NiTX INST. OF STAND & TECH R.I.C. Evaluation of

Alternative In-Flight

Fire Suppressants

FOR

Full-scale Testing

IN Simulated

Aircraft Engine

Nacelles

AND Dry Bays

wluuiam l-. grossh andl.er,

Richard G. Gann ano

William M. Pitts, Editors

HIST

PUBLfCATfONS

Nisr

United States Department

of Commerce

Technology Administration

oe

100

.U57

NO. 861

National Institute of

Standards and Technology

April !994

NtST SP St ?

VALUATION OF

ALTERNATIVE IN-FLIGHT

Fire Suppressants

FOR

Full-scale Testing

IN Simulated

Aircraft Engine

Nacelles

AND Dry Bavs

WiLLIAM l_. GROSSH ANDLER,

Richard G. Gann and

William M . Pitts, Editors

Nisr

Certain commercial products

are identified in this report in order to specify adequately the

equipment used. Such identifi- cation does not imply recommendation by the National

Institute of Standards and

Technology, nor does it imply

that this equipment is the best

available for the purpose.

Building and Fire Research Laboratory

National Institute of Standards

and Technology

Gaithersburg, MD 20899

U.S. Department of Commerce

Ronald H. Brown, Secretary

Technology Administration

Mary L. Good,

Under Secretary for Technology

National Institute of Standards

and Technology

Arati Prabhakar, Director

April 1994

NIST SP 861
National Institute of Standards

and Technology

Special Publication 861

Natl. Inst. Stand. Techno!.

Spec. Publ. 861 852 pages (May 1994)

CODEN: NSPUE2

U.S. Government Printing Office

Washington: 1994

For sale by the Superintendent

of Documents

U.S. Government Printing Office

Washington, DC 20402
EXECUTIVE SUMMARY

William L. Grosshandler, Richard G. Gann and William M. Pitts Building and Fire Research Laboratory

Bromotrifluoromethane (halon 1301 or CF3Br) has been the fire-fighting agent of choice for decades to protect inaccessible spaces aboard aircraft in flight because of its inherent ability to inhibit combustion while possessing a high liquid density and stability, and low boiling point, electrical conductivity, corrosiveness, toxicity and cost. The bromine atom, which is credited with giving halon

1301 its strong inhibition character, reacts readily in a catalytic cycle to destroy ozone molecules. By

international agreement, commercial production of CF3Br has ceased in order to protect the

stratospheric ozone layer. No alternative agents are available which have all the positive

characteristics of halon 1301. Hence, there is an urgency to finding a replacement chemical in order

to maintain aircraft safety. The aff"ected government agencies and industries have a large stake in minimizing the impact of making a transition because of the potentially huge costs involved and

military readiness implications.

The US Air Force, Navy, Army and the Federal Aviation Administration pooled their resources

to support a comprehensive 17 month study at NIST to help them select the best chemicals to include

in a full-scale fire suppression test program at Wright-Patterson Air Force Base. The sponsors were

concerned with two specific applications: protection of commercial and military aircraft engine

nacelles, and protection of military aircraft dry bay areas. The engine nacelle is the portion of the

airframe which surrounds the main jet engines. In the event of a component failure, fuel and other

flammable materials can accumulate, ignite from a hot surface, and threaten the integrity of the

nacelle. The fire extinguishers are activated remotely following an alarm signal to the crew. Dry

bays are located in the wings and fuselage of the aircraft and vary considerably in geometry. They

contain a variety of electronic and mechanical components, and are either adjacent to fuel tanks or have fuel, hydraulic or lubricating fluid lines passing through them. The threat is from an antiaircraft projectile which ruptures a fuel tank or flammable fluid line, necessitating rapid and automatic

detection, agent release and suppression to avoid possible loss of the aircraft. The specific objectives of the research program are listed below. New metrics have been

developed along the way which can be used to rate other fire fighting compounds for a variety of

applications.

1. Determine the best two candidates for each application from the following list of gaseous agents

provided by the sponsors:

C2F6 (FC-1 16), C3F8 (FC-218), C4F10 (FC-31-10), and cyclo-C4F8 (FC-318);

CH2F2 (HFC-32), C2HF5 (HFC-125), CH2F2 (60%)/C2HF5 (40%) (HFC-32/HFC-125), C2H2F4

(HFC- 134a), C3HF7 (HFC-227ea), and C3H2F6 (HFC-236fa); and

CHF2CI (HCFC-22), and C2HF4CI (HCFC-124).
iv Executive Summary

2. Identify tiie best practical alternative from among chemicals not on the list of twelve.

3. Ascertain whether or not sodium bicarbonate powder is a viable option.

The resources of three NIST organizations [the Building and Fire Research Laboratory (BFRL),

the Chemical Science and Technology Laboratory (CSTL) and the Materials Science and Engineering

Laboratory (MSEL)] were enlisted to conduct the research, involving close to forty NIST staff members. Subcontracts were made to Lawrence Livermore Laboratory, the University of California

at San Diego and Yale University in support of the coordinated effort. Tasks were organized around

the following topics: thermodynamic properties of alternative agents; fluid dynamics of agent

discharge; flame suppression effectiveness; flame inhibition chemistry and the search for additional

fire fighting chemicals; agent stability under storage and discharge residue; metals corrosion;

elastomer seal compatibility; and human exposure and environmental impact.

New experimental methods were developed to screen the multitude of chemicals in a manner

which would allow extrapolation to the full-scale test. The agents were rated according to their

following innate physical properties which are independent of the release mechanism or performance

in the fire, with the desired magnitude given in parentheses: saturation liquid density {high}, normal

boiling point {low}, thermal expansion {low}, thermal stability {high}, solubility of nitrogen in agent

{low}, atmospheric lifetime {short}, agent residue {low}, corrosion of metals {low}, swelling of

polymers {low} and durability of elastomers and greases {high}. The performance criteria for evaluating fire suppression effectiveness were the concentrations {low} necessary to extinguish a laminar opposed-flow diffusion flame (OFDF), a small pool fire (cup burner), a turbulent spray

flame, a high speed premixed deflagration, and the amount of acid gases (HE and HCl) formed during

the suppression process. Discharge performance was evaluated from the time required {short} for the

agent to be discharged, dispersed, mixed and evaporated upon sudden release at high pressure. Agent

toxicity and potential exposure to humans was also rated on a relative basis. The likely availability of an alternative (and indirectly, its cost) were important considerations as well because the Air Force

testing schedule dictated that enough be available for them to perform a full-scale test matrix in 1994,

and to begin installing the agent in the fleet beginning in 1996.

The research uncovered deficiencies in a number of compounds which were of enough concern to recommend against testing them in the full-scale test program. EC- 116 required the largest storage

volume of the twelve agents on the original list. This is because its critical temperature is only

20 °C, leading to uncertainty in the design of the storage vessel and the rate at which material could

be discharged. The mixture of HFC-32 and HFC-125 required the second highest storage volume;

and in addition, high over-pressures in the detonation/deflagration tube were experienced. HFC-236fa

could not be recommended for the present aircraft applications because it has the highest boiling point

and is the most difficult to vaporize at the low temperatures which are likely to be encountered.

FC-318 and FC-31-10 also suffer from high boiling points and poor evaporation. In addition they are compatible with fewer materials and have very long atmospheric lifetimes. The HCFC-22 was

compatible with the fewest materials, and has the highest ODP of the core agents evaluated.

The positive attributes of an agent needed to suppress one kind of fire are not necessarily the same needed to suppress a different kind of fire. As a result, a different set of recommendations is made for the simulated full-scale engine nacelle and dry bay testing. HCFC-124 required the least amount of material to extinguish the nonpremixed flames and HFC-125 vaporizes well at low

temperatures; both are recommended for testing in the engine nacelle. EC-218 is recommended as the

first alternate. The most efficient agent of all is the sodium bicarbonate powder. A partial full-scale

test matrix is recommended to determine if possible operational difficulties can be overcome. One

additional gaseous candidate was identified that was more effective than halon 1301 in suppressing the
Executive Summary V

nonpremixed flames, iodotrifluoromethane (CF3I). It is recommended that it also be included in the

engine nacelle test program, but limited availability may prevent the test matrix from being

completed. Atmospheric chemistry, stability, compatibility and toxicity testing is still underway, so

that the recommendation for full-scale testing of CF3I comes with a caution.

The primary candidate for simulated full-scale dry bay testing is FC-218. It vaporizes relatively

easily, which is critical for quickly dispersing the agent through the dry bay in cold weather, and is relatively effective in attenuating shock wave pressure and speed in the premixed closed tube

experiments. HFC- 125 should also perform well in cold weather, and is the second recommended

candidate for dry bay testing. However, high over-pressures were noticed in the premixed closed

tube experiments. Should these occur during full-scale tests, HFC-227ea is recommended as a backup. Sodium bicarbonate is not recommended for this application because, as a powder, rapid

dispersion is hindered by clutter within the dry bay. The CF3I suppressed the premixed flame at lower concentrations than the other agents, but was not as efl'ective as halon 1301. A partial test matrix is recommended for CF3I, with the previously mentioned precautions holding.

Details of all aspects of the research program are provided in diff'erent sections of this special

publication. The rationale for arriving at critical conclusions and the scientific basis supporting the

above recommendations are discussed. A significant amount of resources went into selecting the best

candidates for full-scale evaluation, and work is continuing at NIST to determine long term

compatibility of the recommended agents with metals and polymers. A method for predicting the acid

gases produced when alternative agents pass through the fire is under development, parameters which

can increase the effectiveness of the agents by optimizing their release are being investigated, and

techniques are being developed for measuring the concentration of the agent once it is discharged.
I

L

l|

t

I
vii TABLE OF CONTENTS

Page

EXECUTIVE SUMMARY iii ACKNOWLEDGMENTS xv

1. INTRODUCTION 1 1.1 Background 1 1.1.1 Halons and the Ozone Layer 1 1.1.2 Aircraft Engine Nacelles and Dry Bays 1 1.1.3 Full-scale Simulations 2

1.1.4 Agents Selected for Screening 5

1.2 Objectives and Task Summary 8

1.2.1 Fire Suppression Effectiveness Screening 9

1.2.2 Compatibility with Systems and People 11 1.3 References 12

2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 13

2.1 Introduction 13

2.2 Experimental Apparatus and Procedures 15

2.2.1 Pure Agents 17

2.2.2 N2/Agent and CHF3/Agent Mixtures 19

2.2.3 Solubility of and CHF3 in Agents 20

2.3 Results and Discussion 20

2.3.1 Pure Agents 20

2.3.2 N2/Agent and CHF3/Agent Mixtures 23

2.3.3 Solubility of N2 and CHF3 in Agents 24

2.3.3.1 Experimental Results 24

2.3.3.2 Theoretical Predictions 24

2.4 Summary and Recommendations 32

2.5 References 34

3. FLUID DYNAMICS OF AGENT DISCHARGE 37

3.1 Introduction 37

3.2 Model of Discharging Vessel 38

3.2.1 The Problem and the Objective 38

3.2.2 The Experiment and the Model Assumptions 40

3.2.2.1 The Experimental Arrangement 40

3.2.2.2 The Procedure Prior to An Experimental Run or a FieldDeployed Discharge 40
viii Table of Contents

Page

3.2.2.3 During an Experimental Run or Field-Deployed Discharge .... 41

3.2.2.4 Further Assumptions 41

3.2.3 The Model Equations 41

3.2.4 Definition of P^p and Comments on Equation (8) 43

3.2.5 Reduced Dimensionless Equation Set - The Initial Value Problem 47

3.2.6 N2 Jet-Driven Mixing in the Discharge Vessel - Estimating the

Resulting Disturbance of the Liquid/Gas Interface 47

3.2.7 Solving the Model Equations 49

3.2.8 Example Calculations 49

3.2.8. 1 An Experimental Procedure to Simulate the Discharge of FieldDeployed Vessels 49

3.2.8.2 Criteria for Experimental Discharges to Closely Simulate

Field-Deployed Discharges 50

3.2.8.3 Model Input Parameters 50

3.2.9 Relatively Low-Pressure Large-Volume Holding Tank 52

3.2.9.1 Initiating the Pressurization/Discharge Sequence from the

Agent's Saturation State; No N2 in the Discharge

Vessel at t - 0 53

3.2.9.2 Initiating the Pressurization/Discharge Sequence From a P[)vi

Slightly Below Pguj^^ 61

3.2.9.3 Comparing the Discharges of a Small-Orifice Test Configuration and a Field-Deployed System 62

3.2.9.4 Summary of Results of Simulations Involving the Relatively

Low-Pressure Large-Volume Holding Tank 62

3.2. 10 Relatively High-Pressure Small-Volume Holding Tank 62

3.2.10.1 Pressurization and Discharge for the Small Orifice,

Dq - 0.0005 m 67

3.2.10.2 Pressurization and Discharge for the Large Orifice,

Dq = 0.005 m 73

3.2.11 Summary and Conclusions 73

3.2.12 Nomenclature 78

3.2.13 Appendix: The Dimensionless Initial Value Problem 79

3.3 Design of Experimental Release Vessel 82

3.3.1 Introduction 82

3.3.2 Release Vessel Design 82

3.3.3 Design Validation 83

3.4 Vessel Discharge Rate Measurements 86

3.4.1 Introduction 86

3.4.2 Experimental Method 87

3.4.3 Results and Discussion 89

3.4.3.1 Standard Discharge 89

3.4.3.2 Effect of an Orifice Plate 124

3.4.3.3 Eff"ect of an Extension Tube 124

3.4.3.4 Effect of Initial Vessel Pressure 135

3.4.3.5 Effect of Vessel Temperature 135

1
Table of Contents ix

Page

3.4.3.6 Vertically Upward Discharge 145

3.4.3.7 Horizontal Discharge 146

3.4.4 Conclusions 156

3.4.5 Recommendations 156

3.5 Measurements of Agent Dispersion 160

3.5.1 Introduction 160

3.5.2 Experimental Methods 163

3.5.2.1 High-Speed Filming 165

3.5.2.2 Laser Attenuation Measurements 165

3.5.2.3 Dynamic Pressure Measurements 167

3.5.2.4 Aspirated Hot-Film Probe 171

3.5.2.5 Data Acquisition System 172

3.5.2.6 Test Conditions 172

3.5.3 Experimental Results 172

3.5.3.1 High-Speed Filming 172

3.5.3.2 Near-Field Dynamic Pressure Measurements 193

3.5.3.3 Laser Attenuation Measurements 205

3.5.3.4 Far-Field Dynamic Pressure Measurements 232

3.5.3.5 Aspirated Hot-Film Measurements 244

3.5.3.6 Effects of Orifice Diameter 246

3.5.3.7 Effect of Vessel Pressurization Level 253

3.5.3.8 Effect of Attaching a 0.5 m Extension Tube to the Vessel 263

3.5.3.9 Effect of Cooling the Agent Prior to Release 263

3.5.3.10 Effect of Orientation-Upward Releases 272

3.5.4 Discussion 291

3.5.4.1 Downward Release of Pressurized Superheated Liquids 291

3.5.4.2 Effects of Orifice Diameter on Mixing and Evaporation 302

3.5.4.3 Effects of Vessel Pressurization Level on Mixing and

Evaporation 302

3.5.4.4 Effects of an Extension Tube on Mixing and Evaporation 303

3.5.4.5 Effects of Agent Cooling on Mixing and Evaporation 303

3.5.4.6 Effects of Release Orientation on Mixing and Evaporation 303

3.5.5 Summary and Agent Ranking 305

3.6 Transient Spray Computations 307

3.6.1 Introduction 307

3.6.2 Numerical Simulation of Discharge Sprays 309

3.6.2.1 Models Considered and General Capabilities 309

3.6.2.2 Calculational Approach and Code Modification 311

3.6.2.3 Sample Calculations and Results 312

3.6.3 Establishing Initial/Boundary Conditions from the Vessel Discharge

Simulations 317

3.6.3.1 Overview 317

3.6.3.2 Modeling the Early Development of the Jet: From the Exit

Section of the Discharge Vessel to a Nearby Section in a State of Thermodynamic Equilibrium and Completed

