 3-26

combustion of methane and this enhancement is not reflected in the calculated LFL. However, an 

understanding of the role that other constituents play in the combustion efficiency (such as 

amounts of olefin, propane, etc., and better quantification of this hydrogen-methane effect) are 

needed before this could be considered. 

Also, there could be issues with the testing of the combustion efficiencies for these runs. 

Information was not available to judge the accuracy for the FHR test runs. The SDP EPF test 

report cited the possibility of inaccurately high combustion efficiency measurements during 

times of unstable combustion conditions. However, the test runs were not identified in the report 

where this may have occurred. Without further testing it is difficult to determine why these 

45 test runs had good combustion efficiencies with LFLCZ greater than 15.3 percent. 

3.1.5 Excluding Pilot Gas

Our analyses have not included the heat value or flammability properties of the pilot gas. 

The pilot gas serves as a source of ignition for the combustion zone gas. The pilot gas does not 

significantly contribute to the combustion zone gas mixture, or the final output or operating limit 

value of the LFLCZ. In theory, the pilot gas is already fully combusted at the combustion zone, 

and therefore would contribute a small amount of inert (i.e., carbon dioxide and water vapor) to 

the combustion zone gas mixture. However, the amount of inert that could be attributed to the 

pilot gas is expected to be negligible — relative to the volume of flare vent gas that would 

actually participate in the combustion zone gas mixture (e.g. for 304 of the 312 test runs, the 

pilot gas contributed less than 3% of the volumetric flow rate of the flare vent gas; and the pilot 

gas contributed to less than 1% for 224 of these test runs). 

In addition, pilot gas is needed to maintain the pilot flame which is a requirement of 

current flare regulations. Including the pilot gas in the calculation of the combustion zone 

parameters, does not add any accuracy in identifying flares with (or without) good combustion. 

Also, by leaving pilot gas out of the calculations, the equations are simplified. 
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Section 3.0: Steam and Flare Performance Page 3-27

3.2 Combustible Gas Concentration in the Combustion Zone 

If there are not enough combustible components in the combustion zone, then the flare 

vent gas will not burn, or will burn inadequately, resulting in lower combustion efficiency. The 

addition of assist steam dilutes the combustible components of the flare vent gas, which can 

make it more difficult for the flare vent gas to burn. To simplify consideration of the operating 

conditions for good combustion, we looked at the minimum quantity of combustibles present in 

the combustion zone that would maintain good combustion for the chemical compositions tested. 

Figure 3-9 shows the combustion efficiency versus the fraction of combustibles in the 

combustion zone gas (CCZ). The vertical dotted line in Figure 3-9 identifies the threshold where 

all, but one test run with a fraction of combustibles in the combustion zone gas of 0.32 or greater 

achieved a combustion efficiency of 96.5 percent or greater. All data sets had test runs with a 

fraction of combustibles in the combustion zone gas at greater and less than 0.32, except TCEQ. 

Therefore, this trend was observable for a variety of flare vent gas compositions and constituents. 

All TCEQ test runs had enough inert (either nitrogen or steam) to cause the fraction of 

combustibles in the combustion zone gas to be less than 0.32. Although Figure 3-9 illustrates a 

strong trend (because all test runs except for FHR LOU-A 2.0(1) performed with at least 

32 percent combustibles in the combustion zone gas achieved greater than 96.5 percent 

combustion efficiency – see Section 3.1.3 for a discussion about test run FHR LOU-A 2.0(1)), 

there are a significant number of test runs that achieved good combustion efficiency that did not 

have at least 32 percent combustibles in the combustion zone gas. For these situations, a more 

detailed review of the operating conditions would have to be considered, such as consideration of 

the LFLCZ. 
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Section 3.0: Steam and Flare Performance Page 3-28

Figure 3-9. Combustion Efficiency vs. CCZ

3.3 Heat Content Based Limit for Steam-Assisted Flares 

A white paper (Evans and Roesler, 2011) was reviewed that describes a method to define 

flare operating conditions using the net heating value and the lower flammability limit. The 

method is nearly identical mathematically (observations with test data reveal some differences 

that are discussed later in this section) to the reciprocal of the LFLCZ parameter discussed in 

Section 3.1 (provided you apply “combustion factors” as described in the white paper); however, 

it may be easier for flare owners and operators to understand and implement because it uses a 

familiar parameter (net heating value). The white paper uses Le Chatelier’s principle and 

provides a methodology for determining the ratio of the net heating value of the combustion zone 

gas (NHVCZ) to the net heating value of the flare vent gas if diluted to the lower flammability 

limit (NHVVG-LFL). 

Figure 3-10 shows that as the ratio of NHVCZ to NHVVG-LFL decreases, the combustion 

efficiency of a flare deteriorates. The vertical dotted line in Figure 3-10 marks the threshold 

0%

20%

40%

60%

80%

100%

0.000 0.100 0.200 0.300 0.400 0.500 0.600

Combustion Efficiency (%)

CCZ (volume fraction) 0.32
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Section 3.0: Steam and Flare Performance Page 3-29

where all test runs with a ratio of NHVCZ to NHVVG-LFL greater than or equal to 10 achieved a 

combustion efficiency of 96.5 percent or greater. The vertical solid line in Figure 3-10 shows 

another threshold where most test runs with a ratio of NHVCZ to NHVVG-LFL greater than or equal 

to 6.5 achieved a combustion efficiency of 96.5 percent or greater. There are 12 test runs with a 

ratio of NHVCZ to NHVVG-LFL between 6.54 and 10.0 that did not achieve a combustion 

efficiency of 96.5 percent or greater. Eleven of these 12 runs are those listed in Table 3-2. This 

observation was to be expected since this ratio is mathematically equal to the inverse of the 

LFLCZ; however, this method has two more test runs (having a NHVCZ to NHVVG-LFL ratio 

greater than 6.54) that do not achieve a combustion efficiency of 96.5 percent or greater. Out of 

the 312 data points available for this analysis, 164 test runs achieved a combustion efficiency of 

96.5 percent or greater; Figure 3-10 shows 61 of these runs (or 37%) have a NHVCZ to NHVVGLFL ratio less than 6.54 and achieve greater than 96.5 percent combustion efficiency. 

Figure 3-10. Combustion Efficiency vs. Ratio of NHVCZ to NHVVG-LFL 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

2.00 4.00 6.00 8.00 10.00 12.00 14.00 16.00 18.00

Combustion Efficiency (%)

Ratio of NHVCZ to NHVVG-LFL

6.54

10.0
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Section 3.0: Steam and Flare Performance Page 3-30

Although the ratio of NHVCZ to NHVVG-LFL is mathematically equivalent to the inverse of 

the LFLCZ, there are several test runs where differences are observed between the two 

parameters. Figure 3-11 is a log-log plot of LFLCZ versus the ratio of NHVCZ to NHVVG-LFL for 

each of the 312 test runs. The data points marked in red have the greatest difference between the 

two parameters. 

Figure 3-11. LFLCZ vs. Ratio of NHVCZ to NHVVG-LFL

One possible explanation for the observed differences between the two methods is the 

y-axis is a LFLCZ value adjusted for nitrogen equivalency. A “nitrogen equivalency” adjustment 

was incorporated for determining the ratio of NHVCZ to NHVVG-LFL. However, because the 

procedure considers the LFL of the flare vent gas (as well as the NHVCZ) and not the actual LFL 

of the combustion zone gas, the adjustment is hardly noticeable. The adjustment applies to any 

water and carbon dioxide that is found in the flare vent gas (so it excludes the steam in the 

combustion zone). The data points marked in red have some of the greatest differences between 

the LFL not adjusted for nitrogen equivalency and the adjusted LFL. This means that these data 

y = 1.2432x-1.08

R2 = 0.9952

0.010

0.100

1.000

1.00 10.00 100.00

LFLCZ

NHVCZ / NHVVG-LFL
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Section 3.0: Steam and Flare Performance Page 3-31

points had significant amounts of water and/or carbon dioxide in the combustion zone relative to 

nitrogen. 

3.4 Other Operating Parameters Considered for Steam-Assisted Flares 

Numerous other operating parameters were examined to study the effects of steam on 

flare combustion efficiency; however, these did not prove as promising as the LFLCZ operating 

parameter or ratio of NHVCZ to NHVVG-LFL when compared to the test data. A brief review of 

each of these parameters or concepts is provided in Sections 3.4.1 and 3.4.2. 

3.4.1 Net Heating Value 

The general provisions of 40 CFR 60.18(b) and 40 CFR 63.11(b) limit the net heating 

value of the gas being combusted in a steam-assisted flare (or air-assisted flare) to 300 Btu/scf or 

greater. Figure 3-12 graphs combustion efficiency versus net heating value of the flare vent gas 

(NHVVG) for the steam-assisted flares tested; it shows that poor combustion efficiency can occur 

with flare vent gases that have high heat content (greater than 300 Btu/scf — depicted by the 

vertical blue line) and good combustion may occur in flare vent gas with lower heat content. 
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Section 3.0: Steam and Flare Performance Page 3-32

Figure 3-12. Combustion Efficiency vs. NHVVG

The figure shows a general trend, as the heat content increases, the combustion efficiency 

increases. However, there are a significant number of test runs with flare vent gas heat content 

much higher than 300 Btu/hr that do not achieve good combustion efficiency. Using the NHVVG 

as an indicator of good combustion ignores any effect of steaming. Therefore, to incorporate 

steaming, a net heating value of the combustion zone gas was calculated to include the assist 

steam. 

Figure 3-13 presents the combustion efficiency versus NHVCZ. A much clearer trend 

forms when the steam is factored into the operating parameter; as the NHVCZ increases, the 

combustion efficiency increases. For each test run, the NHVCZ is a lower value than the NHVVG

shown in Figure 3-12 because the steam dilutes the gas mixture; each cubic foot of combustion 

zone gas has less heating value than a cubic foot of the flare vent gas. 

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Section 3.0: Steam and Flare Performance Page 3-33

Figure 3-13. Combustion Efficiency vs. NHVCZ

If Figure 3-13 were used to determine appropriate operating conditions for flares, 

operators would need to maintain a NHVCZ between approximately 300 and 350 Btu/scf to 

ensure good combustion. However, evident from Figure 3-13, there are a significant number of 

test runs that are showing good combustion, but have a NHVCZ less than 300 Btu/scf. Out of the 

312 data points available for this analysis, 164 test runs achieved a combustion efficiency of 

96.5 percent or greater; 80 of these runs (or 49%) have a NHVCZ less than 300 Btu/scf and 

achieve greater than 96.5 percent combustion efficiency. 

In addition, NHV may not be the best operating parameter for determining flare 

performance because the net heating value of a mixture can vary significantly at certain constant 

lower flammability limits (and vice versa); Figure 3-14 shows this phenomenon. Figure 3-14 

shows the NHVCZ versus the LFLVG of each test run. Based on Figure 3-14, for example, flare 

vent gas with a LFL of 0.04, may actually have a net heating value of anywhere between 

500 Btu/scf and 2,000 Btu/scf; flare vent gas with a net heating value of 400 Btu/scf, may 

actually have a LFL of anywhere between 0.06 and 0.25. In fact, Azatyan, et al (2002) observes 

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Section 3.0: Steam and Flare Performance Page 3-34

the same relationship stating that “...in many cases, different gases, even with identical 

combustion heats and amounts of oxygen consumed per mole of fuel, differ essentially in 

concentration limits of flame propagation.” Azatyan supports this statement by providing an 

example of how ethylene and ethanol have similar heats of combustion (~1400 kJ/mol), but also 

have very different upper and lower flammability limits (ethylene: 32.0 to 3.1%; and ethanol: 

19.0 to 4.3%). 

Figure 3-14. NHVVG vs. LFLVG 

3.4.2 Steam Ratios 

The amount of steam added to flares with respect to the amount of flare vent gas was 

considered as a possible flare operating condition. Figure 3-15 shows the relationship between 

combustion efficiency and actual steam-to-vent gas ratio (S/VG) by weight as it varied across the 

different data sets. In general, as the S/VG increases, the combustion efficiency deteriorates. 

However, there appears to be multiple trends that are defined by the different data sets and test 

run series within a data set. The MPC TX data show several data points above 90 percent 

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Section 3.0: Steam and Flare Performance Page 3-35

combustion efficiency until a S/VG of about 3 and then the combustion efficiency begins to 

decline more quickly with increasing S/VG. Many of the MPC-Detroit and FHR AU data points 

seem to follow this trend. The FHR LOU test runs seem to show an ability to accept more steam 

without adversely affecting the combustion efficiency, since a S/VG of 4 yields a 90 percent 

combustion efficiency for most test runs. The Shell, EPA, and some TCEQ data appear to follow 

the FHR LOU trend. Much of the TCEQ data; however, show a rapid degradation of combustion 

efficiency, dropping below 90 percent with S/VG greater than 0.5. 

Figure 3-15. Combustion Efficiency vs. S/VG by weight 

Figure 3-16 is a similar plot to Figure 3-15, except that it shows the S/VG by volume 

instead of by mass. Putting the steam ratio in units of volume consolidates the perceived trends. 

The MPC TX, MPC Detroit, FHR AU and FHR LOU test data seem to all follow the same 

general trend. The TCEQ, EPA, and some of the Shell and INEOS data points seem to define 

two additional trends; one of rapid decline in combustion efficiency with increasing S/VG by 

volume and one that is much slower. 

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Steam-to-vent-gas ratio by weight (S/VGwt) [lb steam/lb vent gas]

TCEQ

INEOS

MPC TX

MPC Detroit

FHR (AU)

FHR (LOU)

Shell (EPF)

EPA 2-83-052

EPA 2-85-106
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Section 3.0: Steam and Flare Performance Page 3-36

Figure 3-16. Combustion Efficiency vs. S/VG by volume 

Figure 3-17 presents the steam-to-hydrocarbon ratio (S/HC) by volume. This figure does 

seem to show further improvement in the variation between different data sets and test run series 

within a data set. However, this parameter does not take into account the amount of inert 

components in the flare vent gas. 

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Steam-to-vent-gas ratio by volume (S/VGvol) [scf steam/scf vent gas]

TCEQ

INEOS

MPC TX

MPC Detroit

FHR (AU)

FHR (LOU)

Shell (EPF)

EPA 2-83-052

EPA 2-85-106
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Section 3.0: Steam and Flare Performance Page 3-37

Figure 3-17. Combustion Efficiency vs. S/HC by volume 

The variation in trends in Figures 3-15 through 3-17 is likely due to the variation in flare 

vent gas components and compositions between each data set. For example, the TCEQ test data 

consists of flare vent gas streams of propylene or propane with nitrogen (and some natural gas 

used as a supplemental gas), while the INEOS tests were conducted with flare vent gas 

containing various mixtures of 1,3-butadiene, natural gas, and nitrogen. It appears that S/VG or 

S/HC alone cannot fully address over steaming because neither considers the variability of 

chemical properties within the flare vent gas. As discussed in Section 3.1, unlike S/VG or S/HC, 

the LFLCZ is a good candidate for evaluating flare performance because it actually considers the 

variability of chemical properties. 

Another parameter investigated was the ratio of actual total mass flow rate of steam to a 

recommended total mass flow rate of steam for a steam-assisted flare. This steam ratio parameter 

was based on Table 11 of the American Petroleum Institute (API) Standard 521 for pressurerelieving and depressuring systems. Table 11 of the API document suggests amounts of steam 

that should be injected into certain gases being flared in order to promote smokeless burning. 

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Steam-to-hydrocarbon ratio by volume (S/HCvol) [scf steam/scf HC]

TCEQ

INEOS

MPC TX

MPC Detroit

FHR (AU)

FHR (LOU)

Shell (EPF)

EPA 2-83-052

EPA 2-85-106
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Section 3.0: Steam and Flare Performance Page 3-38

Because these steam recommendations in API 521 are intended for only smoke suppression, and 

not meant to be correlated to combustion efficiency performance, this steam ratio parameter was 

not investigated any further. In addition, the suggested injection steam rates are provided for only 

certain flared gases that are most common to petroleum and natural gas industries (they do not 

cover a complete list of the chemicals that may be flared). 
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Section 4.0: Air and Flare Performance Page 4-1

4.0 AIR AND FLARE PERFORMANCE 

Similar to how steam is used (See Section 3.0 for steam discussion), air injection can 

promote smokeless burning in a flare. Air adds momentum and turbulence to the combustion 

zone, which improves mixing and reduces the possibility of smoke formation. Because the 

additional air is induced into the waste gas, it also provides oxygen necessary to augment 

smokeless capacity (Castiñeira and Edgar, 2006). However, just as in the use of steam, using too 

much air in a flare (excess aeration) can actually result in a flare operating outside its stable 

flame envelope, decreasing the combustion efficiency (Zeeco Company, 2003). Assist air can 

dilute the flare vent gas, making the flare vent gas too lean to burn in the combustion zone. 

To identify excess aeration situations that may occur on air-assisted flares, the data 

suggest that the stoichiometric air ratio (SR) (the actual mass flow of assist air to the theoretical 

stoichiometric mass flow of air needed to combust the flare vent gas) is the most appropriate 

operating parameter. Specifically, the data suggest that, in order to maintain good combustion 

efficiency, the SR must be 7 or less for an air-assisted flare. Furthermore, the data suggest that 

the lower flammability limit of the flare vent gas (LFLVG) should be 15.3 percent by volume or 

less to ensure the flare vent gas being sent to the air-assisted flare is capable of adequately 

burning when introduced to enough air. This section documents the analysis supporting these 

observations. Note: the LFLVG is different from the LFLCZ described in Section 3.1.2 of this 

report because the LFLVG does not consider any assist media; however, in general, the 

methodology described in Section 3.1.2 for determining LFL still applies. 

4.1 Stoichiometric Air Ratio 

The stoichiometric air ratio (SR) is the ratio of actual mass flow of total assist air to the 

theoretical stoichiometric mass of air needed to combust the flare vent gas (this relationship is 

shown in Equation 4-1). The SR increases as more assist air is added to the flare vent gas. 

 

	


(Eq. 4-1) 
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Section 4.0: Air and Flare Performance Page 4-2

Where: 

SR = Ratio of actual mass flow of total assist air to the theoretical stoichiometric mass 

of air needed to combust the flare vent gas, unitless. 

mAir = Actual mass flow of total assist air, pounds per hour, lb/hr. 

mStoic = Theoretical stoichiometric mass of air needed to combust the flare vent gas, lb/hr. 

For flares that utilize air as an assist medium, the stoichiometric amount of assist air 

(mStoic) represents the theoretical amount of air needed to obtain complete combustion of a fuel 

gas comprised of combustible compounds. This stoichiometric amount is based on the specific 

chemical composition and quantity of the combustible compounds combusted in the flare. The 

stoichiometric amount of assist air needed is determined from only the combustible portion of 

the flare vent gas. While some flare vent gas streams may include non-combustible compounds, 

such as carbon dioxide or nitrogen, the determination of the stoichiometric air amount ignores 

the contributions of these compounds to the total flow and composition of the fuel gas stream. 

The theoretical stoichiometric molar amount of air required for combustion varies 

between different combustible hydrocarbons, and can be determined by a balanced combustion 

reaction equation. Equations 4-2 and 4-3 below are balanced combustion reaction equations for 

methane and propane, respectively. 


  2    2
 (Eq. 4-2)


  5  3  4
 (Eq. 4-3) 

As seen in Equation 4-2, two moles of oxygen are required for the complete combustion 

of one mole of methane. Similarly, Equation 4-3 shows that five moles of oxygen are required to 

completely combust one mole of propane. Assuming ambient air contains approximately 

21 percent oxygen, and using the molecular weight for methane, propane, and air, the 

stoichiometric ratio of air to fuel can be calculated as shown in Equation 4-4 for methane and 

Equation 4-5 for propane. For methane, the stoichiometric ratio of air to fuel is 17.1 pounds of 

air per pound of methane; and for propane, it is 15.6 pounds of air per pound of propane. These 

values are used with the actual amount of assist air and fuel to calculate the SR. All data sets and 
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Section 4.0: Air and Flare Performance Page 4-3

the methodologies used in calculating the SR are outlined in Section 2.0, and Appendices C and 

D of this report. 

  

   !

"

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&.  

"

. ( #$%

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"

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) (  !



*. ( #$%

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(Eq. 4-4) 

+  

  , -

"

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"

. ( #$%

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"

  , -

 ( , -



+.) ( #$%

 ( , -

(Eq. 4-5) 

4.2 TCEQ Test Data 

The Texas Commission on Environmental Quality (TCEQ) 2010 flare study final report 

(Allen and Torres, 2010) examined the effects of SR on an air-assisted flare’s combustion 

efficiency. The results of the TCEQ study show that, in general, as more assist air is added to the 

flare vent gas and the SR increases, the flare combustion efficiency deteriorates. Figure 4-1 plots 

35 test runs from the TCEQ study that clearly shows this trend. 

Figure 4-1. Combustion Efficiency vs. SR (using TCEQ data) 

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Section 4.0: Air and Flare Performance Page 4-4

4.3 Other Test Data 

The trend displayed in Figure 4-1 (as more assist air is added to the flare vent gas and the 

SR increases, the flare combustion efficiency deteriorates) also exists with the nine specific airassist test runs extracted from the EPA-600/2-85-106 study (Pohl and Soelberg, 1985). 

Figure 4-2 shows this trend. It was not possible to include, in Figure 4-2, the 13 air-assisted test 

runs extracted from the EPA-600/2-83-052 study (McDaniel, 1983) because the air flow and air 

flow velocity were considered proprietary information and not included in the test report. 

Figure 4-2. Combustion Efficiency vs. SR (using EPA-600/2-85-106 data) 

4.4 Analysis of Stoichiometric Air Ratio 

Figure 4-3 overlays Figures 4-1 and 4-2. It shows that although the trend exists for both 

data sets, the threshold (or cliff) where combustion efficiency begins to deteriorate is much 

different between the two data sets. The two studies were reviewed carefully in an attempt to 

identify differences between the EPA-600/2-85-106 and TCEQ test runs. The test data from the 

EPA-600/2-85-106 study used mixtures of propane and nitrogen, while TCEQ used propylene or 

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Section 4.0: Air and Flare Performance Page 4-5

propane with natural gas as a supplemental gas. The air-assisted flare from the EPA-600/2-85-

106 study had a 1.5 inch diameter and the air-assisted flare from the TCEQ tests had a diameter 

of 24 inches. Most of the test runs from the EPA-600/2-85-106 study did not use a pilot during 

testing but the TCEQ testing did include pilots. In addition, the flare vent gas volumetric or mass 

flow rates for the EPA-600/2-85-106 test runs were not reported, so assumptions were made 

using the reported flare vent gas velocity to determine flow rates. 

Figure 4-3. Combustion Efficiency vs. SR (combined TCEQ and EPA-600/2-85-106 data) 

The EPA-600/2-85-106 test report noted that smaller flares did not produce data that 

were comparable to larger flares (greater than 3 inches). Gogolek et al. (2010b) supports this 

finding by concluding that results from pipe flares smaller than 3 inches are not scalable to larger 

diameter pipe flares (i.e., 6 to 12 inch diameter pipe flares). This conclusion seems logical with 

respect to all of the available air assist data, especially if the various sized flares are operated 

with similar SR values. There is less distance across the volumetric flare vent gas flow of a small 

flare at the tip exit than the distance across the volumetric flare vent gas flow of a larger flare at 

the tip exit; therefore, more assist air would penetrate and mix with the flare vent gas if a smaller 

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TCEQ

EPA-600/2-85-106 
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Section 4.0: Air and Flare Performance Page 4-6

flare were used. Because of these differences between the EPA-600/2-85-106 and TCEQ test 

runs, the EPA-600/2-85-106 data were removed and only the TCEQ data were analyzed. 

Figure 4-4 is the same as Figure 4-1, but with the axes narrowed to a shorter range. The 

vertical green line identifies the SR threshold where all test runs achieve a combustion efficiency 

of 96.5 percent or greater. (See Section 2.8 of this report for a discussion of why 96.5 percent 

was selected as a measure of good combustion efficiency for reviewing the flare test data.) This 

line shows a transition between good and poor combustion somewhere between a SR of 7.0 and 

9. 

Figure 4-4. Combustion Efficiency vs. SR, zoomed (using TCEQ data) 

Given the differences shown between the TCEQ and EPA-600/2-85-106 data sets (see 

Figure 4-3), we have concerns about the possible differences in SR threshold (transition between 

good and poor combustion) between typical commercial flare sizes. For example, is an SR value 

of 7.0 appropriate for a 54 inch flare? Although Gogolek et al. (2010b) concluded that results 

from flares smaller than 3 inches are not scalable to 6 to 12 inch diameter sized flares; the report 

does not determine whether 6 to 12 inch diameter sized flares are scalable to full-scale industrial 

84%

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Combustion Efficiency, %

SR

TCEQ

SR = 7.0

96.5% CE
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Section 4.0: Air and Flare Performance Page 4-7

sized flares. The observed variation in SR thresholds between a 1.5 inch and 24 inch flare is not 

expected to be as exaggerated for typical commercial flare sizes. Given the steam-assisted flare 

data encompass flare tip sizes (in terms of the effective diameter of the flare tip) from 5.86 inch 

to 54 inches, and the apparent comparability of those results (see Section 3.0 of this report), it is 

likely that this issue is simply inherent with the much smaller flares (i.e., less than 3 inches). 

For these reasons, the data suggest that a SR of 7.0 is the transition between good and 

poor combustion for air-assisted flares. 

4.5 Considering LFLVG for Air-Assisted Flares 

Operating a flare with a SR of 7.0 or less does not tell an owner or operator if there are 

enough combustible components in the combustion zone to burn adequately. If there are not 

enough combustible components in the combustion zone, then the flare vent gas will not burn, or 

will burn inadequately, resulting in lower combustion efficiency. Therefore, it is logical that an 

owner or operator should be conscious of whether the flare vent gas being sent through the 

air-assisted flare is capable of burning. 

It seems reasonable to assume that the LFLCZ analysis in Section 3.0 of this report could 

apply to an air-assisted flare. However, for an air-assisted flare, the “combustion zone gas” 

would simply reduce to the “flare vent gas” because air is the only assist media being added to 

the combustion zone of an air-assisted flare. As discussed in Section 3.0 of this report, for steamassisted flares, the LFL is an indicator of how well a mixture can burn. For steam-assisted flares, 

the LFL in the combustion zone was used to incorporate the effect of the steam in the 

determination of the LFL; however for air-assisted flares, the mixture being introduced with air 

is contained in the flare vent gas. Although, there is air being added as an assist media, it is no 

different than the air in the atmosphere that the flare vent gas would be mixed with, except that 

the mixing is occurring much faster with air-assist. Therefore, for air-assisted flares, instead of 

using the term LFLCZ as an indicator of how well a mixture can burn, we use the lower 

flammability limit of the flare vent gas (LFLVG). 
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Section 4.0: Air and Flare Performance Page 4-8

An observation made regarding the EPA-600/2-85-106 and TCEQ air-assist data is that 

the LFLVG for each of the 44 test runs (i.e., 35 TCEQ test runs and 9 EPA-600/2-85-106 test 

runs) is 15.0 percent by volume or less (when adjusted for nitrogen equivalency as described in 

Section 3.1.2 of this report), but only nine test runs achieved greater than 96.5 percent 

combustion efficiency. This means that each test run (using the 15.3% LFLCZ threshold analysis 

provided in Section 3.0 of this report) had the capability of burning at high combustion 

efficiency, provided the mixture was not diluted with too much assist air (i.e., the SR was 7.0 or 

less). More details regarding the LFL are presented in Section 3.0 of this report; however, there 

is simply not enough air-assisted test data to determine whether a new LFLCZ threshold would be 

warranted for air-assisted flares (i.e., a LFLVG threshold different than 15.3 percent). 

For these reasons, the data seem to suggest that, for an air-assisted flare, a LFLVG

(adjusted for nitrogen equivalency as described in Section 3.1.2 of this report) of 15.3 percent by 

volume or less also indicates that a mixture can combust and maintain a good combustion 

efficiency, assuming that the air-assist flow is not too high to deteriorate efficiency.
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Section 5.0: Wind and Flare Performance Page 5-1

5.0 WIND AND FLARE PERFORMANCE 

The EPA Air Pollution Control Cost Manual - Sixth Edition (U.S. EPA, 2002) states that 

in most flares, combustion occurs by means of a diffusion flame. A diffusion flame is one in 

which air diffuses across the boundary of the fuel/combustion product stream toward the center 

of the fuel flow, forming an envelope of combustible gas mixture around a core of fuel gas. On 

ignition, this combustible gas mixture establishes a stable flame zone around the gas core above 

the flare tip. The inner gas core is heated by diffusion from the hot combustion products 

produced in the flame zone. Leahey et al. (2001) suggest that predicting the size (length, 

diameter, area, and volume) of a diffusion flame could be useful in estimating its combustion 

efficiency because as a flame size is reduced, less oxygen entrains into the flame. Leahey et al. 

(2001) also point out that if a flame’s dimensions are dependent on meteorological variables 

(e.g., wind), a flame’s combustion efficiency would also be dependent on meteorological 

variables (e.g., wind). 