Droplet Formation 317
X Table of Contents

Page

3.6.3.3 Algorithm for Determining the Initial State 324

3.6.3.4 Example Applications of the Initial State Algorithm 326

3.6.3.5 Nomenclature 329

3.6.4 Concluding Remarks on Modeling the Discharge Spray 329

3.7 Summary and Recommendations 332

3.8 References 339

4. FLAME SUPPRESSION EFFECTIVENESS 345

4.1 Overview 345

4.1.1 Comparison of Experimental Configurations 346

4.1.2 Fuels and Agents Tested 348

4.1.3 Agents 348

4.2 Opposed Flow Diffusion Flames 349

4.2.1 Introduction 349

4.2.2 Experimental Apparatus 350

4.2.3 Experimental Procedure 356

4.2.4 Experimental Results 356

4.2.4.1 Impact of Fuel Type 360

4.2.4.2 Impact of Air Temperature 360

4.2.4.3 Powder Agent: Impact of Size 360

4.2.5 Interpretation of the Experimental Results 369

4.2.6 Conclusions from OFDF Study 375

4.3 Coflowing Nonpremixed Flames 377

4.3.1 Apparatus 377

4.3.2 Experimental Method 377

4.3.2.1 Gaseous Agents 377

4.3.2.2 Liquid Agents 382

4.3.2.3 Sodium Bicarbonate Powder 382

4.3.3 Assessment of Operating Parameters 385

4.3.4 Experimental Results 391

4.3.4.1 Core Agents 391

4.3.4.2 Sodium Bicarbonate Powder 395

4.3.4.3 Additional Agents 396

4.3.5 Discussion of Results 397

4.3.6 Summary and Recommendations 397

4.4 Turbulent Spray Flames 401

4.4.1 Background 401

4.4.2 Experimental Facility 401

4.4.3 Experimental Parameters and Procedures 404

4.4.3.1 Gaseous Agents 404

4.4.3.2 Powder Agents 412

4.4.4 Experimental Results 412

4.4.4.1 Characterization of Facility 412

4.4.4.2 Gaseous Agents/JP-8/Ambient Air Test Series 416

4.4.4.3 Effect of Air Temperature and Fuel Type 416

4.4.4.4 Sodium Bicarbonate Powder Experiments 418
Table of Contents xi

Page

4.4.5 Discussion of Results 418

4.4.6 Conclusions from Turbulent Spray Burner Study 419

4.5 High Speed Premixed Flames and Quasi-detonations 424

4.5.1 Introduction 424

4.5.2 Experimental Facility 424

4.5.3 Operating Procedure 425

4.5.5 Results and Analysis 433

4.5.5.1 Lean Mixtures at 100 kPa 436

4.5.5.2 Stoichiometric Mixtures 439

4.5.5.3 Ranking the Agents 452

4.5.6 Conclusions and Recommendations from the Premixed

Deflagration/Detonation Study 452

4.5.6.1 Conclusions 460

4.5.6.2 Recommendations 461

4.6 Summary of Conclusions on Flame Suppression Effectiveness 461

4.7 References 462

5. FLAME INfflBITION CHEMISTRY AND THE SEARCH

FOR ADDITIONAL FIRE FIGHTING CHEMICALS 467

5.1 Kinetics of Fluorine-Inhibited Hydrocarbon Flames 467

5.1.1 Project Objectives 467

5.1.2 Background 468

5.1.3 Reaction Engineering Approach 470

5.1.3.1 Species Thermochemistry 470

5.1.3.2 Reaction Kinetics 477

5.1.3.3 Reactor Models 497

5.1.4 Fluorinated Hydrocarbon Chemistry and Flame Suppression 499

5.1.4.1 Reaction Path Analysis 499

5.1.4.2 Freely Propagating, Premixed Flame Simulations 502

5.1.5 Discussion 505

5.1.6 Recommendations 506

5.2 Acid Gas Formation in Halogenated Hydrocarbon-Inhibited Flames 507

5.2.1 Objective and Background 507

5.2.2 Experimental Approach 508

5.2.3 Model for Acid Gas Formation 510

5.2.4 Results and Discussion 510

5.2.5 Future Work 513

5.3 Alternative Low ODP Chemicals with High Fire Suppression Effectiveness 530

5.3.1 Background 530

5.3.2 Methods 531

5.3.2.2 Exploratory List 539

5.3.3 Results 539

5.3.3.1 Development of the Short List 544

5.3.3.2 Recommendations 545

5.3.4 Conclusions 546

5.4 References 548
xii Table of Contents

Page

Appendix A. Tabulation of Ab Initio Thermochemical Data for Fluorinated

Hydrocarbons 561

Appendix B. Fluorinated Hydrocarbon Reaction Rate Constants 621

6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 643

6.1 Agent Stability Under Storage 643

6.1.1 Selection of Experimental Technique 643

6.1.2 Experimental 645

6.1.2.1 Test Preparation 646

6.1.2.2 Filling Procedure 646

6.1.2.3 Conduct of the Tests 648

6.1.2.4 Exposed metals 649

6.1.3 Results 649

6.1.3.1 FTIR Spectra 649

6.1.3.2 Examination of Metals 660

6.2 Agent Residue Study 660

6.2.1 Experimental Method 660

6.1.2 Results of residue analyses 666

6.3 Conclusions 666

6.4 References 667

7. CORROSION OF METALS 669

7.1 Introduction 669

7.2 Experimental 672

7.2.1 Aircraft Storage System Corrosion 672

7.2.1.1 Materials 672

7.2.1.2 Mass Change (Smooth Coupon) Exposure Experiments 672

7.2.1.3 Localized Corrosion (Weld/Crevice Coupon) Exposure

Experiments 676

7.2.1.4 Environmental Induced Fracture (Slow Strain Rate Tensile

Tests) 676

7.2.2 Post Deployment Experiments 677

7.2.2.1 Materials and Sample Preparation 677

7.2.2.2 Combustion Product Surface Layer Emulation 679

7.2.2.3 Exposure Environment 679

7.2.2.4 Testing Procedure 680

7.2.3 Development of an Electrochemical Measurement Technique 680

7.2.3.1 Materials 683

7.2.3.2 Procedure 683

7.3 Results 683

7.3.1 Aircraft Storage and Distribution System Corrosion 683

7.3.1.1 Mass Change (Smooth Sample) Exposure Experiments 683

7.3.1.2 Localized Corrosion (Crevice/Weld Coupon) Exposure

Experiments 687
Table of Contents xiii

Page

7.3.1.3 Environmental Induced Fracture Experiments (Slow Strain Rate Tensile Tests) 691

7.3. 1 .4 Summary of Aircraft Storage and Distribution System

Corrosion Experiments 696

7.3.2 Post-Deployment Corrosion 698

7.3.2.1 Literature Review of Corrosion of Metals in Halide Ions 698

7.3.2.2 Post-deployment Corrosion Test Results 700

7.3.3 Results of Electrochemical Measurements 706

7.4 Conclusions 718

7.5 References 718

8. ELASTOMER SEAL COMPATIBILITY 729

8.1 Introduction 729

8.2 Thermodynamic Compatibility 730

8.2.1 Theoretical Considerations 730

8.2.1.1 Flory-Huggins Theory 731

8.2.1.2 Flory's Equation-of-State Theory 732

8.2.2 Swelling Measurements 734

8.2.2.1 Experimental Setup 734

8.2.2.2 Experimental Materials 734

8.2.2.3 Experimental Methods 734

8.2.3 Results and Discussion 737

8.3 Durability 748

8.3.1 High Temperature Exposure Methods 748

8.3.2 Elastomers 750

8.3.2.1 Experimental Materials 750

8.3.2.2 Compression Set Measurements 750

8.3.2.3 Tensile Testing 751

8.3.3 Lubricants 756

8.3.3.1 Viscosity Measurements 756

8.4 Summary and Conclusion 757

8.5 References 761

9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 765

9.1 Potential Human Exposure 765

9.1.1 Full-Scale Tests 765

9.1.2 Instrumentation : 765

9.1.2.1 Temperature 766

9.1.2.2 Gas Concentration 766

9.1.3 Experimental Results 769

9.1.3.1 Temperature 769

9.1.3.2 Concentration 780

9.1.4 Modeling Results 794

9.1.5 Exposure Limits 797

9.2 Environmental Requirements for Candidate Replacements for Halon 1301 805

9.2.1 SNAP Protocol - 806
xiv Table of Contents

Page

9.2.2 Discussion of Key Results in Database 807

9.2.2.1 Physical/Chemicallnformation 807

9.2.2.2 Ozone Depletion Potential 808

9.2.2.3 Atmospheric Lifetime 808

9.2.2.4 Flammability 809

9.2.2.5 Replacement Ratio 809

9.2.2.6 SNAP Alternative 809

9.2.2.7 Required Technological Changes 809

9.2.2.8 Availability 809

9.2.2.9 Combustion Products 810

9.2.2.10 Price 810

9.2.2.11 Toxicity. 810

9.3 Conclusions . , 811

9.4 References 811

Appendix A. Halon Replacement Properties 813

10. SUMMARY AND RECOMMENDATIONS 829

10.1 Rationale for Selection 829

10.1.1 Dispersion and Evaporation 830

10.1.2 Storage Volume 831

10.1.3 Environmental Factors 832

10.1.4 Operational Issues 833

10.2 Ranking of Agents 835

10.2.1 Agents Recommended Against Full-Scale Engine Nacelle and

Dry Bay Testing 835

10.2.1 Ranking of Remaining Agents 835

10.2.1.1 For Full-scale Simulated Engine Nacelle Testing 835

10.2.1.2 For Full-scale Simulated Dry Bay Testing 837

10.3 Ongoing Studies 838

10.3.1 By-Product Formation 838

10.3.1.1 Instability Under Storage 838

10.3.1.2 Combustion By-ProductF 840

10.3.2 Metals Corrosion 840

10.3.3 Compatibility with Seals and Elastomers 841

10.4 Complementary Research Underway 841

10.4.1 Discharge Optimization 841

10.4.2 Alternative Agent Concentration Monitors 842

10.4.3 Engine Nacelle Simulations 842

10.5 Recommendations for Further Research 843
ACKNOWLEDGMENTS

XV

The contributors to this Special Publication were many. Specific authorship is indicated at the

beginning of each major section, and is acknowledged here in alphabetical order: Mr. Carlos

Beauchamp, Mr. Emil Braun. Mr. Brett Breuel, Dr. Donald Burgess, Mr. Thomas Cleary,

Mr. William Cleveland, Dr. Leonard Cooper, Mr. James Dante, Mr. James Fink, Dr. Glenn Forney,

Dr. Richard Gann, Prof. Grzegorz Gmurczyk, Dr. William Grosshandler, Dr. Anthony Hamins,

Mr. Richard H. Harris, Jr., Dr. Ferenc Horkay, Dr. Barbara Levin, Dr. Gregory Linteris, Dr. Gregory McKenna, Dr. Thomas Moffat, Dr. George Mulholland, Dr. Marc Nyden, Mr. Richard

Peacock, Dr. Cary Presser, Dr. William Pitts, Mr. R. Gregory Rehwoldt, Dr. Richard Ricker,

Prof. Kalyanasundaram Seshadri, Mr. Mark Stoudt, Dr. Wing Tsang, Ms. Isaura Vazquez,

Dr. William Waldron, Jr., Prof. Phillip Westmoreland, Dr. Jiann Yang, and Dr. Michael Zachariah.

A number of additional NIST statf played critical roles in support of the large number of

research activities. These include Mr. James Allen, Mr. Michael Glover, Dr. Robert Huie,

Mr. Darren Lowe, Mr. William Rinkinen and Dr. Kermit Smyth. The assistance of Ms. Paula

Garrett in organizing and editing the final report is gratefully acknowledged.

The research described in this publication was supported by the following agencies of the federal

government:

Wright Laboratory

Flight Dynamics Directorate

Vehicle Subsystems Division

Survivability Enhancement Branch

Mr. Michael Bennett

Naval Air Systems Command

Subsystems Branch

Fuel and Fluid Power Systems Section

Mr. James Homan

Army Aviation and Troop Command

Mr. Daniel McEneany

Department of Transportation

Federal Aviation Administration

Technical Center

Mr. Larry Curran

In addition to the program managers listed above, the authors wish to acknowledge the active

role played by the staff of each of the sponsors: Lt. Kurt Lee, Lt. Gretchen Brockfeld, Capt. Rene

Little and Maj. Samuel Carbaugh of the U.S. Air Force; Mr. David Thurston, of the U.S. Navy ; Mr. CP Sarkos and Ms. Louise C. Speitel of the Federal Aviation Administration. All helped

ensure that the research remained appropriately focused and timely.

Special mention is in order of the tireless effort maintained by Mr. Bennett of Wright

Laboratory, who coordinated this cross-services effort from concept to completion. His skills as a manager, engineer and visionary were all tapped, and each was equally essential to the success of the

project. The assistance provided by Lt. Gregg Caggianelli and Lt. Brett Poole on the details of the

operation of the Aircraft Engine Nacelle Test Facility and the Aircraft Survivability Research Facility are gratefully acknowledged.

To make a significant contribution to a problem as complex as aircraft fire suppression in so

short a period of time required the cooperation of the airframe makers, equipment manufacturers and

chemical suppliers. The information and assistance provided by the following companies was

appreciated and is hereby acknowledged: Pacific Scientific, Systron-Donner and Walter Kidde
xvi Acknowledgements

Aerospace; McDonnell Douglas, Lockheed, and Boeing; Allied Signal, E.I. DuPont, Great Lakes

Chemical, PGR and 3M.

Interactions with researchers in other laboratories were essential to obtain additional insight and

to avoid wasteful duplications of effort. Dr. Don Wuebbles of Lawrence Livermore Laboratories

provided key calculations on stratospheric ozone depletion; Dr. Andrezej Miziolek of the Army

Research Laboratory, Dr. Ronald Sheinson of the Naval Research Laboratory and Dr. Thomas

Jackson of Wright Laboratory were involved in numerous technical discussions; Dr. Garl Melius of

Sandia Laboratory and Prof. Mitch Smooke of Yale University provided critical data and numerical

calculations for the kinetics study. The contributions of all these scientists and engineers are acknowledged.
1

1. INTRODUCTION

William L. Grosshandler, Richard G. Gann, William M. Pitts Building and Fire Research Laboratory

1.1 Background

1.1.1 Halons and the Ozone Layer. The Montreal Protocol of 1987 identified halon 1301 (CFjBr)

as one of a number of halogenated chemicals that were sufficiently deleterious to stratospheric ozone

that their continued production and use required limitation. An amendment to the Protocol caused

commercial halon production to cease at the begirming of 1994 (Harrington, 1993). The focus of concern is that the halon molecule is sufficiently stable in the lower atmosphere that it will eventually

be transported to the stratosphere unaltered. There the carbon-bromine bond is vulnerable to photodissociation by solar ultraviolet radiation. The liberated bromine atom then enters into a

catalytic cycle which has the net effect of consuming the shielding ozone. Much has been written in the scientific and popular literature on how bromine and chlorine atoms

lead to depletion of the ozone layer. Conferences are held annually (e.g., The 1993 International

CFC and Halon Alternatives Conference, 1993; Halon Alternatives Technical Working Conference,

1993) to discuss progress in finding replacements for technologically important ozone depleting

chemicals. An exhaustive study of the atmospheric chemistry of fluorinated compounds being

considered as alternatives to halons and chlorofluorocarbons (CFCs) was commissioned by the World

Meteorological Organization (1989). Background on the halon/ozone issue is also provided by

Andersen (1987) and Grant (1989). Additional articles are cited in the report by Pitts et al. (1990).

Ford (1975) has traced the development of CFjBr for firefighting purposes since its original

formulation in the 1940s. Halon 1301 has been the agent of choice for fire suppression on board

aircraft for decades because of its inherent ability to inhibit flames at low concentrations while

exhibiting a number of additional strongly positive attributes, including a high level of stability, high

liquid density, low boiling point, low electrical conductivity, low corrosiveness, low toxicity, and low

cost. These properties are desirable for many applications, but the penalty placed upon excessive

weight or volume on board commercial and military aircraft tightly constrains the choices of suitable

alternatives. The elimination of new production of halon 1301 has forced the manufacturers, owners, and

users of aircraft to search for an alternative. Since it may take years to retrofit the large fleet of

current aircraft, there is an urgency to identify one or more chemicals soon. The retrofit process will

likely be expensive. The bromine atom in halon 1301 which is responsible for the ozone destruction

is also largely responsible for the molecule's high fire suppression effectiveness. If the alternative

chemicals do not contain a moiety like bromine and are less efficient suppressants, then new, larger

agent storage and delivery components will need to be designed. To avoid costly design mistakes,

extensive performance data regarding the amount of agent required for flame extinction and the

compatibility of the alternatives with the delivery system and aircraft construction materials is essential.

1.1.2 Aircraft Engine Nacelles and Dry Bays. There are a number of distinct areas in need of

protection against an in-flight fire in an airplane, including the cockpit, passenger compartment,

baggage areas, auxiliary power units, fuel tank ullages, engine nacelles, and dry bays. The research
2 SECTION 1. INTRODUCTION

program described in this publication focussed on the latter two. The requirements and constraints of

the engine nacelle and dry bay areas are quite different, and the methods used to arrive at the best

choice for an alternative to halon 1301 for each application dilTer accordingly.