A high crosswind velocity can have a strong effect on the flare flame dimensions and 

shape. When the flame is bent over on the downwind side of a flare and is imbedded in the wake 

of the flare tip, it is said to be in a wake-dominated regime. Many experts believe a wakedominated flame can lead to poor flare performance. The data suggest that flare performance is 

not significantly affected by crosswind velocities up to 22 miles per hour (mph). There are 

limited data for flares in winds greater than 22 mph. However, a wake-dominated flame in winds 

greater than 22 mph may affect flare performance. The data available indicate that the wakedominated region begins at a momentum flux ratio (MFR) of 3 or greater. The MFR considers 

whether there is enough flare vent gas and center steam (if applicable) exit velocity (momentum) 

to offset crosswind velocity. Because wake-dominated flames can be identified visually, 

observations could be conducted to identify wake-dominated flames during crosswind velocities 

greater than 22 mph at the flare tip. This section documents the analysis supporting these 

observations. 
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Section 5.0: Wind and Flare Performance Page 5-2

5.1 Introduction 

Smoot et al. (2009), who examined operating and design issues for pipe flares, provide an 

illustration on how flare performance is affected by crosswinds that shear the combustion zone. 

These researchers state that wind shortens the flame, causing soot formation on the downwind 

side of the flame. They also warn that high crosswinds can result in air egress into the flare tip, 

which creates internal burning and potential explosive conditions (Figure 5-1). 

Figure 5-1. Air Egression Into Flare Stack 

Source: (Smoot et al., 2009) 

5.2 Flare Flow Mixing Regimes 

Depending on crosswind and flare tip velocity, the appearance of a flare flame can be 

quite different. In a report from a Joint International Combustion Symposium (Seebold et al., 

2004), the authors summarize and provide images (Figure 5-2) of three flare flow mixing 

regimes: “wake-dominated”, “buoyancy-dominated”, or “inertia-dominated”. At low crosswind 

velocity and high flare tip velocity, the flare flame is generally inertia-dominated, with the flame 
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Section 5.0: Wind and Flare Performance Page 5-3

positioned directly above the jet stack and the flame curling straight upwards. In this case, the 

strength of the jet momentum flux in relation to the strength of the wind ensures a stable 

combustion zone and combustion efficiencies are invariably higher (Seebold et al., 2004). 

Figure 5-2. Images of Flow Mixing Regimes 

Source: (Seebold et al., 2004) 

As crosswind velocities increase, “downwash” occurs as the flame lengthens horizontally 

and a portion of the combustion gases are drawn into the low pressure region on the downwind 

side of the stack (Johnson and Kostiuk, 2000). The flame zone diminishes into detached pockets 

of combustion interspersed beneath a non-reacting mixing layer containing unburned fuel 

(Johnson and Kostiuk, 2000; Johnson et al., 2001). In these high crosswind (and low flare vent 

gas velocity) situations, the flare flame is referred to as wake-dominated, which means the flame 

is bent over on the downwind side of a flare pipe and is imbedded in the wake of the flare tip 

(Figure 5-2), protected from the crosswind (Johnson et al., 2001). Under these conditions, the 

main tail of the flame may be extinguished, and the primary combustion zone occurs as a 

standing vortex on the downwind side of the stack (Johnson et al., 2001; Johnson and Kostiuk, 

1999, 2000). 
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Section 5.0: Wind and Flare Performance Page 5-4

5.3 Efficiency Studies 

Seebold et al. (2004) state that the wake-dominated mixing regime may lead to low 

combustion efficiency eddies. Several efficiency studies conducted at high crosswind velocities 

suggest that poor combustion efficiencies in wake-dominated flare flames result from fuel 

stripping where a portion of unburned fuel is stripped from the flame stream prior to reaching a 

flame zone (Johnson et al., 1999; Johnson and Kostiuk, 2000; Johnson et al., 2001; and 

Castiñeira and Edgar, 2006). The unburned fuel is primarily ejected into the underside of the 

flame. Johnston et al. (2001) provide a sketch of the unburned hydrocarbons detected downwind 

in a wake-dominated flare with a crosswind velocity of 8 meters/second (m/s) or about 18 mph 

(Figure 5-3). 

Figure 5-3. Fuel Detection Downwind of Wake-Dominated Flare 

Source: (Johnston et. al, 2001) 
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Section 5.0: Wind and Flare Performance Page 5-5

Researchers from the University of Alberta have published a series of studies that 

describe their investigations on the significance of wind on the fluid mechanics of a flare flame 

at low MFR (Bourguignon et al., 1999; Johnson et al., 1998, 1999a, 1999b, 2001; and Johnson 

and Kostiuk, 1999, 2000). The MFR is a measure of momentum strength of the flare vent gas 

relative to the crosswind (the product of flare vent gas density and velocity squared divided by 

the product of air density and crosswind velocity squared). The MFR at the flare tip can be 

calculated using Equation 5-1.

./ 01234506	75



819:86;<=:

 (Eq. 5-1)

Where: 

MFR = Momentum flux ratio, unitless. 

ρvgcs = Density of flare vent gas including center steam if applicable, lb/scf. 

Vtip = Actual flare tip velocity (i.e., actual flare vent gas velocity plus center steam 

velocity, if applicable), ft/s. The flare tip velocity is dependent on how the 

unobstructed cross sectional area of the flare tip is calculated. John Zink provided 

details for determining the unobstructed cross sectional area of several flare tip 

designs (see Appendix I). 

ρair = Density of ambient air, lb/scf. 

Vwind = Crosswind velocity, ft/s. 

The experiments in these studies were conducted on about a one-inch diameter simple 

pipe flare using a wind tunnel. The maximum crosswind velocity during these experiments was 

31 mph. These studies concluded that increased crosswind velocity can adversely affect the 

combustion efficiency of a flare; while increased flare tip velocity makes the flame less 

susceptible to the effects of crosswind. One report (Johnson and Kostiuk, 2000) provides images 

(Figure 5-4) of the flame, showing that as MFR decreases, the combustion efficiency of the 

flame also decreases. Mathematical models presented by Castiñeira and Edgar (2006) determined 

similar results for about 0.5-inch diameter simple pipe flares. Further modeling (Castiñeira and 

Edgar, 2008) showed a simulated flame almost completely extinguished at crosswind velocities 

of 22 mph. 
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Section 5.0: Wind and Flare Performance Page 5-6

Figure 5-4. Flame Images Relating to Momentum Flux Ratio and Combustion Efficiency 

Source: (Johnson and Kostiuk, 2000) 1 

 

1

 Images from a one-inch diameter simple pipe flare using a wind tunnel. The “R” term is MFR, the “ηc” term is the combustion efficiency of the 

flame, the “U” term is crosswind velocity, and the “Vj” term is flare tip velocity. 
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Section 5.0: Wind and Flare Performance Page 5-7

Seebold et al. (2004) present a graph (Figure 5-5) that shows a potentially strong 

correlation between combustion efficiency and MFR, revealing that as the MFR falls below 0.1, 

the combustion efficiency decreases significantly (for some data). However, we are not able to 

confirm the source of this data or anything about these test runs. Because the reference that 

Seebold et al. (2004) provide (shown in the reproduced Figure 5-5 as “Reaction Efficiency of 

Industrial Flares – The Perspective of the Past.” September 2003) does not contain details about 

how Figure 5-5 was created, we have not used Figure 5-5 in our analysis. 

Figure 5-5. Combustion Efficiency vs. Momentum Flux Ratio, Seebold Data 

Source: (Seebold et al., 2004) 

Gogolek et al. (2010b) investigated the transition from a vertical jet flame to a 

wake-dominated flame by performing tests on a 2 inch and 3 inch flare tip (and also on a 3 inch 

flare tip that had a flame retention ring). Gogolek et al. (2010b) determined visually that the 

transition to a wake-dominated flame occurred at a MFR of around 3 for all three pipes tested. 

However, the study also revealed that there is no sharp change in flare performance with the 

establishment of the wake-dominated flame. Additionally, Gogolek et al. (2010b) found that two 

sets of operating conditions with the same MFR produce drastically different flare performance 

(one with high crosswind and moderate flare tip velocity, and another with moderate crosswind 
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Section 5.0: Wind and Flare Performance Page 5-8

and low flare tip velocity), indicating that MFR may have limitations for correlating the 

performance of flares. The study concluded that a decrease in combustion efficiency is more 

likely to occur when the transition to a wake-dominated flame is caused by an increase in 

crosswind velocity rather than by a reduction in flare tip velocity. 

In lieu of the MFR, other parameters such as a plume buoyancy factor or a power factor 

may be considered. It was through the research performed by the University of Alberta 

(Bourguignon et al., 1999; Johnson et al., 1998, 1999a, 1999b, 2001; and Johnson and Kostiuk, 

1999, 2000) that these parameters were developed. Gogolek et al. (2010a) summarize the 

methodologies for determining a plume buoyancy factor or power factor. The plume buoyancy 

factor is calculated using Equation 5-2; and it considers crosswind velocity and actual flare vent 

gas velocity plus center steam velocity. The power factor is calculated using Equation 5-3; it is 

the ratio of the power of the crosswind to the power of combustion of the flare vent gas. (See 

Gogolek et al. (2010a) for further detail on these two parameters.) Center steam should be 

incorporated into the flare vent gas properties because center steam adds momentum (i.e., center 

steam exits the flare tip in the same direction as the flare vent gas).

>?  6;<=

0@AB7A6	75

C

,

D

 (Eq. 5-2) 

Where: 

BP = Plume buoyancy factor, unitless. 

Vwind = Average cross-sectional wind velocity, ft/s. 

g = Acceleration due to gravity, ft/s2. 

Dp = Effective diameter of flare pipe, ft. 

Vtip = Actual flare tip velocity (i.e., actual flare vent gas velocity plus center steam 

velocity, if applicable), ft/s. The flare tip velocity is dependent on how the 

unobstructed cross sectional area of the flare tip is calculated. John Zink provided 

details for determining the unobstructed cross sectional area of several flare tip 

designs (see Appendix I). 

?/  E 19A6;<=AB7



1FAG7A6	7AH 6I

J





D

 (Eq. 5-3) 
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Section 5.0: Wind and Flare Performance Page 5-9

Where: 

PF = Power factor, unitless. 

ρa = Density of air, lb/scf. 

Vwind = Average cross-sectional wind velocity, ft/s. 

Dp = Effective diameter of flare pipe, ft. 

ρf = Density of flare vent gas including center steam, lb/scf. 

Ap = Unobstructed cross sectional area of the flare tip, ft2. John Zink provided details 

for determining the unobstructed cross sectional area of some of several flare tip 

designs (see Appendix I). 

Vtip = Actual flare tip velocity (i.e., actual flare vent gas velocity plus center steam 

velocity, if applicable), ft/s. The flare tip velocity is dependent on how the 

unobstructed cross sectional area of the flare tip is calculated (See Appendix I). 

LHVm = Lower heating value (mass basis) of flare vent gas including center steam, 

Btu/scf. 

Gogolek et al. (2010b) tested the power factor using natural gas on 3, 4, and 6 inch flare 

tips, and concluded that the power factor appears to be a useful dimensionless parameter for 

correlating flare performance data. However, Gogolek et al. (2010b) conclude that the power 

factor would have to be augmented to correlate performance of different flare vent gas mixtures. 

Gogolek et al. (2010b) also indicates that crosswind velocity thresholds for indicating good 

combustion efficiency may differ depending on the composition of the flare vent gas. The study 

concluded (based on tests using 3” flare retention ring flare tips) that destruction efficiency 

remains better than 98 percent for ethylene at crosswind velocities tested up to 24.6 mph; but fall 

below 98 percent at crosswind velocities of 17.9 mph for propylene. The study also reports that 

destruction efficiency remained above 97 percent for all crosswind velocities tested for 

propylene. These tests using 3” flare retention ring flare tips produced results similar to the 2” 

pipe flare, which are not scalable to industrial sized flares. All tests conducted by Gogolek et al. 

(2010b) on the flares found to be scalable to industrial sized flares (i.e., 3”, 4”, and 6” pipe flares, 

and 6” with flame retention rings) showed combustion efficiencies greater than 96.5 percent for 

all fuels tested for winds of 22 mph and less. 

Gogolek et al. (2010b) incorporate a “fuel factor” into their analysis as an attempt to 

incorporate the effects of flare gas combustion properties, wind speed, and flare vent gas flow 
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Section 5.0: Wind and Flare Performance Page 5-10

rate; however, the “fuel factor” uses the upper flammability limit (UFL), which is difficult to 

calculate for mixtures, especially ones with inert. Also, Gogolek et al. (2010b) report that the 

"fuel factor" did not correlate well when flaring natural gas; and provide an explanation that the 

"fuel factor" does not consider the reaction kinetics of combustible mixtures. 

Excluding the data shown in Figure 5-5 (because flare size is unknown for this data) as 

well as Gogolek’s et al. (2010b) power factor analysis, all of the studies mentioned in this section 

(i.e., Section 5.3 of this report) were conducted on 3 inch flare tips and smaller, which are not of 

typical size for an industrial flare. Gogolek et al. (2010b) states that results of pipes smaller than 

3 inch do not scale-up to larger pipes; and it has not been determined whether results for 3 inch 

to 6 inch pipes can successfully be applied to full-scale industrial flares. However, it does appear 

crosswind velocity influences the size of a diffusion flame (e.g., length, diameter, area, and 

volume) and should be considered in flare performance no matter the size of a flare tip. It seems 

reasonable to assume that combustion efficiency of an industrial sized flare will also likely be 

sensitive to crosswind velocity. 

5.4 Test Data Analysis 

Although Gogolek et al. (2010b) indicates that MFR may have limitations for correlating 

the performance of flares, we reviewed available MFR data as provided in raw test data 

spreadsheets for 245 steam-assisted flare test runs from five of the eight studies discussed in 

Section 2 of this report (i.e., TCEQ, MPC TX City, MPC Detroit, FHR AU, and FHR LOU). 

Unfortunately, the data provided in these reports were not used consistently to calculate MFR 

(i.e., some reports incorporated total assist steam, while others only center steam when using an 

actual flare tip velocity in their MFR methodology). Therefore, we recalculated MFR for all 

available test runs to incorporate only the center steam volume into the flare vent gas exit 

velocity because it exits the flare tip in the same direction as the flare vent gas. Steam and/or air 

nozzles associated with the upper and lower ring locations were not incorporated into the flare 

vent gas exit velocity when calculating MFR because it is less clear whether steam contribution 

from these nozzles should be included. Steam velocity and nozzle angle for upper and lower ring 

locations was not available for all of the test reports. Upper and lower ring nozzles are generally 
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Section 5.0: Wind and Flare Performance Page 5-11

directed up at about 30 to 45 degrees from horizontal (jetting into but also in the direction of the 

flare vent gas momentum), so only a portion of steam from upper and lower ring nozzles could 

enhance the momentum of the flare vent gas flow. 

These test data represent 245 test runs with flare tip velocities (including center steam 

velocity) of less than 9 feet per second (ft/s) and wind speeds typically below 12 mph 

(approximately 20% of the test runs were performed in wind speeds above 12 mph; and no test 

runs were performed in winds any higher than 22 mph). Most of these test runs inevitably have 

low MFR values because the runs were performed at relatively low flare tip velocities. 

Nevertheless, the MFR from these test runs was plotted against the combustion efficiency to 

investigate whether a trend was discernable (Figures 5-6 and 5-7). 

In general, Figure 5-6 shows that all runs were tested at a MFR less than about 7.0. 

However, it is assumed that most of these data (149 steam-assisted test runs) had poor 

flammability characteristics because the LFLCZ for these test runs was greater than 15.3 percent 

by volume (see Section 3.0 of this report for an explanation of this LFLCZ level). Also, there are 

10 test runs highlighted in red in Figure 5-6 that had a LFLCZ less than 15.3 percent by volume 

and did not achieve a combustion efficiency of 96.5 percent or greater. The poor combustion 

efficiency reported for these 10 test runs are discussed in Section 3.1.3 of this report. 
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Section 5.0: Wind and Flare Performance Page 5-12

Figure 5-6. Combustion Efficiency vs. Momentum Flux Ratio 

Figure 5-7 presents a subset of the data in Figure 5-6—those with MFR values less than 

3.0—because Gogolek et al. (2010b) suggests that a MFR of 3.0 is the observed boundary for the 

transition to the wake-dominated mixing regime. 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0 1 2 3 4 5 6 7 8

Combustion Efficiency (%)

MFR

LFL<15.3%, CE >96.5%

LFL<15.3%, CE < 96.5%

LFL>15.3%
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Section 5.0: Wind and Flare Performance Page 5-13

Figure 5-7. Combustion Efficiency vs. Momentum Flux Ratio, zoomed 

(MFR < 3.0; wake-dominated mixing regime) 

These plots indicate that good flare performance can exist in a wake-dominated mixing 

regime. However, it is difficult to establish a MFR limit to define a wake-dominated regime with 

good combustion efficiency because it is not clear from these plots whether low MFR, high 

LFLCZ, or some combination of both contribute to poor flare performance. Also, the available 

test data represent only a narrow range of wind and flare vent gas velocities (primarily <12 mph 

and <8 ft/s, respectively). For example, although the test data show that good flare performance 

can occur at a MFR as low as 0.003 (identified on Figure 5-8 with solid blue vertical line), this 

conclusion is based on only a single data point. 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0.0 0.5 1.0 1.5 2.0 2.5 3.0

Combustion Efficiency (%)

MFR

LFL<15.3%, CE >96.5%

LFL<15.3%, CE < 96.5%

LFL>15.3%
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Section 5.0: Wind and Flare Performance Page 5-14

Figure 5-8. Combustion Efficiency vs. Momentum Flux Ratio, further zoomed 

(MFR < 0.1) 

The test runs shown in Figures 5-6 through 5-8 are of moderate crosswind and low flare 

tip velocity. It appears that the data confirms what Gogolek et al. (2010b) observed which was 

that low MFRs created because of the low flare tip velocity (and less so because of higher 

crosswind) do not have the degraded combustion efficiency that low MFRs created by high 

crosswind (and less so because of low flare tip velocity) show. In other words, in the absence of 

higher crosswind velocities, a lower MFR does not necessarily indicate poor flare performance. 

Because we do not have test data of crosswind speeds greater than 22 mph, we are unable to 

investigate this hypothesis further using the industrial flare test data. It should be noted that 

Gogolek et al. (2010b) performed tests as high as about 27 mph; no test runs from that study that 

were performed below 22 mph reflected poor flare performance. 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0.00 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09 0.10

Combustion Efficiency (%)

MFR

LFL<15.3%, CE >96.5%

LFL<15.3%, CE < 96.5%

LFL>15.3%
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Section 5.0: Wind and Flare Performance Page 5-15

For these reasons, it seems reasonable to conclude that the momentum flux ratio (MFR) 

should be 3 or greater in crosswinds that are greater than 22 miles per hour (mph) at the flare tip. 

As an alternative to monitoring the MFR, flare operators could perform observations to identify 

wake-dominated flames during crosswind velocities greater than 22 mph at the flare tip and flare 

vent gas velocity could be increased if a wake-dominated flame is observed. Observation 

monitoring for wake-dominated flames could be established similarly to the current visible 

emissions monitoring requirements for flares. For example, the length of the observation period 

as well as the frequency could be triggered by different wind speed thresholds or categories. 

There are two main concerns with relying on observations: (1) that some flames may simply be 

too difficult to see, and (2) it may be difficult to recognize when a flame is wake-dominated. 

We also reviewed the power factor for the same 245 steam-assisted flare test runs that 

were analyzed for MFR. Figure 5-9 is a plot of combustion efficiency versus power factor. The 

same could be said about power factor as was said about MFR; it is assumed that most of these 

data (149 steam-assisted test runs) had poor flammability characteristics because the LFLCZ for 

these test runs was greater than 15.3 percent by volume (see Section 3.0 of this report for an 

explanation of this LFLCZ threshold). Also, there are 10 test runs highlighted in red in Figure 5-9 

that had a LFLCZ less than 15.3 percent by volume and did not achieve a combustion efficiency 

of 96.5 percent or greater. The poor combustion efficiency reported for these 10 test runs are 

discussed in Section 3.1.3 of this report. Although it appears that a power factor of 0.15 or less 

could be used as a level for indicating good combustion efficiency, it is less clear whether this 

level (i.e., 0.15 or less) would be appropriate for all flare vent gas mixtures. In addition, we did 

not attempt to incorporate a “fuel factor” as Gogolek et al. (2010b) did with their data, because 

the UFL is difficult to calculate for mixtures, especially ones with inert. 
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Section 5.0: Wind and Flare Performance Page 5-16

Figure 5-9. Combustion Efficiency vs. Power Factor 

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0.05 0.15 0.25 0.35 0.45 0.55

Combustion Efficiency (%)

Power Factor

LFL<15.3%, CE >96.5%

LFL<15.3%, CE < 96.5%

LFL>15.3%
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Section 6.0: Flare Flame Lift Off Page 6-1

6.0 FLARE FLAME LIFT OFF 

Flare flame lift off can cause a flame to become unstable, which can negatively affect the 

combustion efficiency of a flare. To avoid flame lift off, the data suggest that the actual flare tip 

velocity (i.e., actual flare vent gas velocity plus center steam velocity, if applicable) should be 

less than an established maximum allowable flare tip velocity calculated using an equation that is 

dependent on combustion zone gas composition, the flare tip diameter, density of the flare vent 

gas, and density of air. This section documents the analysis supporting this observation. 

6.1 Literature Review and Vmax Calculation 

Flare flame lift off is a condition where a flame separates from the tip of the flare and 

there is space between the flare tip and the bottom of the flame due to excessive air induction as 

a result of the flare gas and center steam exit velocities (Gogolek et al., 2010a describes this 

situation as flame “blow-off”). Flame stability can exist when flame lift off occurs; however, as 

flame stability decreases (which can be caused by a variety of reasons, such as changes in flare 

tip velocity, combustion zone gas composition, or wind direction and speed), a flare operating at 

flame lift off can progress to a condition where the flame is extinguished; at a minimum, the 

flame will be inconsistent (Kalghatgi, 1981a; and Shore, 2007). Therefore, flame lift off should 

be monitored and avoided. A primary contributor to flame lift off is flare tip velocity, especially 

its relationship to the flame speed of the constituents in the flare vent gas (Shore, 2007). 

Gogolek et al., 2010a provide three different equations that correlate flame lift off and 

flare tip velocity. Each of the equations presented by Gogolek et al. use a different combination 

of variables. However, this section primarily focuses on an equation presented by Shore (2007). 

Shore’s equation was one of the equations presented by Gogolek et al. (2010a). This equation 

was selected because it considers lower flammability limit, which would already be known by a 

source if the recommendation of Section 3.0 of this report were followed. Also, other equations 

presented by Gogolek et al. (2010a) use parameters that are not readily available for mixtures 

(i.e., flame speed and upper flammability limit). 
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Section 6.0: Flare Flame Lift Off Page 6-2

In order to calculate the boundary velocity between an unstable and stable discharge 

velocity (or maximum velocity) for any gas or mixture exiting the flare tip, Shore’s equation 

(Equation 6-1) considers the flare tip diameter, lower flammability limit and density of the flare 

vent gas, and density of air. Center steam should be incorporated into the flare vent gas 

properties because center steam exits the flare tip in the same direction as the flare vent gas. 

Shore developed the equation using published experimental results on lift off using various 

mixtures from several researchers. 

K#L

2MN

OP

M



Q

R

E

100 U V/VW@XY

V/VW@XY J Z[W@XY[#$%

\

6.85 _

`

+

(Eq. 6-1) 

Where: 

Vmax = Maximum flare tip velocity including, if applicable, center steam at which 

flame lift off is not expected to occur, ft/sec. 

π = 3.14, constant. 

Au = Unobstructed cross sectional area of the flare tip, ft

2

. John Zink provided 

details for determining the unobstructed cross sectional area of several flare 

tip designs (see Appendix I). 

LFLvgcs = The lower flammability limit of flare vent gas including, if applicable, center 

steam, volume %. A methodology for calculating lower flammability limits is 

included in Section 3.0 of this report. 

ρvgcs = Density of flare vent gas including center steam if applicable, lb/scf. 

ρair = Density of ambient air, lb/scf. 

6.2 Test Data Analysis 

Test run data were evaluated against Equation 6-1 to see how well they would correspond 

to the equation’s ability to predict a maximum flare tip velocity for flame lift off. Shore explains 

that at the equation’s predicted maximum flare tip velocity, stream eddies from the flare tip have 

reached a critical vertical shedding frequency and flame lift off may occur (Shore, 2007). 
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Section 6.0: Flare Flame Lift Off Page 6-3

Three hundred thirty (330) of the 356 steam-assisted and air-assisted flare test runs were 

considered for this analysis. Twenty-six (26) test runs from the Shell Deer Park East Property 

Flare study (Shell Global Solutions (US) Inc., 2011a) were not included in the analysis because 

there were limited steam information for these data points and a center steam rate could not be 

determined (which is necessary to calculate the “LFLvgcs” term in Equation 6-1). It was assumed 

that most of these data (183 steam-assisted test runs) had poor flammability characteristics 

because the LFLCZ for these test runs was greater than 15.3 percent by volume (see Section 3.0 

of this report for an explanation of this LFLCZ threshold). It was also assumed that 30 air-assisted 

test runs had too much assist air because the SR was greater than 7 (see Section 4.0 of this report 

for an explanation of this SR threshold). Therefore, 213 of the 330 test runs were disregarded 

because they were expected to have poor combustion efficiency regardless of the flare tip 

velocity. Finally, all nine air-assist data points from the EPA-600/2-85-106 test report (Pohl and 

Soelberg, 1985) were removed from the Equation 6-1 evaluation because these tests were 

performed on a flare with a 1.5 inch diameter, which is not expected to be scalable to larger 

industrial type flares (see Section 4.0 of this report for more details). 

The analysis was performed on the remaining 108 test runs. Figure 6-1 is a plot of 

Equation 6-1 against the 108 test runs that have either a LFLCZ of less than or equal to 

15.3 percent by volume (for steam-assist data), or a SR of less than or equal to 7 (for air-assist 

data). Ten (10) steam-assist test runs with a LFLCZ less than 15.3 percent by volume that did not 

achieve a combustion efficiency of 96.5 percent or greater, are highlighted in red in Figure 6-1. 

The poor combustion efficiency reported for these 10 test runs is discussed in Section 3.1.3 of 

this report. 

The left side of the green line in Figure 6-1 marks the region where vertical stability of 

the flame may be compromised (the green line is the linear relationship of Equation 6-1 and 

represents the maximum flare tip velocity including, if applicable, center steam at which flame 

lift off is not expected to occur). Although actual flame lift off data were available for only two 

of these 108 test runs (lift off information is only provided in the EPA-600/2-85-106 study (Pohl 

and Soelberg, 1985)), it is reasonable to assume that all of the 108 test runs did not exhibit flame 

lift off because the tests were done at high turndown ratios, which means the actual flare vent gas 
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Section 6.0: Flare Flame Lift Off Page 6-4

flow rate was much lower than what the flare is designed to handle (even when center steam is 

incorporated into the calculation). Therefore, all 108 test runs should fall on the right side of the 

green line in Figure 6-1 because one would expect all 108 test runs to not have flame lift off. 

 

It is difficult to draw any conclusions on whether the three data points on the left side of 

the Equation 6-1 threshold (see Figure 6-1) should be considered outliers. Using Equation 6-1 as 

the indicator, it is expected that these three data points would have exhibited flame lift off during 

their test run. However, one of these data points came from the EPA-600/2-85-106 study (Pohl 

and Soelberg, 1985), which noted that no lift off occurred during this particular test run (i.e., test 

run 206). The other two test runs came from the EPA-600/2-83-052 study (McDaniel 1983), 

which does not provide any information on lift off. 

Figure 6-1. Conditions for Stable Flare Flame 

Equation 6-1

0.010

0.100

1.000

10.000

100.000

1.000 10.000 100.000

>96.5% CE

>96.5% CE (Air Assist)

<96.5% CE

Note:

VVG-S is the actual flare vent gas velocity including center steam
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Section 6.0: Flare Flame Lift Off Page 6-5

Even though these three test runs fall on the left side of the Equation 6-1 threshold, they 

each achieved a combustion efficiency of 96.5 percent or greater. This observation is not 

surprising because good flare performance can still exist when flame lift off occurs. There are 

four other data points in Figure 6-1 that are from the 1980s EPA studies, but these test runs fall 

on the right side of the Equation 6-1 threshold and achieve a combustion efficiency of 

96.5 percent or greater (one of these runs is from the EPA-600/2-85-106 study, which noted that 

no lift off occurred during the test run; and the other three runs came from the EPA-600/2-83-

052 study). 