The engine nacelle is the portion of the airframe which surrounds the main jet engines. It varies

in shape and size but is typically annular with a length and diameter of the same order as the engine it encases. Fuel and hydraulic lines, pumps and lubrication systems are located within the nacelle

volume. Ventilation is provided to prevent the build-up of combustible vapors, and drain holes on the

underside are used to reduce the amount of fluid that could pool following a leak. The engine's

combustion chamber liner is isolated from the nacelle wall, but surface temperatures up to 700 °C can

exist during operational extremes. The firefighting agent can be located directly adjacent to the

nacelle or removed by tens of meters and piped to the fire zone, depending upon the availability of

space. An example of a nacelle and fire suppression system is shown schematically in Figure 1. The operational requirements for commercial and military engine nacelle protection are similar. Thermal sensors are used to determine when an abnormal, overheated condition exists in the nacelle

region. In such a case, the flight of the airplane is leveled and the designated crew member arms the

appropriate halon 1301 bottle and fires the release mechanism. The halon is stored at pressures in the

range of 2 to 4 MPa, and the certification process requires that enough agent be available to maintain

a minimum concentration (about 6% by volume) throughout the nacelle for a minimum time interval (about 0.5 s) to ensure that the fire will be extinguished and not relight. Many systems have a

duplicate bottle to be used as a back-up should the first shot be unsuccessful.

Dry bays refer to normally closed spaces, often adjacent to flammable liquid storage areas and in which a combustible mixture and an ignition source could co-exist following penetration by an antiaircraft projectile. The fire threat is unique to military aircraft. Dry bays vary considerably in volume, typically being in the range of 0.2 to 3.0 m-^. They are located in the wings and fuselage,

and their shape is most often irregular, as can be seen in Figure 2. Aspect ratios up to 10:1 are not uncommon. The bays may or may not be ventilated, and are usually cluttered with electronic,

hydraulic and mechanical components. Compared to the events leading to engine nacelle fire suppression, the required timing is two orders-of-magnitude faster for dry bay protection. High speed

infrared detectors sense the initial penetration of the projectile and automatically arm and fire the

halon bottle. The storage bottles are located directly in or adjacent to the protected space to minimize

the time needed to flood the volume totally. The entire suppression sequence occurs in less than

100 ms and requires no crew intervention.

1.1.3 Full-scale Simulations. Fire suppression is afl'ected by a combination of the agent's physical

properties, chemistry and transport phenomena, making it a highly nonlinear process and difficult to

scale. Full-scale tests are needed to extend bench-scale experiments and validate theoretical models

before aircraft safety systems can be reliably designed. The Survivability Enhancement Branch of

Wright Laboratory, located at Wright-Patterson Air Force Base, Ohio, operates unique full-scale

aircraft engine nacelle and dry bay fire test facilities. How these facilities are to be used to identify

the best near-terra alternative to halon 1301 is outlined by Bennett (1992). The Design of Experiments (DOX) methodology which has been adopted allows the minimum number of full-scale tests to have the maximum impact on identifying the critical parameters which influence an engine nacelle or dry bay fire. The range of parameters chosen was determined by the Air Force after surveying

existing military aircraft manufacturers and users (Mascarella, 1993).

The Aircraft Survivability Research Facility (ASRF) at Wright-Patterson AFB consists of a gun

range in which an anti-aircraft round can be directed at a precise location in a full-size test article in a

repeatable, controlled manner while maintaining the safety of the operating persoimel. Rectangular

boxes with internal volumes between 0.28 and 2.8 m^ have been used to simulate actual aircraft dry
SECTION 1. INTRODUCTION 3

Fire Extinguishing Agent Botties

Figure 1. Schematic diagram of wing-mounted engine nacelle and fire extinguishing system

(based upon C5-A and provided by M.L. Kolleck, Booz-Allen & Hamilton, Inc.).
SECTION 1. INTRODUCTION

gure 2. Schematic diagram of typical dry bays in an aircraft (taken from Figure 1 of

Heinonen a/., 1990).
SECTION 1. INTRODUCTION 5

bays. A fuel tank is located behind the dry bay so that it will rupture following the impact of a fired projectile. Dry bay geometry (aspect ratio), internal obstructions, location and timing of the release

of the firefighting agent, the type and amount of fuel released, and the characteristics of the projectile are some of the parameters which can be varied. Jet engine exhaust streams can be directed past the

test article to simulate the high speed flows experienced in flight. A typical test article is shown

in Figure 3. The Aircraft Engine Nacelle (AEN) facility (Figure 4) was designed to simulate a 120° arc segment of a full size engine nacelle. The air flow, air temperature, and air pressure can be varied to cover the range experienced by an actual aircraft in flight. A fire from a fuel spray or ruptured

hydraulic fluid line can be simulated, and the amount of clutter, location and timing of agent

injection, and hot-surface temperature are some of the parameters which can be precisely controlled.

The AEN is used to determine the amount of agent necessary to extinguish a realistic engine nacelle

fire, and can compare the relative eff"ectiveness of two different agents in suppressing a nearly

identical fire. A full 360° simulated engine nacelle with adjustable diameter core is a new addition to the facility.

1.1.4 Agents Selected for Screening. In a study by Pitts et al. (1990), an exploratory list of over 100 possible halon alternatives was constructed. These included both high boiling point and low

boiling point compounds for replacements of halon 1211 and halon 1301, respectively. The chemicals

embodied a range of chemical and physical principles thought to aff'ect flame suppression capability.

Examination of these chemicals, then, provides a basis for the search for alternatives to the current commercial halons. This list included the following classes of compounds:

• saturated and unsaturated halocarbons,

• halogenated hydrocarbons containing single and double-bonded oxygen,

• sulfur halides,

• phosphorous compounds,

• silicon and germanium compounds,

• metallic compounds, and

• inert gases.

Zallen (1992) reviewed past investigations of firefighting agents, starting with the seminal work

by Purdue University (1950). In that early investigation, the concentrations necessary to inhibit a mixture of heptane and air from igniting were tabulated for over 50 chemicals. An objective was to identify general purpose agents which were effective, clean, stable and less toxic than CCI4. A large

number of halocarbons were scrutinized, along with a few compounds containing sulfur, phosphorous

and silicon. Zallen (1992) summarized the extensive work done by the U.S. Navy (e.g., Sheinson et

al., 1989) and the New Mexico Engineering Research Institute (e.g., Tapscott et al., 1990) for shipboard and ground-based aircraft fire suppression, respectively, before arriving at his recommendations

to the U.S. Air Force for their in-flight applications. An important consideration was the likely

availability of the chemical in large quantities for full-scale testing in 1994, and fleet installation by

1996.

Zallen recommended four perfluorocarbons (FCs), six hydrofluorocarbons (HFCs), and two

hydrochlorofluorocarbons (HCFCs). The core candidate compounds actually selected by the Air

Force for screening are shown in Table 1. Several compounds in the table were not on Zallen's list: a 60%/40% mixture by mass of HFC-32 and HFC- 125, sodium bicarbonate powder, and HFC-236fa.

The HFC-32/ 125 mixture and sodium bicarbonate were added by the Air Force; HFC-236fa was

added following a press release soliciting additional chemicals that met the criteria discussed above
SECTION 1. INTRODUCTION

I— INSERT

ADJUSTABLE VDLUME

FIRE EXTINGUISHER

EXTINGUISHER

DRYRAY TEST APTILLE

Figure 3. Example of test article used in live-fire dry bay simulator at WPAFB Aircraft

Survivability Research Facility (provided by B. Poole, WPAFB).
Figure 4. Aircraft Engine Nacelle test facility at WPAFB (provided by G. Caggianelli,

WPAFB).
8 SECTION 1. INTRODUCTION

Table 1 . Core chemicals examined

Chemical Formula Designation lUPAC Name

C2F6 FC-116 hexafluoroethane

C3F8 FC-218 octafluoropropane

C4F10 FC-31-I0 decafluorobutane

cyclo-C4F8 FC-318 octafluorocyclobutane

CH2F2 HFC-32 difluoromethane

C2HF5 HFC- 125 pentafluoroethane

CH2F2 (60%)/C2HF5 (40%) HFC-32/ 125 (60/40)

C2H2F4 HFC- 134a 1,1,1 ,2-tetrafluoroethane

C3HF7 HFC-227ea 1,1,1 ,2,3,3,3-heptafluoropropane

C3H2F6 HFC-236fa 1,1,1 ,3,3,3-hexafluoropropane

CHF2CI HCFC-22 chlorodifluoromethane

C2HF4CI HCFC-124 2-chloro- 1,1,1 ,2-tetrafluoroethane

NaHC03/Si02 (1%) sodium bicarbonate/silicon dioxide

(Kosko, 1993). On Zallen's list but not included among the core candidates were HFC-23 and

HFC- 143a. HFC-32 was dropped early in the NIST research program because testing confirmed that

it was flammable at concentrations likely to be encountered. A full test program was therefore

carried out for eleven gaseous potential alternatives and sodium bicarbonate.

1.2 Objectives and Task Summary

This research program began m 1992, funded jointly by the Air Force, Army, Navy and the Federal

Aviation Administration because of their mutual interest in aircraft safety. The program entailed:

• establishing a comprehensive experimental program to screen the performance of the agents

listed in Table 1 , plus other promising chemicals, as a means to identify the best candidates

for subsequent full-scale aircraft fire extinguishment evaluation at Wright Laboratory; and

• addressing the compatibility of these agents with flight systems, people, and the environment.
SECTION 1. INTRODUCTION 9

Based on these data, three candidates for each of the two applications (dry bay and engine nacelle) were to be recommended for full-scale evaluation at Wright-Patterson AFB. These did not need to be

the same sets. Ideally, each set would include the top two candidates from Table 1. If the search for additional chemicals uncovered other promising candidates, then the best of

these was to be recommended for full-scale testing as well. Otherwise, a third chemical from Table 1 would be identified. These recommendations were to be presented to the sponsors in September,

1993. Following the sponsors' selection of chemicals for full-scale testing, longer-term compatibility

experiments to increase confidence in the preliminary results were to continue, with a report on that phase of the research to be delivered by September 30, 1995.

Sections 2-9 of the current report document the research activities related to the fire fighting

eff'ectiveness of possible alternatives, and research activities dealing with various aspects of the

agents' compatibility with aircraft construction materials and associated systems. Establishing a

scientific basis upon which to choose the best alternatives has been an overriding objective of the

program. The facilities developed and knowledge gained in meeting this objective go well beyond the

specific applications to aircraft, and advance fire suppression technology in general. A summary of

the major conclusions are given and significant recommendations are highlighted in Section 10. In the course of searching for alternative agents and in developing these facilities, certain

products were used and are identified by trade name in this report to adequately document the

research performed. Such identification does not imply recommendation by the National

Institute of Standards and Technology, nor does it imply that this commercial equipment is the

best available for the piu-pose intended.

1.2.1 Fire Suppression Effectiveness Screening. Because of storage weight and volume constraints,

the new suppressant should demonstrate optimal efficiency for quenching flames. The thermodynamic

state, transport properties, and combustion chemistry of an agent are critical to this ability. An

experimental design was chosen that allowed these three aspects of the different agents to be examined

under conditions which isolated the influence of the many independent parameters. Considerable

thought was given to the most appropriate means to relate the results of the individual tests to overall performance criteria. The uncertainty in each experiment was estimated by replication when

appropriate. In other cases, analogy to similar experimental procedures, experience with related test methods, and theoretical calculations were used to establish error limits. The thermodynamic state of a pure substance is defined by its density over a range of pressures

and temperatures. In mixtures, the solubility of one material into another influences the state as well. Quantitative thermodynamic state relations are incomplete for many of the pure alternative agents, and

very little data on the solubility of pressurizing gases into the liquid agent existed at the start of this investigation. An objective of this task was to measure the pressure of the agent at -60 °C, 25 °C

and 150 °C in a constant volume vessel, with the agent pure and in mixtures with CHF3 and N2. The

temperatures chosen span the maximum anticipated range of exposure of a storage bottle in the field. The solubility of the pressurizing gas was measured at room temperature and 4 MPa, a typical halon

1301 bottle pressure. Section 2 presents the results of the thermodynamic study.

The release and delivery of the agent is particularly critical for dry bay protection, where events leading to a threatening over-pressure occur in less than 50 ms. The behavior of the agent during

conditions of vessel blow-down and the characterization of the dispersion, mixing and evaporation of

the resulting two-phase turbulent jet are the objects of investigation in Section 3. Four different flame suppression experimental facilities were designed with the objective of

examining the flame extinction properties of the agents over the whole range of conditions likely to be

encountered by aircraft in flight. The facilities, briefly mentioned in the following paragraphs, are described and the suppression concentrations are tabulated in Section 4.
10 SECTION 1. INTRODUCTION

The first suppression facility, an opposed-flow diff\ision flame (OFDF) burner, is an excellent

test-bench for evaluating an agent's extinguishing efficiency over a broad range of flame conditions.

The reasons for this are five-fold:

1 . the geometry of the flow is eliminated as variables in the system;

2. the flow field strain rate is an independent variable;

3. the system is steady, one-dimensional, and accessible to probing;

4. the reactant flow rates are relatively small, enabling extensive testing with limited

quantities of chemicals; and

5. the concentrations and temperatures of the agent, fuel, and oxidizer are easily controlled

and repeatable.

Cup burner measurements, which have been relied upon almost exclusively in the past to determine

suppression concentrations, were made in the second facility and compared to the OFDF results. The

cup burner is also flexible, and was used to assess the impact of multiple fuels on the ability of all the

agents, including the sodium bicarbonate powder, to suppress the flames.

In the third facility, a turbulent spray burner was built to simulate a fire created by a ruptured

fuel or hydraulic fluid line in an engine nacelle, or the splash of fuel into a dry bay. This is because

threatening fires are invariably turbulent, and the turbulence influences the mixing of the fuel and air, the heat released, and the products of combustion. The amount of agent (gaseous or powder) and the

rate of injection were independent parameters which could be controlled with this facility.

The speed of an accelerating turbulent flame near the detonation limit can reach magnitudes two

orders higher than that of the spray flame. The interaction of the agent with a detonation could not be predicted from the other flame results. Thus, a detonation tube (the fourth facility) was used to simulate the threat posed by an impending explosion. The reduction in velocity and pressure build-up

of the high speed, turbulent flame was used as a quantitative measure of the agent's effectiveness in suppressing the detonation.

Significant improvements to suppression performance are likely to occur only if the agent is able

to inhibit key chain-branching reactions in the combustion process. It would be beneficial to understand the mechanisms by which alternative agents extinguish a flame. If such chemical intervention were possible with materials that were tailored to degrade before threatening the atmosphere, then

additional environmental constraints might be overcome, and opportunities for enhancing the

performance of advanced aircraft through the reduction in bulk of the onboard fire suppression system

might exist. To assist in gaining the objective of superior chemical inhibition, the elementary

reactions of diff'erent classes of chemicals were examined, and a partial understanding of how the

molecular structure influences the suppression process and the atmospheric reactivity of the neat agent was obtained. Section 5 outlines the progress made to date on the kinetics of fluorine inhibited flames

and possible routes to acid gas formation. Alternative chemicals beyond the core list from Table 1 are discussed, with the rationale for choosing them and our current assessment of their potential. The

development of an understanding of the acid gas formation process is ongoing, and a method of

estimating the amount formed during suppression of a given fire threat will be provided in the

September, 1995 report.
SECTION 1. INTRODUCTION 11

1.2.2 Compatibility \^1th Systems and People. A desirable suppressant also needs to have minimal

harmful interactions with aircraft materials, people, and the environment. A set of tasks were

designed to provide essential data on the stability of the chemicals, compatibility of the chemicals and

their by-products with both metallic and polymeric materials, and definition of possible limitations for exposure of people and the environment. The approach followed from an earlier study (Garm et al.,

1990), in which preliminary screening procedures and criteria were developed for the general evaluat- ion of alternatives to both halons 1211 and 1301. The advanced procedures developed here are more

directly applicable to the conditions in engine nacelles and dry bays.

The compatibility studies covered a range of storage conditions. Based on information supplied

by the Air Force, it was presumed that the agent storage would be initially at 25 °C and 4.1 MPa. In

flight, these could rise to 150 °C and 5.8 MPa. Thus, the stability of a chemical and its potential

damage to the storage vessel were examined under these conditions. Lower temperatures and pressures are also possible in service. However, lesser damage would occur under these conditions, and we deemed reduced temperature measurements unnecessary.

At present, the storage times on board aircraft are of the order of five years. In order to enable

a selection of the most promising chemicals by the September 30, 1993 deadline, only much shorter

exposures were possible. Phenomenological modeling, based on exposures at various time intervals,

has been pursued to increase confidence in the extrapolation to longer times. Were severe non-linear

eflFects observed for some chemicals, NIST would recommend against their use or would suggest that shorter storage times be implemented. For those chemicals selected for further testing, longer

exposures (e.g., one year) will be performed to improve confidence in the extrapolations. The

procedures and data are presented in Section 6. The results of the long-term experiments will be

summarized in the September, 1995, report.

During storage under elevated pressure and perhaps temperature, it is desirable that the agent not harm the storage container. This could result in leakage of the chemical over time or even in-flight

rupmre of the bottle. With the assistance of the sponsors, a list of current and potential materials for storage bottles was compiled. This included eight metal alloys, six elastomer o-rings, and three

lubricants. Approximately one-month exposures at ca. 150 °C and 5.8 MPa provided data on the

potential for afi'ecting the storage materials. The metals corrosion research is discussed in Section 7. Section 8 details the study of the compatibility of polymeric materials which could be in contact with

the agent under the above range of conditions.

Following dispensing, the suppressant chemical spends time in and near the flames, leading to some decomposition and the production of potentially harmful by-products. It is known that the fluids

in Table 1 produce HF and, in two cases, HCl. To a first order, it was presumed that the yields of

HF from the various chemicals would pose the same threat to aircraft materials, thus providing little difl'erentiation among the chemicals. Testing of this will be reported in the September, 1995 report.

Pilot examination of the potential for NaHC03 and its by-products was performed, since there was

reason to expect a serious interaction with the aluminum used in aircraft bodies. This is presented in Section 7. There are two other key components affecting the selection of alternative suppressants. First, one must be able to predict the potential exposure of people during an accidental discharge. The

simplest assumption is that the distribution of the chemical in a confined space is uniform. Experiments were conducted to verify the validity of this hypothesis. Second, with the rapidly-changing

regulations for protecting the earth's environment, it is necessary to know the pertinent requirements

for testing of these families of chemicals. Both of these topics are discussed in Section 9.
12 SECTION 1. INTRODUCTION

1.3 References

Andersen, S.O., "Halons and the Stratospheric Ozone Issue," Fire Journal 81, 56-62 and 116-

125, May/June (1987).

Bennett, M., "Halon Replacement for Aviation Systems," The 1992 International CFC and

Halon Alternatives Conference Proceedings , pp. 667-670, The Alliance for Responsible CFC Policy,

Frederick, MD, September 29-October 1, 1992.

Ford, C.L., "An Overview of Halon 1301 Systems," in Halogenated Fire Suppressants , (Gann,

R.G., editor), pp. 1-63, American Chemical Society, Washington, DC, 1975.

Gann, R.G., Barnes, J.D., Davis, S., Harris, J.S., Harris, R.H., Herron, J.T., Levin, B.C.,

Mopsik, F.I., Notarianni, K.A., Nyden, M.R., Paabo, M., and Ricker, R.E., Preliminary Screening

Procedures and Criteria for Replacements for Halon 1211 and 1301, NIST Technical Note 1278,

August 1990.

Grant, C.C., "Halons and the Ozone Layer: An Overview," Fire Journal 83, 58-62 and 78-81,

September (1989).

Halon Alternatives Technical Working Conference Proceedings, New Mexico Engineering

Research Institute, Albuquerque, May 11-13, 1993.

Harrington, J.L., "The Halon Phaseout Speeds Up," NFPA Journal, pp. 38-42, March/April,

1993.

Heinonen, E.W., Moore, T.A., Nimitz, J.S., Skaggs, S.R., Beeson, H.D., and Moussa, N.A.,

Fire/Explosion Protection Characterization and Optimization: Phase II - Alternative Dry Bay Fire

Suppression Agent Screening, NMERI OC 90/20, Report JTCG/AS-90-T003.

Kosko, J.E., press release. Public Affairs, NIST, February 22, 1993.

Mascarella, S., Booz-AUen & Hamilton Inc., Survivability/Vulnerability Information Analysis

Center, personal communication, October 6, 1993.

Pitts, W.M., Nyden, M.R., Gann, R.G., Mallard, W.G., and Tsang, W., Construction of an

Exploratory List of Chemicals to Initiate the Search for Halon Alternatives, NIST Technical Note

1279, National Institute of Standards and Technology, Gaithersburg, MD, August 1990.

Purdue Research Foundation and Department of Chemistry, Final Report on Fire Extinguishing

Agents, AD 654322, Purdue University, West Lafayette, IN, July 1950.

Sheinson. R.S., Penner-Hahn, J.E., and Indritz, D., "The Physical and Chemical Action of Fire

Suppressants," Fire Safety Journal 15, 437-450 (1989).

Tapscott, R.E., Moore, J. P., Lee, M.E., Watson, J.D., and Morehouse, E.T., Next Generation

Fire Extinguishing Agent Phase III, ESL-TR-87-03, Air Force Engineering and Services Center,

Tyndall AFB, FL, April 1990.

The 1993 International CFC and Halon Alternatives Conference Proceedings, The Alliance for Responsible CFC Policy, Frederick, MD, October 20-22, 1993.

World Meteorological Organization, Scientific Assessment of Stratospheric Ozone: 1989, Report

No. 20, Alternative Fluorocarbon Environmental Acceptability Study (AFEAS), 1989.

Zallen, D.M., Halon Replacement Study, SBIR Report No. ZIA-92-001 to Aeronautical Systems

Division, Wright-Patterson AFB, Zallen International Associates, Albuquerque, February 28, 1992.
13

2. THERMODYNAMIC PROPERTIES OF

ALTERNATIVE AGENTS

Jiann C. Yang and Brett D. Breuel

Building and Fire Research Laboratory

2.1 Introduction

Depending upon their applications, current halon 1301 (CF3Br) bottles are normally filled to about

half of the bottle volume, and the bottle is then pressurized with nitrogen to 4.1 MPa (600 psig) at room temperature. The purpose of using the pressurization gas is to expedite the discharge of the

agent and to increase the penetration distance of the agent during discharge. However, this driving

force, i.e., the total pressure in the bottle, will vary depending on the ambient temperature because

the vapor pressure of the agent and the solubility of the pressurization gas in the liquid agent vary

with temperature. The current military specification (MIL-C-22284A) for halon 1301 containers

stipulates that the container should have a proof pressure of 9.62 MPa (1400 psig), a burst pressure of

12.37 MPa (1800 psig), and a frangible disc that will rupture at a pressure between 8.59 MPa (1250

psig) and 9.62 MPa (1400 psig) at 70 °C. In order to explore the possibility of using existing halon

1301 bottles for "drop-in" replacement agents or to provide safety guidelines on bottle design for the

alternative agents, two important tasks are to determine (1) the solubility of the pressurization gas in the liquid agent and (2) the final pressure of the vessel when exposed to different ambient temperatures.In the first task, some thermodynamic properties of the core agents listed in Section 1 were

measured. Halon 1301 was also included in this study for the purposes of reference and comparison

with the alternative agents. In particular, the objective was to determine the final equilibrium

pressure in the bottle when the bottle, filled with either a pre-determined amount of pure agent or with agent and pressurization gas, was exposed to diff'erent temperatures. Experiments on pure agents were performed in order to obtain the vapor pressure data when documented data were not available. Two pressurization gases, nitrogen and CHF3 (HFC-23), were used in the study. Trifluoromethane

(CHF3) was explored as a pressurization gas because it itself could be used as a fire suppressing

agent. Another property that makes CHF3 a potential candidate as a pressurization gas is its vapor

pressure of 4.16 MPa (606 psia) at 20 °C, which is high enough to pressurize the vessel to the

appropriate level. The second task of this study was to determine the solubility of the pressurization gas in the

agent at room temperature. Specifically, the objective was to measure the amount of the gas initially

used to pressurize the bottle to 4. 1 MPa. This part of the study, in essence, determined the initial condition of the vessel. For the purpose of reference, selected thermo-physical properties of pure agents are tabulated in Table 1 where MW is the molecular weight, is the normal boiling point (at 0.101 MPa), is the

critical temperature, is the critical pressure, p^^^ is the saturation vapor pressure at 25 °C, is the

critical density, p; is the saturated liquid density at 25 °C, Cpj is the isobaric liquid heat capacity at

Tfj, is the isobaric liquid heat capacity at 25 °C, and h^^ is the latent heat of vaporization at T^. The properties of CHF3 are also included in the table. One mixture is included which is 60% (by

mass) HFC-32 and 40% HFC- 125. Table 2 lists the characteristic constants used to estimate the
14 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Table 1 . Tabulated values of selected thermo-physical properties of agents

Agent MW

(kg/mol)

Tt T Pc Psa, Pc Pi Cp2

(kJ/kg) CO (MPa) (kg/m-^) (kJ/kg K)

HFC-236fd 0.152 -1.5 1 jU.d J . 1 / u.z / A ir\^ 4jU 1 JJO 1 1 1 y1 1.114 1.1/9 150''

FC-31-10 0.238 -Z.U 1 i J . Z l.iL u.z / ouu A 0< 1 u.yj 1

1 A

1

1 .Ul / 96'

FC-318 0.200 1 nil 1 1 ^ 1 13.Z z . / / U . J 1 a 1 q' oiy 1 AAQh 1 .uyo 112"

HCFC-124 0.137 1 ZZ . 3 3 .03 U . JO jOj 1 JD 1 1 .uou 1 111'' 1.111 162"

HFC-227ea 0.170 -16. 4-' 1 n 1 lUl . /J z .yir U.4 / oZ P 1 j95 1 A"7 A h

1 .0 /4 1 . 177" 131^

CF3I 0.196 -22.0' 122.0'" 4.04'" 0.49" 871'" 2106'' 0.542" 0.592" 106''

HFC- 134a 0.102 -26.2'' 101.2'^ 4.05'' 0.67'' 508" 1209'' 1.269" 1.395" 217"

FC-218 0.188 -36.8" 72.0'^ 2.67'' 0.88'' 629' 1321'' 0.977" 1.151" 101"

HCFC-22 0.087 -40.9'^ 96.2'' 5.04'' 1 .05'' 525' 1192'' 1.116" 1.239" 229"

HFC- 125 0.120 -48.6'^ 66.3'' 3.62" 1.38'' 571f 1190'' 1.107" 1.358" 160"

HFC-32/125 0.067 -52.54 73.21 5.041 1.674 4791 10404 1.449'' 1.763" 287*1

CFjBr 0.149 -57.8'^ 67.0'' 3.95'^ I.6I'' 745' 1551" 0.670" O.88I" Ul"

FC-116 0.138 -78.2' 19.7' 2.98' 608' 0.956' 117''

CHF3 0.070 -82.

r

25.6' 4.82' 4.69' 516' 0.685' 1.269" 240"

'From Du Pont Chemicals Data Sheet, personal communication (1993)

''Estimated by the modified Rackett method (Reid et al., 1987)

"Estimated by the method of Rowlinson (Reid et al., 1987)

''Estimated by Pitzer acentric factor correlation (Reid et al., 1987) "From Daubert and Danner (1993)

'^From Braker and Mossman (1980) ^From Brown and Mears (1958) "From Gallagher et al. (1993)

'From Shankland et al. (1990) Jprom Great Lakes Chemical Corporation (1993)

''From Hoechst AG, personal communication (1992) 'From Kudchadker et al. (1979) "From Sladkov and Bogacheva (1992)

"^Estimated by the method of Lee and Kesler (Reid et al., 1987) "From Piao et a/. (1991) Pprom Allied Signal Inc. (1990) iFrom Allied Signal Inc. (1991) 'From ASHRAE Inc. (1969) 'From Platzer et al. (1990)

isobaric ideal-gas heat capacities (C^^) of all the agents used in this study. The sources of these data

are also documented in Tables 1 and 2.
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 15

Table 2. Characteristic constants used to calculate C^" (kJ/kg K) as a function of temperature T

(K); Cp"" = A + ET + Cr^ + DT^ + (E/T)

Agent A B c D E

HFC-236fe^ -8.05 X

10-2 4.01 X

10-3 -3 82 X

10-^ 1.33 X

10-'=* 0

FC-31-10^ -1.59 X

10-' 4.24 X

10-3 -4 41 X10-^ 1.60 X lO-*^ 0

FC-318'' 4.51 X

10-2 3.33 X

10-3 -3 22 X

10-^ 1.14 X

10-'' 0

HCFC-124'^ 1.76 X

10-' 2.12 X

10-3 -1 00 X

10-6 0 0

HFC-227ea* -4.61 X

10-2 3.70 X

10-3 -3 54 X io-<' 1.22 X

10-' 0 CF3P 3.81 X

10-2 1.53 X

10-3 -1 70 X

10-6 6.60 X10-'° 0

HFC-134a'^ -5.26 X

10-3 3.30 X

10-3 -2 02 X10-6 0 15.82

FC-218^ 6.88 X

10-2 3.31 X10-3 -3 41 X 10-6 1.28 X

10-' 0

HCFC-22'' 1.61 X

10-' 2.08 X

10-3 -1 70 X10-6 5.28 X10-'° 0

HFC-125'^ 1.90 X 10' 2.47 X

10-3 -1 44 X10-6 0 0

HFC-32/125^ 3.82 X

10-' 1.62 X

10-3 -3 85 X10-^ 0 0

CFjBff 1.47 X

10-' 1.45 X

10-3 -1 42 X10-6 5.01 X lO 'O 0

FC-116'' 1.03 X

10-' 3.04 X10-3 -2 98 X 106 1.05 X

10-' 0

CHF3'' 5.49 X

10-2 2.97 X 103 -2 62 X

10-6 8.97 X 10 '° 0

^Estimated by the method of Joback (Reid et ai, 1987)

''From Yaws et at. (1988)

"^From Allied Signal Inc. (1990)

''From Fang and Joffe (1966)

''From Allied Signal Inc. (1991)

'^From Property Data Bank (Reid et al. 1987)

2.2 Experimental Apparatus and Procedures

The pressure-temperature relationship for pure agents and agent/pressurization gas mixtures at one fill condition {i.e., the volume of the liquid agent in the vessel was fixed in all the experiments) was measured. Except for the experimental procedure, the same apparatus was used in both studies. For

pure agents and agent/nitrogen mixtures, pressure measurements were determined at three temperatures, -60 °C, 22 °C, and 150 °C. For agent/CHF3 mixtures, pressure measurements were obtained

at -60 °C, 22 °C , and at the maximum temperature which did not cause the pressure to exceed the

safe operating limit of the vessel (for reasons to be discussed).

A fill condition of 1/3 of the vessel volume was selected because the so-called liquid-full

condition could be avoided when the vessel was heated to high temperatures. To illustrate the

significance of the liquid-full condition, let us consider a PT (pressure-temperature) phase diagram for a pure substance for the purpose of simplicity and clarity. Figure 1 shows the general behavior of a pure substance. The saturation curve (or the vapor pressure curve), which represents the existence of
16 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Figure 1 . PT diagram showing the saturation curve for a pure substance and isochores.
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 17

two phases (vapor and liquid) in equilibrium, separates the vapor and liquid regions. The curve

terminates at the critical point C. If one puts a fixed amount (M) of a liquid-vapor mixture of a pure

substance in a closed vessel with a volume V at 7} (represented by point A in Figure 1) and raises the

temperature of the vessel, two things will happen: the liquid vaporizes and at the same time expands.

The heating process at first follows the saturation curve until a single phase state is reached. This

state depends on the ratio of M to V, or the fill density (p^^/) of the vessel. If pj^ is greater than the

critical density (p^), continuous heating initially from point A causes the liquid-vapor interface (the

meniscus) to rise because the rate of thermal expansion of the liquid is faster than its vaporization

rate. Eventually, a state (point B) on the saturation curve will be reached where the liquid completely

fills the vessel. This condition is termed liquid-full. The location of point B can be determined by

identifying the temperature at which the saturation liquid density is equal to p^//. Further heating of

the vessel beyond point B can only occur in the liquid region along the liquid constant density line (or

isochore), BD, corresponding to the constant p^^n. If p^n is less than p^, heating initially from point A

causes the meniscus to recede due to the slower rate of thermal expansion of the liquid than its vaporization rale. The liquid eventually will completely vaporize, and the vessel is filled with vapor. The state at which this happens is represented by point E, which can be located by finding the

temperature at which the saturated vapor density is equal to the fill density. Further heating beyond

point E will only occur along the vapor isochore, EE, corresponding to the fixed p^n. If p^^y is equal

to p^, heating will cause little change in the level of the meniscus because the rate of thermal

expansion of the liquid is more or less balanced by its vaporization rate. The heating process will

follow the saturation curve until the critical point C is reached where the meniscus disappears.