The two test runs from the EPA-600/2-85-106 study did not use a pilot during testing but 

all other data points in Figure 6-1 had pilot operations. Also, the flare vent gas volumetric or 

mass flow rates for the EPA-600/2-85-106 test runs were not reported so assumptions were made 

using the reported steam-to-fuel ratio and fuel composition to determine flow rates. The data 

from the EPA-600/2-83-052 study came from tests using a steam-assisted flare that was the 

smallest sized flare of all test data included in Figure 6-1 (i.e., the flare tip diameter was 

8.625 inches, and the effective diameter was less than 6 inches). 

It appears nearly all the available test data fits Equation 6-1; all but three test runs 

evaluated had velocities less than the maximum velocity allowed by Equation 6-1. However, this 

analysis does not test Equation 6-1 over a very large range of possible flare tip velocities. All of 

the recent flare test data were collected during high turndown ratios and consequently have 

relatively low flare tip velocities. For the steam-assisted flares that were tested, the turndown 

ratio ranged from 504.1 to 13.8 (or 0.20 to 7.24 percent utilization); and tip velocities were less 

than 10 feet per second with center steam excluded. For the air-assisted flare (from the TCEQ 

data set), the turndown ratio ranged from 410.2 to 155.3 (or 0.24 to 0.64 percent utilization); and 

tip velocities were less than 2 feet per second. Maximum flare capacity design is unknown for 

the flares used in this analysis that are from the EPA-600/2-83-052 and EPA-600/2-85-106 

studies; therefore, the turndown ratios are unknown. Flare tip velocities for these studies ranged 

from approximately 2 to 50 feet per second (for those test runs included in Figure 6-1). 
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Section 6.0: Flare Flame Lift Off Page 6-6

6.3 Other Operating Parameters Considered for Flame Lift Off

There are other methods studied by researchers. For example, Gogolek et al. (2010a) 

discuss two additional equations (other than Equation 6-1) that were developed by Nobel et al. 

(1984) and Kalghatgi (1981a and b); however, these two equations were not analyzed against the 

available flare test data. One of these equations that Gogolek et al. (2010a) discuss uses the upper 

flammability limit (UFL), which is difficult to calculate for mixtures, especially ones with inerts. 

Le Chatelier’s equation could be used to determine UFLs, but it is much less accurate than 

calculating a LFL. The other equation Gogolek et al. (2010a) mention uses flame velocity, which 

is a difficult parameter to acquire, because there are relatively few flame velocities that have 

been tabulated and no accepted method is known for estimating flame velocities for mixtures. 
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Section 7.0: Other Flare Type Designs to Consider Page 7-1

7.0 OTHER FLARE TYPE DESIGNS TO CONSIDER 

7.1 Non-Assisted Flares 

A non-assisted flare does not use any auxiliary steam, air, or vent stream pressure to 

promote mixing at the flare tip. We are unable to verify whether any of the analyses presented in 

this technical report could apply to non-assisted flares because there are minimal test data 

available for non-assisted flares. It seems reasonable to assume that the LFLCZ analysis in 

Section 3.0 of this report could apply to a non-assisted flare. For a non-assisted flare, the 

“combustion zone gas” would simply become the “flare vent gas” because no assist media is 

added to the combustion zone of a non-assisted flare. Therefore, a determination that LFLCZ is 

the most appropriate operating parameter to monitor for steam-assisted flares, could also apply to 

non-assisted flares (i.e., the LFLCZ must be 15.3 percent by volume or less when regulated 

material is being routed to a non-assisted flare in order to maintain good combustion efficiency). 

Nine of the 312 test runs that were performed on a steam-assisted flare did not use any steam 

during testing and could be considered a ‘non-assisted’ flare test run. Of these nine test runs, 

eight achieved greater than 96.5 percent combustion efficiency, but only three had a LFLCZ less 

than 15.3 percent. The remaining test run did not achieve greater than 96.5 percent combustion 

efficiency and the LFLCZ was greater than 15.3 percent. There is simply not enough non-assisted 

test data to determine whether a new LFLCZ threshold would be warranted (i.e., a LFLCZ

threshold different than 15.3 percent); however, a LFLCZ level of 15.3 percent and less appears to 

adequately predict good combustion for the nine non-assisted flare test runs. 

It is assumed that the wind analysis in Section 5.0 of this report and the flame lift off 

analysis in Section 6.0 of this report apply to a non-assisted flare because these factors (wind and 

flame lift off) affect a diffusion flame from a non-assisted flare in the same way they would 

affect a diffusion flame created by a steam-assisted or air-assisted flare. The MFR and Vmax

methodologies in this report apply to non-assisted, steam-assisted, and air-assisted flares; assist 

media is only included in these methodologies if applicable. 
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Section 7.0: Other Flare Type Designs to Consider Page 7-2

7.2 Pressure-Assisted Flares and Other Flare Designs 

The EPA Air Pollution Control Cost Manual - Sixth Edition (U.S. EPA, 2002) states that 

pressure-assisted flares use the vent stream pressure to promote mixing at the burner tip. Several 

vendors now market proprietary, high pressure drop burner tip designs. If sufficient vent stream 

pressure is available, these flares can be applied to streams previously requiring steam or air 

assist for smokeless operation. Pressure-assisted flares generally (but not necessarily) have the 

burner arrangement at ground level, and consequently, must be located in a remote area of the 

plant where there is plenty of space available. They have multiple burner heads that are staged to 

operate based on the quantity of gas being released. The size, design, number, and group 

arrangement of the burner heads depend on the flare vent gas characteristics. 

The amount of data available for pressure-assisted flares is small, especially considering 

the wide range in pressure-assisted flare designs. In addition, we have no data to analyze other 

new technologies, such as a hybrid steam-and-air-assist flare. We are aware of the Pohl and 

Soelberg (1985) flare study performed on pressure-assisted flares. The 1985 flare study 

published two curves for pressure-assisted flares illustrating a relationship between a given exit 

velocity and the minimum gas heat content that maintains the flame stability. We are also aware 

of pressure-assisted flare testing that was conducted at the John Zink test facility located in 

Tulsa, Oklahoma; this test program was coordinated with TCEQ and the Dow Chemical 

Company (Varner et al., 2007, see Docket ID No. EPA-HQ-OAR-2010-0868). The test report 

concluded that pressure-assisted flares can achieve performance levels at least as good as steamassisted and air-assisted flares; however, the test program only tested two flare burner designs. 

Because of lack of performance test data on pressure-assisted flare designs and other flare 

design technologies, and given the uniqueness in design of pressure-assisted flare designs from 

non-assisted, steam-assisted, and air-assisted flares (and across the population of pressureassisted flares), it seems likely that the observations made in this report for non-assisted, steamassisted, and air-assisted flares, cannot be applied to pressure-assisted flare designs or other flare 

design technologies. Also, test data are not available to form general conclusions on operating 

parameters that represent good combustion for these other flare designs. 
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Section 8.0: Monitoring Considerations Page 8-1

8.0 MONITORING CONSIDERATIONS 

In order to determine if a flare is operated within the operating parameter values 

described in the previous sections, monitoring equipment will be needed. This section describes 

the types of monitoring equipment and methodologies that could be used to determine LFLCZ, 

LFLVG, LFLVG,C, the ratio of NHVCZ to NHVVG-LFL, CCZ, SR, MFR, and Vmax on an ongoing 

basis. 

8.1 LFLCZ, LFLVG, and LFLVG,C

We are only aware of one viable monitoring method that could be used to continuously 

determine the LFLCZ, LFLVG, and/or LFLVG,C. These parameters (i.e., LFLCZ, LFLVG, and/or 

LFLVG,C) require a gas chromatograph to determine the individual component concentrations 

present in the flare vent gas, and flow meters to determine the volumetric flow rates of flare vent 

gas and total steam (if applicable). The individual component concentrations present in the flare 

vent gas, the pure component LFL values, and volumetric flow rates of flare vent gas and total 

steam (if applicable) are applied to the equations and methods described in Section 3.1.2 of this 

report to determine LFLCZ, LFLVG, and/or LFLVG,C. 

There are several types of LFL monitors used for safety purposes, but we are not aware 

of any that provide a direct measure of the LFL of a gas mixture. Instead, these types of LFL 

monitors react to specific gas concentrations in air. There are personal monitoring devices 

available on the market that are used to detect explosive environments based on the LFL (for 

safety purposes); however, these devices are not meant for a continuous flow gases, can become 

saturated quickly if they are in continuous contact with a gas mixture, and may need a specific 

amount of oxygen to work properly. Other types of LFL monitoring devices that are used for 

safety are generally calibrated for only one type of gas, yet flare vent gas can consist of any 

variety of gas mixtures. The EPA is aware of one company2 that makes a LFL monitor that could 

potentially be modified to continuously determine the LFLCZ, LFLVG, and/or LFLVG,C; however, 

the monitor has never been used for this specific purpose. The device uses a continuous 

 

2

 Control Instruments Corporation – PrexEx Flammability Analyzer 
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Section 8.0: Monitoring Considerations Page 8-2

hydrogen flame inside a small chamber. Flammable vapors are drawn from the sample point into 

the chamber, where they are incinerated by the flame. A temperature detector measures the 

resulting change in flame temperature and provides an output reading in percent of the LFL of 

the gas mixture (not actual LFL). This LFL monitor provides a response time of less than 

1 second, and can be accessed remotely. By combining this LFL monitor with metered air 

injection and a feedback control signal to adjust the air until the meter reads a preset percent of 

LFL, it seems possible that the meter could be used to measure the LFL. Some integrated 

programming to calculate the LFL from the flow rates and percent of LFL reading would also be 

needed, but this type of set up could potentially provide a LFL monitor without the expense of a 

gas chromatograph. 

8.2 Ratio of NHVCZ to NHVVG-LFL

Because the ratio of NHVCZ to NHVVG-LFL requires the LFLVG to be known as part of the 

calculation (see Equation D.22 of Appendix D of this report), this parameter (i.e., the ratio of 

NHVCZ to NHVVG-LFL) requires a gas chromatograph. The gas chromatograph is used to 

determine the individual component concentrations present in the flare vent gas. Flow meters are 

also required to determine the volumetric flow rates of flare vent gas and total steam (if 

applicable). If there were a LFL monitor that could determine LFLVG as discussed in Section 8.1 

of this report, a calorimeter (or BTU analyzer) could be used in lieu of a gas chromatograph to 

determine the ratio of NHVCZ to NHVVG-LFL. 

8.3 CCZ

Flares being used to control vent streams that have a consistently high combustible 

concentration can benefit from monitoring CCZ in lieu of either LFLCZ, or the ratio of NHVCZ to 

NHVVG-LFL (see Sections 3.2 and 3.3 of this report). The CCZ parameter requires either a gas 

chromatograph or hydrocarbon analyzer to determine the total combustible constituents in the 

flare vent gas, and flow meters to determine the volumetric flow rates of flare vent gas and total 

steam (if applicable). The total combustible constituents in the flare vent gas, and volumetric 
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Section 8.0: Monitoring Considerations Page 8-3

flow rates of flare vent gas and total steam (if applicable) are applied to Equation D.26 of 

Appendix D of this report to determine CCZ. 

8.4 SR 

Flow meters measuring mass flow rates of flare vent gas and assist air (includes air 

controls for manifold and valve instrumentation) would be required to determine the SR 

parameter using the equation and methods described in Section 4.1 and Appendix D of this 

report. For purposes of calculating the denominator of the SR parameter, a gas chromatograph 

would also be needed so as to determine the individual component concentrations present in the 

flare vent gas. 

8.5 MFR 

A meteorological station or anemometer (to measure cross-sectional wind velocity at the 

flare tip and the density of the ambient air) and flow meters (to determine the volumetric flow 

rates, velocity, and density of flare vent gas and total steam (if applicable)) would be required to 

determine the MFR parameter using Equation D.41 of Appendix D of this report. 
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Section 8.0: Monitoring Considerations Page 8-4

8.6 Vmax

The LFLvgcs parameter that is used in the Shore equation methodology discussed in 

Section 6.0 of this report requires a gas chromatograph to determine the individual component 

concentrations present in the flare vent gas, and flow meters to determine the volumetric flow 

rates of flare vent gas and total steam (if applicable). The individual component concentrations 

present in the flare vent gas, the pure component LFL values, and volumetric flow rates of flare 

vent gas and total steam (if applicable) are applied to Equation D.43 of Appendix D of this report 

to determine LFLvgcs. The LFLvgcs and Equation D.45 of Appendix D are then used to determine 

Vmax. The flow meters would also be required to determine the velocity and density of flare vent 

gas and total steam (if applicable). 
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Section 9.0: References Page 9-1

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Shore, D. 2007. Improving flare design, a transition from art-form to engineering science. 

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Smith, S. 2003. Flare system emissions control. Presented at the Texas Technology Conference 

by the Zeeco Company (March). 

Smoot, D. L., J.D. Smith, and R.E. Jackson. 2009. Technical foundations to establish new 

criteria for efficient operation of industrial steam-assisted gas flares. Presented at the 

International Flame Research Foundation 16th International Members Conference, Boston, MA 

(June). 

U.S. EPA. 2002. Air pollution control cost manual, 6th ed. EPA/452/B-02-001. 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not 

been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

Section 9.0: References Page 9-6

Van den Schoor, F., R.T.E. Hermanns, J.A. van Oijen, F. Verplaetsen, and L.P.H. de Goey. 

2008. Comparison and evaluation of methods for the determination of flammability limits, 

applied to methane/hydrogen/air mixtures. Journal of Hazardous Materials. 150:3, 573-81. 

Vidal, M., W. Wong, W.J. Rogers, M.S. Mannan. 2006. Evaluation of lower flammability limits 

of fuel-air-diluent mixtures using calculated adiabatic flame temperatures. Journal of Hazardous 

Materials. 130:1-2, 21-7. 

Wang, J., Z. Huang, C. Tang, H. Miao, and X. Wang. 2009. Numerical study of the effect of 

hydrogen addition on methane–air mixtures combustion. International Journal of Hydrogen 

Energy. 34:2, 1084-1096. 

Wang, T., Chen, C., & Chen, H. 2010. Nitrogen and carbon dioxide dilution effect on upper 

flammability limits for organic compound containing carbon, hydrogen and oxygen atoms. 

Journal of the Taiwan Institute of Chemical Engineers, 41(4), 453-464. Taiwan Institute of 

Chemical Engineers. 

White, A. G. 1925. Limits for the propagation of flame in inflammable gas/air mixtures, Part II: 

Mixtures of more than one gas and air. Journal of the Chemical Society, Transactions. 127:48, 

48. 

Wierzba, I., and V. Kilchyk. 2001. Flammability limits of hydrogen–carbon monoxide mixtures 

at moderately elevated temperatures. International Journal of Hydrogen Energy. 26:6, 639-643. 

Zabetakis, M.G. 1965. Flammability Characteristics of Combustible Gases and Vapors. U.S. 

Bureau of Mines, Bulletin 627. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

APPENDIX A 

Appendix A. Brief Review Summary of 

Each Flare Performance Study and Test Report 
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A-1 

A.1 Data Set A: EPA-600/2-83-052 (McDaniel, 1983)

Tests from data set A were conducted on pilot-scale test flares in June 1982 at the John 

Zink Company flare demonstration facility in Tulsa, Oklahoma. The gases used in this study 

included mixtures of propylene diluted with nitrogen (chemical composition for each test run, by 

test report, used in the steam data analysis described in this section 3.0 of this report is provided 

in Appendix E). The primary objectives for these tests were to determine the combustion 

efficiency and hydrocarbon destruction efficiency for both air- and steam-assisted flares under a 

wide range of operating conditions. The test methodology involved a special 27-foot sample 

probe suspended by a crane over the flare flame. The sample extracted by the probe was 

analyzed by continuous emission monitors to determine concentrations of carbon dioxide, carbon 

monoxide, total hydrocarbon, sulfur dioxide, nitrogen oxides, and oxygen. 

The report generally concluded that when flares operate under conditions representing 

good industrial operating practice, combustion efficiencies at the sampling probe were greater 

than 98 percent. Combustion efficiencies declined under conditions of excessive steam and high 

exit velocities of low heat content gases. 

There are 14 non-assisted, 15 steam-assisted, and 13 air-assisted test runs from this test 

report. The 14 non-assisted data points came from tests performed on the steam-assisted flare, 

but no steam was used during the test. The steam-assisted flare was a John Zink Standard STF-S8 flare tip (effective flare tip diameter of 5.86 inches). The recommended steam flow for this 

flare is approximately 0.4 pounds of steam per pound of propylene. The air-assisted flare was a 

John Zink STF-LH-457-5 flare (flare tip diameter unknown). The mass or volumetric flow rate 

of assist air as well as the assist air velocity were considered proprietary information and not 

included in the test report; therefore, the test data associated with this particular air-assisted flare 

were not used in any analysis that is discussed in this report. Both flares had two constant 

ignition pilots designed to burn a total of 300 standard cubic feet per hour of natural gas. 

 
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A-2 

A.2 Data Set B: EPA-600/2-84-095 (Pohl et al., 1984)

Tests from data set B were conducted at the Energy and Environmental Research 

Corporation’s El Toro, California test site. This study was limited to measuring the combustion 

efficiencies of 3-, 6-, and 12-inch pipe flares burning mixtures of propane diluted with nitrogen 

at steady operating conditions with and without steam injection, in the absence of wind. All 

steam injection was introduced as upper steam. 

A.3 Data Set C: EPA-600/2-85-106 (Pohl and Soelberg, 1985)

Tests from data set C were conducted at the Energy and Environmental Research 

Corporation’s El Toro, California test site as a continuation of the tests conducted under 

EPA-600/2-84-095. The gases used in this study included mixtures of propane diluted with 

nitrogen (chemical composition for each test run, by test report, used in the steam data analysis 

described in this section 3.0 of this report is provided in Appendix E). The primary objective of 

these tests was to determine the influence of flare head design on flare combustion efficiency. 

The study held proprietary the design of the different flare heads; however, some information 

was provided. A “coanda” steam-injected flare head was tested at gas exit velocities ranging 

from 0.2 to 9.9 feet per second based on a 12 inch diameter opening. Steam was injected into the 

flare waste stream through nozzles located above the main flare tip opening (upper steam). All 

measurements on this flare were made with a constant upper steam flow rate of 140 pounds per 

hour. All steam-assisted test runs were performed without pilots. Also, an air-assisted flare head 

was tested at gas exit velocities ranging from 8.5 to 428 feet per second based on a 1.5 inch 

diameter opening. The assist air flow rate was varied. Some air-assisted test runs were performed 

with pilots, but the majority were not. 

There are six steam-assisted and nine air-assisted test runs from this test report. This 

report also includes 16 pressure-assisted test runs (nine runs came from a 1.5 inch diameter 

pressure-assisted flare head and the other seven from a 3.8 inch diameter pressure-assisted flare 

head). 

 
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A-3 

A.4 Data Set D: MPC TX (Clean Air Engineering, Inc., 2010a)

PFTIR performance testing was conducted from September 15, 2009 to September 24, 

2009 on Marathon’s Texas City (MPC TX) refinery’s main flare in response to an EPA Office of 

Enforcement request pursuant to section 114 of the Clean Air Act. Industrial Monitor and 

Control Corporation (IMACC) performed the testing, and Clean Engineering, Inc. wrote the test 

report. Tests were conducted while flare vent gas contained saturates, olefins, nitrogen, and 

hydrogen mixtures (chemical composition for each test run, by test report, used in the steam data 

analysis described in this section 3.0 of this report is provided in Appendix E). For each test 

series, steam was increased from the manufacturer’s recommended minimum cooling steam rate 

to the point of snuffing the flare. For the majority of tests conducted, combustion efficiency 

declined with increasing steam at constant flare vent gas mass loading and constant composition. 

The Marathon Texas City main flare is an elevated steam-assisted flare. The flare has an 

effective tip diameter of 23.25 inches, was manufactured by Callidus Technologies (model BTZIS3/US-24-C), and was installed in December 2000. The tip has three points of steam addition: 

center steam, a lower steam ring, and an upper steam ring. Although the flare tip is equipped 

with upper steam, it was not used during any test runs; and all measurements on this flare were 

made with a constant center steam flow rate of 500 pounds per hour (Dickens 2011) and variable 

lower ring steam addition. The lower steam ring manifold piping has a connection to a small 

sweep gas ring used for shaping the flare at the tip. The flare has a manufacturer minimum total 

steam requirement of 1,250 pounds per hour in order to protect the flare tip. The flare operated 

with approximately 100 standard cubic feet per hour of pilot gas during the test runs. The typical 

base load for flare operation is approximately 1,100 to 2,000 pounds per hour of waste gas flow, 

or less than 0.25 percent of the hydraulic capacity of the flare (approximately a 250:1 turndown 

factor). 

Raw test data (in the form of an Excel worksheet) was available for this data set. There 

are 138 steam-assisted test runs from the raw test data. However, 56 of these test runs were noted 

in the test report as having a relatively low video score indicating that the PFTIR camera was 

lacking aiming accuracy during the test run (see Table A.1). 
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A-4 

Table A.1 Test Runs Explicitly Removed From Consideration 

Condition Run 

Number Reason for Omission 

Data Set D 

A19 

1-2 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

1-3 

3-2 

4-2 

5-1 

6-1 

7-2 

A8 

1-1 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-1 

2-2 

2-3 

3-1 

3-2 

4-1 

5-1 

5-2 

6-1 

7-1 

7-2 

8-1 

8-2 

9-1 

10-1 

B 

1-3 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-3 

3-3 

4-3 

6-3 

7-3 

10-1 

C 

1-3 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-3 

3-3 

4-1 

5-1 

D 7-1 Relatively low “video score” indicating that the PFTIR 

9-1 camera was lacking aiming accuracy during the test run. 

 
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A-5 

Table A.1 (Continued) 

Condition Run 

Number Reason for Omission 

Data Set D 

E 

1-2 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-2 

3-2 

4-2 

6-1 

F 

1-1 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-1 

3-1 

4-1 

5-1 

6-1 

G 

1-1 

Relatively low “video score” indicating that the PFTIR 

camera was lacking aiming accuracy during the test run. 

2-1 

3-1 

4-1 

5-1 

6-1 

LTS 2-7 Relatively low “video score” indicating that the PFTIR 

2-8 camera was lacking aiming accuracy during the test run. 

A.5 Data Set E: INEOS (INEOS ABS (USA) Corporation, 2010)

Pursuant to the consent decree (Civil Action No. 1:09-CV-545), PFTIR performance 

testing was conducted from November 3, 2009 to November 5, 2009 on INEOS ABS (USA) 

Corporation’s flare controlling their P001 process in Addyston, Ohio. IMACC performed the 

testing. The purpose of the evaluation was to determine the appropriate net heating value of flare 

vent gas necessary to assure the flare achieves 99 percent control efficiency. Tests were 

conducted while flaring gases containing various mixtures of 1,3-butadiene, natural gas, and 

nitrogen (chemical composition for each test run, by test report, used in the steam data analysis 

described in this section 3.0 of this report is provided in Appendix E). 

The flare tested is an elevated steam-assisted flare. The flare tip has an effective diameter 

of 16 inches and is manufactured by John Zink (model no. EEF-QS-16). Steam is injected into 
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A-6 

the flare waste stream through nozzles located above the main flare tip opening (upper steam). 

The manufacturer minimum total steam requirement to the flare tip was not provided in the test 

report. The flare operated with ~150 standard cubic feet per hour of pilot gas during the test runs. 

There are 21 steam-assisted test runs from this test report. 

A.6 Data Set F: MPC Detroit (Clean Air Engineering, Inc., 2010b)

PFTIR performance testing was conducted from July 8, 2010 to July 20, 2010 on 

Marathon’s Detroit (MPC Detroit) refinery’s Complex 3 and 4 flare in response to an EPA 

Office of Enforcement request pursuant to section 114 of the Clean Air Act. IMACC performed 

the testing. The main objective of the test was to better understand the effects of steam on the 

overall performance of the flare in terms of combustion efficiency. Tests were conducted while 

flaring gases containing refinery fuel gas, propylene, hydrogen, and nitrogen mixtures (chemical 

composition for each test run, by test report, used in the steam data analysis described in this 

section 3.0 of this report is provided in Appendix E). For each test series, steam was increased 

from the manufacturer’s recommended minimum cooling steam rate to the point of snuffing the 

flare. For the majority of tests conducted, combustion efficiency declined with increasing steam 

at constant flare vent gas mass loading and constant composition. 

Marathon’s Detroit refinery’s Complex 3 and 4 flare is an elevated steam-assisted flare. 

The flare tip has an effective diameter of 16 inches and is manufactured by NAO, Inc. (model no. 

20” NFF-RC) and was constructed in 1961 to 1962. The flare tip was replaced in October 2005. 

The flare has two points of steam addition: center steam and ring steam. The center and ring 

steam each have a manufacturer minimum steam requirement of 300 pounds per hour. All 

measurements on this flare were made with a constant center steam flow rate of 300 pounds per 

hour (and varying ring steam rates from 300 pounds per hour and up). The ring steam has 

alternating high and low points of injection around the flare tip exit. The flare operated with 

~135 standard cubic feet per hour of pilot gas during the test runs. The typical base load for 

typical flare operation is approximately 500 to 600 pounds per hour, or less than 0.25 percent of 

the hydraulic capacity (approximately a 400:1 turndown factor).
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A-7 

Raw test data (in the form of an Excel worksheet) were available for this data set. There 

are 62 steam-assisted test runs from the raw test data. However, six of these test runs were noted 

in the test report as invalid runs (see Table A.2). 

Table A.2 Test Runs Explicitly Removed From Consideration 

Condition Run 

Number Reason for Omission 

Data Set F 

A 

6-1 Invalid run, secondary (road location) PFTIR only, no 

combustion efficiency data. 

8-2 Invalid run, only 4 combustion efficiency readings. 

9-2 Invalid run, only 1 combustion efficiency reading. 

B 1-2 Invalid run, secondary (road location) PFTIR only, no 

combustion efficiency data. 

D 1-1 Invalid due to run instability. 

E 4-1 

Combustion efficiency for the E4-1 (N2≈66%) was 

measured with the secondary PFTIR at the road location. 

Because of hardware issues with this instrument the data 

is not reported and the run was marked invalid. 

A.7 Data Set G: Flint Hills Resources Aromatics Unit (FHR AU) and Light Olefins Unit 

(FHR LOU) (Clean Air Engineering, Inc., 2011)

PFTIR performance testing was conducted in October and November of 2010 at Flint 

Hills Resources, LLC in Port Arthur, TX on their Aromatics Unit and Light Olefins Unit flares in 

response to an EPA Office of Enforcement request pursuant to section 114 of the Clean Air Act. 

IMACC performed the testing. The main objective of the tests was to better understand the 

effects of steam on the overall performance of each flare in terms of combustion efficiency. Two 

additional operating parameters were also examined during this test program. The effect of 

hydrogen on combustion efficiency was studied on the FHR AU flare. The effect of flare vent 

gas flow rate on combustion efficiency was studied on the FHR LOU flare. 

The FHR AU flare is an elevated steam-assisted flare. The current tip has an effective 

diameter of 20 inches and was installed in 1996. This tip was manufactured by Callidus (model 
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A-8 

no. BTZ-US-16/20-C) and has two points of steam addition: center steam and ring steam. The 

center steam is injected in the flare vent gas stack prior to reaching the flare tip, and the ring 

steam is injected with nozzles around the flare tip rim. The center steam has a manufacturer 

minimum steam requirement of 500 pounds per hour. All measurements on this flare were made 

with a constant center steam flow rate of 500 pounds per hour (and varying ring steam rates). 

The flare operated with approximately 100 standard cubic feet per hour of pilot gas during the 

test runs. The typical FHR AU flare vent gas flow rate during normal operation is approximately 

800 pounds per hour, or less than 0.4 percent of the hydraulic capacity (approximately a 250:1 

turndown factor). 

The FHR LOU flare is an elevated steam-assisted flare. The current tip has an effective 

diameter of 54 inches and was installed in June 2010. This tip was manufactured by Callidus 

(model no. BTZ-1S3-54C) and has two points of steam addition: center steam and lower steam. 

The center steam is injected in the flare vent gas stack prior to reaching the flare tip, and the 

lower steam is injected through internal tubes interspersed throughout the flare tip. The center 

steam has a manufacturer minimum steam requirement of 2,890 pounds per hour; the lower 

steam has a manufacturer minimum steam requirement of 500 pounds per hour. All 

measurements on this flare were made with a constant center steam flow rate of 2,890 pounds 

per hour (and varying lower steam rates from 500 pounds per hour and greater). The flare 

operated with approximately 200 standard cubic feet per hour of pilot gas during the test runs. 

The typical FHR LOU flare vent gas flow rate during normal operation is approximately 

3,000 pounds per hour, or less than 0.3 percent of the hydraulic capacity (approximately a 333:1 

turndown factor). 

Raw test data (in the form of an Excel worksheet) were available for this data set. There 

are 45 steam-assisted test runs from the FHR AU flare raw test data. There are 46 steam-assisted 

test runs from the FHR LOU flare raw test data. Chemical composition for each test run, by test 

report, used in the steam data analysis described in this section 3.0 of this report is provided in 

Appendix E.