Further heating then follows the critical isochore, CG. Note that the critical isochore has the same

slope as the saturation curve at the critical point. Figure 1 clearly illustrates the undesirable feature of heating the vessel beyond the liquid-full condition because a small increase in temperature can

create a dangerous over-pressure large enough to rupture the vessel.

2.2.1 Pure Agents. The experimental set-up is shown in Figure 2. A stainless steel (SS 304) vessel was used to measure the pressure-temperature relationship of the eleven agents. The vessel was tested hydrostatically to 20.6 MPa (3000 psia).

A K-type thermocouple (Omega TJ36-CASS-il6U-12) was inserted into the vessel to monitor the

internal temperature of the vessel in order to ensure that thermal equilibrium was reached between the

vessel interior and the ambience. For high temperature (150 °C) measurements, the vessel was immersed in a silicone oil bath heated with an immersion heater (Cole-Parmer Model L-01266-02).

For low temperature measurements, the vessel was submerged in a heat transfer fluid (Dow Corning

Syltherm XLT) which was cooled by two immersion coolers (NESLAB). One cooler (Model

1BC-4A) had a cooling capacity between 20 °C and -30 °C whereas the other (Model CCIOO-II) had

a capacity between -25 °C and - 100 °C. The rationale for using two coolers was that the cooling

time of the heat transfer fluid to the required temperature (-60 °C) was greatly reduced and that together with the immersion heater the entire temperature range, from -60 °C to 150 °C, was

covered.

A pressure relief valve (Nupro Model 177-R3A-K1-E) with a relieve pressure of 15.5 MPa (2250

psig) was used in the experimental set-up to prevent any accidental pressure overshoot when the

vessel was exposed to high ambient temperature. A pressure transducer (DRUCK Model PDCR 330) was used to measure the internal pressure of the vessel. The transducer had an accuracy of 6.9 kPa

(1 psi) and an operating range from 0.101 MPa (14.7 psia) to 13.8 MPa (2015 psia).
18 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

pressure

relief valve

It

/

pressure

transducer

constant

temperature bath

thermocouple

/

vessel

Figure 2. Schematic of temperature-pressure measurement experimental set-up.
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 19

There were two needle valves (Whitey SS-0RS2) in the apparatus. The one that was attached to the pressure vessel was used for filling and the other was used for isolating the apparatus from the vacuum pump line or the pressurization gas supply line. The total internal volume of the experimental apparatus was determined as follows. For the

vessel together with the attached needle valve, the internal volume was measured by filling them with

de-ionized water and weighing them on an electronic scale with an accuracy of 0.1 g. After this volume was determined, the entire system (Figure 2) was evacuated for at least 5 minutes and

pressurized with nitrogen to a fixed pressure at room temperature. The lower needle valve was then

closed, confining nitrogen to the vessel. The nitrogen remaining in the rest of the system was bled

out. With the lower needle valve closed, the apparatus was evacuated for at least 5 minutes by

cormecting a vacuum line to the upper needle valve. The upper needle valve was then closed before

removal of the vacuum line from the apparatus. The lower needle valve was then fully opened,

allowing the nitrogen in the vessel to expand throughout the apparatus. The final pressure of the set- up was recorded after thermal equilibrium with room temperature had been reached. Knowing the

initial pressure (before expansion), the initial volume (lower needle valve together with the vessel),

and the final pressure (after expansion), the total internal volume of the system was calculated by

using the ideal gas law and was found to be 53.2 ± 0.5 cm^.

The experimental procedure was as follows. The vessel together with its needle valve was

initially disconnected from the experimental set-up and evacuated for at least 5 minutes. The vessel was then connected to the agent supply bottle (not shown in Figure 2). By immersing the vessel in dry ice, liquid agent was dispensed through the needle valve into the vessel by condensing gaseous

agent from the supply bottle. No attempt was made to remove noncondensible gases, if there were

any, in the liquid agents. The amount of agent used in each of the experiments was approximately

equal to the amount required to fill one-third of the vessel volume with liquid agent. The vessel was

then weighed on an electronic balance with an accuracy of 0.1 g to determine the actual mass of the

agent. The vessel was re-connected to the experimental set-up, which was then evacuated for at least 5 minutes before the needle valve attached to the vessel was opened fully. The apparatus was then

immersed in the controlled temperature bath. The pressure of the vessel was recorded from the

pressure transducer read-out when the internal temperature had reached thermal equilibrium with the

bath temperature. Note that since FC-116 is a gas at room temperature, the 1/3 fill condition was not applicable to this agent. Although liquid FC-116 could stilled be dispensed by condensing gaseous

FC-116 in the dry-ice-cooled vessel, this approach was not taken because the amount of liquid mass

in the vessel could cause a huge pressure rise when the vessel was warmed back to room temperature.

An average fill density of about 370 kg/m^ was used for FC-116. This fill density was governed by

the pressure of the agent supply bottle, i.e., gaseous FC-116 was dispensed into the vessel directly

from the supply bottle until an equilibrium pressure was reached in both the vessel and the supply

bottle.

2.2.2 Nj/Agent and CHF3/Agent Mixtures. As mentioned in the beginning of this section, the same experimental set-up for evaluating pure agents was used for the mixtures. The experimental

procedure is described in the following paragraph.

After a pure liquid agent was dispensed in the vessel as described above and the vessel was re- connected to the experimental set-up, the pressurization gas supply line was then connected to the upper needle valve. The supply line together with the experimental set-up was evacuated for at least 5 minutes before initiating a slow flow of the pressurization gas. The lower needle valve was then

fully opened to allow the gas to bubble slowly through the liquid agent until an equilibrium pressure

of approximately 4. 1 MPa (600 psia) was reached. Repetitive bubbling of the pressurization gas was

required before the final equilibrium pressure could be obtained. This value of pressure was selected
20 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

in order to simulate a charging pressure commensurate with existing halon 1301 vessel. The vessel was then immersed in the controlled temperature bath. Once the vessel was in thermal equilibrium

with the bath, the pressure of the vessel was recorded.

2.2.3 Solubility of N2 and CHF3 in Agents. Figure 3 shows a schematic of the apparatus used for solubility measurements. The experimental apparatus included the same stainless steel vessel used in the pressure-temperature measurements, a needle valve for dispensing the liquid agent and pressurization gas, a thermocouple, and a pressure transducer (DRUCK Model PDRR 330). The actual volume

of the system was determined by pressurizing it with nitrogen to 4.02 MPa and then weighing it on an

electronic scale with an accuracy of 0.1 g. Knowing the temperature, the pressure, and the mass of

nitrogen, the volume was then calculated by using a generalized correlation for the compressibility

factor (Smith and Van Ness, 1975). The total volume of the apparatus was found to be 53.9 ±

0.6 cm^

The experimental procedure was as follows. The apparatus was evacuated for at least 5 minutes

before the agent was dispensed to the vessel. The same filling and pressurization procedure were

used as in the pressure-temperature measurements for mixtures. After pressurization with nitrogen or CHF3, the amount of pressurization gas required to pressurize the vessel to the specified final pressure was then obtained by weighing the apparatus. This amount corresponded to the sum of the mass of nitrogen or CHF3 in the vapor phase and that of the dissolved nitrogen or CHF3 in the liquid

agent.

2.3 Results and Discussion

2.3.1 Pure Agents. Table 3 summarizes the results (averages of at least two runs ± half the range)

of temperature-pressure measurements for pure agents at 150 + 1 °C. Fill density is defined as the

ratio of the mass of the agent to the volume of the apparatus (53.2 cm-^). The third column shows the

measured final pressures at 150 + 1 °C. For all the agents evaluated, the final pressures at 150 °C

were less than 11 MPa (1600 psia). At -60 °C, the final pressures for all the agents, except

HFC-27ea and FC-116, were below 0.101 MPa (14.7 psia) which is below the measuring range of

the pressure transducer. For FC-116, an average final pressure of 0.25 MPa was found at -61 °C.

For HFC-227ea, an average final pressure of 0.18 MPa was measured at -60 °C. Subsequent

measurements of HFC-227ea from a different agent bottle showed an average value below 0.101

MPa. The results of HFC-227ea reported herein were obtained by using the agent supply bottle that gave a final pressure of 0.25 MPa at -60 °C, which represented the worst case. Since the amount of agent and the volume of the apparatus (53.2 cm-^) were constant, an existing

equation of state was used to predict the final pressure at the final temperature along an isochore or

constant density line. The Peng-Robinson equation of state (Peng and Robinson, 1976) has recently

been used to correlate thermodynamic properties of refrigerants and their mixtures (Abu-Eishah,

1991). It is given by

p ^ RT a{T)

V - b v(v + b) + b{v - b)

where v is the molar volume, R is the universal gas constant, and T is the absolute temperature (in

K). The parameters a and b are given in terms of the reduced temperature (7^ = T/T^), and the

acentric factor, cj:

i
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 21

Table 3. Comparison of measured pressures at 150 + 1 °C with calculated values from REFPROP

and the Peng-Robinson equation of state

Agent Fill density

(kg/m3) Pf (MPa)

measured

Pf (MPa)

REFPROP

erroi^ (%)

Pf (MPa) P-R EOS

error'' (%)

HFC-236fe 467 ± 2.0 4.50 ± 0.03 - - 4.24 5.8

FC-31-10 523 ± 0 3.45 ± 0.01 - - 3.76 9.0

FC-318 526 ± 1.1 4.05 ± 0.01 4.31 6.7 4.40 8.6 HCFC-124 471 ± 2.4 4.93 ± 0 5.22 5.9 5.27 6.9

HFC-227ea 495 ± 1.3 5.13 ± 0.04 5.27 2.7 5.41 5.5

CF3I 733 ± 0.1 5.19 ± 0.01 - - 6.06 16.8

HFC- 134a 430 ± 0.8 6.80 ± 0.02 7.53 10.7 7.73 13.7

FC-218 500 ±3.5 5.41 ± 0.02 5.89 8.9 6.11 12.9

HCFC-22 436 ± 2.7 8.48 ± 0.13 9.25 9.1 9.48 11.8

HFC-125 456 ± 1.2 7.73 ± 0.07 8.60 11.3 8.90 15.1

HFC-32/125 390 ± 0.5 10.78 ± 0.07 12.67 17.5

CFjBr 605 ± 0.3 8.11 ±0 9.07 11.8

FC-116 368 ± 5.9 6.77 ± 0.10 7.70 13.7

^error(%)= | P. (measured) - P. (REFPROP) | x 100/P, (measured)

''error (%) = | Pj (measured) - P. (P-R EOS) | x 100//) (measured)

a{T) = a(r)a(r^,o)) (2)

a{T) = 0.45724 - P

1/2x12 (4) a = [1 + p(i - n)\

P

- 0.37464 + 1.54226(0 - 0.26992 co^
22 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

pressure

transducer

thermocouple

vessel

Figure 3. Experimental set-up for solubility measurement.
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 23

0) = -log

(p \ - 1.000 (6)

7/7^ = 0.7

RT

0.07780

- (7)

P

In addition, the computer program, REFPROP (Gallagher et al., 1993) was used to compare

with the calculated results obtained by using the Peng-Robinson equation of state. REFPROP was

developed by using the Carnahan-Starling-DeSantis equation of state (Morrison and McLinden, 1986):

Pv _ I + y + y'^ -y^

RT n -

(1 - yy RT{v +b*)

4v

(8)

(9)

a' = a^expiaj + a^T^) (10)

b* = b^ + bj + bj^ (11)

where Cq, fly, bg, bj, and ^2 ^re parameters obtained from fitting of limited experimental data. Table 3 also summarizes the results obtained from REFPROP and the Peng-Robinson equation of

state (P-R EOS). The calculations were based on the conditions (average values) given in the second

and third columns in Table 3. The blank entries in Table 3 imply that those agents are not currently

included in REFPROP

From Table 3, it appears that REFPROP predicts the final pressures much closer to the

experimental values than the Peng-Robinson equation of state. In all cases, except for HFC-236fa,

both equations of state tend to overestimate the final pressures.

2.3.2 N2/Agent and CHF3/Agent Mixtures. Table 4 tabulates the results (averages of at least two

runs + half the range) of N2/agent mixtures. The third column shows the initial pressures of the

vessel at 23 ± 1 °C, the fourth column lists the final pressures of the vessel when the temperature is at — 60 ± 1 °C, and the fifth column tabulates the final pressures of the vessel at 150 + 1 °C.

At —60 °C, there was still significant residual pressure in the vessel. This residual pressure was

mainly due to the nitrogen in the vapor phase because the contribution to the pressure from the agent

vapor was insignificant due to very low vapor pressure (< 0.101 MPa) of the agent at —60 °C.

Therefore, the initial pressurization of the vessel with nitrogen ensures that enough residual driving

force to propel the agent out of the vessel is still present when the vessel is discharged at a cold

ambience. Without the pressurization gas, the vapor pressure of the agent was so low that the only
24 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

driving force was the static head of the liquid agent when the vessel was placed in the vertically

downward position.

At 150 °C, the final pressures of all the agents, with the exception of the HFC-125/HFC-32

mixture, were below 13.7 MPa (2000 psia). For the mixture, the experiment was terminated at 140 °C because the pressure was close to exceeding the operating range of the pressure transducer.

The results of CHF3/agent mixtures are summarized in Table 5 . The explanation of the content

of the table is similar to that of Table 4. The values shown in Table 5 are the averages of at least two runs with their half range. At about —60 °C, the final residual pressures in all cases were very low. This would make

CHF3 as a pressurization gas unfavorable if the vessels were to be discharged at low temperatures.

The results also imply that a significant amount of CHF3 was dissolved in the liquid agent at -60 °C.

When the agent/CHF3 mixtures (except FC-II6/CHF3) were heated, the pressures exceeded

13.7 Pa by the time a temperature of 70 °C had been reached. Because CHF3 is highly soluble in all the liquid agents and the initial amount of CHF3 required to pressurize the vessel is very significant

(see Section 2.3.3), an increase in pressure at moderately elevated temperatures was caused by a decrease in solubility of CHF3 in the liquid agent and a subsequent increase in the amount of CHF3 in the ullage. As a result, if CHF3 were to be selected as a pressurization gas, extreme care should be

taken in the design of the bottle for the replacement agent, especially if the bottle is to be exposed to a high temperature environment.

2.3.3 Solubility of N2 and CHF3 in Agents.

2.3.3.1 Experimental Results. Table 6 summarizes the experimental results (averages of at least two runs ± half the range) for nitrogen. In all cases, the total amount of nitrogen required to pressurize the system to 4.1 MPa was less than 8% of the total mass of the agent.

Experimental results (averages of at least two runs + half the range) from CHF3/agent mixtures

are shown in Table 7. The amount of CHF3 needed to pressurize the vessel to ~ 4.1 MPa is as much as or greater than that of the agent in the vessel. This is not surprising because CHF3 is readily

soluble in all the agents. Since a significant amount of CHF3 is dissolved in the liquid agent,

subsequent heating of the vessel causes a steep rise in pressure due to decreasing solubility of CHF3

in the liquid agent at high temperatures (see Section 2.3.2).

2.3.3.2 Theoretical Predictions. In order to determine the solubility of nitrogen in the agent,

measurements of nitrogen and agent concentrations in both phases are required. Solubility is normally defined as the mole fraction of nitrogen dissolved in the liquid agent. The present

experimental set-up does not have the provision for sampling the liquid and gas phases, but the mass

of nitrogen that is dissolved in the liquid agent can be estimated by applying mass balances on the

agent and nitrogen. The calculation procedure is as follows.