 
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A-9 

A.8 Data Set H: Shell Deer Park East Property Flare (SDP EPF) (Shell Global Solutions 

(US) Inc., 2011a)

PFTIR performance testing was conducted in March and April of 2010 on the Shell Deer 

Park refinery’s East Property Flare in response to an EPA Office of Enforcement request 

pursuant to section 114 of the Clean Air Act. IMACC performed the testing. The primary 

objective of the SDP EPF testing was to obtain a better understanding of how the ratio of steam 

to vent gas at the flare combustion zone and the hydrogen content of the flare vent gas affect the 

combustion efficiency. The composition of the flare vent gas during testing was consistent with 

normal operation of the SDP EPF; flare vent gas constituents consisted of hydrogen, nitrogen, 

and methane, with minor contributions from other low-molecular weight paraffinic, olefinic, and 

aromatic hydrocarbons (chemical composition for each test run, by test report, used in the steam 

data analysis described in this section 3.0 of this report is provided in Appendix E).

The SDP EPF is an elevated steam-assisted flare. The current tip has an effective 

diameter of 36 inches and was installed in May 1990. This tip was manufactured by John Zink 

(model no. EEF-QA-36-C) and has two points of steam addition: center steam and upper steam. 

The center steam has a manufacturer minimum steam requirement of 1,500 pounds per hour; the 

upper steam has a manufacturer minimum steam requirement of 500 pounds per hour. The flare 

operated with 150 standard cubic feet per hour of pilot gas during the test runs. Typical flows at 

the SDP EPF are approximately 4 percent of the smokeless capacity of the flare and 1 percent of 

the maximum relief capacity. 

Raw test data (in the form of an Excel worksheet) were available for this data set. There 

are 53 steam-assisted test runs from the raw test data. However, 11 of these test runs were noted 

in the test report as being part of an experiment for comparing PFTIR detectors and were 

removed from consideration (see Table A.3). 

 
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A-10 

Table A.3 Test Runs Explicitly Removed From Consideration 

Condition Run 

Number Reason for Omission 

Data Set H 

[EP-D-2.0(MCT1)] “Without a beam splitter IMACC could not use a MCT 

detector and an InSb detector at the same time. Since SDP 

has very limited means to control the composition of the 

SDP EPF vent gas, it was also impractical to recreate test 

conditions identical to those that occurred when SDP 

collected data with the InSb detector. Given these 

constraints, the only way to compare the MCT detector to 

the InSb detector was to gather as many data points as 

practical with the MCT detector at similar, but not 

identical, conditions as occurred when data were gathered 

with the InSb detector, and then compare the two data 

sets.” 1

[EP-D-3.0(MCT1)] 

[EP-D-4.0(MCT1)] 

[EP-D-5.0(MCT1)] 

[EP-D-4.0(MCT2)] 

[EP-D-5.0(MCT2)] 

[EP-D-2.0(MCT2)] 

[EP-D-3.0(MCT2)] 

[EP-D-4.0(MCT3)] 

[EP-D-5.0(MCT3)] 

[EP-D-4.5(MCT)] 

1 See page 19 of Section 2.1.4 of Shell Global Solutions (US) Inc., 2011a. 

2.9 Data Set I: Shell Deer Park Ground Flare (SDP GF) (Shell Global Solutions (US) 

Inc., 2011b)

AFTIR performance testing was conducted in March and April of 2010 on the Shell Deer 

Park refinery’s Olefin Plant 3 Ground Flare in response to an EPA Office of Enforcement 

request pursuant to section 114 of the Clean Air Act. These tests were conducted with assistance 

from both Shell Global Solutions (US) Inc. and IMACC personnel. The primary objective of the 

SDP GF testing was to obtain a better understanding of how the ratio of steam to vent gas at the 

flare combustion zone and the hydrogen content of the flare vent gas affect combustion 

efficiency. The primary components of the SDP GF flare vent gas are hydrogen, nitrogen, 

methane, ethylene, propylene, and small quantities of higher molecular weight olefins, di-olefins, 

and aromatics. 

The SDP GF is a multistage steam-assisted enclosed ground flare with three different 

stages, which become active at successively higher flows. The SDP GF was installed in 1978 and 

has 92 burners (basically a refractory lined steel shell into which 92 raw flare vent gas burners 

discharge). The flare has three stages. The first stage has eight, horizontally-mounted burners and 

eight pilots. The second stage has 24 horizontally-mounted burners and 24 pilots, while the last 
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A-11 

stage has 60, horizontally-mounted burners and 30 pilots. The enclosed ground flare has eight 

faces. The horizontally-mounted burners are arranged in identical vertical drops, with four 

burners to a drop. The flare operated with a total of 150 standard cubic feet per hour pilot gas 

during the test runs. 

Each of the 92 burners is manufactured by John Zink (model no. ZTOF SM-10 burners) 

and equipped with a steam injection system to inject high-pressure steam for cooling, to aid 

combustion, and to reduce smoke formation. A small amount of cooling steam is always routed 

to each of the SDP GF’s 92 burners whether each burner has flare vent gas going to it or not. 

Steam enters the center of each burner via a “1” steam tip. The cooling steam is controlled by a 

restriction orifice in a by-pass line around the steam valve to each stage. This results in 

approximately 870 pounds per hour of cooling steam to the 92 burners, or approximately 

9.46 pounds per hour per burner. At the time of testing, typical flow at the SDP GF was 

approximately 5,000 pounds per hour, or 5 percent of the maximum smokeless capacity. The 

flare vent gas flows used for the test runs required only the eight first stage burners. 

Raw test data (in the form of an Excel worksheet) were available for this data set. There 

are 21 steam-assisted test runs from the raw test data. 

2.10 Data Set J: TCEQ (Allen and Torres, 2011)

In May 2009, the TCEQ contracted with The University of Texas at Austin to conduct a 

comprehensive flare study project on full-scale steam- and air-assist flares at the John Zink 

Company flare demonstration facility in Tulsa, Oklahoma. The gas used in this study included 

mixtures of propylene or propane supplemented with natural gas and diluted with nitrogen 

(chemical composition for each test run, by test report, used in the steam data analysis described 

in this section 3.0 of this report is provided in Appendix E). The purpose of the project was to 

conduct field tests to measure flare emissions and collect process and operational data in a semicontrolled environment to determine the relationship between flare design, operation, flare vent 

gas lower heating value and flow rate, destruction efficiency, and combustion efficiency. 
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A-12 

Aerodyne Research, Inc., (ARI) was contracted to directly measure flare emissions at the 

end of the flare plume using extractive techniques, and calculate destruction and combustion 

efficiency based on those measurements. The study also evaluated the performance of remote 

sensing technologies against the extractive technique. These remote sensing technologies 

included Infrared Hyper-Spectral Imaging Technology (Contractor: Telops Inc.), PFTIR and 

AFTIR (Contractor: IMACC), and Forward Looking Infrared (FLIR) GasFindIR Passive 

Infrared (IR) Cameras (Contractor: Leak Surveys Inc.). The test report for data set J concludes 

that the mean difference and standard deviation of the reported AFTIR and PFTIR combustion 

efficiency values increase as the reported extractive combustion efficiency values decreases; 

however, both the AFTIR and PFTIR methods actually compare very well to the extractive test 

results for combustion efficiencies reported as 90 percent or greater. For combustion efficiencies 

reported as 90 percent or greater, the test report for data set J states that the mean difference of 

combustion efficiency values is less than 3.2 percent between extractive and AFTIR, and less 

than 3.5 percent between extractive and PFTIR. 

The steam-assisted flare used in this study was a John Zink model EE-QSC-36” flare tip 

(flare tip effective diameter of 36 inches) with three EEP-503 pilots, and a maximum capacity of 

937,000 pounds per hour. The center steam has a manufacturer minimum steam requirement of 

300 pounds per hour; the upper steam has a manufacturer minimum steam requirement of 

525 pounds per hour. The air-assisted flare was a John Zink model LHTS-24/60 flare tip (flare 

tip effective diameter of 24 inches) with three pilots, and a maximum capacity of 144,000 

pounds of propylene per hour. Both flares operated with 225 standard cubic feet per hour of pilot 

gas during the test runs. 

Raw test data (in the form of an Excel worksheet) were available for this data set. There 

are 131 steam-assisted test runs and 89 air-assisted test runs from the raw test data. 
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APPENDIX B 

Appendix B is an Excel workbook that combines all data sets. The Excel workbook 

identifies each specific test run by the exact test condition and run identification used in 

each individual report. Appendices C and D provide explanation and methodologies for 

how certain data fields in the workbook were determined. 
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guidelines. It has not been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency 

determination or policy.

APPENDIX C 

Test Report Nomenclature Matrix 
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represent and should not be construed to represent any Agency determination or policy.

C-1 

Table C-1. Test Report Nomenclature Matrix 

KEY: “N/A” – not applicable. The specific parameter was not considered in the data set; or the specific parameter was 

“Unknown” –unable to calculate the specific parameter because there was not enough information provided in thdata set. “Eq. D.##” –details from the test report and associated data set used to calculate the specific parameter. See App

methodology; refer to the specific Equation identified in the matrix. 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Basic Run Info Reported Combustion Efficiency (CE) [%] "Combustion Efficiency (%)" "Combustion Efficiency (Percent)" "Comb Eff. (%)" "CE" Run Avg. Relative Efficiency using Carbon Count "CE (weighted) %" "CE (weighted) %" "CE (weighted) %" "CE" General Flare Vent Gas Data 

Flare Vent Gas 

Net Heating 

Value (NHVVG)

[Btu/scf] 

"Lower 

Heating 

Value 

(Btu/SCF)" 

"Low 

Heating 

Value 

(Btu/ft3)" 

"Low Htg 

Val 

(Btu/ft3)" 

"Vent Gas 

HV" Eq. D.1 

"Flare Gas 

NHV 

BTU/scf" 

"VG NHV (calc) 

BTU/scf" 

"VG NHV 

(calc) 

BTU/scf" 

"East Property 

Gas Net Hea

Value 

(BTU/SCF

Flare Vent Gas 

Flow Rate 

(mvg) 

[lb/hr] 

Eq. D.2 N/A Eq. D.3 

"Vent Gas 

Flow Rate 

(lb/h)" 

"Total 

Panametrics 

Flow in lbs/hr" 

"Flare Gas 

Flow lb/hr" 

"Flare Gas 

Flow (DCS) 

lb/hr" 

"Flare Gas 

Flow (DCS) 

lb/hr" 

”East Prop

Flare- Calcu

Mass Flow (l

Flare Vent Gas 

Flow Rate 

(Qvg) 

[scf/hr] 

Eq. D.4 N/A Eq. D.5 

"Vent Gas 

Flow Rate 

(scfh)" 

Eq. D.6 

"Std Flare 

Gas Flow 

scf/hr" 

Eq. D.7 "Std VG Flow (calc) scf/hr" Eq. D.6 

Flare Vent Gas 

Velocity 

(Vvg) [ft/s] 

Eq. D.8A 

"Actual Exit 

Velocity 

(ft/sec)" 

"Actual 

Exit 

Velocity 

(ft/sec)" 

"Flare Tip 

Velocity" Eq. D.8A "Flare Tip Velocity ft/s" "VG Exit Velocity (calc) ft/s" "Tip Velocity (no steam) (calc) ft/s" Vent gas vel

report in test 
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represent and should not be construed to represent any Agency determination or policy.

C-2 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

General Flare Vent Gas Data 

Flare Vent Gas 

Velocity w/ 

Center Steam 

(Vvg-s) 

[ft/s] 

Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C Eq. D.8C 

Flare Vent Gas 

Molecular 

Weight 

(MWvg) 

[lb/lb-mol] 

Eq. D.9 N/A Eq. D.9 "MWvg" "Panametrics Measured MW" "MWvg (GC calc) lb/lbmol" "MWvg (calc) lb/lb-mol" "MWvg (calc) lb/lb-mol" "Vent Gas M

(lb/lb-mol

Reported Pilot 

Gas Flow Rate 

(Qpg) 

[scf/hr] 

Reported as 

26.3 lb/hr in 

Test Report 

No pilot gas 

Reported 

as 2.1 

scf/min for 

some runs 

in Test 

Report 

Reported as 

100 scf/hr in 

Test Report 

Reported as 

150 scf/hr in 

Test Report 

Reported as 

135 scf/hr in 

Test Report 

"Pilot Gas Flow 

Rate (scf/hr)" 

"Pilot Gas 

Flow Rate 

(scf/hr)" 

No pilot g

Molar Percentage of Components in Flare Vent Gas [mol %] 

1-Butene 

[mol %] N/A N/A N/A "Butenes mol%" N/A "i-Butene, Butene-1 mol %" "Butene-1 (bag)" N/A "HRVOC E

FLARE_Buten

obutane

1,2-Butadiene 

[mol %] N/A N/A N/A N/A N/A N/A "1,2-Butadiene (bag)" N/A N/A 

1,3-Butadiene 

(%1,3-B) 

[mol %] 

N/A N/A N/A "Butadiene mol%" Eq. D.10 "1,3- Butadiene mol %" "1,3-Butadiene (bag)" "Butadiene (GC) mol%" "HRVOC E

FLARE_13B

Acetylene 

[mol %] N/A N/A N/A "Acetylene mol%" N/A "Acetylene mol %" "Acetylene (bag)" N/A "HRVOC E

FLARE_Acety

Benzene [mol %] N/A N/A N/A N/A N/A N/A "Benzene (bag)" "Benzene (GC) mol%" N/A 

Carbon Dioxide 

[mol %] 

N/A N/A N/A "CO2 mol%" N/A 

"Carbon 

Dioxide mol 

%" 

"Carbon 

Dioxide (bag)" 

"Carbon 

Dioxide (GC) 

mol%" 

"HRVOC E

FLARE_Car

Dioxide

Carbon 

Monoxide 

[mol %] 

N/A N/A N/A "CO mol%" N/A 

"Carbon 

Monoxide 

mol %" 

N/A 

"Carbon 

Monoxide 

(GC) mol%" 

"HRVOC E

FLARE_Car

Monoxide
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represent and should not be construed to represent any Agency determination or policy.

C-3 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Molar Percentage of Components in Flare Vent Gas [mol %] 

Cis-2-Butene 

[mol %] N/A N/A N/A 

"Cis-2-

Butene 

mol%" 

N/A "Cis-Butene- 2 mol %" "Cis-Butene-2 (bag)" "2-Butene (GC) mol%" "HRVOC E

FLARE_Cis2B

Cyclopropane 

[mol %] N/A N/A N/A N/A N/A N/A "Cyclopropane (bag)" N/A N/A 

Ethane 

[mol %] N/A N/A N/A "Ethane mol%" N/A "Ethane mol %" "Ethane (bag)" "Ethane (GC) mol%" "HRVOC E

FLARE_Eth

Ethyl Benzene

[mol %] N/A N/A N/A N/A N/A N/A N/A 

"Ethylbenzene (GC) 

mol%" 

N/A 

Ethylene 

[mol %] N/A N/A N/A "Ethylene mol%" N/A "Ethylene mol %" "Ethylene (bag)" "Ethylene (GC) mol%" "HRVOC E

FLARE_Ethy

Hydrogen 

[mol %] N/A N/A N/A "H2 mol%" N/A "Hydrogen mol %" "Hydrogen (bag)" "Hydrogen (GC) mol%" "HRVOC E

FLARE_Hydr

Hydrogen 

Sulfide 

[mol %] 

N/A N/A N/A N/A N/A N/A N/A N/A 

"HRVOC E

FLARE_Hyd

Sulfide" 

Iso-Butane 

[mol %] N/A N/A N/A "Iso-Butane mol%" N/A "Iso-Butane mol %" "Isobutane (bag)" "i-Butane (GC) mol%" "HRVOC E

FLARE_IsoBu

Iso-Butylene 

[mol %] N/A N/A N/A N/A N/A N/A "Isobutylene (bag)" "iso-butylene (GC) mol%" N/A 

Methane 

[mol %] N/A N/A N/A "CH4 mol%" Eq. D.11 "Methane mol %" "Methane (bag)" "Methane (GC) mol%" "HRVOC E

FLARE_Meth

Methyl 

Acetylene 

[mol %] 

N/A N/A N/A N/A N/A N/A "Methyl Acetylene (bag)" N/A N/A 

n-Butane 

[mol %] N/A N/A N/A "N-Butane mol%" N/A "Normal Butane mol %" "N-Butane (bag)" "n-Butane (GC) mol%" "HRVOC E

FLARE_But
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represent and should not be construed to represent any Agency determination or policy.

C-4 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Molar Percentage of Components in Flare Vent Gas [mol %] 

Nitrogen 

(%N) 

[mol %] 

Eq. D.14 N/A Eq. D.13 "N2 mol%" Eq. D.12 "Nitrogen mol %" "Nitrogen (bag)" "Nitrogen (GC) mol%" "HRVOC E

FLARE_Nito

Oxygen 

[mol %] N/A N/A N/A "O2 mol%" N/A "Oxygen mol %" "Oxygen (bag)" "Oxygen (GC) mol%" "HRVOC E

FLARE_Oxy

Pentane Plus 

(C5+) 

[mol %] 

N/A N/A N/A "C5+ mol%" N/A 

"PentanePlus (C5+) 

mol %" 

"C5& Heavier Hydrocarbons 

(bag)" 

"Pentane 5+ 

(GC) mol%" 

"HRVOC E

FLARE_C5

Propadiene 

[mol %] N/A N/A N/A N/A N/A N/A "Propadiene (bag)" N/A N/A 

Propane 

(%p) 

[mol %] 

N/A N/A 

“% 

Propane 

in 

Nitrogen” 

"Propane 

mol%" N/A "Propane mol %" "Propane (bag)" "Propane (GC) mol%" "HRVOC E

FLARE_Prop

Propylene 

[mol %] Eq. D.14 N/A N/A "Propylene mol%" N/A "Propylene mol %" "Propylene (bag)" "Propylene (GC) mol%" "HRVOC E

FLARE_Prop

Toluene 

[mol %] N/A N/A N/A N/A N/A N/A N/A "Toluene (GC) mol%" N/A 

Trans-2-

Butene 

[mol %] 

N/A N/A N/A 

"Trans-2-

Butene 

mol%" 

N/A 

"TransButene-2 mol 

%" 

"Trans-Butene-2 

(bag)" N/A 

"HRVOC E

FLARE_Trans

e" 

Water 

[mol %] N/A N/A N/A "Water mol%" N/A N/A N/A N/A "HRVOC E

FLARE_Wa

Xylenes 

[mol %] N/A N/A N/A N/A N/A N/A N/A N/A N/A 
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represent and should not be construed to represent any Agency determination or policy.

C-5 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Steam Data 

TOTAL 

Volumetric 

Steam Rate 

(Qs) 

[scf/hr] 

Eq. D.18 N/A Eq. D.18 Eq. D.18 Eq. D.18 Eq. D.18 "Std Steam Flow (calc) scf/hr" "Std Steam Flow (calc) scf/hr" Eq. D.18 

TOTAL 

Mass Steam 

Rate 

(ms) 

[lb/hr] 

“Steam Flow 

(lbs/hr)” N/A 140 lb/hr Constant "Steam Flow (lb/h)" Eq. D.19 "Steam Flow lb/hr" "Steam Flow (DCS) lb/hr" Eq. D.17 Eq. D.17 

Lower 

Volumetric 

Steam Rate 

(Ql) 

[scf/hr] 

N/A N/A N/A Unknown N/A N/A N/A Eq. D.20 N/A 

Lower Mass 

Steam Rate 

(ml) 

[lb/hr] 

N/A N/A N/A Unknown N/A N/A N/A Eq. D.19 N/A 

Center 

Volumetric 

Steam Rate 

(Qc) 

[scf/hr] 

N/A N/A N/A Unknown N/A Eq. D.18 Eq. D.18 Eq. D.18 Unknow

Center Mass 

Steam Rate 

(mc) 

[lb/hr] 

N/A N/A N/A Unknown N/A 300 lb/hr Constant 500 lb/hr Constant 2,890 lb/hr Constant Unknow

Ring Volumetric 

Steam Rate 

(Qr) 

[scf/hr] 

N/A N/A N/A N/A N/A Eq. D.20 Eq. D.20 N/A N/A 

Ring Mass 

Steam Rate 

(mr) 

[lb/hr] 

N/A N/A N/A N/A N/A Eq. D.19 Eq. D.19 N/A N/A 

Upper 

Volumetric 

Steam Rate 

(Qu) 

[scf/hr] 

Eq. D.20 N/A Eq. D.18 N/A Eq. D.18 N/A N/A N/A Unknow
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represent and should not be construed to represent any Agency determination or policy.

C-6 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Steam Data 

Upper Mass 

Steam Rate 

(mu) 

[lb/hr] 

“Steam Flow 

(lbs/hr)” N/A 140 lb/hr Constant N/A "Steam flow in lbs/hr primeter" N/A N/A N/A Unknow

Steam-to-ventgas ratio by 

weight 

(S/VGwt) 

[lb steam/ 

lb vent gas] 

Eq. D.21A N/A Eq. D.21A Eq. D.21A Eq. D.21A Eq. D.21A Eq. D.21A Eq. D.21A Eq. D.21

Steam-to-ventgas ratio by 

volume 

(S/VGvol) 

[scf steam/ 

scf vent gas] 

Eq. D.21B N/A Eq. D.21B Eq. D.21B Eq. D.21B Eq. D.21B Eq. D.21B Eq. D.21B Eq. D.21

Steam-tohydrocarbon 

ratio by 

volume 

(S/HCvol) 

[scf steam/ 

scf HC] 

Eq. D.21C N/A Eq. D.21C Eq. D.21C Eq. D.21C Eq. D.21C Eq. D.21C Eq. D.21C Eq. D.21

Air Data 

Air Assist Flow 

Rate 

(mAir) 

[lb/hr] 

N/A N/A N/A N/A N/A N/A N/A N/A N/A 

Air Assist Flow 

Rate 

(QAir) 

[scf/hr] 

N/A N/A 

"Air Assist 

Flow Rate 

SCFH" 

N/A N/A N/A N/A N/A N/A 

Volumetric Flow 

Rates For Each 

Flare Vent Gas 

Component In The 

Combustion Zone 

(QCZ,i) 

[scf/hr] 

Eq. D.15B N/A Eq. D.15A Eq. D.15A Eq. D.15C and Eq. D.15D Eq. D.15A Eq. D.15A Eq. D.15A Eq. D.15

Concentration Of 

Each Flare Vent 

Gas Component In 

The Combustion 

Zone 

(XCZ,i) 

[volume fraction] 

Eq. D.16A N/A Eq. D.16A Eq. D.16A Eq. D.16A Eq. D.16A Eq. D.16A Eq. D.16A Eq. D.16
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represent and should not be construed to represent any Agency determination or policy.

C-7 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Concentration of 

Each Flare Vent 

Gas Combustible 

Component 

(VGc) 

[volume fraction] 

Eq. 16B Eq. 16B Eq. 16B Eq. 16B Eq. 16B Eq. 16B Eq. 16B Eq. 16B Eq. 16B 

Concentration of 

Each Flare Vent 

Gas and Center 

Steam Only Component 

 (Xvgcs) 

[volume fraction] 

Eq.16C Eq.16C Eq.16C Eq.16C Eq.16C Eq.16C Eq.16C Eq.16C Eq.16C 

Dynamic NHVCZ Calculation 

Ratio of 

NHVcz to 

NHVLFL-VG

(NHVratio) 

[unitless] 

Eq. D.22 N/A Eq. D.22 Eq. D.22 Eq. D.22 Eq. D.22 Eq. D.22 Eq. D.22 Eq. D.22 

Net Heating 

Value of 

Combustion 

Zone Gas 

(NHVCZ)

[BTU/scf] 

Eq. D.23 N/A Eq. D.23 Eq. D.23 Eq. D.23 Eq. D.23 Eq. D.23 Eq. D.23 Eq. D.23 

Net Heating 

Value of Vent 

Gas if Diluted 

to Lower 

Flammability 

Limit 

(NHVVG--LFL) 

[BTU/scf] 

Eq. D.24 N/A Eq. D.24 Eq. D.24 Eq. D.24 Eq. D.24 Eq. D.24 Eq. D.24 Eq. D.24 

Lower 

Flammability Limit of Vent 

Gas Adjusted 

for 

Nitrogen 

Equivalency 

(LFLVG) 

[vol frac]] 

Eq. D.25 N/A Eq. D.25 Eq. D.25 Eq. D.25 Eq. D.25 Eq. D.25 Eq. D.25 Eq. D.25 
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represent and should not be construed to represent any Agency determination or policy.

C-8 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

LFLCZG calculation 

Combustible 

Components in 

the 

Combustion 

Zone 

(CCZ) 

[vol frac] 

Eq. D.26 N/A Eq. D.26 Eq. D.26 Eq. D.26 Eq. D.26 Eq. D.26 Eq. D.26 Eq. D.26 

Coefficient of 

Nitrogen 

Equivalency of 

Water Relative 

to Nitrogen 

(Ne,H2O) 

Eq. D.29 N/A Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 

Coefficient of 

Nitrogen 

Equivalency of 

Carbon Dioxide 

Relative to 

Nitrogen 

(Ne,CO2) 

Eq. D.29 N/A Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 Eq. D.29 

Lower 

Flammability 

Limit of 

Combustion 

Zone Gas 

Adjusted for 

Nitrogen 

Equivalency 

(LFLCZ) 

[vol frac] 

Eq. D.27 N/A Eq. D.27 Eq. D.27 Eq. D.27 Eq. D.27 Eq. D.27 Eq. D.27 Eq. D.27 
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represent and should not be construed to represent any Agency determination or policy.

C-9 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Stoichiometric Air Ratio Calculation 

Theoretical 

Stoichiometric 

Mass of Air 

Needed to 

Combust the 

Flare Vent Gas 

(mStoic) [lb/hr] 

N/A N/A 

Eq. D.31 

through 

Eq. 37 

N/A N/A N/A N/A N/A N/A 

Calculated 

Ratio of Actual 

Mass Flow of 

Total Assist 

Air to the 

Theoretical 

Stoichiometric 

Mass of Air 

Needed to 

Combust the 

Flare Vent Gas 

(mAir / mStoic) 

(SR) 

N/A N/A Eq. D.30 N/A N/A N/A N/A N/A N/A 

Reported 

(mAir / mStoic) 

(SR) 

N/A N/A "Air- Assist SR" N/A N/A N/A N/A N/A N/A 
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C-10 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

MFR Calculation 

Reported MFR 

[unitless] N/A N/A N/A "MFR" N/A "Momentum Flux Ratio" "Momentum Flux Ratio (calc)" "Momentum Flux Ratio (calc)" N/A 

Flare Vent Gas 

Flow Rate 

Including 

Center Steam 

(mvg-s) 

[lb/hr] 

Eq. D.38 N/A N/A Eq. D.38 Eq. D.38 Eq. D.38 Eq. D.38 Eq. D.38 N/A 

Density of 

Flare Vent Gas 

Including 

Center Steam 

(pvg-s) 

[lb/scf] 

Eq. D.39 N/A N/A Eq. D.39 Eq. D.39 Eq. D.39 Eq. D.39 Eq. D.39 N/A 

Wind Velocity 

(Vwind) 

[ft/s] 

N/A N/A N/A Eq. D.40 N/A Eq. D.40 Eq. D.40 Eq. D.40 N/A 

Calculated 

MFR 

[unitless] 

N/A N/A N/A Eq. D.41 N/A Eq. D.41 Eq. D.41 Eq. D.41 N/A 
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C-11 

Table C-1. Test Report Nomenclature Matrix (Continued) 

Data Set ID A: EPA 2-83-052 B: EPA 2-84-095 C: EPA 2-85-106 D: MPC TX E: INEOS F: MPC Detroit G1: FHR AU G2: FHR LOU H: SDP EPF 

Shore Equation Analysis 

Shore 

Equation 

Velocity-Jet 

Periphery 

(UO/∏/DO) 

[sec-1] 

Eq. D.42 N/A Eq. D.42 Eq. D.42 Eq. D.42 Eq. D.42 Eq. D.42 Eq. D.42 N/A 

Lower 

Flammability 

Limit of Flare 

Vent Gas 

Including 

Center Steam 

Adjusted for 

Nitrogen 

Equivalency 

(LFLvgcs) 

[vol frac] 

Eq. D.43 N/A Eq. D.43 Eq. D.43 Eq. D.43 Eq. D.43 Eq. D.43 Eq. D.43 N/A 

Shore 

Equation 

(RAM) 

[lb flare vent 

gas including 

center steam / 

lb of air] 

Eq. D.44 N/A Eq. D.44 Eq. D.44 Eq. D.44 Eq. D.44 Eq. D.44 Eq. D.44 N/A 

Maximum 

Allowable 

Flare Vent Gas 

Velocity 

Including 

Center Steam 

(Vmax) 

[ft/s] 

Eq. D.45 N/A Eq. D.45 Eq. D.45 Eq. D.45 Eq. D.45 Eq. D.45 Eq. D.45 N/A 
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APPENDIX D 

Appendix D is a detailed list of all equations used in the Excel file. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-1

APPENDIX D – CALCULATIONS 

INTRODUCTION 

Appendix D provides detailed calculation methodologies for the specific parameters identified in 

the Appendix C matrix. If a cell is labeled as “Eq. D.##” in the Appendix C matrix, the specific 

calculation methodology can be found here in Appendix D. As such, the equations presented in 

Appendix D were used only for the indicated data sets. See Table 1 in Section 2.0 of the memo 

to identify the test report associated with each data set. If a particular data set is not identified in 

the equation in this appendix, it means that the corresponding parameter was included in the test 

report (and no calculation was necessary). 

 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-2

 

a
K6b  c E%W@,$

100 J a
K$

f

$g

Equation D.1 

The flare vent gas net heating value (NHVVG), in Btu/scf, was calculated using Equation 

D.1.

This equation was only used for data set E. The reported NHVVG from the associated test report 

was used for all other data sets. It was assumed that the reported values used this equation as a 

basis. The net heating value of the flare vent gas is the sum of the net heat of combustion of each 

individual combustible component. 

Where: 

NHVVG = Flare vent gas net heating value, BTU/scf. 

i = Individual combustible component in flare vent gas. 

n = Number of individual combustible components in flare vent gas. 

%vg,i = Percentage of combustible component i in flare vent gas (i.e., 1,3-Butadiene and 

Natural Gas), volumetric percentage. (See Equations D.10 through D.12). 

100 = Constant, percentage. 

NHVi = Reported net heating value of combustible component i, BTU/scf.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-3

hW@  hi  hj

hW@  .kW@ A E lW@

379.48J

Equations D.2 and D.3 

If not reported, the flare vent gas flow rate (mvg), in lb/hr, was calculated using either 

Equation D.2 or D.3. 

Equation D.2:

This equation was used for data set A. 

Where: 

mvg = Flare vent gas flow rate, lb/hr. 

mp = Reported mass flow rate of combustible components (i.e., Propylene Flow) in 

flare vent gas, lb/hr. 

mN2 = Reported mass flow rate of nitrogen (i.e., Nitrogen Flow) in flare vent gas, lb/hr. 

Equation D.3:

This equation was used for data set C. 

Where: 

mvg = Flare vent gas flow rate, lb/hr. 

MWvg = Molecular weight of flare vent gas, lb/lb-mole. (See Equation D.9). 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.4). 

379.48 = Constant, lb-mole/scf at dry standard conditions. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-4

lW@  60 A 0li  lj5

lW@  379.48 A o

Z

hY

/\

E

%i

100J  Z%j100\

p

Equations D.4 through D.7 

If not reported, the flare vent gas flow rate (Qvg), in scf/hr, was calculated using one of the 

Equations D.4 through D.7. 

Equation D.4:

This equation was used for data set A. 

Where: 

Qvg = Flare vent gas flow rate, scf/hr. 

60 = Constant, min/hr. 

Qp = Reported volumetric flow rate of propylene (i.e., Propylene Flow), scf/min. 

QN2 = Reported volumetric flow rate of nitrogen (i.e., Nitrogen Flow), scf/min. 

Equation D.5:

This equation was used for data set C. 

Where: 

Qvg = Flare vent gas flowrate, scf/hr. 

379.48 = Constant, lb-mole/scf at dry standard conditions. 

ms = Reported total steam rate (i.e., Steam Flowrate), lb/hr. 

SF = Reported steam ratio (i.e., Steam Ratio), lb steam / lb fuel. 

%p = Reported percentage of propane in flare vent gas (i.e., %Propane in Nitrogen), 

molar percentage. 

%N = Percentage of nitrogen in flare vent gas, molar percentage. (See Equation D.13). 

100 = Constant, molar percentage. 
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D-5

 

lW@  379.48 A q hW@

.kW@r

lW@  lW@,%si%tsu A 1,000

Equations D.4 through D.7 (Continued) 

Equation D.6:

This equation was used for data sets E, H, and J. 

Where: 

Qvg = Flare vent gas flow rate, scf/hr. 

379.48 = Constant, lb-mole/scf at dry standard conditions. 

mvg = Reported flare vent gas flow rate, lb/hr. 

 Reported for data set E as “Total Panametrics Flow in lbs/hr.” 

 Reported for data set H as “East Property Flare – Calculated Mass Flow 

(lb/hr).” 

 Reported for data set J as “Actual Vent Gas (VG) Flow Rates, Total (lb/hr).”

MWvg = Reported molecular weight of flare vent gas, lb/lb-mole. 

 Reported for data set E as “Panametrics Measured MW.” 

 Reported for data set H as” East Property Flare Gas Molecular Weight 

(lb/lbmole).” 

 Reported for data set J as “Actual Vent Gas Mol Wt (lb/lb-mole).”

Equation D.7:

This equation was used for data set G1. 

Where: 

Qvg = Flare vent gas flow rate, scf/hr. 

Qvg.reported = Reported flare vent gas flow rate (i.e., Vol Flare Gas Flow (DCS)), kscf/hr. 

1,000 = Constant, scf/kscf. 
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D-6

 

KW@ 

KW@,%si%tsu

60

KW@  KW@vY U

Q

w

R lX

Z

M

4

\ A Exyt

12J



A 60 A 60_

z

`

Equation D.8A and D.8B 

If not reported, the flare vent gas velocity (Vvg), in ft/s, was calculated using Equation D.8A 

or Equation D.8B. 

Equation D.8A:

This equation was used for data sets A and E. 

Where: 

Vvg = Flare vent gas velocity, ft/s. 

Vvg,reported = Reported flare vent gas velocity, ft/min. 

 Reported for data set A as “Relief Gas Flow (SCFM).” 

 Reported for data set E as “Exit Velocity in ft/min.”

60 = Constant, s/min. 

Equation D.8B:

This equation was used for data set J. 

Where: 

Vvg = Flare vent gas velocity, ft/s. 

Vvg-s = Reported flare vent gas velocity with center steam (i.e., Actual Vent Gas Exit 

Velocity (fps), Includes Center Steam), ft/s. 

Qc = Center steam rate, scf/hr. (See Equation D.18 for data sets F, G1 and G2). 

π/4 = Constant, used to calculate area of flare tip. 

dft = Reported effective diameter of flare tip, in. (See tip sizes identified in Section 2.0 

of report). 

12 = Constant, in/ft. 

60 = Constant, s/min. 

60 = Constant, min/hr. 
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D-7

 

KW@vY  KW@ 

Q

w

R lX

Z

M

4

\ A Exyt

12J



A 60 A 60_

z

`

Equation D.8C 

If not reported, the flare vent gas velocity with center steam (Vvg-s), in ft/s, was calculated 

using Equation D.8C. For data sets A, B, C, E, and I that are without a center steam rate 

(Qc) or for data set H that has an unknown center steam rate (Qc), the flare vent gas 

velocity with center steam (Vvg-s) is equal to flare vent gas velocity (Vvg). 

Equation D.8C:

This equation was used for data sets A, B, C, D, E, F, G1, G2, H and I. 

Where: 

Vvg-s = Flare vent gas velocity with center steam, ft/s. 

Vvg = Flare vent gas velocity, ft/s. (See Equation D.8A for data sets A and E). 

Qc = Center steam rate, scf/hr. (See Equation D.18 for data sets F, G1 and G2). 

π/4 = Constant, used to calculate area of flare tip. 

dft = Reported effective diameter of flare tip, in. (See tip sizes identified in Section 2.0 

of report). 

12 = Constant, in/ft. 

60 = Constant, s/min. 

60 = Constant, min/hr. 
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D-8

.kW@  c E%W@,{

100 J .k{

f

{g

Equation D.9 

If not reported, the flare vent gas molecular weight (MWvg), in lb/lb-mol, was calculated 

using Equation D.9.

This equation was used for data sets A and C.

Where: 

MWvg = Flare vent gas molecular weight, lb/lb-mole. 

j = Individual combustible or inert component in flare vent gas. 

n = Number of individual components in flare vent gas. 

%vg,i = Percentage of component j in flare vent gas, molar percentage. 

 (See Equation D.14 for data set A and Equation D.13 for data set C). 

100 = Constant, percentage. 

MWj = Molecular weight of component j, lb/lb-mole. 
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D-9

 

%,v|  100 q l,v|

l,v|  lj  ljbr

%jb  100 q ljb

l,v|  lj  ljb r

%j  100 q lj

l,v|  lj  ljbr

Equations D.10 through D.12 

For data set E, the percentages of 1,3-butadiene (%1,3-B), natural gas (%NG), and nitrogen 

(%N) in the flare vent gas, in volumetric percentage, were calculated using Equations D.10 

through D.12, respectively.

Equation D.10:

Where: 

%1,3-B = Percentage of 1,3-Butadiene in flare vent gas, volumetric percentage. 

100 = Constant, percentage. 

Q1,3-B = Volumetric flow rate of 1,3-Butadiene, scf/hr. (See Equation D.15C). 

QN = Volumetric flow rate of nitrogen, scf/hr. (See Equation D.15D). 

QNG = Reported volumetric flow rate of supplemental natural gas (i.e., Nat Gas flow in 

SCFH), scf/hr. 

Equation D.11:

Where: 

%NG = Percentage of natural gas in flare vent gas, volumetric percentage. 

100 = Constant, percentage. 

Q1,3-B = Volumetric flow rate of 1,3-Butadiene, scf/hr. (See Equation D.15C). 

QN = Volumetric flow rate of nitrogen, scf/hr. (See Equation D.15D). 

QNG = Reported volumetric flow rate of supplemental natural gas (i.e., Nat Gas flow in 

SCFH), scf/hr. 

Equation D.12:

Where: 

%N = Percentage of nitrogen in flare vent gas, volumetric percentage. 

100 = Constant, percentage. 

Q1,3-B = Volumetric flow rate of 1,3-Butadiene, scf/hr. (See Equation D.15C). 

QN = Volumetric flow rate of nitrogen, scf/hr. (See Equation D.15D). 

QNG = Reported volumetric flow rate of supplemental natural gas (i.e., Nat Gas flow in 

SCFH), scf/hr. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-10

 

%j  100 U %}

Equation D.13 

For data set C, the percentage of nitrogen in the flare vent gas (%N), in molar percentage, 

was calculated using Equation D.13.

Where: 

%N = Percentage of nitrogen in flare vent gas, molar percentage. 

100 = Constant, percentage. 

%P = Reported percentage of propane in flare vent gas (i.e., %Propane in Nitrogen), 

molar percentage. 
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D-11

 

%{  100 q h{

h{  h{r

Equation D.14 

For data set A, the percentage of component (propylene or nitrogen) in the flare vent gas 

(%j), in molar percentage, was calculated using Equation D.14.

Where: 

%j = Percentage of component in flare vent gas, molar percentage. 

j = Individual component in flare vent gas (i.e., propylene or nitrogen). 

100 = Constant, percentage. 

mj1 = Reported mass flow rate of component 1 (i.e., Proplylene Flow or Nitrogen Flow), 

lb/hr. 

mj2 = Reported mass flow rate of component 2 (i.e., Proplylene Flow or Nitrogen Flow), 

lb/hr. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-12

l~,$  lW@ E

%W@,$

100 J

l~,$  l~,$ %si%tsu A 60

Equation D.15A and D.15B 

The volumetric flow rate for each flare vent gas component in the combustion zone (QCZ,i), 

in scf/hr, was calculated using Equation D.15A and Equation D.15B.

Equation D.15A:

This equation was used for data sets C, D, F, G1, G2, H, and J.

Where: 

QCZ,i = Volumetric flow rate of each flare vent gas component i in combustion zone, 

scf/hr. 

i = Individual flare vent gas component in combustion zone. 

Qvg = Flare vent gas flow rate, scf/hr. (If not reported, see Equations D.4 through D.7). 

%vg,i = Percentage of component i in flare vent gas, molar percentage. 

100 = Constant, percentage. 

Equation D.15B:

This equation was used for data set A. 

Where: 

QCZ,i = Volumetric flow rate of each flare vent gas component i in combustion 

zone, scf/hr. 

i = Individual flare vent gas component in combustion zone. 

QCZ,i ,reported = Reported volumetric flow rate of each flare vent gas component i in 

combustion zone (i.e., Propylene Flow or Nitrogen Flow), scf/min. 

60 = Constant, s/min. 
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D-13

 

l,v|  Zh,v|,%si%tsu

0.159 \

lj  lW@ U l,v| U ljb

Equation D.15C and D.15D 

For data set E, the volumetric flow rates of 1,3-butadiene (Q1,3-B) and nitrogen (QN) in the 

flare vent gas, in scf/hr, were calculated using Equations D.15C and D.15D, respectively.

Equation D.15C:

Where: 

Q1,3-B = Volumetric flow rate of 1,3-Butadiene, scf/hr. 

m1,3-B,reported = Reported mass flow rate of 1,3-Butadiene (i.e., BD flow in lb/hr), lb/hr. 

0.159 = Constant, density of 1,3-Butadiene, lb/scf. 

Equation D.15D:

Where: 

QN = Volumetric flow rate of nitrogen, scf/hr. 

Qvg = Flare vent gas flow rate, scf/hr. (See Equations D.6). 

Q1,3-B = Volumetric flow rate of 1,3-Butadiene, scf/hr. (See Equation D.15C). 

QNG = Reported volumetric flow rate of supplemental natural gas (i.e., Nat Gas flow in 

SCFH), scf/hr. 
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D-14

 

~,$ 

l~,$

∑ l~,$

Equation D.16A 

The concentration of each flare vent gas component in the combustion zone (XCZ,i) was 

calculated using Equation D.16.

This equation was used for data sets A, C, D, E, F, G1, G2, H, and J.

Where: 

XCZ,i = Concentration of component i in combustion zone gas, volume fraction. 

i = Individual flare vent gas component in combustion zone. 

QCZ,i = Volumetric flow rate of component i in combustion zone gas, scf/hr. (See 

Equation D.15A or D.15B). 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-15

 

KX 

l~,X

∑ l~,X

Equation D.16B 

The concentration of each flare vent gas combustible component (VGC) was calculated 

using Equation D.16B.

This equation was used for all data sets.

Where: 

VGc = Concentration of combustible component c in the flare vent gas, volume fraction. 

c = Individual combustible component in the flare vent gas. 

QCZ,c = Volumetric flow rate of combustible component c in the flare vent gas, scf/hr. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-16

 

W@XY 

l~,X

∑ l~,X

Equation D.16C 

The concentration of each flare vent gas and center steam only component (Xvgcs) was 

calculated using Equation D.16C.

This equation was used for all data sets.

Where: 

Xvgcs = Concentration of combustible component c in the flare vent gas and center steam 

mixture, volume fraction. 

c = Individual combustible component in the flare vent gas and center steam mixture. 

QCZ,c = Volumetric flow rate of combustible component c in the flare vent gas and center 

steam mixture, scf/hr. 
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D-17

 

hY  hY,%si%tsu A 1,000

Equation D.17 

If not reported, the total mass steam rate (ms), in lb/hr, was calculated using Equation D.17. 

This equation was used for data sets G2 and H. 

Where: 

ms = Total steam rate, lb/hr. 

ms,reported = Reported total steam rate, klb/hr. 

 Reported for data set G2 as “Steam Flow (DCS).” 

 Reported for data set H as “EPF Steam.”

1,000 = Constant, lb/klb. 
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D-18

 

lL  379.48 E hL

.kY

J

Equation D.18 

The volumetric steam rate (Qx), in scf/hr, was calculated using Equation D.18. 

This equation was used for data sets A, C, D, E, F, G1, G2, H, and J. 

Where: 

Qx = Type x steam rate, scf/hr. 

x = Type of steam rate (i.e., Total, Lower, Center, Ring, or Upper). 

379.48 = Constant, lb-mole/scf at dry standard conditions. 

mx = Type x steam rate, lb/hr. (See Equation D.19). 

MWs = Molecular weight of steam, 18 lb/lb-mole. 
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D-19

 

hY  h  hX  h%  hP

Equation D.19 

The total mass steam rate (ms), lower mass steam rate (ml), and ring mass steam rate (mr), 

in lb/hr, were calculated using Equation D.19. 

This equation was used for data sets E, F, G1, and G2. 

Where: 

ms = Total steam rate, lb/hr. (See Equation D.17 for data sets G2 and H). 

 Total steam rate was a reported value for data sets R and G1. 

ml = Lower steam rate, lb/hr. 

mc = Center steam rate, lb/hr. 

 Center steam rate was a reported constant value for data sets F, G1 and G2. 

mr = Ring steam rate, lb/hr. 

mu = Upper steam rate, lb/hr. 

 Upper steam rate was a reported value for data set E.
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D-20

 

lY  l  lX  l%  lP

Equation D.20 

The ring volumetric steam rate (Qr) and lower volumetric steam rate (Ql), in scf/hr, were 

calculated using Equation D.20. 

This equation was used for data sets F, G1, and G2. 

Where: 

Qs = Total steam rate, scf/hr. 

Ql = Lower steam rate, scf/hr. 

Qc = Center steam rate, scf/hr. (See Equation D.18 for data sets F, G1 and G2). 

Qr = Ring steam rate, scf/hr. 

Qu = Upper steam rate, scf/hr. 
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D-21

 

/Kt 

hY

hW@

Equation D.21A 

The steam-to-vent-gas ratio by weight (S/VGwt), in lb steam/lb vent gas, was calculated 

using Equation D.21A. 

This equation was used for data sets A, C, D, E, F, G1, G2, and H.

Where: 

S/VGwt = The steam-to-vent-gas ratio by weight, lb steam/lb vent gas. 

ms = Total steam rate, lb/hr. (See Equation D.17 for data sets G2 and H. See 

Equation D.19 for data set E). 

mvg = Flare vent gas flow rate, lb/hr. (See Equation D.2 for data set A. See 

Equation D.3 for data set C). 
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D-22

/KW 

lY

lW@

Equation D.21B 

The steam-to-vent-gas ratio by volume (S/VGvol), in scf steam/scf vent gas, was calculated 

using Equation D.21B. 

This equation was used for data sets A, C, D, E, F, G1, G2, H, I and J. 

Where: 

S/VGvol = The steam-to-vent-gas ratio by volume, scf steam/scf vent gas. 

Qs = Total steam rate, scf/hr. (See Equation D.18 for data sets A, C, D, E, F, and 

H). 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.8A for data sets A and E. 

See Equation D.8B for data set J). 
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D-23

 

/
W 

lY

∑ l~, $

Equation D.21C 

The steam-to-vent-hydrocarbon ratio by volume (S/HCvol), in scf steam/scf hydrocarbon, 

was calculated using Equation D.21C. 

This equation was used for data sets A, C, D, E, F, G1, G2, H, I and J. 

Where: 

S/HCvol = The steam-to-vent-gas ratio by weight, lb steam/lb vent gas. 

Qs = Total steam rate, scf/hr. (See Equation D.18 for data sets A, C, D, E, F, and 

H). 

QCZ,HCi = Volumetric flow rate of component i in combustion zone gas that is a 

hydrocarbon, scf/hr. (See Equation D.15A or D.15B). 
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D-24

a
K%#t$ 

a
K~

a
KHHv6b

a
K~ 

lW@a
K6b

0lW@  lY5

Equations D.22 through D.25 

The ratio of the net heating value of the combustion zone gas to the net heating value of the 

flare vent gas if diluted to the lower flammability limit (NHVratio), was calculated using 

Equations D.22 through D.25. 

Equations D.22 through D.25 were used for data sets A, C, D, E, F, G1, G2, H and J. 

Equation D.22:

Where: 

NHVratio = The ratio of the net heating value of the combustion zone gas to the net 

heating value of the flare vent gas if diluted to the lower flammability 

limit. 

NHVCZ = The net heating value of the combustion zone gas, BTU/scf (See Equation 

D.23). 

NHVLFL-VG = The net heating value of the flare vent gas if diluted to the lower 

flammability limit, BTU/scf (See Equation D.24). 

Equation D.23:

The net heating value of the combustion zone gas is calculated by correcting the net heating 

value of the vent gas for the inclusion of steam in the combustion zone gas, as shown below. 

Where: 

NHVCZ = The net heating value of the combustion zone gas, BTU/scf. 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.4 for data set A. See Equation 

D.5 for data set C. See Equation D.6 for data sets E, H and J. See Equation D.7 for 

data set G1). 

NHVVG = Flare vent gas net heating value, BTU/scf (See Equation D.1 for data set E). 

Qs = Total steam rate, scf/hr. (See Equation D.18 for data sets A, C, D, E, F, H and J). 
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D-25

a
KHHv6b  a
KW@ A V/VW@

Equations D.22 through D.25 (Continued) 

Equations D.22 through D.25 were used for data sets A, C, D, E, F, G1, G2, H and J. 

Equation D.24:

The net heating value of the flare vent gas if diluted to its lower flammability limit is calculated 

by multiplying the actual net heating value of the flare vent gas by its lower flammability limit, as 

shown below. 

Where: 

NHVLFL-VG = The net heating value of the flare vent gas if diluted to the lower 

flammability limit, BTU/scf. 

NHVVG = Flare vent gas net heating value, BTU/scf. (See Equation D. 1 for data set 

E). 

LFLvg = Lower flammability limit of the flare vent gas adjusted for nitrogen 

equivalency, volume fraction (See Equation D.25). 

Equation D.25:

This equation uses the Le Chatelier Principle with a nitrogen equivalency adjustment; it allows 

you to estimate the lower flammability limit of any mixture. For this particular analysis, the 

lower flammability limit of the flare vent gas is of interest. The calculation uses the weighted 

average of the LFLs of the individual compounds (combustible and inert) weighted by their 

volume percent of the flare vent gas. All inerts, including nitrogen, are assumed to have an 

infinite lower flammability limit (e.g., LFLN2 = ∞). 

( )( ) ( )( )

2 2 2 2

1 1

%

1

, ,

1

e H O H O e CO CO

n

i i

i

vg

N x N x

LFL

LFL

− − − − 















=

∑=

Where: 

LFLvg = Lower flammability limit of the flare vent gas adjusted for nitrogen equivalency, 

volume fraction 

n = Number of individual combustible components in flare vent gas. 

i = Individual combustible component in flare vent gas. 

%i = Percentage of combustible component i in flare vent gas, molar percent. (See 

Table C-1 in Appendix C, most values are reported.) 

LFLi = Lower flammability limit of combustible component i in flare vent gas, volume 

percent. Zabetakis M.G. “Flammability Characteristics of Combustible Gases and 

Vapors”, U.S. Bureau of Mines, Bulletin 627, 1965 was used to determine the 

lower flammability limit for each component in each test run. Table D.1 lists these 

values. 

Ne,H2O = Coefficient of nitrogen equivalency for water relative to nitrogen, unitless (See 

Equation D.29). 

xH2O = Reported concentration of water in flare vent gas, volume fraction. 

Ne,CO2 = Coefficient of nitrogen equivalency for carbon dioxide relative to nitrogen, 

unitless (See Equation D.29). 

xCO2 = Reported concentration of carbon dioxide in flare vent gas, volume fraction. 
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D-26

Table D.1 – Lower Flammability Limit of Select Compounds (LFLi) – Flare Vent Gas 

Compound LFLi (vol %) Compound LFLi (vol %) 

1-Butene 1.8 Iso-Butylene 1.8 

1,2-Butadiene 2 Methane 5 

1,3-Butadiene 2 Methyl Acetylene 1.7 

Acetylene 2.5 n-Butane 1.8 

Benzene 1.3 Nitrogen ∞

Carbon Dioxide ∞ Oxygen ∞

Carbon Monoxide 12.5 Pentane Plus (C5+) 1.4 

Cis-2-Butene 1.7 Propadiene 2.16 

Cyclopropane 2.4 Propane 2.1 

Ethane 3 Propylene 2.4 

Ethyl Benzene 1 Toluene 1.2 

Ethylene 2.7 Trans-2-Butene 1.7 

Hydrogen 4 Water ∞

Hydrogen Sulfide 4 Xylenes 1.1 

Iso-Butane 1.8 

 
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D-27

 

Equation D.26 

The combustible components in the combustion zone (CCZ), in volume fraction, was 

calculated using Equation D.26. 

This equation was used for data sets A, C, D, E, F, G1, G2, H and J.

( )

1

vg s

h vg

n

i

i vg

cz Q Q

Q Q

C

+

 +











=

∑=

χ χ

Where: 

CCZ = Combustible components in the combustion zone gas, volume fraction. 

n = Number of individual combustible organic components in flare vent gas. 

i = Individual combustible organic component in flare vent gas. 

xi = Concentration of combustible organic component i in flare vent gas, volume 

fraction. 

xh = Concentration of hydrogen in flare vent gas, volume fraction. 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.4 for data set A. See Equation 

D.5 for data set C. See Equation D.6 for data sets E, H and J. See Equation D.7 for 

data set G1.). 

Qs = Total steam flow, scf/hr. (See Equation D.18 for data sets A, C, D, E, F, H and J.). 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-28

Equations D.27 through D.29 

The lower flammability limit of combustion zone gas adjusted for nitrogen equivalency 

(LFLCZ), in volume fraction, was calculated using Equations D.27 through D.29. 

Equations D.27 through D.29 were used for data sets A, C, D, E, F, G1, G2, H and J.

Equation D.27:

This equation uses the Le Chatelier Principle with a nitrogen equivalency adjustment; it allows 

you to estimate the lower flammability limit of any mixture. For this particular analysis, the 

lower flammability limit of the combustion zone gas is of interest. The calculation uses the 

weighted average of the LFLs of the individual compounds (combustible and inert) weighted by 

their volume percent of the combustion zone gas. All inerts, including nitrogen, are assumed to 

have an infinite lower flammability limit (e.g., LFLN2 = ∞). 

( )( ) ( )( )

2 2 2 2

1 1

%

1

, ,

1

e H O H O e CO CO

n

i i

i

cz

N x N x

LFL

LFL

− − − − 















=

∑=

Where: 

LFLCZ = Lower flammability limit of combustion zone gas adjusted for nitrogen 

equivalency, volume fraction. 

n = Number of components in combustion zone gas. 

i = Individual component in combustion zone gas. 

%i = Percentage of component j in combustion zone gas, volume percentage. (See 

Equation D.28). 

LFLi = Lower flammability limit of component j, volume percent. Zabetakis M.G. 

“Flammability Characteristics of Combustible Gases and Vapors”, U.S. Bureau of 

Mines, Bulletin 627, 1965 was used to determine the lower flammability limit for 

each component in each test run. Table D.2 lists these values. 

Ne,H2O = Coefficient of nitrogen equivalency for water relative to nitrogen, unitless (See 

Equation D.29). 

xH2O = Concentration of water in combustion zone gas, volume fraction. 

Ne,CO2 = Coefficient of nitrogen equivalency for carbon dioxide relative to nitrogen, 

unitless (See Equation D.29). 

xCO2 = Concentration of carbon dioxide in combustion zone gas, volume fraction. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-29

 

%$  q l~,$

0lW@  lY5

r A 100

Equations D.27 through D.29 (Continued) 

The coefficients of nitrogen equivalency of water and carbon dioxide relative to nitrogen, 

Ne,H2O and Ne,CO2, was calculated using Equation D.29. 

Equations D.27 through D.29 were used for data sets A, C, D, E, F, G1, G2, H and J.

Equation D.28:

Where: 

%i = Percentage of component i in combustion zone gas, volume percentage. 

QCZ,i = Volumetric flow rate of each flare vent gas component i in combustion zone, 

scf/hr. (See Equation D.15A for data sets C, D, E, F, GI, G2,H and J. See Equation 

D.15B for data set A). 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.4 for data set A. See Equation 

D.5 for data set C. See Equation D.6 for data sets E, H and J. See Equation D.7 for 

data set G1.). 

Qs = Total steam rate, scf/hr. (See Equation D.18 for data sets A, C, D, E, F, H and J.).

Equation D.29:

∑= ∑ 















=

n

i CZ i

CZ i

e k k i Q

Q

N K

1 ,

,

, ,

Where: 

Ne,k = Coefficient of nitrogen equivalency for water or carbon dioxide relative to 

nitrogen, unitless. 

k = Individual inert component (i.e. water and carbon dioxide). 

n = Number of individual combustible components in flare vent gas. 

i = Individual combustible component in flare vent gas. 

Kk,i = Values for inert component k (i.e. water or carbon dioxide) for each combustible 

component i in flare vent gas. Table D.3 lists these values. 