Initially, the vessel is filled with a pre-determined amount of agent, ^ at temperature T (K) (in

this case room temperature); which is expressed as m = m +w, = p V + p ,V, (12)

a,t a,v a,l r a,v v " a,l I

where m is the mass, subscript t represents total, subscript a represent agent, subscripts v and / represent vapor and liquid phases respectively, p is the saturation density of the agent, and V is the

volume. Since ^ V^, ^, and ^ are known and V/ = - V^, the masses of agent in both

liquid and vapor phases can be calculated by substituting V, - for and solving for in Equation (12).
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 25

Table 4. Initial pressures (P,) at 23 ± 1 °C and final pressures (P^j) and (Pf 2) at -60 + 1 °C

and 150 ± 1 °C, respectively, for N2/agent mixtures

Agent Fill density

(kg/m^) (MPa) (MPa) (MPa)

HFC-236fe 468 + 0.2 4.16 ± 0 2.72 ± 0 9.66 ± 0.02

FC-31-10 523 + 0.1 4 15 ± 0.02 2.68 ± 0.01 8.82 ± 0.03

FC-318 524 ± 2.3 4 16 ± 0.04 2.80 ± 0.18 9.32 ± 0.28 HCFC-124 471 ± 0.5 4 16 ± 0.05 2.65 + 0 9.96 ± 0.02

HFC-227ea 492 ± 2.0 4 14 ± 0.02 2.55 + 0.02 9.99 ± 0.10

CF.I 736 ± 1.0 4 17 ± 0.01 2.67 ± 0 10.25 ± 0.01

HFC-134a 430 ± 0.5 4 16 ± 0.01 2.50 ± 0.01 11.59 ± 0.02

FC-218 497 ± 2.9 4 24 ± 0.05 2.40 ± 0.06 10.16 ± 0.44

HCFC-22 433 ± 0.3 4 20 ± 0.02 2.46 ± 0.01 13.00 + 0.03

HFC-125 458 ± 1.2 4 16 ± 0.05 2.11 ± 0.04 11.59 ± 0.71

HFC-32/125 390 ± 2.9 4 13 + 0.02 2.04 ± 0.01 13.82' ± 0

'

CFjBr 601 ± 1.8 4.20 ± 0 2.07 + 0 12.30 ± 0.04

FC-116 371 ± 0.9 4.24 ± 0.04 1.24 ± 0.02 9.07 + 0.08

^at 140 °C

y __ ^a, - P../^

(13)

V - n

Pa,v Pa,/

The vessel is then pressurized with nitrogen until an equilibrium pressure Pj at T is reached.

Assuming that the dissolved nitrogen would not cause a significant change in the saturation liquid

density of the agent and that the liquid agent is incompressible, the liquid mixture density can be

approximated by the saturation liquid density of the pure agent. This approximation is reasonable if the amount of nitrogen dissolved in the liquid agent is much smaller than the mass of the liquid agent.

If the vapor phase is further assumed to be ideal, then the total pressure P^ is given by

P=p + p (14)

where and p^ are the partial pressures of the agent and nitrogen respectively. From Raoult's law

(Denbigh, 1984),
26 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Table 5. Initial pressures (P,) at 23 + 1 °C and final pressures (Pf j) and (fy2) at -60 ± 1 °C

and Tf2, respectively, for CHF3/agent mixtures

Agent Fill density

(kg/m^)

P. (MPa) ^ IVl r a ) rc)

Pf.2 (MPa)

HFC-236ra 467 ± 1.0 4 19 + 0 07 n T A 1 n A

1

U.jU + U.Ul 48 ± 1 13 51 ± 0 01

FC-31-10 523 ± 0.4 4 31 + 0 01 A TJ _L A A

1

U.JJ + U.Ul 56 ± 0 13 50 ± 0 01

FC-318 525 + 0.6 4 21 ± 0 01 0.34 + 0 53 + 1 13 53 ± 0 02

HCFC-124 472 + 0.5 4 11 ± 0 01 0.29 ± 0 48 + 1 13 55 ± 0 02

HFC-227ea 495 ± 0.6 4 22 ± 0 02 0.29 ± 0 53 ± 1 13 51 ± 0 03

HFC- 134a 426 ± 0.4 4 18 ± 0 05 0.27 ± 0 47 ± 1 13 55 ± 0 03

FC-218 498 + 1.8 4 19 ± 0 02 0.44 ± 0.04 62 ± 0 13 54 ± 0 01

HCFC-22 433 ± 0.1 4 21 ± 0 05 0.29 ± 0.01 49 ± 1 13 53 ± 0 03

HFC-125 455 ± 0.5 4 23 ± 0 04 0.28 ± 0.01 52 ± 1 13 58 ± 0 01

HFC-32/125 391 + 1.1 4 19 + 0 02 0.27 ± 0.01 51 ± 3 13 59 ± 0 11

FC-116 376 ± 1.5 4 21 ± 0 08 0.39 ± 0.02 150 ± 0 11 22 ± 0 70

n Pa = ^a,l Psat - p . (15) n , + n ,

where / is the mole fraction of the agent in the liquid phase, p^^^ is the saturation vapor pressure of

the agent at temperature T, j is the number of moles of dissolved gas in the liquid phase, and j is the number of moles of agent in the liquid phase. Furthermore, can be calculated by

n RT (n - n ,) RT

n = = -1^:^ ^(16) ^ F V V V

where ^ is the number of moles of pressurization gas in the vapor phase, , is the total moles of

pressurization gas required to pressurize the vessel to Pj at a temperature T, and R is the universal gas

constant. Substituting Equations (15) and (16) into Equation (14), a quadratic equation for

j (unknown) is obtained.

n'l + - « . + ^) - n , - ^ - n ) = 0 d^)

Solving for j(only one root is meaningful) and knowing n^ ^, = n^j - the amount of

nitrogen in both phases can be calculated. When no nitrogen is dissolved in the liquid agent

{i.e., n I = 0), Pj = p + p^^j is recovered from Equation (17).
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS \ 27

Table 6. Amount of nitrogen required to pressurize mixture to total pressure, at 23 + 1 °C

Agent P, (MPa) Nitrogen (g) Agent (g)

HFC-236fa 4.27 ± 0.08 1.9 + 0.1 24.8 ± 0.1

FC-31-10 4.15 ± 0.02 2.0 ± 0 27.8 ± 0.1

FC-318 4.16 ± 0.03 2.0 ± 0.1 28.0 + 0 HCFC-124 -

4.16 ± 0.01 1.9 ± 0.1 25.2 ± 0.1

HFC-227ea 4.16 ± 0 1,8 + 0 26.3 ± 0.1

CF3I 4.28 ± 0.05 1.9 ± 0 39.0 ± 0 HFC- 134a 4.18 ± 0.02 1.8 ± 0.1 22.8 ± 0.1

FC-218 4.15 ± 0.02 1.7 ± 0.1 26.6 + 0.1 HCFC-22 4.12 ± 0.01 1.4 ± 0.1 21.4 ± 0.1

HFC-125 4.21 ± 0.14 1.5 ± 0.1 24,3 + 0

HFC-32/125 4.17 ± 0.01 1.3 ± 0 20.7 ± 0,1

CFjBr 4.17 + 0.01 1.4 ± 0.1 32.0 ± 0.1

The results and the parameters that are required to perform the calculation are shown in Table 8. The conditions used for the calculations correspond to Table 6. The calculated solubilities (x^ /) and mass fractions (w p are also tabulated in Table 8. Solubility is defined as follows:

n X

8-1

and the mass fraction of nitrogen in the liquid agent is calculated by

(18)

'riA (19)

g.l g ^ gJ' a

where and are the molecular weights of agent and pressurization gas (nitrogen) respectively.

For all the agents studied, the amount of nitrogen dissolved in the liquid agent was calculated to be less than 25% of the total amount of nitrogen added, and less than 2% of the mass of the agent.

The solubility of nitrogen in liquid agent can also be estimated if the solution can be assumed

regular. A correlating scheme based on the theory of regular solutions and the law of corresponding

states to predict solubility of a gas in liquid was developed by Prausnitz and Shair (1961). The

following equations form the basis of the correlation of gas solubilities:
28 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Table 7. Amount of CHF3 required to pressurize mixture to total pressure, at 23 + 1 °C

Agent P, (MPa) CHF3 (g) Agent (g)

HFC-236fe 4.29 ± 0.04 32.8 ± 0 24.7 ± 0.1

FC-31-10 4.16 ± 0 28.8 ± 0,2 27.9 ± 0.1

FC-318 4.17 ± 0.01 30.3 ± 0 27.9 ± 0.1

HCFC-124 4.21 ± 0.05 32.0 + 0.1 25,1 ± 0.1

HFC-227ea 4.16 + 0.04 30,4 ± 0 26.3 + 0.1

HFC- 134a 4.17 + 0.03 31.9 + 0,1 22.7 ± 0

FC-218 4.23 ± 0.03 25 ± 0.1 26.4 ± 0

HCFC-22 4.33 ± 0.21 30.8 ± 0.4 22.8 ± 0

HFC- 125 4.16 ± 0.01 28.2 ± 0 24.4 ± 0.1

HFC-32/125 4,20 ± 0.01 29.9 ± 0.1 20.7 ± 0.1

CFjBr 4.12 ± 0.01 23.9 ± 0,1 31.8 ± 0.3

_1_

with

L = (21)

pure,g f

exp[

RT

(20)

(22)
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 29

Table 8. Parameters used in the solubility calculations, calculated mass of nitrogen in liquid agents,

calculated nitrogen solubility, and calculated nitrogen mass fractions in liquid agents

(conditions shown in Table 6)

Agent Pa.l (kg/m^)

Pa.v (kg/m^)

Pscu (MPa) (g)

HFC-236fe 1371 15.8^ 0.26 0.2 0.05 0.01

FC-31-10 1505 27.3 0.25 0.4 0.11 0.01

FC-318 1504 27.0 0.30 0.4 0.10 0.02 HCFC-124 1363 22.5 0.36 0.3 0.06 0.01

HFC-227ea 1400 36.3 0.45 0.3 0.07 0.01

CF3I 2102 36.7^ 0.46'' 0.3 0.06 0.01

HFC-134a 1214 30.2 0.62 0.3 0.05 0.01

FC-218 1333 83.4 0.84 0.3 0,08 0.01 HCFC-22 1206 40.7 0.96 N.S.

HFC-125 1201 84.0 1.31 0.3 0.05 0.01

HFC-32/125 1045 59.3 1.63 0.2 0.02 0.01

CF3Br 1556 126.6 1.53 0.3 0.05 0.01

^Estimated by using the ideal gas law "Estimated by the method of Lee and Kesler (Reid et al., 1987)

N.S. : not significant

and

6. -

\

RT

(23)

where /is the fugacity, is the molar "liquid" volume, 5^ is the solubility parameter of solute

(nitrogen), 6^ is the solubility parameter of solvent (liquid agent), h^^ is the heat of vaporization,

superscripts L and G represent liquid and vapor phases respectively, and subscripts a and g

represent

agent and pressurization gas respectively. In order to use Equation (20) to estimate solubility, the

three parameters for the dissolved gas as a hypothetical liquid if^ure ^"d need to be

specified. These parameters are all dependent on temperature; nowever, according to the theory of

regular solutions the term (6^ - b^)^ is not temperature-dependent. Therefore, any convenient

reference temperature can be used for and 6^ as long as the same reference temperature is also

used for 6^ and v^. Since the "liquid" molar volume and solubility parameters for gaseous solutes are normally tabulated at 25 °C in the literature, this temperature is used in the following calcuiatiof! of
30 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

(6^ - bgf- 'i'^. However, the dependence of the fugacity of the hypothetical liquid on temperature

must be considered in the solubility estimation.

By applying the theorem of corresponding states, it can be shown that the reduced fugacity

^ure J^c hypothetical pure liquid solute can be correlated as a universal function of the

reduced temperature, TIT^ ^ (Prausnitz and Shair, 1961), where P^^, Tand ^ are the solute critical pressure, the solution temperature and the solute critical temperature respectively. The function can be represented by the following two equations (Walas, 1985):

f

^ 8 06 r T- 7- \nCJ^) = 7.81 - ^ - 2.94 hi(—) 0.7< —<2.5 (24)

P T T T c,g c,g

Ce^~~e.O 2.5<-^<3.0 (25)

The correlation given by Prausnitz and Shair (1961) is for a total pressure (P^) of 0.101 MPa. To

account for higher pressure, the Poynting correction (Prausnitz et al., 1986) should be applied; thus

fpure,, (® - fpure,s (@ 0.101 MPa) exp [ ' '^^ ] (26)

where P^ is in MPa.

Since contains i, iteration is required to solve for Xg i in Equation (20). However, the

iteration converges rapidly. If the solubility is small, a first approximation of x^ j can be obtained by

setting equal to unity.

The fugacity, J^, is approximated by the partial pressure of nitrogen in the vapor phase and can be estimated as follows. From Equation (15),

f ^ p ^ P - p (27)

- P- {\ - x ,)p , (28)

Table 9 lists the parameters used in the solubility estimation and the calculated solubilities (Xg /). The solubility parameters of the agents were obtained from Daubert and Banner (1992) or were

calculated by using Equation (22) with assumed to be the molar liquid volume of pure agent at 25 °C. The solubility parameter of nitrogen and were obtained from Prausnitz et al. (1986).

The third method for predicting solubility is to perform a detailed vapor-liquid equilibrium

calculation by using an appropriate equation of state. Specifically, an isothermal flash calculation is required (Prausnitz et al., 1980, 1986, and Walas, 1985). Such calculations involve the determination

of vapor and liquid phase compositions, given the total system composition (in this case, the amount

of pressurization gas and the agent in the vessel, total pressure, and temperature). The computer

program PROZPER developed by Ely and Huber (1993), was used to calculate the solubility of

nitrogen in agents. The Peng-Robinson equation of state was used in the calculation although other
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 31

Table 9. Parameters used in the estimation of solubilities, calculated solubilities of nitrogen in agents, and calculated mass fractions of nitrogen in agents

Agent (cm^/mole) 5 (J/cm^)"^ @ 25°C @ 25°C

HFC-236ta 111.4 12.9 0.18 0.04

FC-31-10 159.0 11.9 0.18 0.03

FC-318 133.4 13.0 0.18 0.03 HCFC-124 100.6 13.1 0.17 0.04

HFC-227ea 121.9 11.5 0.17 0.03

CF3I 93.2 13.0 0.17 0.03 HFC-134a 84.6 13.5 0.16 0.05

FC-218 142.1 12.3 0.16 0.03 HCFC-22 72.4 17.3 0.14 0.05

HFC-125 100.9 10.2 0.14 0.04

HFC-32/125 64.7 13.0 0.12 0.05

CFjBr 96.8 14.5 0.13 0.03

Nitrogen 32.4 5.3

equations of state could be used with some modification. Calculations were performed without taking

the binary interaction coeflftcient into account. Table 10 shows the input parameters used for the calculation by PROZPER and the calculated

solubility Agents that are not currently included in PROZPER iiave blank entries in the table. The

column, represents the overall mole fraction of nitrogen in the vessel:

X = (29)

where , and ^ are the total moles of pressurization gas required and agent in the vessel, respectively, (given in Table 6), and ^ {= 1 - is the overall mole fraction of agent.

Table 1 1 compares the three methods used to estimate nitrogen solubility in the agents studied.

The solubilities obtained from the correlation scheme appear to be higher than those obtained by

PROZPER, and the values calculated by the mass balance method are generally lower than those

obtained by correlations and vapor-liquid equilibrium calculations. However, all three methods

predict solubility on the same order of magnitude, and the calculated mass fractions of nitrogen in liquid agents in all cases are less than or equal to 0.05. The discrepancy could be due to the

assumptions invoked in both mass balance and correlation techniques. In the mass balance approach,

the density of the mixture is approximated by the density of pure agent. This assumption is only true

if ; ^ 0. On the other hand, the correlation scheme may not work well if the pressure is high or if the solution is not regular.
32 2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Table 10. Experimental parameters, calculated solubilities of nitrogen, and mass fraction of nitrogen

in liquid agents

Agent r(K) P, (MPa)

g' HFC-236fa 296 4.27 0.29

FC-31-10 296 4.15 0.38

FC-318 297 4.16 0.34 0.13 0.02 HCFC-124 296 4.16 0.27 0.09 0.02

HFC-227ea 297 4.16 0.29 0.12 0.02

CF3I 296 4.28 0.25 0.09 0.01 HFC- 134a 296 4.18 0.22 0.09 0.03

FC-218 296 4.15 0.30 0.15 0.03 HCFC-22 295 4.12 0.17 0.07 0.02

HFC-125 296 4.21 0.21 0.11 0.03

HFC-32/125 297 4.17 0.13

CFjBr 296 4.17 0.19 0.09 0.02

The first two approximate methods, discussed above, for estimating solubility are not appropriate

for calculating solubility of CHF3 in liquid agents because the assumptions made in the calculations

become invalid. Given the amount of CHF3 and that of the agent at a specified temperature and

pressure, the solubility of CHF3 in the liquid agent can, in principle, be estimated by performing an

isothermal flash calculation. However, since the critical point of CHF3 is very close to the experimental condition (at room temperature, and -4.1 MPa), attempts to use PROZPER to estimate the

solubility of CHF3 in the agents were unsuccessful, the reason being that the Peng-Robinson equation

of state may not be able to represent the region near the critical point well.