QCZ,i = Volumetric flow rate of each flare vent gas component i in combustion zone, 

scf/hr. (See Equation D.15A or D.15B). 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-30

Table D.2 – Lower Flammability Limit of Select Compounds (LFLi) – 

Combustion Zone Gas 

Compound LFLi (vol %) Compound LFLi (vol %) 

1-Butene 1.8 Methane 5 

1,2-Butadiene 2 Methyl Acetylene 1.7 

1,3-Butadiene 2 n-Butane 1.8 

Acetylene 2.5 Nitrogen ∞

Benzene 1.3 Oxygen ∞

Carbon Dioxide ∞ Pentane Plus (C5+) 1.4 

Carbon Monoxide 12.5 Propadiene 2.16 

Cis-2-Butene 1.7 Propane 2.1 

Cyclopropane 2.4 Propylene 2.4 

Ethane 3 Toluene 1.2 

Ethyl Benzene 1 Trans-2-Butene 1.7 

Ethylene 2.7 Water ∞

Hydrogen 4 Xylenes 1.1 

Hydrogen Sulfide 4 Steam ∞

Iso-Butane 1.8 Pilot Gas 5 

Iso-Butylene 1.8 

Table D.3 – Values for inert component k (i.e. water or carbon dioxide) for each 

combustible component i in flare vent gas 

Combustible Component i in 

Flare Vent Gas KWater

KCarbon 

Dioxide

Methane 1.87 2.23 

Ethane 1.40 1.87 

Propane 1.51 1.93 

Ethylene 1.68 1.84 

Propylene 1.36 1.92 

Hydrogen 1.35 1.51 

All Other Combustibles 1.50 1.87 
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D-31



hG$%

ht$X

ht$X  c8h$

:

f

$g

8O/$:

O/$ 

28.84 Ea,$

0.21J

.k$

Equations D.30 through D.37 

For air-assisted flares, the stoichiometric ratio (SR) was calculated using Equations D.30 

through D.37. 

Equations D.30 through D.32 were used for data sets C and J. 

Equation D.30:

Where: 

SR = Ratio of actual mass flow of total assist air to the theoretical stoichiometric mass 

of air needed to combust the flare vent gas, unitless. 

mAir = Actual total assist air mass flow rate, lb/hr. (See Equation D.34 for data set C). 

mStoic = Theoretical stoichiometric mass of air needed to combust the flare vent gas, lb/hr 

(See Equation D.31). 

Equation D.31:

Where: 

mStoic = Theoretical stoichiometric mass of air needed to combust the flare vent gas, lb/hr. 

n = Number of individual combustible components in flare vent gas. 

i = Individual combustible component in flare vent gas. 

mi = Reported mass flow rate of combustible component i in the flare vent gas, lb/hr. 

(See Equation D.35 for data set C). 

AFi = Stoichiometric air to fuel ratio of combustible component i, lb of air/lb of 

combustible component i (See Equation D.32). 

Equation D.32:

Where:

AFi = Stoichiometric air to fuel ratio of combustible component i, lb of air/lb of 

combustible component i. 

NO2,i = Stoichiometric amount of oxygen needed for one mole of combustible component 

i, moles (See Equation D.33). 

i = Individual combustible component in flare vent gas. 

MWi = Molecular weight of combustible component i in the flare vent gas, lb/lb-mole. 

28.84 = Constant, molecular weight of ambient air, lb/lb-mole. 

0.21 = Constant, fraction of ambient air comprised of oxygen. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-32

h#$%  l#$% %si%tsu A 60 A 0.07517

h$  .k$ A E l$

379.48J

Equations D.30 through D.37 (Continued) 

Equation D.33:

This equation was used for data sets C and J. 

4

,

,2 ,

H i

O i C i

N

N = N +

Where: 

NO2,i = Stoichiometric amount of oxygen needed for one mole of combustible component 

i, moles. 

i = Individual combustible component in flare vent gas. 

NC,i = Number of carbon atoms per molecule of combustible hydrocarbon i. 

NH,i = Number of hydrogen atoms per molecule of combustible hydrocarbon i. 

4 = Constant. 

Equation D.34:

This equation was used for data set C. 

Where: 

mair = Actual total assist air mass flow rate, lb/hr. 

Qair,reported = Reported air assist flow rate, scf/min. 

60 = Constant, min/hr.

0.07517 = Constant, assumed density of air, lb/scf. 

Equation D.35:

This equation was used for data set C. 

Where: 

mi = Mass flow rate of combustible component i in the flare vent gas, lb/hr. 

i = Individual flare vent gas component in combustion zone. 

MWvg = Molecular weight of combustible component (i.e., propane) in the flare vent gas, 

lb/lb-mole. 

Qi = Volumetric flow rate of each flare vent gas component i, scf/hr. (See Equation 

D.36). 

379.48 = Constant, lb-mole/scf at dry standard conditions. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-33

l$  lW@ E

%W@,$

100 J

lW@  KW@ A 60 A 60 E1.77

144J

Equations D.30 through D.37 (Continued) 

Equation D.36:

This equation was used for data set C. 

Where: 

Qi = Volumetric flow rate of each flare vent gas component i (i.e., propane or 

nitrogen), scf/hr. 

i = Individual component in flare vent gas (i.e., propane or nitrogen). 

Qvg = Flare vent gas flow rate, scf/hr. (See Equation D.37). 

%,i = Percentage of component i (i.e., propane or nitrogen) in flare vent gas, molar 

percentage. 

%p = Reported percentage of propane in flare vent gas (i.e., %Propane in Nitrogen), 

molar percentage. 

%N = Percentage of nitrogen in flare vent gas, molar percentage. (See Equation D.13). 

100 = Constant, percentage. 

Equation D.37:

This equation was used for data set C. 

Where: 

Qvg = Flare vent gas flow rate, scf/hr. 

Vvg = Reported flare vent gas velocity, ft/s. 

60 = Constant, s/min. 

60 = Constant, min/hr. 

1.77 = Constant, area of fuel exit port, in2. 

144 = Constant, ft2/in2. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-34

hW@,Y  hW@  hX

[W@vY  q hW@vY

lW@  lX

r

Equations D.38 through D.41 

The calculated momentum flux ratio (MFR) was calculated using Equations D.38 through 

D.41. 

Equation D.38:

This equation was used for data sets A, D, E, F, G1, G2, and J. 

Where: 

mvg,s = Flare vent gas flow rate including center steam, lb/hr. 

mvg = Flare vent gas flow rate, lb/hr. (See Equation D.2 for data set A. See Equation D.3 

for data set C). 

mc = Center steam rate, lb/hr. 

 Center steam rate was a reported constant value for data sets F, G1 and G2. 

Equation D.39:

This equation was used for data sets A, D, E, F, G1, G2, and J. 

Where: 

[ vg-s = Density of flare vent gas including center steam, lb/scf. 

mvg,s = Flare vent gas flow rate including center steam, lb/hr. 

Qvg = Flare vent gas flow rate, scf/hr. 

Qc = Center steam rate, scf/hr. (See Equation D.18 for data sets F, G1 and G2). 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-35

K$fu 

K$f@,%si%tsu A 5280

60 A 60

./ q [W@vY A KW@vY



0.07492 A 8K$fu:



r

Equations D.38 through D.41 (Continued) 

Equation D.40:

This equation was used for data sets D, F, G1, G2, and J. 

Where: 

Vwind = Wind velocity, ft/s. 

Vwind, reported = Reported wind velocity, mph. 

5280 = Constant, ft/mile. 

60 = Constant, s/min. 

60 = Constant, min/hr. 

Equation D.41:

This equation was used for data sets A, D, E, F, G1, G2, and J. 

Where: 

MRF = Calculated momentum flux ratio, unitless. 

[ vg-s = Density of flare vent gas including center steam, lb/scf. (See Equation 

D.39). 

Vvg-s = Flare vent gas velocity with center steam, ft/s. (If not reported, see 

Equation D.8C). 

Vwind = Wind velocity, ft/s. (See Equation D.40). 

0.07492 = Constant, assumed density of air, lb/scf.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-36

/M/ 

KW@vY

M A Z x

12\

Equations D.42 through D.45 

The shore equation analysis was performed using Equations D.42 through D.45. 

These equations were used for data sets A, C, D, E, F, G1, G2, and J. 

Equation D.42:

Where: 

/M/ = Shore Equation Velocity-Jet Periphery, sec-1

. 

Vvg-s = Flare vent gas velocity with center steam, ft/s. (If not reported, see 

Equation D.8C). 

M = Constant pi used to calculate area of flare tip. 

d = Reported effective diameter of flare tip, in. (See tip sizes identified in 

Section 2.0 of report). 

12 = Constant, in/ft.
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D-37

 

Equations D.42 through D.45 (Continued) 

Equation D.43:

This equation uses the Le Chatelier Principle with a nitrogen equivalency adjustment; it allows 

you to estimate the lower flammability limit of any mixture. For this particular analysis, the 

lower flammability limit of the flare vent gas including center steam is of interest. The 

calculation uses the weighted average of the LFLs of the individual compounds (combustible and 

inert) weighted by their volume percent of the flare vent gas plus center steam. All inerts, 

including nitrogen, are assumed to have an infinite lower flammability limit (e.g., LFLN2 = ∞). 

( )( ) ( )( ) 2 2 2 21 1

1

, ,

1

,

e H O H O e CO CO

n

i i

vgcs i

vgcs

N x N x

LFL

X

LFL

− − − − 















=

∑=

Where: 

LFLvgcs = Lower flammability limit of the flare vent gas including center steam adjusted for 

nitrogen equivalency, volume fraction. 

n = Number of individual combustible components in flare vent gas. 

i = Individual combustible component in flare vent gas. 

Xvgcs,i = Concentration of combustible component i in the flare vent gas and center steam 

mixture, volume fraction. (See Equation D.16C). 

LFLi = Lower flammability limit of combustible component i in flare vent gas, volume 

percent. Zabetakis M.G. “Flammability Characteristics of Combustible Gases and 

Vapors”, U.S. Bureau of Mines, Bulletin 627, 1965 was used to determine the 

lower flammability limit for each component in each test run. Table D.4 lists these 

values. 

Ne,H2O = Coefficient of nitrogen equivalency for water relative to nitrogen, unitless. (See 

Equation D.29). 

xH2O = Reported concentration of water in flare vent gas, volume fraction. 

Ne,CO2 = Coefficient of nitrogen equivalency for carbon dioxide relative to nitrogen, 

unitless. (See Equation D.29). 

xCO2 = Reported concentration of carbon dioxide in flare vent gas, volume fraction. 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

D-38

Table D.4 – Lower Flammability Limit of Select Compounds (LFLi) – Flare Vent Gas 

Compound LFLi (vol %) Compound LFLi (vol %) 

1-Butene 1.8 Iso-Butylene 1.8 

1,2-Butadiene 2 Methane 5 

1,3-Butadiene 2 Methyl Acetylene 1.7 

Acetylene 2.5 n-Butane 1.8 

Benzene 1.3 Nitrogen ∞

Carbon Dioxide ∞ Oxygen ∞

Carbon Monoxide 12.5 Pentane Plus (C5+) 1.4 

Cis-2-Butene 1.7 Propadiene 2.16 

Cyclopropane 2.4 Propane 2.1 

Ethane 3 Propylene 2.4 

Ethyl Benzene 1 Toluene 1.2 

Ethylene 2.7 Trans-2-Butene 1.7 

Hydrogen 4 Water ∞

Hydrogen Sulfide 4 Xylenes 1.1 

Iso-Butane 1.8 

 
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D-39

G  q100 U 0V/VW@XY A 1005

V/VW@XY A 100 r Z [W@XY0.07492\

KG 

Q

w

R

2M A 

M

4

A Z x

12\



M

_

z

`

A EG

6.85J

+

Equations D.42 through D.45 (Continued) 

The shore equation analysis was performed using Equations D.42 through D.45. 

These equations were used for data sets A, C, D, E, F, G1, G2, and J. 

Equation D.44:

Where: 

G = Shore Equation, lb flare vent gas including center steam / lb of air. 

LFLvgcs = Lower flammability limit of the flare vent gas including center steam 

adjusted for nitrogen equivalency, volume fraction. (See Equation D.43). 

100 = Constant, percentage. 

[ vgcs = Density of flare vent gas including center steam, lb/scf. (See Equation 

D.39). 

0.07492 = Constant, assumed density of air, lb/scf.

Equation D.45:

Where: 

KG = Maximum allowable flare vent gas velocity including center steam, ft/s. 

G = Shore Equation, lb flare vent gas including center steam / lb of air. (See 

Equation D.44). 

M = Constant pi used to calculate area of flare tip. 

d = Reported effective diameter of flare tip, in (See tip sizes identified in 

Section 2.0 of report). 

12 = Constant, in/ft. 

6.85 = Empirical constant, may contain elements of viscosity and other transport 

properties of the gas and air.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

Appendix E 

Type And Amount Of Components In Each Test Run By Test Report. 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally diss

represent and should not be construed to represent any Agency determination or policy.

E-1 

100.0%

99.9%

99.5%

99.9%

99.8%

99.7%

99.8%

99.8%

95.1%

99.3%

99.5%

98.7%

99.7%

99.9%

99.9%

82.2%

99.4%

0%

10%20%30%40%50%60%70%80%90%

100%

1 5 11a 11b 11c 16a 16b 17 59a 59b 50 51 56 57 60 61 202 

EPA-600 2-83-052 and EPA-600 2-85-106 

Propane Propylene Steam Nitrogen
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represent and should not be construed to represent any Agency determination or policy.

E-2 

99.5%

99.5%

97.5%

96.3%

94.8%

89.7%

92.6%

85.9%

82.8%

98.4%

99.0%

97.7%

97.4%

92.7%

93.9%

97.1%

89.2%

99.1%

99.0%

99.4%

99.7%

99.1%

97.8%

95.8%

92.5%

88.2%

91.7%

97.7%

98.9%

96.6%

0%

10%20%30%40%50%60%70%80%90%

100%

A 2.0 (1)

A 2.0 (2)

A 3.0 (1)

A 3.0 (2)

A 4.0 (1)

A 4.0 (2)

A 4.0 (3)

A 5.0 (1)

A 5.0 (2)

B 1.0 (1)

B 1.0 (2)

B 2.0 (1)

B 2.0 (2)

B 2.5 (1)

B 2.5 (2)

B 2.5 (3)

B 3.5 (1)

C 1.0 (1)

C 1.0 (2)

C 1.0 (3)

C 1.0 (4)

C 2.0 (1)

C 2.0 (2)

C 3.0 (1)

C 3.0 (2)

C 3.7 (1)

C 3.7 (2)

D 2.0 (1)

D 2.0 (2)

D 3.0 (1)

Flint Hills Resources (AU)

1,3-Butadiene Ethane Pentane Plus (C5+) Ethylene Propane Hydrogen Propylene Methane Other Iso-Butane 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally diss

represent and should not be construed to represent any Agency determination or policy.

E-3 

95.9%

98.6%

98.6%

97.2%

99.1%

96.5%

98.6%

95.5%

96.5%

88.6%

95.1%

78.7%

97.5%

99.2%

98.8%

98.4%

99.4%

98.7%

99.3%

98.3%

98.7%

96.8%

97.5%

92.0%

95.0%

89.0%

92.8%

88.4%

99.3%

98.8%

98.7%

98.9%

97.8%

95.8%

94.7%

0%

10%20%30%40%50%60%70%80%90%

100%

A 2.0 (1)

A 2.0 (2)

A 2.0 (3)

A 3.0 (1)

A 3.0 (2)

A 4.0 (1)

A 4.0 (2)

A 5.0 (1)

A 5.0 (2)

A 6.0 (1)

A 6.0 (2)

A 8.5 (1)

A MIN (1)

A MIN (2)

A MIN (3)

B 1.0 (1)

B 1.0 (2)

B 2.0 (1)

B 2.0 (2)

B 3.0 (1)

B 3.0 (2)

B 4.0 (1)

B 4.0 (2)

B 5.0 (1)

B 5.0 (2)

B 5.7 (1)

B 5.7 (2)

B 6.4 (2)

C 1.0 (1)

C 1.0 (2)

C 2.0 (1)

C 2.0 (2)

C 3.0 (1)

C 3.0 (2)

C 4.0 (1)

Flint Hills Resources (LOU)

1,3-Butadiene Ethane Pentane Plus (C5+) Ethylene Propane Hydrogen Propylene Methane Other Iso-Butane 
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represent and should not be construed to represent any Agency determination or policy.

E-4 

86.2%

31.8%

75.5%

97.8%

84.9%

97.0%

99.0%

95.9%

99.3%

99.1%

99.5%

95.8%

97.2%

88.3%

93.6%

94.5%

93.2%

99.3%

938%

0%

10%20%30%40%50%60%70%80%90%

100%

1 1A 1B 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 1

INEOS

1,3-Butadiene Methane Steam Nitrogen
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E-5 

99.0%

98.8%

97.7%

98.5%

98.0%

98.6%

96.9%

97.7%

95.2%

96.7%

95.2%

92.0%

95.2%

81.3%

88.7%

98.9%

98.4%

99.1%

98.3%

99.3%

98.2%

99.2%

96.7%

98.6%

94.6%

96.9%

99.4%

98.4%

99.1%

96.8%

97.4%

92.7%

93.7%

93.1%

91.5%

94.6%

97.5%

88.6%

91.6%

98.4%

83.5%

96.2%

98.6%

81.8%

91.6%

88.3%

93.5%

97.2%

0%

10%20%30%40%50%60%70%80%90%

100%

A 1-1

A 1-2

A 2-1

A 2-2

A 3-1

A 3-2

A 4-1

A 4-2

A 5-1

A 6-2

A 7-1

A 8-1

A 8-3

A 9-1

A 9-3

B 1-1

B 2-1

B 2-2

B 3-1

B 3-2

B 4-1

B 4-2

B 6-1

B 6-2

B 8-1

B 8-2

C 1-2

C 2-1

C 2-2

C 3-1

C 3-2

C 4-1

C 4-2

C 5-1

C 5-2

D 2-1

D 3-1

D 4-1

D 5-1

D 6-1

D 7-1

D 8-1

D 9-1

D 10-1

E 1-1

E 2-1

E 3-1

E 5-1

E 61

MPC Detroit

1,3-Butadiene Ethane Pentane Plus (C5+) Ethylene Propane Hydrogen Propylene Methane Other Iso-Butane 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally diss

represent and should not be construed to represent any Agency determination or policy.

E-6 

95.6%

97.4%

96.7%

97.8%

95.5%

93.8%

92.5%

87.0%

81.2%

74.4%

76.6%

67.9%

61.6%

61.9%

93.7%

89.6%

99.3%

98.7%

99.0%

98.4%

98.5%

97.8%

97.9%

97.6%

95.8%

97.6%

89.1%

96.7%

98.9%

99.2%

98.4%

98.2%

96.0%

98.6%

98.3%

98.1%

97.7%

96.8%

97.5%

0%

10%20%30%40%50%60%70%80%90%

100%

A11 2-1

A11 2-2

A113-1

A113-2

A114-1

A115-1

A116-1

A117-1

A118-1

A119-1

A1110-1

A1111-1

A1112-1

A1113-1

A19 4-1

A19 7-1

B 4-2

B 5-1

B 5-2

B 6-1

B 6-2

B 7-1

B 7-2

B 8-1

B 8-2

B 9-1

B 9-2

B 10-2

C 3-1

C 3-2

D4-1

D4-2

D7-2

E 3-1

E 3-3

E 4-1

E 4-3

E 5-1

E 5-2

MPC TX

1,3-Butadiene Ethane Pentane Plus (C5+) Ethylene Propane Hydrogen Propylene Methane Other Iso-Butane 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally diss

represent and should not be construed to represent any Agency determination or policy.

E-7 

96.9%

97.5%

97.3%

98.6%

99.0%

97.3%

79.8%

98.7%

98.9%

97.5%

57.1%

98.8%

98.0%

90.4%

95.7%

99.7%

99.3%

94.1%

90.9%

99.3%

98.0%

98.0%

0%

10%20%30%40%50%60%70%80%90%

100%

4-13, 12:51-13:11

3-30, 10:40-11:30

3-30, 9:55-10:38

4/1 08:48- 09:16 

4/1 09:17- 09:46 

4-6, 10:00-10:16

4-6, 10:43-10:53

4-6, 11:16-11:24

4-6, 11:32-11:49

4-6, 11:50-11:58

4-6,10:17-10:42

3-22, 17:14-17:38

3-22, 17:39-19:16

3-22, 19:17-20:08

3-22, 20:09-20:39

3-24, 14:08-14:18

3-24, 14:26-14:35

3-24, 16:20-16:31

3-24, 16:41-16:52

3-24, 16:56-17:07

3-24.15:47-15:58

3-24.16:01-16:11

A B B B B B B B B B B C C C C C C C C C C C 

Shell Deer Park (EPF)

Ethane Pentane Plus (C5+) Ethylene Propane Hydrogen Propylene Methane Other Iso-Butane St
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally diss

represent and should not be construed to represent any Agency determination or policy.

E-8 

99.9%

99.9%

99.8%

98.5%

94.7%

94.7%

94.9%

92.9%

84.7%

95.4%

93.9%

91.7%

82.7%

98.5%

96.6%

88.3%

48.4%

98.4%

97.8%

93.9%

71.3%

90.9%

99.1%

94.0%

99.4%

45.8%

83.5%

23.5%

88.0%

96.7%

99.1%

98.5%

94.4%

65.3%

48.6%

97.7%

11.6%

97.2%

92.6%

96.1%

86.5%

78.4%

70.0%

95.8%

48.1%

85.2%

66.8%

90.2%

92.9%

98.9%

64.3%

96.4%

83.1%

98.7%

90.3%

93.9%

86.7%

75.7%

38.6%

66.5%

0%

10%20%30%40%50%60%70%80%90%

100%

S1.5 ,1

S1.6 ,1

S1.7 ,1

S1.8 ,1

S1.9 ,1

S10.1 ,1

S10.2 ,1

S10.3 ,1

S10.4,1

S11.1,1

S11.2,1

S11.3,1

S11.4,1

S12.1,Averages

S12.2,Averages

S12.3,Averages

S12.4,1

S13.1,1

S13.2,1

S13.3,1

S13.4,Averages

S13.5,1

S14.1,1

S14.4,1

S2.1,Averages

S2.2,Averages

S2.3,Averages

S3.1,1

S3.2,2

S3.5,Averages

S3.6,1

S3.7,1

S4.1,Averages

S4.10,1

S4.11,1

S4.2,Averages

S4.3,Averages

S4.4,1

S4.5,1

S4.6,1

S4.7,1

S4.8,1

S4.9,1

S5.1,Averages

S5.2,1

S5.3,Averages

S5.4,Averages

S5.5,1

S5.6,Averages

S6.1,Averages

S6.2,Averages

S6.3,Averages

S6.4,Averages

S6.5,1

S6.6,1

S7.1,Averages

S7.2,Averages

S7.3,Averages

S7.4,1

S7.5,1

TCEQ

Propane Propylene Methane Steam Nitrogen
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Appendix F 

Charts Of Calculated And Measured LFL For Various Combustible Gases 

In Nitrogen And Carbon Dioxide 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-1 

0%

10%

20%

30%

40%

50%

60%

70%

80%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Hydrogen and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1932)

0%

10%

20%

30%

40%

50%

60%

70%

80%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Hydrogen and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones, 1929)
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not 

been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-2 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methane and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006a)

Experimental Data - (Vidal et al., 2006)

Pure component LFL at 5.0 for Jones and Kennedy (1933) and Vidal et al. (2006) experiments, but 4.9 for Kondo et al. (2006a) experiments.

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006b)

Experimental Data (Vidal et al., 2006)

Pure component LFL at 5.0 for Jones and Kennedy (1933) and Vidal et al. (2006) experiments, but 4.9 for Kondo et al. (2006b) experiments.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-3 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ethane and Nitrogen

Calculated using Le Chatelier Equation assuming 

an LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ethane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-4 

0%

10%

20%

30%

40%

50%

60%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ethylene and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006a)

0%

10%

20%

30%

40%

50%

60%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ethylene and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006b)

Experimental Data (Vidal et al., 2006)

Pure component LFL at 2.74 for Kondo et al. (2006b) experiments, but 2.70 for 

Vidal et al. (2006) experiments.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-5 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Propane and Nitrogen

Calculation using Le Chatelier Equation assuming an

LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006a)

Pure component LFL at 2.3 for Jones and Kennedy (1933) 

experiments, but 2.03 for Kondo et al. (2006a) experiments.

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Propane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006b)

Pure component LFL at 2.3 for Jones and Kennedy (1933) 

experiments, but 2.03 for Kondo et al. (2006b) experiments.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-6 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Propylene and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1938)

Experimental Data (Kondo et al., 200ba)

Pure component LFL at 2.4 for Jones and Kennedy (1938) 

experiments, but 2.16 for Kondo et al. (2006a) experiments.

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Propylene and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1938)

Experimental Data (Kondo et al., 2006b)

Pure component LFL at 2.4 for Jones and Kennedy (1938) 

experiments, but 2.16 for Kondo et al. (2006b) experiments.
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-7 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Butane and Nitrogen

Calculated using Le Chatelier Equation assuming 

an LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Butane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-8 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

1,3 Butadiene and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1943)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

1, 3 Butadiene and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1943)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-9 

0%

10%

20%

30%

40%

50%

60%

70%

80%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Carbon Monoxide and Nitrogen

Calculated using Le Chatelier Equation assuming 

an LFL for Nitrogen of Infinity

Experimental Data (Jones, 1929)

0%

10%

20%

30%

40%

50%

60%

70%

80%

90%

100%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Carbon Monoxide and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones, 1929)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-10 

0%

10%

20%

30%

40%

50%

60%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Dimethyl Ether and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006a)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Dimethyl Ether and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006b)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-11 

0%

10%

20%

30%

40%

50%

60%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methyl Formate and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006a)

0%

10%

20%

30%

40%

50%

60%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methyl Formate and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006b)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-12 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

1,1-Difluoroethane and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006a)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

1,1-Difluoroethane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006b)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

F-13 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ammonia and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006a)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Ammonia and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Kondo et al., 2006b)
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

Appendix G 

The following Appendix contains details about inerts and further explanation for 

including an equivalency adjustment to correct for different inert behavior. Section G.1 

explores the accuracy of Le Chatelier with varying chemical composition in the mixture. 

Section G.2 presents research on nitrogen composition and its effect on calculated LFL 

with respect to measured LFL; and Section G.3 discusses the effect of other types of inerts. 

There are additional uncertainties associated with the calculation of UFL; however, these 

are not discussed in this report because UFL is not presented as an operating parameter to 

indicate adequate combustion in flares. 
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G-1 

G.1 LFL Calculation Accuracy Associated with No-inert in Gas Mixture 

Researcher’s Observations of Measured and Calculated LFL 

Some researchers have observed that certain mixtures do not obey the Le Chatelier 

principle and Le Chatelier’s equation does not accurately predict the LFL for these mixtures. 

Heffington and Gaines (1981) noted that large amounts of hydrogen sulfide can cause deviations 

between calculated and experimental LFLs, and White (1925) specifically identified hydrogen 

sulfide and methane mixtures as not following the rule. White (1925) also found that hydrogen 

and ethylene, acetylene and hydrogen, and mixtures containing carbon disulfide do not follow Le 

Chatelier. Coward and Jones (1952) specifically noted that the LFL calculated using Le 

Chatelier’s equation for mixtures containing at least 2 of the compounds methane, hydrogen, or 

carbon monoxide are “approximately accurate”. Choudhuri (2005) described variations in the 

measured and calculated LFLs (using Le Chatelier’s equation) for methane and propane, and 

methane and ethane mixtures; and for hydrogen and carbon monoxide mixtures (Choudhuri and 

Subramanya, 2006). 

Coward and Jones (1952) provided a qualitative discussion of the variation in accuracy of 

the LFL calculated for mixtures using Le Chatelier’s equation with different chemical 

composition of the combustible gases. Table G.1 provides a summary of the researcher’s 

observations and of other researchers’ findings on the differences in LFLs determined through 

measurement and calculations using Le Chatelier’s equation. Although it is difficult to judge the 

extent of deviations shown in Table G.1 from the researcher’s qualitative remarks (e.g., “Agreed 

Closely” versus “Deviated Somewhat”) between measured and calculated LFLs, it would appear 

that mixtures containing carbon disulfide, acetylene, and paraffin-hydrocarbon halides give the 

largest deviations. Carbon disulfide and paraffin-hydrocarbon halides seem to effect a broader 

group of mixtures, while it is less clear whether acetylene mixed with non-hydrogen 

combustibles will cause calculated LFL values to vary significantly from the measured values. 

Some variation was seen with methane and acetylene mixtures, but only slight; and carbon 

monoxide and acetylene mixtures had measured and calculated LFLs that were in close 

agreement. 

Coward and Jones (1952) provide graphs (Figures 56 and 57 of Coward and Jones) for 

the LFL of different mixtures of hydrogen sulfide and methane, and hydrogen sulfide and 
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G-2 

hydrogen. Both figures indicate a variation of the calculated LFL up to 1 percentage point below 

the measured value for these mixtures. For hydrogen sulfide and hydrogen mixtures the variation 

does not occur until there is about 22 percent H2S in the mixture. For hydrogen sulfide and 

methane the variation occurs in almost all mixtures, but is less significant for mixtures with 

2 percent or less H2S. 

Paraffin-hydrocarbon halides cause variations of the measured LFL from the calculated 

value; however, current regulations do not allow flaring of fuel gases with significant amounts of 

halides, so these compounds should not affect the calculation of LFLCZ or its effectiveness as a 

parameter to indicate flare combustion performance. It is assumed that the amount of hydrogen 

sulfide and carbon disulfide that is burned in regulated flares is also small because of the 

emission limits of previous rules, such as the NSPS for petroleum refineries (40 CFR 60, 

subparts J and Ja), Onshore Natural Gas Processing (40 CFR 60, subpart LLL), and New Source 

Review permit limits, although this may require further investigation. Therefore, hydrogen 

sulfide and carbon disulfide are also not expected to be present in large enough quantities to 

affect the accuracy of the calculated LFLCZ. 

 
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G-3 

Table G.1. Le Chatelier Accuracy Observations by Coward and Jones (1952) 

Mixture Constituents Differences Between Measured LFL and Values 

Calculated Using Le Chatelier’s Equation1

Hydrogen and carbon monoxide Just exceeded experimental error

Hydrogen and hydrogen sulfide Diverged widely; calculated LFL lower than measured

Hydrogen, carbon monoxide, and methane Approximately accurate

Hydrogen, methane, and ethane Approximately accurate

Hydrogen and ethylene Calculated LFL slightly lower than measured 

Hydrogen and acetylene Calculated LFL lower than measured by a significant 

amount for some mixtures.2 

Hydrogen sulfide and methane Calculated LFL lower than measured

Hydrogen sulfide and acetylene Deviated by maximum of 0.3%

Carbon disulfide and ether, benzene, acetone, or 

acetaldehyde 

No agreement

Carbon monoxide and methane Fair agreement

Carbon monoxide and acetylene Close agreement

Any two of: methane, ethane, propane, and butane Close agreement

Methane and pentane in air Close agreement

Natural gas Fair agreement

Methane and ethylene Calculated LFL slightly lower than measured

Methane and acetylene Calculated LFL slightly lower than measured

Two or more of: cyclohexane, benzene, ethanol Fair agreement at temperatures of 100 to 250°C 

Two or more of: toluene, ethanol, ethyl acetate Fair agreement

Two or more of: methyl cyclohexane, alcohol, 

ether 

Fair agreement

Ethanol and gasoline; ethanol, gasoline and ether Fair agreement

Isopropanol and gasoline Nearly agreed

Benzene and toluene Close agreement

Benzene and ethanol Deviated somewhat3

Benzole, methanol, ethanol Agreed within 11% at temperatures of 100 to 250°C 

Methanol and ethanol Agreed (at 75°C) 

Methanol and ether; methanol and acetone; ethyl 

acetate and benzene; acetaldehyde and toluene; 

ethyl nitrite and ether 

Close agreement

Ethanol and ether; acetone and ether Good agreement

Ethanol and furfural Deviated somewhat 

Ethanol and acetone Nearly agreed

Ether and acetaldehyde Good agreement

Acetone and methyl ethyl ketone Close agreement4

Paraffin-hydrocarbon halides: methyl chloride 

and ethyl chloride; methyl chloride and methyl 

bromide; methyl bromide and ethyl chloride 

Differed appreciably

 

1

 As observed by Coward and Jones (1952) 

2

 For mixtures up to 43% acetylene calculated values significantly lower than measured values. Between 43 to 50% 

acetylene, the difference between calculated and measured LFL decreased and greater than 50% acetylene the 

mixtures follow Le Chatelier’s equations. 

3

 However, at temperatures from 100 to 250 °C the values agreed fairly well. 

4

 However, the LFL measured using an upward propagation method were slightly higher than calculated. 
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G-4 

Table G.1 also indicates that several of the alcohol mixtures had some deviation between 

calculated and measured LFLs, indicating that mixtures of benzene and ethanol, and furfural and 

ethanol deviated “somewhat;” and measured LFL values for ethanol and acetone, and 

isopropanol and gasoline mixtures “nearly agreed” with the calculated values. 

The researchers (Coward and Jones 1952) also noted small differences in calculated and 

measured LFL values of various mixtures of hydrogen, carbon monoxide, methane, ethane, and 

ethylene. The qualitative observations in Table G.1 do not provide enough specific information 

to judge the accuracy of using Le Chatelier’s equation for these mixtures, but they do raise some 

questions regarding how effective the LFLCZ can be as an indicator of good combustion for a 

wide variety of flare gas compositions. 

Research on Reaction Kinetics and the Effect in LFL

Several other researchers (Azatyan et al., 2002, 2005a, 2005b, and 2007; Degges et al. 

2010; Laskin et al., 2000; Wang et al., 2009) have investigated the reaction kinetics of burning 

gases, which is a possible explanation for these observations. In the flame of a flare, numerous 

chemical reactions are taking place with the reaction priority dependent on the temperature and 

chemical composition in the burning mixture. Changes in composition can shift the reaction 

kinetics to other primary pathways that can have consequences on the ability to achieve good 

combustion. The LFL does a reasonable job of “summarizing” (in one number) the reactions 

going on during combustion of a specific chemical or mixture. However, LFL of mixtures 

calculated using Le Chatelier’s equation is only as accurate as how closely the individual 

component’s LFL represent the combustion pathways the chemical follows when it is part of a 

mixture. This is because pure component LFLs used in the Le Chatelier’s equation to estimate 

LFL for the mixture are based on the combustion pathways when the chemical is burned alone. If 

mixture components follow reaction pathways that are different from their pure components, the 

calculated LFLCZ will vary from the actual. However, if the mixture’s reaction pathways result in 

similar heats of combustion, the variation in the calculated LFL and the actual may not be 

meaningful with respect to the LFLCZ of 15.3 percent. Also, other factors may effect the 

accuracy of the calculated LFL which could mitigate or exacerbate the error, either making the 

deviation between calculated and actual LFL more, or less, of a concern. 
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G-5 

The reaction kinetic researchers describe reactions that are considered “chain branching” 

and “chain terminating.” The chain terminating reactions are ones that lead to products that are 

less active or that participate in reaction chains that recreate the original chemical and may not 

result in complete combustion. Degges et al. (2010) describes the addition of steam and its 

chemical influence on the kinetics of a flame including the effect of the steam cooling the flame. 

Azatyan et al. (2002) provides an analysis on the reactions taking place with hydrogen and 

methane combustion (pure gases not a mixture). Wang et al., (2009) analyzes the combustion 

reactions when hydrogen is added to methane in small amounts. The researcher shows an 

enhancement of combustion with the addition of hydrogen to methane and found that in mixtures 

with hydrogen concentrations lower than 20 percent, hydrogen participates predominately in 

intermediate reactions. Azatyan et al. (2005a, 2005b, and 2007) describes the reaction kinetics of 

hydrogen combustion with the addition of small amounts of propylene or isopropanol and shows 

that these compounds can inhibit the combustion reactions of hydrogen to some degree. Laskin 

et al. (2000) investigates the reaction kinetics of 1,3-butadiene combustion. 

Compiled Measured and Calculated LFL 

To judge whether the differences between measured and calculated LFLs are significant, 

experimental LFL values for mixtures were collected and compared to calculated LFL values. 

Experimental LFL values were collected from eight previous studies. These studies were 

identified from the numerous research papers collected for the investigation of flare operating 

parameters that indicate performance level. An exhaustive search for every measured mixture 

LFL was not conducted, but using a targeted search of papers describing calculational or 

modeled methods of determining LFL, a large group of measured values was collected. (Papers 

presenting analytical methods for determining LFL often have experimental values of LFL to 

compare against the computational method.) 

From the eight studies, 221 measured LFL values for mixtures were identified. The 

values were limited to mixtures that had at least two combustible compounds. (Mixtures with 

one combustible and an inert are discussed in Sections G.2 and G.3). Values were found for 

mixtures containing two or more of 16 combustible constituents (hydrogen, methane, ethane, 

ethylene, propane, propylene, butane, methyl ethyl ketone, toluene, dimethyl ether, methyl 
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G-6 

formate, 1,1-difluoroethane, 1,2-dichloroethane, hydrogen sulfide, ammonia, and carbon 

monoxide). Some mixtures also had inert compounds including nitrogen, carbon dioxide and 

carbon tetrachloride. 

Of the 221 measured values, 146 did not have inert compounds and 65 were mixtures 

with inert compounds (mixtures with inerts are discussed in Section G.3). In calculating the LFL 

for the mixture, the pure component LFL values used in Le Chatelier’s equation were taken from 

the research paper that presented the measured values. This was done so the pure component 

LFL was consistent with the same measurement method, technique, apparatus, and environment 

that the mixture LFL was measured. Therefore, any significant variation in measured and 

calculated LFL that was identified was more likely due to the chemical constituents not 

following Le Chatelier’s equation and not a variation due to the test method. LFL was also 

calculated using pure component LFLs from Zabetakis (1965) for comparison. 

Appendix H provides the list of pure component LFL values used in the LFL calculations 

for each research report and the measured LFLs collected, the mixture composition, the 

calculated LFL using Zabetakis referenced pure component LFLs, the calculated LFL using the 

specific researcher’s pure component LFLs, the source of each measured value, and the 

difference between the measured and calculated values. The mixtures with and without inerts are 

presented in two different tables in Appendix H. 

The variation between measured and calculated LFL values is small for those mixtures 

without inert components. The difference between measured and calculated values was between 

-0.30 and 0.88 percentage points and the average deviation for the 146 measured and calculated 

pairs was 0.04 percentage points. Averaging the absolute value of the difference gave an average 

deviation just above this of 0.08 percentage points. There were only six measured LFL values 

that varied from the calculated LFL by ±0.3 percentage points or greater. These are shown in 

Table G.2. 

Table G.2 does not reveal an obvious trend for the six measured LFL values with the 

largest deviation from the Le Chatelier calculated values. One could speculate on what these six 

points imply; however, the overall trend is that for most of the data for mixtures without inerts, 
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G-7 

the Le Chatelier’s equation does a good job of predicting the LFL. With all but six data points 

showing agreement within a third of a percentage point of the measured value, it would be 

difficult to parse out experimental error from any situations where the components do not obey 

Le Chatelier’s principal. Even the mixtures with higher deviations (Table G.2) are very close, all 

but one having a deviation less than 10 percent of the experimental value. 

Furthermore, very few, and possibly none, of the variations between measured and 

experimental LFL values that could occur in gas mixtures containing only combustion gases 

(with no steam or air added) are important from the perspective of flare vent gas. In using the 

LeChatelier equation to estimate the LFLCZ, all mixtures will have a LFLCZ of less than the 

maximum pure component LFL of any component in the mixture. For example, ammonia has the 

highest LFL (15%) of any compound in the Zabetakis (1965) LFL reference list; any mixture of 

a combustible gas with ammonia (and no inert added) will have a LFLCZ less than 15 percent. 

This is dictated by the use of Le Chatelier’s equation. A value greater than 15 percent is not 

mathematically possible. Consequently all mixtures of 100 percent combustible gases, including 

mixtures with ammonia, will have a LFLCZ less than 15.3 percent. Therefore, any flare vent 

stream with no added assist air or steam containing no inerts will have a LFL that meets 

15.3 percent and any variations between calculated and measured are likely unimportant unless 

there are specific chemical interactions that can cause a mixture of only combustibles to have a 

much higher LFL than calculated. The comparison of the Le Chatelier equation to experimental 

LFL values shown in the Appendix H data does not indicate large variations for 100 percent 

combustible mixtures. However, it should be noted that none of these mixtures with measured 

LFLs contained acetylene, carbon disulfide, hydrogen sulfide, or alcohol, which were the 

chemicals specifically cited as not following Le Chatelier’s rule as closely as most chemicals. 

Table G.2. Largest Deviations Found Between Measured and Calculated LFLs 

for Mixtures with No Inerts 

Hydrogen (H2) 

(Volume %) 

Dimethyl ether 

(C

H2 O) 6

(Volume %) 

Ammonia (NH3) 

(Volume %) 

Other 

(Volume %) 

Experimental 

LFL 

(Volume %) 

Researcher Pure Component 

LFLs 

Zabetakis Pure Component 

LFLs 

Source 

Calculated 

LFL 

(Volume %) 

Difference 

Between 

Experimental 

and 

Calculated 

Calculated 

LFL 

(Volume %) 

Difference 

Between 

Experimental 

and 

Calculated 

50 

50% methyl 

ethyl ketone 

(C4H8O) 

3.15 2.75 0.40 2.58 0.57 Loehr et al. 

(1997) 
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G-8 

Table G.2. Largest Deviations Found Between Measured and Calculated LFLs 

for Mixtures with No Inerts (Continued) 

Hydrogen (H2) 

(Volume %) 

Dimethyl ether 

(C

H2 O) 6

(Volume %) 

Ammonia (NH3) 

(Volume %) 

Other 

(Volume %) 

Experimental 

LFL 

(Volume %) 

Researcher Pure Component 

LFLs 

Zabetakis Pure Component 

LFLs 

Source 

Calculated 

LFL 

(Volume %) 

Difference 

Between 

Experimental 

and 

Calculated 

Calculated 

LFL 

(Volume %) 

Difference 

Between 

Experimental 

and 

Calculated 

50 

50% 1,2 

Dichloroethane 

(C2H4C12) 

5.35 4.92 0.43 4.38 0.97 Loehr et al. 

(1997) 

75 

25% 

methyl formate 

(C2H4O2) 

10.94 10.31 0.63 10.00 0.94 Kondo et al. 

(2008) 

75 25% Carbon 

Monoxide (CO) 15.2 14.32 0.88 14.29 0.91 Kondo et al. (2008) 

75 25 3.8 4.10 -0.30 4.21 -0.41 Kondo et al. 

(2008) 

50 50% Methane 

(CH4) 4.25 3.94 0.31 4.05 0.20 Kondo et al. 

(2008) 

G.2 Accuracy of Le Chatelier Principle for Mixtures with Nitrogen 

As stated in Section 3.1.2, the calculation of LFLCZ shown in Figure 3-3 of this report 

includes an assumption that inerts have an infinite LFL. This can be understood for nitrogen by 

reviewing the same Zabetakis nose plot that is shown in Figure 3-1 of this report (for ease, refer 

to Figure G.1 as it is the same as Figure 3-1 of this report). The plot shows several different 

inerts in methane and air mixtures. Looking specifically at the LFL portion of the nitrogen curve 

in Figure G.1, the methane concentration stays relatively constant as nitrogen is added (moving 

along the line at the underside of the nose). The LFL boundary is relatively horizontal as 

nitrogen is added and the concentration of the combustible stays constant in the combustiblenitrogen-air mixture. Additional combustion gas (methane in Figure G.1) is not needed to offset 

the addition of nitrogen. The addition of nitrogen displaces air in the methane-nitrogen-air 

mixture. Because the chemical properties of nitrogen and air are very similar and the amount of 

oxygen in air displaced by the nitrogen is not significant (since oxygen is not a limiting factor at 

the LFL), the addition of nitrogen does not affect the measured methane concentration at the 

LFL. 
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G-9 

Figure G.1. Zabetakis nose plot for methane and inert in air

In the research conducted on calculating the LFL, mixtures with inert gases have 

specifically been investigated. In general, the more inert (for flares, these are generally nitrogen, 

carbon dioxide, and steam) added to a combustible gas, the higher the LFL will be for the 

mixture. Le Chatelier’s principle was investigated regarding its accuracy to estimate LFL values 

for combustion gases and nitrogen. It was found that the LFL calculated using Le Chatelier’s 

principle deviates somewhat from experimental values with higher concentrations of nitrogen in 

the fuel mixture (and as the fuel mixture’s LFL approaches the UFL at the point of maximum 

inert – at the tip of the nose on a Zabetakis plot). In addition, the amount the calculated value 

deviates from the experimental value is dependent on the type of combustible compound in the 

fuel mixture. 

Experimental LFL data for individual combustible compounds (i.e. hydrogen, methane, 

ethane, ethylene, propane, propylene, butane, butadiene, carbon monoxide, dimethyl ether, 

methyl formate, 1,1-difluoroethane, and ammonia) in nitrogen were graphed against the nitrogen 

volume percent used in the experiments. The experimental data were extracted from a series of 
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G-10 

tests conducted by Jones and Kennedy (1932, 1938, and 1943), Kondo et al. (2006a and 2006b), 

and Vidal et al (2006). Figure G.2 shows the effect of increasing nitrogen on the LFL of 

methane-nitrogen mixtures. Appendix F provides these graphs for all 13 combustible gases in 

nitrogen. 

Figure G.2 illustrates the variation that can be found between the calculated LFL and the 

measured LFL for methane and nitrogen mixtures. For most compounds considered (see 

Appendix F for graphs showing other combustible compounds with nitrogen) there is little to no 

variation identified between the calculated and experimental LFL values. The deviations that 

were found generally occur at greater than a LFL of 15.3 percent. Other than ammonia, methane 

showed the greatest deviation at the lowest concentration for the compounds examined. Looking 

at a nitrogen concentration of 66 percent on Figure G.2, which crosses the LFL lines near the 

LFL value of 15.3 percent shown in Figure 3-4 of this report (where combustion efficiency 

begins to degrade), the LFL is 15.5 percent for the experimental methane data and 14.8 percent 

for the calculated results. 

Figure G.2. Effect of Nitrogen on LFL of Methane 

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methane and Nitrogen

Calculated using Le Chatelier Equation assuming an LFL 

for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006a)

Experimental Data - (Vidal et al., 2006)

Pure component LFL at 5.0 for Jones and Kennedy (1933) and Vidal et al. (2006) experiments, but 4.9 for Kondo et al. (2006a) experiments.
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G-11 

This example illustrates that there could be situations where a source would be in 

compliance with the proposed 15.3 percent LFLCZ threshold for a given gas composition in the 

combustion zone based on the calculated value, but a measured LFL for the mixture would not 

support this conclusion. However, most combustion gases showed a smaller deviation near 

15.3 percent LFL (see Appendix F) between the calculated and measured LFL and the deviation 

for methane is relatively small and would not affect a wide range of flare gas compositions. Also 

the uncertainties in using Le Chatelier’s equation are built into the analysis that resulted in the 

proposed 15.3 percent maximum LFL. The limit was established using the Le Chatelier equation 

and pure component LFLs from Zabetakis (1965). Therefore, the determination of compliance 

should occur using Le Chatelier’s equation and Zabetakis pure component LFLs. 

In reviewing the information in Figures G.1 and G.2 it is important to consider the 

differences in these graphs. Figure G.1 shows the percent of methane in a mixture of air and 

added nitrogen (the LFL of the methane in a “nitrogen-enriched” atmosphere); from the 

perspective of a flare, Figure G.1 represents the gas mixture immediately above the flare. The 

x-axis of Figure G.1 shows the composition of added nitrogen in the “nitrogen-enriched” 

atmosphere. Therefore, the quantity of air can be calculated from the figure by adding the LFL of 

methane and the amount of added nitrogen, and subtracting from 100 percent; or for example, for 

Figure G.1, added nitrogen of 20 percent shows a 5 percent concentration of methane, which 

means there is 75 percent air in the mixture at the LFL (100-20-5=75). Figure G.2 shows the 

same data but gives the LFL of a mixture containing methane and nitrogen (instead of the LFL 

for methane in nitrogen-enriched air). The LFL value still reflects the percent of “combustible” 

in air (“combustible” for Figure G.2 represents the mixture of methane and added nitrogen); 

however, the x-axis of Figure G.2 is the composition of the nitrogen in the methane-nitrogen 

mixture, prior to mixing with air. Using the same example for illustration, the mixture of 

5 percent methane, 20 percent added nitrogen and 75 percent air on Figure G.2 would be shown 

with a nitrogen content of 80 percent because the ratio of methane to added nitrogen in the air 

would be the same as in the mixture without air (5% methane to 20% nitrogen would be 

20/(5+20) or 80% nitrogen). From Figure G.2, 80 percent nitrogen in the mixture shows about a 

25 percent LFL which means 25 percent of the methane-nitrogen mixture in air is the LFL for 

that mixture. This Figure G.2 example also indicates that there would be 75 percent air at this 

LFL which is consistent with the Figure G.1 Zabetakis plot. 
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G-12 

G.3 Using Le Chatelier’s Equation for Non-Nitrogen Inerts 

The amount of deviation between a calculated LFL (using the Le Chatelier equation and 

assuming infinity as the LFL for inerts) from experimental LFL values is also dependent on the 

type of inert in the gas mixture. For example, Figure G.1 shows the variation of the LFL for 

methane in different inert-air mixtures. As can be seen, it takes a concentration lower than 

nitrogen for some non-nitrogen inerts to render the mixture non-flammable. Also the bottom of 

the nose is less horizontal. For carbon dioxide, water, carbon tetrachloride, and methylene 

bromide, Figure G.1 shows that more methane must be added as the inert is increased to maintain 

a combustible gas mixture. The carbon dioxide curve has the least horizontal nose underside of 

the three inerts that are most often found in flare gas (steam, nitrogen, and carbon dioxide). For 

example, at a carbon dioxide concentration in the methane, air, and carbon dioxide mixture of 

20 percent, the methane concentration (LFL) would have to increase an additional percent, up to 

about 6 percent, in order to maintain a flammable mixture. Based on Figure G.1, the effect of 

water in a mixture is greater than nitrogen, but not as great as carbon dioxide. Because of this, 

we have included a ‘nitrogen equivalency’ adjustment in the Le Chatelier’s equation (see 

Equation 3-3 in Section 3.1.2 of this report). 

Figure G.3 provides the same data for methane and carbon dioxide mixtures in 

Figure G.1 but in the same format as Figure G.2. The figure shows three sets of experimental 

data for methane and carbon dioxide mixtures and two sets of calculated values. One set of 

calculated LFLs (outline of red squares) is based on using Le Chatelier’s equation assuming a 

LFL of infinity for the carbon dioxide (no adjustment) and the other set of calculated LFL values 

(outlines of blue circles) is the Le Chatelier’s equation with values adjusted for nitrogen 

equivalency. The nitrogen adjusted calculated LFL values are very close to the experimental data 

from all three sources. Comparing the two calculated sets of values shows the effect of the 

nitrogen adjustment. Left unadjusted, the calculated LFL for methane in CO2 would deviate a 

significant amount even for mixtures with a LFL less than 15 percent. Appendix F provides 

similar graphs to Figures G.2 and G.3 (with nitrogen and carbon dioxide) for 13 combustible 

gases. 
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G-13 

Figure G.3. Effect of Carbon Dioxide on LFL of Methane 

As mentioned in Section 3.1.2 of this report, the nitrogen equivalency methodology is 

based on a method investigated by Molnarne et al. (2005) for taking into account non-nitrogen 

inerts which is used in ISO 10156 and described by Besnard (1996). This nitrogen equivalency 

adjustment is also described by Shore (2007) and Gogolek et al. (2010a). The Molnarne et al. 

(2005) recommended values of “nitrogen equivalency” are tabulated in Table G.3 for water (or 

steam) and carbon dioxide, the inerts that are likely to be included in the flare combustion zone 

(other than nitrogen). 

Table G.3. Recommended Values of Nitrogen Equivalency for Water and Carbon 

Dioxide 

Combustible Component i in Flare Vent Gas KWater KCarbon Dioxide

Methane 1.87 2.23

Ethane 1.40 1.87

Propane 1.51 1.93

Ethylene 1.68 1.84

Propylene 1.36 1.92

Hydrogen 1.35 1.51

All Other Combustibles 1.50 1.87

Source: Molnarne et al. (2005)

0%

5%

10%

15%

20%

25%

30%

35%

40%

45%

50%

0% 10% 20% 30% 40% 50% 60% 70% 80% 90%

Lower Flammability Limit of the Mixture (%)

Inert in the Mixture (%)

Methane and Carbon Dioxide

Calculated using Le Chatelier Equation with no equivalency 

adjustment and assuming an LFL for Inert of Infinity

Calculated using Le Chatelier Equation adjusting CO2 for Nitrogen 

equivalency and assuming and LFL for Nitrogen of Infinity

Experimental Data (Jones and Kennedy, 1933)

Experimental Data (Kondo et al., 2006b)

Experimental Data (Vidal et al., 2006)

Pure component LFL at 5.0 for Jones and Kennedy (1933) and Vidal et al. (2006) experiments, but 4.9 for Kondo et al. (2006b) experiments.
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G-14 

As can be seen in Figure G.2 and in Appendix F for other combustibles, the assumption 

that the LFL is equal to infinity works well for nitrogen, but not as well for other inerts (see 

Figure G.3 and Appendix F for carbon dioxide). The intent of the adjustment is to normalize the 

effect of non-nitrogen inerts in the gas mixture and to adjust the LFL to be on the same basis 

with respect to inerts. After making the adjustment, all inerts are assumed to have an infinite LFL 

and the Le Chatelier equation can be used. The flare data sets used in the final analysis have gas 

mixtures that contain various amounts of nitrogen, carbon dioxide, and water. 

Compiled Measured and Calculated LFL for Mixtures with Inerts 

As previously described, Appendix H provides measured LFLs for 65 mixtures that 

contain inert components. The variation between the measured and the calculated value 

(calculated using the specific researcher’s pure component LFL values and adjusting for nitrogen 

equivalency) ranged from -1.56 and 4.88 percentage points, and an average of 0.73 percentage 

points. The information for these mixtures is shown in Table H.3 of Appendix H. The mixtures 

in Table H.3 include the inerts carbon tetrachloride, nitrogen, carbon dioxide, or both nitrogen 

and carbon dioxide; combustibles include hydrogen, methane, ethane, methyl ethyl ketone, 

toluene, 1,2-dichlorethane, and carbon monoxide. Table G.4 summarizes the variation between 

measured and calculated LFLs for the different inerts and shows a summary for the mixtures 

with no inerts. Clearly the addition of inert compounds has increased the variation between 

calculated and measured LFLs for mixtures. The average of the absolute value of the differences 

between calculated and measured LFLs for mixtures with inerts is roughly ten times that seen 

with mixtures with no inerts. 

Table G.4. Summary of Differences Between Calculated and Measured LFLs for Mixtures 

Type of Inert Average 

Difference

Average of the 

Absolute 

Value of 

Difference

Maximum Minimum

Number 

of 

Mixtures

No Inert 0.04 0.08 0.88 -0.30 145

All Inerts 0.73 0.94 4.88 -1.56 65

Carbon Tetrachloride 0.74 0.74 2.52 0.08 18

Carbon Dioxide -0.78 0.78 -0.19 -1.56 5

Nitrogen 0.71 0.91 4.88 -0.70 24

Carbon Dioxide and Nitrogen 1.16 1.21 2.97 -0.21 18

Carbon Dioxide and/or Nitrogen 0.72 1.01 4.88 -1.56 47
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G-15 

The differences summarized in Table G.4, as well as shown in Appendix H, are 

calculated as the measured minus the calculated values. Therefore, a negative value indicates that 

the calculated value is higher than the measured value. In less than 25% of the mixtures with 

inerts (15 out of 65) the difference was negative. Generally for mixtures with inert 

concentrations close to the maximum amount of inert possible, while still flammable, the 

deviation between measured and calculated LFL is at its greatest, with the measured LFL being 

greater than the calculated value. This can be quickly seen for mixtures with only one 

combustible and one inert in Figures G.2 and G.3 and in Appendix F; note the difference in 

acceleration of the increase in slope between the measured and calculated values as inert 

increases. Look at the termination of the calculated data in Figure G.2. The calculated data ends 

at a nitrogen content of about 86 percent and a LFL of 35 percent; the measured data from Jones 

and Kennedy (1933) ends at a LFL of about 39.5 percent, giving a maximum difference between 

measured and calculated data of 4.5 percentage points near the point of maximum inert 

concentration.5 As previously discussed regarding Figure G.2, the difference in measured and 

calculated LFL at 66 percent nitrogen concentration is about 1 percentage point. Comparing this 

to the 4.5 percentage point difference shows that the greater difference between measured and 

calculated LFLs are at maximum inert concentrations. Looking at Figure G.2 also shows why the 

difference is expected to be a positive difference; Le Chatelier’s equation does not account for 

the same level of increasing slope that appears to happen as mixtures approach their lower limit. 

Except for two data points with carbon tetrachloride as the inert and one negative difference, the 

difference between measured and calculated LFL is less than 1 percentage point for LFLs 

measured less than 25%. Therefore, the larger positive differences seen in Appendix H and 

Table G.4 are at least partly due to the mixture being closer to the maximum inert concentration. 

Most of the experimental data found were for mixtures of hydrogen, methane, carbon 

monoxide, nitrogen, and/or carbon dioxide, except for mixtures with carbon tetrachloride as the 

inert. The calculated LFL for the carbon tetrachloride mixtures were adjusted for nitrogen 

equivalency, but because a k value was not available for carbon tetrachloride, the k values for 

carbon dioxide were used. All of the measured and calculated differences for these mixtures 

were less than 1 except for two mixtures that were two of the highest LFLs (9.7 and 10.1%). 

 

5 The Kondo, 2006a, data point in Figure G.2 at about 88 percent nitrogen and a LFL of 44.5 percent is an 

inconsistency between the measurement methods used by Kondo and Jones and Kennedy in determining the 

maximum concentration of nitrogen that will still allow flame propagation.
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G-16 

Although these are not high values with respect to the 15.3% that appears to be important for 

good flare combustion, these could be high values for carbon tetrachloride. Carbon tetrachloride 

is a stronger inert than either carbon dioxide or nitrogen. 