2.4 Summary and Recommendations

A vessel filled 1/3 with liquid agent initially at room temperature increased in pressure as it was

heated to 150 °C to a value below 11 MPa (1600 psia), for all of the pure compounds studied. The

HFC-32/125 mixture had the highest final pressure, 10.78 MPa. When cooled to -61 °C, the

pressure of all the agents except FC-1 16 dropped below the ambient (0.101 MPa). The final pressure

of FC-116 was 0.25 MPa (37 psia).

The amount of nitrogen required to pressurize the vessel to 4.1 MPa (600 psia) accounted for

less than 8% of the total mass of the agent. The calculated mass fractions of nitrogen dissolved in the

liquid agents in all cases was less than 0.05. For N2/agent mixtures, the final pressure of the vessel

at 150 °C was less than 14 MPa (2040 psia), with the HFC-32/125 mixture again yielding the highest

final pressure (13.8 MPa). The nitrogen in the mixture maintained an appreciable residual pressure
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS 33

Table 11. Comparison of the three methods used to calculate solubilities of nitrogen in agents

(conditions shown in Table 6)

Agent

(Mass balance) (Correlation) (PROZPER)

HFC-236fe 0.05 0.18 -

FC-31-10 0.11 0.18 -

FC-318 0.10 0.18 0.13 HCFC-124 0.06 0.17 0.09

HFC-227ea 0.07 0.17 0.12

CF3I 0.06 0.17 0.09

HFC-134a 0.05 0.16 0.09

FC-218 0.08 0.16 0.15 HCFC-22 0.14 0.07 HFC-125 0.05 0.14 0.11

HFC-32/125 0.02 0.12

CFjBr 0.05 0.13 0.09

of 1.24 MPa (180 psia) even when the vessel was cooled to -60 °C. This residual pressure is essential when the vessel is discharged at a cold ambience.

Significant amounts of CHF3 were needed to pressurize the vessel to 4.1 MPa because CHF3 is readily soluble in all of the liquid agents studied. For all of the CHFj/agent mixtures, pressures

above 13.7 MPa (2000 psia) were reached for temperatures as low as 70 °C. At -60 °C, the final pressures dropped below 0.45 MPa (65 psia). This low pressure would make CHF3 an unfavorable

choice for a pressurization gas if discharge of the vessel happened at very cold temperatures.

The following recommendations are made to ensure the integrity of the storage vessel:

1. The current pressure vessels described in military specification MIL-C-22284A (proof

pressure of 9.62 MPa and minimum burst pressure of 12.37 MPa) should not be used as drop-ins for any of the core alternative agents when the vessels are to be exposed to or stored

at temperatures up to 150 °C.

2. Trifluoromethane, CHF3, should not be used as a pressurization gas. If discharge of the

vessel is to occur at cold ambience, a very low residual pressure will exist, and at temperature in excess of 70 °C, very high pressures could lead to rupture of the vessel.
34 2. I III.KMODYNAMK I'KOI'I'.RI IIiS ()!• Al .ri'.RNATIVI'; AGENTS

2.5 krloroiKcs

Ahii I'jsIkiIi, si,, "(

';ik m hi oI Vipor I iiiiii Dal.i loi Hiiuiry ( 'lilorolliioroca- Mixliiics Usui)' (lie IVii)- Kohiiison I ol Sl.ilc," l liiid I'hasc lujiiilihria 62, 41 (1991).

Allied Ni,-ii;il Inc., Sricclnl riivsinil I'lopriltrs : IK IC /2J, IKl'C I2.kl. HCFC-I24. IIFC-125,

IIK I >'•/, NIC l.)4(i, (iiul lll'C I4.i(i, Allied Si)',ii;il Inc., HLiHalo Research Laboratory, Ikillalo, New

Allied Nij'ii.il liK , Cciiciioii " rKxIiicls I'cchtiical lliillctiti: R J2/I25 Azeotrope, Allied Signal

liK , (ieiielioii riddiu lN, Mori islowii. New .Icrsey, .Inly 1991.

AMIKAI': Inc., ASIIRAl'. I'liciiiuxlyiKiiitn I'lopcrlics oj Rcfri^cniiils, ASIIRAIi Inc., New York,

I'X)') Itiakcr, W., and Mos.siiian, A.I,., Malhcsoii (his Data Hook, ed., Matheson Gas Products,

New leisey, I9S(). Brown, .I.A., and Meats, W.ll., "Physical I'ropeilies ol n IVrlluorobulaiic, " ./. Phys. Chem. 62,

')(.() ( l')SK). Danheil, T. I',

, and Daiinei, R.I'., I'hxsiail and Tlicniiot/vntiniic Properties of Pure Chemicals,

Diilii < '(>ni/>iliifion , I leniisplu-ii' Riihlishiii)' ( 'oi poial ion. New York, 1992.

I )eiil)i)'ji, K ., ///(' l'rin( i[>l('s oj Chcinudl lu/iiilihriiini, Cambridge University Press, Cambridge,

l*),SI

I ly. .1 , and lliihei. M l ., I'RO/I'I R, personal comnumicalion with M.I.. Ilnber (1993).

l aii)'., I

' , ami .lollc, .1., "

I herinodynamic Properties of Perlluoropropane," ./. Chem. Eng. Data

//, yiu ( \')uu).

( iallai'liei .1 , Mcl inden, M., Ilnber, M., anil Ply, .1., NIST StaiuUird Reference Datahase 23:

ihi'i inoilxiiomic rropcriics of Refriy^enmts aiul Refrit>eniiif Mixtures Daldhdse {Rill PROP), Version I (). \IS Di-paiinieiii ol ( oninicrce, Washington, DC 1993.

(ileal I aki- ( lieniical Corporalit)n, l<'M-200'^: hjiviromiieiitally Superior Fire Extitiguishant for

Ibldl i'loiuliN}: A/>/>li( iitions, (ileal l ake Chemical Corporaliitn, West Pafayetle, Indiana, May 1993.

Kiulcliadkei, A. P., Kndcliadkei, S A., Shnkla, R P. , and Palnaik, PR., "Va|)or Pressures and

Moiling Ponils ol Nelecleil I laU)inellianes, " ./. /V/\'.v. Clieni. Ref Data .S', 499 (1979).

Morrison, (i., and McliiKien, M.O. , Applicatio)! of a Hard Sphere lu/iiation of State to Refrigeranfs (ind R(iri}'cnini Mixtures. NHS Technical Note 1226, U.S. Department of C(Miimerce,

Washini'lon, DC P)S()

Penr,, D. ^ . , and Robinson, D.H , "A New Two Conslani l'A|nalion of Slate, " Ind. Fug. Chem.

Mint him. /.^, .S9 (197(»).

Piao, ('., Salo, II., and Walanabe, Is..,

"

I hermodynamic Charts, Tables, and litiualions tor Refrigerant IIIC 13 la," ASIIRAF iniiisactioiis V7, 26S (1991).

Platzer, B., Poll, A., and Manrer, (i., Tlieniioplix'sicul Properties of Refrigerants, SpringerVerlag, Berlin, 199t). Piansnil/., .I.M., and Shair, T'.ll., "A Therinodynanuc Correlation ol Gas Solubilities," AlCliE

.loiinuil 7, 6S2 tP)(>l).

Prausnit/, .I.M., Anderson, T.F., Grens, p:.A., Pxkert. C.A., llsieh,R., and O'Connell, J.P,

Computer Ciilciilutions for Miilticomponeni Mipor-lM/iiid and iM/iiid Im/iikI lu/iiilihria . Prentice-Hall,

New Jersey, 1980.

Piansnil/, .I.M., I ichtenthaler. R.N., and Gomes de Azevedo, li., MoleciiUir Ihermodxiuimics of

I'liiul Phase i'ljiiilibnn, 2nd ed.. Prentice Hall, New Jersey, 1986.

Reitl, R.C., Piansnil/, J.M., and Poling, B.Li.. Fhe Properties of Cases and Liquids, 4th ed.,

McGraw Hill, New York, 1987.
2. THERMODYNAMIC PROPERTIES OF ALTERNATIVE AGENTS

Shankland, I.R., Basu, R.S., and Wilson, D.P., "Thermophysical Properties of HCFC-124: An

Environmentally Acceptable Refrigerant," ASHRAE Transactions 96, 317 (1990).

Sladkov, I.B., and Bogacheva, A.V., "Critical Parameters of Mixed Carbon Halides," Zhurnal

Prikladnoi Khimii 64, 2435 (1991); 64, 2276 (1992) (English translation).

Smith, J.M., and Van Ness, H.C., Introduction to Chemical Engineering Thermodynamics, 3rd

ed., McGraw-Hill. New York, 1975.

Walas, S.M., Phase Equilibrium in Chemical Engineering, Butterworth, Boston, 1985.

Yaws, C.L., Ni, H.M., and Chiang, P.Y., "Heat Capacities for 700 Compounds," Chemical

Engineering 95(7), 91 (1988). -
I
37

3. FLUID DYNAMICS OF AGENT DISCHARGE

William M. Pitts, Jiann C. Yang, Grzegorz Gmurczyk,

Leonard Y. Cooper, William L. Grosshandler

Building and Fire Research Laboratory

William G. Cleveland, Cary Presser Chemical Science and Technology Laboratory

3.1 Introduction

The extinguishment of a fire using gaseous agents is a very complicated process which is not completely understood. Current fire-fighting agents such as halon 1301 and halon 1211 are believed

to function by a combination of chemical (catalytic removal of hydrogen atoms at the flame front due

to the presence of bromine atoms) and physical (cooling and dilution of flame gases) actions. All of

the proposed alternative agents are known to be less efl'ective (i.e., considerably higher molar

concentrations of the agent are required) than halon 1301. This reduction in efl'ectiveness is attributed

to the absence of bromine atoms in these chemicals and thus the absence of a highly effective

chemical means of fire extinguishment.

An aspect of fire extinguishment using halons which has not received a great deal of attention is the role that the dispersion and evaporation characteristics of the agent play in fire-fighting efl'ective- ness. The reason for this lack of interest in the past was that these properties have been nearly

optimized by the use of halon 1301 for total flooding applications and halon 1211 for spray applications. With regard to total flooding applications, which are of most interest for this study, the low

boiling point of halon 1301 has ensured that, for the majority of applications, it can be dispersed and

vaporized very rapidly into a space. Design of total flooding systems generally have focused on the use and placement of hardware designed to ensure all areas of an enclosure quickly receive concentrations of the agent sufficient to extinguish the fire. The fire fighting applications-nacelle and dry bay fires on aircraft-which are the focus of the current investigation have different requirements with respect to these agent characteristics. Due to the nature of dry bay fires, detection and extinguishment is generally required within tens of

milliseconds. On the other hand, nacelle fires tend to have burned for longer periods before

extinguishment and a premium is placed on the achievement of extinguishing concentrations which

can be maintained for a period of time following extinguishment to prevent relight on heated surfaces. Effective dispersion throughout the volume is required, but the dispersion can take place over longer

periods of time than for the dry-bay application. Due to the more stringent requirements for dry bay

fire extinguishment, much of the work which follows has focused on rapid release and dispersion of

agents. It should be kept in mind, however, that the findings also have applications for nacelle fire- fighting.

Even though the physical mixing processes involved are not generally considered in designs of

halon fire-fighting systems, they are crucial. The important processes can be characterized into three

major categories-release rate, dispersion and mixing, and evaporation. The release rate determines

the minimum time required to extinguish a fire. If the release is very slow it is possible that the fire may never be extinguished.
38 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

The release rate for an agent also generally, but not always, affects the subsequent dispersion and

mixing behavior of the agent. Much of the dispersion and mixing is the result of the momentum

imparted to the agent by its release. If sufficiently high, this momentum causes the development of

turbulent flows which can effectively transport the released agent throughout a volume.

In order to be fully effective, gaseous agents must interact with the flame as a gas. All of the

alternative agents considered here, with the exception of FC-116, are liquids when stored under

pressure at room temperature. As will be shown shortly, the room-temperature agents are released

from pressurized bottles into ambient atmospheres as superheated liquids, i.e., the liquid has a vapor

pressure which is above the ambient pressure. Therefore, the vaporization behavior of these liquids

following release becomes an important characteristic for fire fighting effectiveness.

This chapter discusses experimental and modeling studies which were designed to rank the

proposed alternative agents with respect to release rate, dispersion and mixing, and evaporation. The

following section (Section 3.2) describes the development of a model for predicting the release

behavior of pressurized liquids from a containment vessel. This effort provided guidance for the

design of an experimental system for investigating the release behavior and subsequent mixing and

evaporation of the proposed alternative agents and halon 1301. The vessel design is discussed in Section 3.3. Experimental characterizations of the release rates from the vessel and characterization

of vessel interior conditions are summarized in Section 3.4 for all of the agents and halon 1301.

Section 3.5 describes the development and use of an experimental system for characterizing the

dispersion and evaporation behavior of the alternative agents and halon 1301 following release from

the pressurized vessel. During the course of this study initial attempts were made to modify existing

computer models for two-phase evaporating flows to allow computation of agent release behaviors.

These efforts are discussed in Section 3.6. The final section of the report summarizes the major

findings of the chapter and provides recommendations for future work as well as a ranking for the

alternative agents based solely on their release rate, dispersion and mixing, and evaporation behaviors.

3.2 Model of Discharging Vessel

3.2.1 The Problem and the Objective. This work formulates a mathematical model to simulate the

discharge of halon and halon-alternative fire extinguishment agents from Nj-pressurized vessels. The

objective is to develop a mathematical model which simulates agent-discharge experiments.

The model is expected to have three applications. First, it was used to guide the development of

the experimental design and procedure (see Section 3.3) which closely simulates discharge of field- deployed vessels while allowing for acquisition of data, including high speed photography, to characterize adequately the discharge process. Second, the model will be used to evaluate the

discharge characteristics of a wide range of alternative-agent/pressure-vessel configurations, thereby

extending the slow and relatively costly experimental method of making such evaluations. Finally, it will be used to determine the discharge vessel exit-flow conditions for use in the simulation of agent

dispersal outside of the vessel. [See Section 3.6.3 and Cooper (1993b).] After presenting the mathematical model, this section will include results of example calculations which addressed the first of

these applications. Descriptions of the model have been presented previously elsewhere (Cooper,

1993a).

The analysis is based on the experimental arrangement depicted in Figure 1

.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 39

HIGH PRESSURE N 2 HOLDING TANK

Figure 1. The experimental arrangement.
40 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

3.2.2 The Experiment and the Model Assumptions

3.2.2.1 The Experimental Arrangement. Refer to Figure 1 . This represents generically the

arrangement associated with the downward discharge experiments which were carried out at NIST.

The Figure 1 arrangement can also be used to describe the phenomena of discharge from vessels

under field conditions.

The experimental arrangement involves a right cylinder discharge vessel at pressure Pj^y, which

contains N2-pressurized test agent, and a holding tank filled with mass Mfjj-j^2 of at pressure f^.

The discharge vessel is of height Zj^y and cross-sectional area A^y. The volume of the holding taiik

is VfjT.

The holding tank is connected via an orifice of area Aq to the discharge vessel. It is assumed

that the flow path through the orifice can be opened or closed with a relatively fast-acting solenoid

valve. If the path is open it is assumed that Pf^^ — ^dV' always from the holding

tank to the discharge vessel. In the case of a field-deployed system there is no holding tank and Aq is zero. At the bottom of the vessel is a short-nozzle/orifice-type opening (note that the containment

vessel exit will be referred to as the "nozzle/orifice" for the purposes of this section and Section

3.6.3) of area A^. The agent will be discharged through this to the outside ambient environment

which is at pressure P^^b ^dv- discharge flow path is originally closed ofl" by a cap or diaphragm. When the cap is removed, the test agent liquid is driven out of the discharge vessel by

virtue of the cross-nozzle/orifice pressure difl'erence.

3.2.2.2 The Procedure Prior to An Experimental Run or a Field-Deployed Discharge. The

orifice and nozzle/orifice flow paths are closed and the discharge vessel is evacuated. The vessel is then filled completely with a known mass of test agent. There will be a volume of liquid agent below

and a volume of gaseous agent above and the pressure will be Psat^ the saturation pressure at the

agent temperature. The vessel is then pressurized with N2. This flows into the vessel from the

holding tank or from some other relatively high-pressure N2 source. In general, the upper gas

volume is now a two-component mixture of N2 and test agent gas. Although some N2 may be

dissolved in the liquid volume, it is assumed that the amount is always so small that throughout

subsequent sequential (additional) pressurization and discharge processes to be studied here the

properties of the liquid in the discharge vessel are well approximated by the liquid properties of the

pure test agent.