The mixtures with inert other than carbon tetrachloride all have hydrogen in them except 

for eight with methane and ethane. The methane and ethane mixtures all have LFLs of 21% or 

less, and the range of differences between measured and calculated is -0.15 to 0.43 with an 

average of 0.14 percentage points. These appear very close differences. The second highest 

difference is 0.33 percentage points, so most of the differences are less than what it may appear 

from the range. 

There are six additional mixtures containing ethane. Three of these also include 

hydrogen, methane, and nitrogen. All of these have high differences between measured and 

calculated LFL ranging from 2.47 to 4.88 percentage points. They also have very high LFLs of 

37 to 42%. Therefore, the mixtures are likely close to their maximum inert concentration and any 

additional cause for the high differences cannot be determined given only 3 data points. 

The other three mixtures containing ethane also include hydrogen and nitrogen. The 

range of differences between measured and calculated values is 0.15 to 2.74 percentage points. 

The higher difference (2.74) is for a mixture with a high LFL or 25.8, but it is difficult to form 

any conclusions with these three points. However, because of ethane’s similarity to methane, 

these points could be compared to the results for hydrogen and methane. 

The remainder of the mixtures (33) with inert other than carbon tetrachloride, contain 

methane, hydrogen and nitrogen and/or carbon dioxide, with seventeen of the mixtures also 

containing carbon monoxide. 

Numerous researchers (for example Karim et al., 1996, Wang et al. 2009 and 2010, Hu et 

al. 2009, and White, 1925) have shown that hydrogen enhances the combustion of methane. For 

mixtures that are not premixed, this effect levels off somewhat at about 32 or 40 percent 

hydrogen in the fuel mixture (Wang et al., 2010). From the perspective of LFL values for 

mixtures of hydrogen and methane, one would expect that the LFL values would be lower than 
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been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

G-17 

what would be calculated using Le Chatelier’s equation, because the equation does not take into 

account the enhancement behavior. However, for mixtures where there is more hydrogen than 

methane, hydrogen still enhances the methane combustion, but the primary combustible is 

hydrogen; and from the perspective of hydrogen, methane is inhibiting hydrogen combustion. 

Because, methane has a higher LFL than hydrogen, the use of Le Chatelier’s equation results in a 

“dilution” of the hydrogen LFL by an amount proportional to methane’s LFL and the 

concentration of methane in the mixture. This calculation may be enough to represent the 

inhibition of hydrogen combustion by methane or there would be positive differences between 

the measured and calculated LFL values. 

Table G.5 shows all the mixtures with hydrogen, methane, and an inert. The values in 

Table G.5 are shown in order of the relative amount of hydrogen in total combustibles [hydrogen 

concentration/(hydrogen plus methane concentrations)]. As expected the difference between the 

measured and calculated LFL values is negative for mixtures where hydrogen is 33% or less 

showing the enhancement in combustion is likely due to the presences of hydrogen. For mixtures 

with a relative amount of hydrogen of 66% or greater, the difference between the measured and 

calculated LFLs is between -0.66 and 0.55 percent. There are four of these mixtures and two 

have negative difference and two with positive differences. These correspond with the inert in 

the mixture; carbon dioxide is part of the mixtures with negative differences, and nitrogen is the 

inert in the mixtures with positive differences. Gogolek et al. (2010a) recommended different kvalues from those provided in Molnarne et al. (2005) and shown in Table G.3. The Gogolek et al. 

recommended values are shown in Table G.6. Using the Gogolek et al. recommended k values, 

the differences between the measured and calculated LFLs is between -0.07 to 0.55. This is 

closer to what one would expect for hydrogen and methane mixtures when hydrogen makes up 

66 percent or more of the combustibles. 
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should not be construed to represent any Agency determination or policy.

G-18 

Table G.5: Hydrogen, Methane Mixtures With Either Only Nitrogen or Carbon Dioxide as th

Hydrogen 

(H2) 

(Volume %) 

Methane (CH4) 

(Volume %) 

Nitrogen (N2) 

(Volume %) 

Carbon Dioxide 

(CO2) 

(Volume %) 

Experimental 

LFL 

(Volume %) 

Calculated 

LFL 

(Volume %) 

Difference 

Between 

Calculated 

Experimental 

LFL Values 

% Error 

Fraction 

Hydrogen 

Total 

Combustib

(volume 

fraction

3.6 20.6 75.8 20.85 21.55 -0.70 -3.23 05.6 29.7 64.7 14.37 14.75 -0.38 -2.57 06.9 26.5 66.7 15 15.38 -0.38 -2.48 05.1 18.7 76.3 21.17 21.58 -0.41 -1.88 07.3 23.6 69.1 16.07 16.43 -0.36 -2.20 0

10.4 29.0 60.6 13.2 13.91a -0.71a -5.13 010.5 21.1 68.4 15.93 17.49b -1.56b -8.94 016.6 16.7 66.7 14 13.58 0.42 3.06 010.9 10.9 78.2 21.5 20.74 0.76 3.65 015.1 15.1 69.8 16.41 17.20c -0.79c -4.57 012.8 12.8 74.5 18.55 18.46 0.09 0.51 025.3 25.2 49.5 9.5 8.95 0.55 6.13 014.7 7.6 77.6 20.35 20.15 0.20 0.98 016.5 4.4 79.1 22.75 23.41d -0.66d -2.83 013.5 2.8 83.7 27 26.45 0.55 2.09 018.3 3.6 78.1 21.59 21.78e -0.19e -0.87 0

a Using k values recommended by Gogolek et al. (2010a), the calculated LFL would be 13.39 % with a difference of -0.19 % from b Using k values recommended by Gogolek et al. (2010a), the calculated LFL would be 16.65 % with a difference of -0.72 % from c Using k values recommended by Gogolek et al. (2010a), the calculated LFL would be 16.56 % with a difference of -0.15 % from d Using k values recommended by Gogolek et al. (2010a), the calculated LFL would be 22.82 % with a difference of -0.07 % from e Using k values recommended by Gogolek et al. (2010a), the calculated LFL would be 21.37 % with a difference of +0.22 % from 
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G-19 

Table G.6. Nitrogen Equivalency Values from 

Molnarne et al. (2005) and Gogolek et al. (2010a) 

Species 

Molnarne et al. 

Nitrogen 

Equivalency 

Values for CO2

Gogolek et al 

Nitrogen 

Equivalency 

Values for CO2

Methane 2.23 1.6 

Propane 1.93 2.5 

Ethylene 1.84 1.8 

Ethane 1.92 2.4 

Hydrogen 1.51 1.5 

There are five mixtures that have 50 percent of the combustibles as hydrogen. The 

differences between the measured and calculated values are 0.09 to 0.76% for the mixtures with 

nitrogen, and -0.79% for the mixture with carbon dioxide. Using the Gogolek et al. (2010a) 

recommended k values, the difference for the mixture with carbon dioxide between the measured 

and calculated LFL is -0.15 percent. The Jones (1929) values have a much higher level of 

uncertainly than the newer Karim et al. (1985) data has. For the higher inert mixtures, the 

uncertainty is at least ±0.5 percent for the Jones data. Also, the highest difference for hydrogen 

of 66% or greater for the Karim et al data set (0.55%) is for the highest inert concentration, 

which one would expect to result in a higher positive difference. Taking the high uncertainty in 

the Jones data and the high inert concentration mixture into account and using Gogolek et al. 

recommended k values, the mixtures with hydrogen of 66 percent and greater have close 

calculated and measured LFLs; the range of differences is from -0.15 and 0.20 percent. 

There are 17 mixtures with hydrogen, methane, and carbon monoxide. Small amounts of 

hydrogen in carbon monoxide mixtures have also been found to enhance combustion (Wierzba 

and Kilchyk, 2001). Wierzba and Kilchyk showed that 20 percent or less hydrogen with carbon 

monoxide, will result in the calculated LFL values to be greater than the measured values. This 

behavior is not identifiable by the measured LFL found for the 17 mixtures of hydrogen, 

methane, and carbon monoxide. These data all came from Jones (1929) and have a large 

variability. Eleven of the mixtures have an inert concentration of 79 percent or greater; therefore, 

the mixtures are close to its LFL where the variation between measured and calculated are at 

their greatest. The difference between measured and calculated LFL from the 17 mixtures is 
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G-20 

between -0.21 and 2.97 percent. Just looking at those mixtures with an inert concentration less 

than 79 percent the variation is -0.21 and 0.71 percent. All but two of the mixtures had carbon 

dioxide and nitrogen as inert; therefore, these would be affected by using the Gogolek et al. 

(2010a) k values. Using these k values, the range in differences for the 17 mixtures is -0.21 to 

3.48 and looking only at the mixtures with inert concentration less than 79 percent, the variation 

would be -0.21 to 0.85 percent. In both cases the range widens between the calculated and 

measured differences. 

The differences in the measured and calculated LFLs are relatively close when looking at 

mixtures that are not as close to the maximum inert concentration. For hydrogen concentrations 

of 33% or less of the total combustible, the data appears to show some combustion enhancement 

by hydrogen (although the dataset is small so this can only be considered a preliminary 

observation). Additional observations are less clear or not possible with the data collected. Also, 

it should be noted that no data was found to investigate other chemical interactions reported in 

literature (e.g., propylene inhibition of hydrogen combustion). 
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Appendix H 

Effect Of Nitrogen And Carbon Dioxide On 

The LFL Of Various Components. 

[comparison of Le Chatelier equation to experimental LFL values] 
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represent and should not be construed to represent any Agency determination or policy.

H-1 

Table H.1. List of pure component LFL values used in the LFL calculations for each research re

Component i

LFLi

Zabetakis 

(1965)

Van den 

Schoor 

(2008)

Kondo 

et al. 

(2008)

Vidal et 

al. 

(2005)

Mashuga 

(1999)

Loehr et 

al. 

(1997)

Karim 

(1985)

Jon

(19

Hydrogen (H2) 4 - - - - 5 4.13 4 Methane (CH4) 5 4.4 4.9 5 4.85 - 5.47 5.Ethane (C2H6) 3 - - - - - - - Ethylene (C2H4) 2.7 - 2.74 2.7 2.62 - 3.22 - Propylene (C3H6) 2.4 - 2.16 - - - - - Propane (C3H8) 2.1 - 2.03 - - - 2.4 - Butane (C4H10) 1.8 - - - - - - - Methyl ethyl ketone (C4H8O) 1.9 - - - - 1.95 - - Toluene (C7H8) 1.2 - - - - 1.2 - - Dimethyl ether (C2H6O) 3.4 - 3.3 - - - - - Methyl formate (C2H4O2) 5 - 5.25 - - - - - 1,1 Difluoroethane (C2H4F2) 4.32 a - 4.32 - - - - - 1,2 Dichloroethane (C2H4Cl2) 4.85 b - - - - 4.85 - - Hydrogen Sulfide (H2S) 4 - - - - - - - Carbon Monoxide (CO) 12.5 - 12.2 - - - 13.76 13Ammonia (NH3) 15 - 15.2 - - - - - a Zabetakis (1965) does not include a LFL for 1,1 Difluoroethane. This value is from Kondo et al. (2008).b Zabetakis (1965) does not include a LFL for 1,2 Difluoroethane. This value is from Loehr et al. (1997).
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H-2 

Table H.2. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference between 

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Ethylene (C2H4) 

Propylene (C3H6) 

Propane (C3H8) 

Butane (C4H10) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

Dimethyl ether 

(C2H6O) 

Methyl formate 

(C2H4O2) 

1,1 Difluoroethane 

(C2H4F2) 

1,2 Dichloroethane 

(C2H4Cl2) 

Ammonia (NH3) 

Carbon Monoxide 

(CO) Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated 

and 

Experimental 

Differ

On 

Perc

Bas

(%

 75 25 3.8 4.21 -0.41 -10.

40 60 4 4.55 -0.55 -13.

 50 50 5.2 5.54 -0.34 -6.6 25 75 8.71 9.09 -0.38 -4.3

33.3 33.3 33.4 3.91 3.66 0.25 6.5

60 40 4 4.35 -0.35 -8.7

 50 50 2.65 2.73 -0.08 -3.0

20 80 4.2 4.76 -0.56 -13.

 25 75 6.44 6.55 -0.11 -1.7 75 25 3.35 3.40 -0.05 -1.3

23.1 76.9 5 4.73 0.27 5.4

 50 50 2.7 2.73 -0.03 -1.1

50 50 4.63 4.44 0.19 4.0

3.13 3.13 93.7 12.21 11.23 0.98 8.0

40 40 20 4.25 3.94 0.31 7.330 40 30 4.13 3.76 0.37 9.0

 50 50 4.42 4.44 -0.02 -0.4 33 33 33 3.41 3.56 -0.15 -4.5 50 50 3.54 3.68 -0.14 -4.0 50 50 1.45 1.47 -0.02 -1.4

40 30 30 4.05 3.69 0.36 8.9

 50 50 3.75 4.14 -0.39 -10.

33.3 33.3 33.3 4.05 3.66 0.39 9.7

 33 33 33 3.2 3.27 -0.07 -2.2 33 33 33 3.4 3.57 -0.17 -5.0 33 33 33 3.29 3.34 -0.05 -1.4 50 50 4.62 4.58 0.04 0.9

83.3 8.34 8.33 4.46 4.32 0.14 3.2

 33 33 33 2.53 2.70 -0.17 -6.5

7.15 85.7 7.15 5.56 5.13 0.43 7.7

 33 33 33 4.02 4.07 -0.05 -1.2 25 75 3.74 3.72 0.02 0.6 50 50 2.91 2.96 -0.05 -1.6
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H-3 

Table H.2. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference between 

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Ethylene (C2H4) 

Propylene (C3H6) 

Propane (C3H8) 

Butane (C4H10) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

Dimethyl ether 

(C2H6O) 

Methyl formate 

(C2H4O2) 

1,1 Difluoroethane 

(C2H4F2) 

1,2 Dichloroethane 

(C2H4Cl2) 

Ammonia (NH3) 

Carbon Monoxide 

(CO) Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated 

and 

Experimental 

Differ

On 

Perc

Bas

(%

 50 50 2.4 2.54 -0.14 -5.8 33 33 33 2.63 2.77 -0.14 -5.4 50 50 2.08 2.24 -0.16 -7.6

30 50 20 4.03 3.70 0.33 8.1

 50 50 3.66 4.03 -0.37 -10. 75 25 5.89 6.00 -0.11 -1.8 25 75 3.07 3.05 0.02 0.6

25 75 8.69 8.17 0.52 6.0

 33 33 33 3.26 3.34 -0.08 -2.4 75 25 2.56 2.65 -0.09 -3.5 50 50 2.51 2.60 -0.09 -3.4 50 50 4.05 4.05 0.00 0.0 33 33 33 3.37 3.57 -0.20 -6.0 50 50 2.33 2.36 -0.03 -1.3 33 33 33 1.95 1.94 0.01 0.7 50 50 2.76 2.83 -0.07 -2.4 50 50 3.74 3.81 -0.07 -1.7 33 33 33 2.61 2.66 -0.05 -1.7 50 50 3.48 3.60 -0.12 -3.3 50 50 3.06 3.24 -0.18 -5.9 50 50 2.88 3.09 -0.21 -7.1 50 50 4.74 4.64 0.10 2.2 75 25 2.72 3.01 -0.29 -10. 75 25 2.47 2.70 -0.23 -9.3 75 25 2.57 2.62 -0.05 -1.8 25 75 3.87 3.93 -0.06 -1.6 75 25 2.25 2.32 -0.07 -3.2 50 50 3 3.01 -0.01 -0.3 50 50 7 7.14 -0.14 -2.0 25 75 2.18 2.22 -0.04 -2.0 25 75 4.02 4.05 -0.03 -0.6 75 25 3.65 3.70 -0.05 -1.2 75 25 2.14 2.32 -0.18 -8.2 33 33 33 3.43 3.46 -0.03 -0.8
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H-4 

Table H.2. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference between 

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Ethylene (C2H4) 

Propylene (C3H6) 

Propane (C3H8) 

Butane (C4H10) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

Dimethyl ether 

(C2H6O) 

Methyl formate 

(C2H4O2) 

1,1 Difluoroethane 

(C2H4F2) 

1,2 Dichloroethane 

(C2H4Cl2) 

Ammonia (NH3) 

Carbon Monoxide 

(CO) Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated 

and 

Experimental 

Differ

On 

Perc

Bas

(%

 50 50 3.37 3.32 0.05 1.3 33 33 33 2.63 2.77 -0.14 -5.4

6.24 93.8 12.03 11.04 0.99 8.2

 75 25 3.13 3.05 0.08 2.5 75 25 5 4.81 0.19 3.7 50 50 2.63 2.81 -0.18 -6.9 75 25 2.36 2.41 -0.05 -2.1 25 75 3.16 3.19 -0.03 -1.0

83.3 16.7 4.7 4.51 0.19 4.0

 75 25 2.39 2.59 -0.20 -8.3 33 33 33 3.56 3.71 -0.15 -4.2

80 10 10 3.97 3.74 0.23 5.8

 25 75 4.48 4.47 0.01 0.1 75 25 5.29 5.26 0.03 0.6 50 50 2.9 2.96 -0.06 -1.9

40 40 20 3.98 3.63 0.35 8.7

 25 75 3.81 3.76 0.05 1.4 50 50 3.55 3.51 0.04 1.2

80 10 10 4.15 3.89 0.26 6.2

 25 75 3.63 3.70 -0.07 -1.8

14.3 14.3 71.4 2.83 2.47 0.36 12.6

 50 50 3.64 3.51 0.13 3.6 33 33 33 4.88 4.80 0.08 1.6

33 33 33 2 1.88 0.12 5.8

33.3 33.3 33.3 6.07 5.66 0.41 6.7

 25 75 7.33 7.49 -0.16 -2.1

64.3 35.7 5.55 5.28 0.27 4.8

50 50 6.4 6.06 0.34 5.3

7.7 84.6 7.7 5.12 4.61 0.51 9.9

16.7 83.3 9.96 9.23 0.73 7.38.33 83.3 8.33 4.88 4.40 0.48 9.8

 50 50 3.06 3.24 -0.18 -5.9

50 50 3.1 2.75 0.35 11.1

 50 50 13.6 13.64 -0.04 -0.2
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H-5 

Table H.2. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference between 

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Ethylene (C2H4) 

Propylene (C3H6) 

Propane (C3H8) 

Butane (C4H10) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

Dimethyl ether 

(C2H6O) 

Methyl formate 

(C2H4O2) 

1,1 Difluoroethane 

(C2H4F2) 

1,2 Dichloroethane 

(C2H4Cl2) 

Ammonia (NH3) 

Carbon Monoxide 

(CO) Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated 

and 

Experimental 

Differ

On 

Perc

Bas

(%

 50 50 5.26 5.35 -0.09 -1.6 50 50 4.66 4.64 0.02 0.5 50 50 6.45 6.42 0.03 0.4

28.6 71.4 2.8 2.43 0.37 13.212.5 12.5 75 3.57 2.99 0.58 16.162.5 37.5 3.33 2.99 0.34 10.3

 75 25 3.7 3.72 -0.02 -0.4

37.2 62.8 3.7 3.52 0.18 4.7

 75 25 6.22 5.88 0.34 5.4 50 50 7.44 7.14 0.30 3.9 75 25 3.82 3.93 -0.11 -3.0 25 75 8.48 8.48 0.00 -0.0

33 33 33 3.4 3.08 0.32 9.251 29.4 19.6 3.97 3.58 0.39 9.9

77.8 22.2 3.67 3.33 0.34 9.2

50 50 2.05 1.85 0.20 9.950 50 2.05 1.85 0.20 9.9

42.9 57.2 3.04 2.64 0.40 13.2

25 25 25 25 2.4 2.20 0.20 8.225 25 25 25 2.4 2.20 0.20 8.2

 50 50 7.53 7.50 0.03 0.4

62.8 37.2 3.6 3.30 0.30 8.2

 50 50 2.05 1.92 0.13 6.1 50 50 6.86 6.71 0.15 2.2 50 50 7.94 7.50 0.44 5.5

82.8 17.2 2.9 2.69 0.21 7.1

 50 50 5.22 5.00 0.22 4.2

90.5 7.1 1.6 0.8 4.8 4.61 0.19 3.8

 75 25 5.15 5.00 0.15 2.9

90.7 9.3 4.9 4.71 0.19 3.9

30 40 30 3.89 3.36 0.53 13.6

 25 75 6 5.92 0.08 1.4

33 33 33 2.65 2.35 0.30 11.3

83.5 16.5 4.75 4.50 0.25 5.1
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H-6 

Table H.2. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference between 

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Ethylene (C2H4) 

Propylene (C3H6) 

Propane (C3H8) 

Butane (C4H10) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

Dimethyl ether 

(C2H6O) 

Methyl formate 

(C2H4O2) 

1,1 Difluoroethane 

(C2H4F2) 

1,2 Dichloroethane 

(C2H4Cl2) 

Ammonia (NH3) 

Carbon Monoxide 

(CO) Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated 

and 

Experimental 

Differ

On 

Perc

Bas

(%

74.4 25.6 3.65 3.44 0.21 5.891.2 8.8 4.75 4.46 0.29 6.1

 50 50 3.65 3.51 0.14 3.9 25 75 6.31 6.49 -0.18 -2.8

33.4 33.3 33.3 3.83 3.24 0.59 15.4

89.5 10.5 4.5 4.21 0.29 6.3

 50 50 4.25 4.05 0.20 4.7

50 50 3.15 2.58 0.57 18.250 50 5.35 4.38 0.97 18.0

 25 75 10.94 10.00 0.94 8.5 75 25 15.2 14.29 0.91 6.0
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally disseminated b

H-7 

Table H.3. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference betwe

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

1,2 Dichloroethane 

(C2H4Cl2) 

Carbon Monoxide 

(CO) 

Nitrogen (N2) 

Carbon Dioxide (CO2)

Carbon Tetrachloride 

(CCl3) Total Inert Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Using Experime

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated and 

Experimental 

Difference 

On a 

Percent 

Basis 

(%) 

Calculated 

Le 

Chatelier 

LFL 

C

E

5.2 4.95 4.6 61.6 23.65 85.25 42.5 37.62 4.88 11.48 38.49 

10.48 21.09 68.42 68.42 15.93 14.62 1.31 8.21 15.64 

10.9 10.9 78.2 78.20 21.5 20.39 1.11 5.18 20.74 

6.3 6 5.25 72.45 10 82.45 34 31.30 2.70 7.94 32.01 9.3 10.3 9.5 70.9 70.90 20.5 19.44 1.06 5.19 19.92 

15.11 15.11 69.77 69.77 16.41 14.70 1.71 10.39 15.57 10.38 29.02 60.61 60.61 13.2 11.91 1.29 9.79 12.79 

3.60 20.62 75.78 75.78 20.85 19.90 0.95 4.53 21.55 

16.48 4.40 79.12 79.12 22.75 20.00 2.75 12.09 20.86 

5.05 18.66 76.29 76.29 21.17 20.02 1.15 5.44 21.58 6.87 26.47 66.67 66.67 15 14.27 0.73 4.90 15.38 5.57 29.71 64.72 64.72 14.37 13.63 0.74 5.12 14.75 7.28 23.65 69.07 69.07 16.07 15.27 0.80 4.99 16.43 

4.1 1.35 7.6 79.2 7.75 86.95 58 52.55 5.45 9.40 53.88 

18.30 3.57 78.14 78.14 21.59 18.91 2.68 12.40 19.68 

4.3 0.2 23.7 55.9 15.9 71.80 36 33.21 2.79 7.75 34.54 

14.5 5.7 79.8 79.80 20.95 20.83 0.12 0.56 21.10 

6.3 6.7 6.1 80.9 80.90 31.5 29.39 2.11 6.71 30.10 5.4 0.4 8.5 73.75 11.95 85.70 52.5 47.39 5.11 9.73 48.40 

52.6 18 0.3 29.1 29.40 6.2 6.05 0.15 2.37 6.13 52.3 19 28.7 28.70 6 5.95 0.05 0.75 6.03 

69 15.8 0.4 14.8 15.20 5.4 5.24 0.16 2.87 5.29 

 33 33 33 33.00 2.35 2.23 0.12 5.16 2.25 

12.75 12.75 74.50 74.50 18.55 17.43 1.12 6.04 18.46 

20.8 4.9 74.3 74.30 14.9 14.63 0.27 1.78 14.75 

 33 33 33 33.00 2.45 2.23 0.22 9.03 2.25 

9 9.7 11.3 52 18 70.00 21.5 19.63 1.87 8.69 20.14 

14.74 7.62 77.64 77.64 20.35 19.20 1.15 5.66 20.15 

 25 25 25 25 25.00 2.8 2.55 0.25 8.77 2.58 

27.5 8.4 64.1 64.10 12.4 12.05 0.35 2.84 12.18 

25.3 25.2 49.5 49.50 9.2 8.80 0.40 4.36 8.95 

25 25 25 25 25.00 2.9 2.49 0.41 14.31 2.59 

73.1 12.1 14.8 14.80 5.7 5.36 0.34 5.95 5.40 40.6 7.6 27.6 24.2 51.80 10.05 9.39 0.66 6.60 9.46 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not been formally disseminated b

H-8 

Table H.3. Collected measured LFLs, the mixture composition, the calculated LFL using Zabetakis referen

researcher’s pure component LFLs, the source of each measured value, and the difference betwe

Hydrogen (H2) 

Methane (CH4) 

Ethane (C2H6) 

Methyl ethyl ketone 

(C4H8O) 

Toluene (C7H8) 

1,2 Dichloroethane 

(C2H4Cl2) 

Carbon Monoxide 

(CO) 

Nitrogen (N2) 

Carbon Dioxide (CO2)

Carbon Tetrachloride 

(CCl3) Total Inert Experimental LFL 

Using Zabetakis Pure 

Component LFL Values

Using Experime

Calculated 

Le 

Chatelier 

LFL 

Difference 

Between 

Calculated and 

Experimental 

Difference 

On a 

Percent 

Basis 

(%) 

Calculated 

Le 

Chatelier 

LFL 

C

E

16.6 16.7 66.7 66.70 14 13.35 0.65 4.63 13.58 

26.3 23.6 50.1 50.10 8.2 7.62 0.58 7.10 7.77 

 50 50 50.00 2.9 2.40 0.50 17.24 2.40  33 33 33 33.00 3.45 2.92 0.53 15.50 2.92 

20 20 20 20 20 20.00 3.4 2.75 0.65 19.01 2.85 50 50 50.00 10.8 8.00 2.80 25.93 10.00 

7.25 7.8 9.95 60 15 75.00 26 23.99 2.01 7.73 24.63 

13.52 2.78 83.70 83.70 27 25.41 1.59 5.88 26.45 

 33 33 33 33.00 3.5 2.92 0.58 16.71 2.92  33 33 33 33.00 4.85 4.14 0.71 14.70 4.21 

25.2 23.6 51.2 51.20 7.8 7.06 0.74 9.50 7.19 7.55 3.55 2.4 81.6 4.9 86.50 39.5 35.85 3.65 9.24 36.36 

 50 50 50.00 4.65 3.80 0.85 18.28 3.90 

33 33 33 33.00 3.65 2.80 0.85 23.36 2.93 25 25 25 25 25.00 3.95 3.10 0.85 21.47 3.23 25 25 25 25 25.00 5.15 4.07 1.08 20.95 4.35 

0.95 36.7 3.7 58.2 0.45 58.65 13 12.70 1.30 9.28 13.20 

33 33 33 33.00 5.2 3.90 1.30 24.93 4.25 33 33 33 33.00 5.2 3.90 1.30 24.93 4.25 

3.65 3.55 9.5 67.25 16.05 83.30 48.5 41.97 6.53 13.46 43.30 3.55 1.95 13.8 78.5 2.2 80.70 46 41.99 4.01 8.72 43.48 

2.1 13.75 4 73.4 6.75 80.15 30.5 27.82 2.68 8.80 28.82 

3 0.1 30.65 57.95 8.3 66.25 34 31.04 2.96 8.72 32.53 

6.25 2.4 12.05 73 6.3 79.30 36.5 33.26 3.24 8.87 34.13 

33 33 33 33.00 9.7 6.64 3.06 31.52 7.46 

2.3 9.4 1.5 86.8 86.80 36.5 33.84 2.66 7.29 34.03 

33 33 33 33.00 10.1 6.64 3.46 34.23 7.46 

4.7 9.8 85.5 85.50 25.8 22.51 3.29 12.74 23.06 2.3 8.3 2.1 87.3 87.30 37.7 34.07 3.63 9.62 34.34 2.4 8.95 6.65 69.95 12.05 82.00 40 34.22 5.78 14.44 35.45 2.1 9 1.1 87.8 87.80 42.2 37.15 5.05 11.96 37.32 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not 

been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

I-1 

APPENDIX I 

Appendix I provides methodology for calculating 

unobstructed cross sectional area of several flare tip designs.

 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not 

been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

I-2 

 
This information is distributed solely for the purpose of pre-dissemination peer review under applicable information quality guidelines. It has not 

been formally disseminated by EPA. It does not represent and should not be construed to represent any Agency determination or policy.

I-3