Now consider the system at the time, r = 0, when an experimental run or a field-deployed

discharge is initiated. The liquid/gas interface is a distance Zj above the bottom of the vessel. The

mass of liquid in the vessel is Mj^y^^ (the subscript refers to that portion of Discharge Vessel test Agent in the Liquid state) and, throughout the experimental run or field-deployed discharge, the

temperature and density of the liquid that remains in the vessel will be assumed to maintain their

respective initial values, T^^ and p^^ = Psat(^aiJ- indicated, p^^ is estimated to be the density of

saturated test agent liquid at the saturation temperature T^^. The gas mixture is at an assumed

uniform initial temperature Tj^y j and the initial masses of gaseous test agent and N2 in the discharge

vessel are M^^y^Q and M^y^2 7' respectively. Note that the prepressurization process, prior to r =

0, may have been so rapid that the liquid and gas are not in thermal equilibrium with each other. For

this reason, Tj^yj is not necessarily identical to T^^. The discharge vessel and the holding tank are at the uniform pressures P^yj and Pffj- j, respectively.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 41

3.2.2.3 During an Experimental Run or Field-Deployed Discharge. During discharge of a Figure 1-type system, the flow path through the orifice is either open or closed at r = 0"^. When the

orifice is closed the discharge will simulate discharge of a field-deployed system which does not

involve a holding tank. When the orifice is open the discharge will simulate discharge of the

experimental system where additional N2 pressurization during the discharge is provided by the

holding tank. The flow path through the nozzle/orifice at the exit of the discharge vessel is either

opened at r = O"*" (as in the field-deployed system, which uses an explosive cap to initiate discharge

from an equilibrium state at the specified P^yj) or else it opens if and when Pj^y rises to some

specified diaphragm rupture/burst pressure, Pqjjj^^ < Phtj ^^e experimental system). The

time when this latter flow path is opened is designated as t^yj^^.

3.2.2.4 Further Assumptions. For times and conditions of interest, it is assumed that as long

as liquid remains in the vessel Pj^y > Psat^^aO^ ^^'^ there is no possibility of flashing {i.e.,

spontaneous change from a thermodynamically unstable liquid state to an equilibrium, two-phase,

liquid/gas state) in the liquid volume. Future work will address the problem of removing this latter constraint from the analysis. It is assumed that any N2 dissolved in the liquid that may come out of

solution during the discharge process is negligible.

The purpose of this work is to mathematically model the state of the system at any r > 0 up to Iq, where is the smaller of 1) the time when the discharge vessel is emptied of test-agent liquid,

and 2) the time when P^y is reduced to Psat^ai)- experimental systems of interest here, tgjjj^^

is expected to be of the order of 10 s and the time interval of the discharge process, (to'^burst)-

the order of lO"'^ s. Prior to tj^ there is no gas mixture discharge from the vessel. The mass of

gaseous agent in the vessel therefore remains constant at its initial value M^y^Q.

In view of the relatively short times of interest it is assumed that during an experimental run or

field-deployed discharge there is no heat or mass transfer across the liquid/gas interface. It is also

assumed that there is no heat or mass transfer across the interface that contains the nitrogen and

gaseous test agent in the combined holding-tank/discharge-vessel system, e.g., there is negligible heat

transfer to the walls of the vessel and holding tank. This is the basis of the assumption that both

^DV.AG ^"^^ vc\2i^% of N2 in the combined system, designated as M^2- constant for all t, 0

< t < t^. M^2 determined from the initial conditions.

It is assumed that the gaseous test agent and the N2 can be modeled as perfect gases with

constant specific heats and gas constants, Cy^Q Cp^Q, R^q and Cyjj2 ^p,n2- ^N2- values of the

specific heats, which depend on temperature, are taken to be those which correspond to Tj^yj.

3.2.3 The Model Equations. Together with specified parameters, the unknown time-dependent

variables of the problem that define the state of system at an arbitrary time are: P^^j, P^y, M^j^2-

^DVN2' ^DVAL- ^HT- ^DV' ^^'^ ^ Known/spccificd parameters would include: the initial values of

these variables, P^j, Ppyj, Mf^j^2,i- ^DV.Ni.i- ^dv,al,1' '^htj- T^dv.i; ^1- respectively; the

geometric parameters of the system, Aq, A^^, Aj^y, Z^^y, and V^^ material properties of the test agent

and of N2; the constant value Mj^y^Q, and nozzle/orifice discharge coefficients, to be introduced

below. The equations governing the variables are:
42 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Equations of State in the Holding Tank and in the Discharge Vessel and the Law of Partial Pressure in the Discharge Vessel :

( M \ HT V

V ^ HT I

(1)

and

DV (2)

Conservation of Mass of N^:

^Dv,m + ^HT.N2 = constant = M^^ = M^^^^.i ^HT,m,\ (3)

Relation of Z to Mj^y ^^i. MDV,AL

'AL

Reversible Adiabatic Expansion for in Holding Tank:

(4)

' P C R

= constant = ; y = —— = 1 ^^HT,N2,0''

c c

(5)

First Law of Thermodynamics for Entire Gas Svstem :

^\-i^DV,AG^VAG ^DKN2^V,N2^'^DV ^HT,N2^V^Nl'^HT^ ^ p A ^

dt

'^DV^DV^^ (6)
SECTION 3. FLUID DYNAMICS

Flow Across the Orifice:

OF AGENT DISCHARGE 43

dMHT,N2

dt

CAP

P

^ DV

(Y - 1)

(Y + 1)

Y/(Y - 1)

(Y + 1)/(Y - 1) 1/2

r A p

DV

HT

1/Y

P (y - 1)/Y 2y 1 - ^ DV

P^ HT

(Y

-

{i.e., choked flow)

1/2

DV

Y/(Y - 1) {i.e., un-choked flow)

(7)

where q is the (compressible) flow coefficient for the orifice (Shapiro, 1953).

Flow Across the Nozzle Orifice:

dMDV,AL

dt - Co^M'^PAPdv - Psp)T ^fPAMB < PSP

(8)

where C^, is the flow coefficient for the nozzle/orifice and P^p is defined below. Assuming that A^jlAjyy < < 1 in discharge vessels of interest here, in Equation (8) the kinetic energy of the liquid

upstream of the exit nozzle is neglected compared to the kinetic energy at. and immediately downstream of the nozzle.

3.2.4 Definition of P^p and Comments on Equation (8). In Equation (8) it is assumed that the

liquid in the discharge vessel, at state [7". P] = [T^j^, Pj^y > Psat^aiJ^- ^^^'^ through the

exit nozzle/orifice while moving along state paths of constant entropy. When the pressure of the

liquid drops below its saturation pressure, the liquid is assumed to initiate its movement into the

"vapor dome" as a metastable superheated liquid.

For the generic pure material, the relevant P-V diagram which depicts the metastable states is sketched in Figure 2. This includes the region of metastable superheated liquid states and the

corresponding region of metastable subcooled vapor states. For a pure material the metastable liquid or vapor state can only be maintained where (dP/dV)\j- < 0 (Model! and Reid. 1983), i.e., only

outside the locus of points, referred to as the spinoidal curve, where {dP/dV)\j- = 0. The spinoidal

curve, sketched in Figure 2, passes through the critical state and has the superheated liquid region to

its left and the subcooled vapor region to its right.

At any time during the discharge. P^p is defined as the particular pressure along the liquid-leg of

the spinoidal curve of the test agent associated with the intersection of the spinoidal curve and th ?
44 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 2. P-V diagram for a generic material; the spinoidal curve and regions of metastable

liquid and vapor.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 45

above-mentioned, instantaneous, constant-entropy paths. Since states along the spinoidal curve are

states of unstable equilibrium where spontaneous nucleation will occur (Modell and Reid, 1983;

Kim-E and Reid, 1983). i.e., the constant-entropy metastable-state path of the liquid cannot be

sustained at the state defined by the intersection point, it is conjectured that if P^p > P^^g violent

flashing of the liquid jet occurs where and when its pressure reaches P^p. This would be upstream of

a physically unachievable, i.e., unstable, vena contracta at pressure P^j^g.

If Pamb achieved prior to the P^p state then the top portion of Equation (8) is used. For this circumstance the constant-entropy state of the possibly metastable superheated liquid at P^j^p is a physically admissible endpoint to the liquid-jet development process. The situation is consistent with

the idea that: 1) the endpoint state coincides with the vena contracta of the jet; 2) the jet is convected

for some distance downstream of this endpoint without substantial changes and at the vena contracta diameter; and 3) the jet eventually breaks apart due to fluid-dynamic and/or thermodynamic (flashing)

instabilities. Note that the upper Equation (8)-description of the nozzle flow rate of a superheated

liquid as an incompressible, nonflashing fluid, is consistent with results reported in the literature, e.g., (Benjamin and Miller, 1941) for flow through sharp-edge orifices and short nozzles.

If P^p is achieved prior to P^j^p then flashing of the metastable liquid will be initiated immediately downstream of the position in space where the intercept occurs. In the latter case the incompressible flow calculation methodology is still applicable at and upstream of the spinoidal-curve

intercept point. This is indicated by the use of the bottom portion of Equation (8). The above discussion is illustrated with CO2. For stable and metastable liquid CO2, a sketch

based on a drawing by Kim-E (1981, see his Figure 4) of constant-entropy paths in a P-T diagram is presented in Figure 3. In preparing his figure, Kim-E used the Peng-Robinson (1976) equation of

state to describe the metastable liquid CO2.

First assume that liquid CO2 is in a Figure 1-type discharge vessel at Pp^y =70 x 10^ Pa and = 288 K. In Figure 3 this initial state is seen to lie on the s = s2 constant-entropy line, which

has no positive P intercept with the spinoidal curve. As the liquid CO2 flows toward and then out of

the vessel's nozzle/orifice its state is assumed to move downward along the s2 curve of Figure 3. As

can be seen, the P = Py^j^ = 1 x 10^ Pa intercept on this curve, corresponding to T = 277 K,

represents an achievable metastable liquid state for the material. This is the metastable liquid state that is predicted by the present model. It would be expected at, and for some distance downstream

of, the jet's vena contracta. Now assume that liquid CO2 in a discharge vessel is at f* = Pp)y = 60 X 10^ Pa and T = T^p = 291 K. In Figure 3 this initial state is seen to lie on the s = s3 constant-entropy line, which

intersects the spinoidal curve at approximately P^p = 10 x 10^ Pa > > P^n^g- In this case, as the

liquid CO2 flows through and out of the vessel nozzle/orifice the model predicts that it will flash explosively when P = P^p. This will occur upstream of the position at which a fluid jet vena

contracta would otherwise occur; well within a distance of one nozzle/orifice diameter downstream of

the vessel exit. This nozzle/orifice flow model is consistent with a discharge process involving a smoothly timevarying Pp,y, a constant T^p, and an abrupt change in nozzle/orifice exit flow from a simple

incompressible-fluid-jet type flow to an explosively flashing two-phase flow. Note that the idea of

using the condition of the existence of P^p > P/^j^g as a criterion for violent flashing of liquids

flowing through sharp-edged orifices or very short nozzle-like openings does not seem to have been

proposed previously. Nor has it been validated quantitatively.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 47

3.2.5 Reduced Dimensionless Equation Set - The Initial Value Problem. Equations (l)-(8) are made dimensionless by introducing a dimensionless time, t, dimensionless mass of and liquid

agent in the discharge vessel, X2 and x^, respectively, and a combined variable, x^.

(1 + X^^)Pj^y

(1 - MDV.N2_ MDV.AL MN2 M

IPDVA

1/2 DV,ALA

(9)

^DV,AG ^AG

HT.l DVA AG 1) MHT.N2A MN2

(Y - 1) K

= (Y AG HT

(Y - l)ZA

(10)

The initial values of the x's are functions of the various parameters of the particular problem of

interest. The dimensionless initial value problem for an arbitrary Figure 1 -configuration is

dx. j7

o^; Xj(0) specified, f = 1, 2, 3; (11)

= o^(x^; X,; x^; parameters of the problem)

The reader is referred to Section 3.2.13 (Appendix) for definitions of the a- and for explicit equations

to extract values of time-dependent dimensional variables from a solution of Equation (11).

3.2.6 N2 Jet-Driven Mixing in the Discharge Vessel - Estimating the Resulting Disturbance of

the Liquid/Gas Interface. The above model assumes that throughout the pressurization and

discharge process the agent gas and the N2 in the discharge vessel are fully mixed and in a state of

thermodynamic equilibrium. It is also assumed that there is no significant heat or mass transfer interactions at the gas/liquid interface, i.e., it is assumed that the interface is relatively quiescent.

In cases where there is no N2 flow from the holding tank (e.g., the orifice is closed) the initial fully-mixed state of the gases will persist throughout depressurization, the gas volume will be

relatively quiescent, and the assumption of negligible interface interactions is expected to hold. This

is the case for actual field-deployed systems.

When there is N2 flow from the holding tank, it is the N2 jet from the orifice that will drive gas mixing in the discharge vessel. Refer to Figure 4. If jet velocities are too large, the liquid/gas

interface disturbance of the jet can be violent enough to invalidate the quiescent interface assumption.

For the configuration of Figure 1 and Figure 4, the significance of the eflFect of the N2 jet

impinging on the liquid surface can be determined from an estimate of the axial velocity of the orifice

jet at the liquid/gas interface. Such an estimate is obtained here from the characteristics of an

incompressible submerged jet (i.e., analogous to the N2 orifice jet) in an unconfined space (analogous

to the gas volume of the discharge vessel). The estimate is also useful for other orifice orientations.
Figure 4. Sketch of the orifice-driven N2 jet flow in the discharge vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 49

Let the velocity on the jet axis a distance Zj^ = Zj-^y - Z from the orifice {i.e., at the elevation

of the liquid/gas interface) be denoted by f/y^^-and assume a uniform velocity, Uq, across the orifice opening,

dMDV,N2

dt

^O^HT,N2

(12)

From Abramowitz (1963),

HT

UJET 0.96

0.132Z

for

•JET + 0.29

^JET > 5

(13)

When the time-dependent solution for MQyj^2 available, Uj^j can now be estimated from

Equations (12) and (13) and the significance of the jetting phenomenon can be assessed.

3.2.7 Solving the Model Equations. A computer program was developed to solve the initial value

problem of Equation (11) corresponding to an arbitrary choice of geometric parameters, material

properties, and initial conditions, and to determine the value of Uj^jixom Equations (12) and (13).

The method of solution is based on the differential equation solver RKQC presented by Press (1986).

The program was used in the example calculations to follow.

3.2.8 Example Calculations

3.2.8.1 An Experimental Procedure to Simulate the Discharge of Field-Deployed Vessels.

As mentioned in Section 3.2.1, the example calculations to be presented here focused on providing

guidance for the experimental design and procedure such that a vessel discharge of the experimental

system closely simulates the discharge of a field-deployed vessel, while allowing for acquisition of

data, including high-speed photography, to characterize the discharge process. Field-deployed systems use an explosive device to remove, "on demand," a cap covering the exit nozzle/orifice of a prepressurized discharge vessel. Thus, the deployed system is characterized by a Figure 1-type configuration with no holding tank component, i.e., with = 0. The nature of the experimental program at NIST precluded the use of an explosive cap device.

The experimental procedure to be evaluated therefore involved a vessel discharge process which was

initiated, instead, by the rupture of a nozzle/orifice diaphragm cap at a high cross-diaphragm pressure

difference. The pressurization of the discharge vessel from P^y j to the diaphragm burst pressure, PbursT' was assumed to be achieved with the use of the holding-tank/orifice-flow feature of Figure 1. As discussed earlier, an experimental run involves sequential processes of vessel pressurization

and discharge. Initially, the diaphragm cap prevents flow from the vessel to the outside environment.

At r = 0"^ the orifice connecting the vessel and the holding tank is opened and pressurization of the

vessel is initiated. At the instant the pressure in the vessel reaches the diaphragm burst pressiire, the

diaphragm ruptures. Nozzle/orifice flow and discharge of the vessel are then initiated. A.s mentioped
50 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

earlier, the pressurization and discharge processes of interest here are expected to occur over time

intervals of the order of 10 s and 10"- s, respectively.

Prior to any particular test run, Pbur^ is only known to an accuracy of approximately 10%.

Also, the high-speed camera used to photograph the discharge process during the test procedure could

record a total time interval no longer than the order of a few seconds, including an initial interval of

approximately 1 s required to bring the camera up to its operating speed.

3.2.8.2 Criteria for Experimental Dischai^es to Closely Simulate Field-Deployed Dischai^es.

One would hope to determine that an experimental discharge would closely simulate a field-deployed

system discharge by obtaining good agreement between results of model simulations of the two

processes. The basic criterion would be good reproduction of the predicted time-dependent values of

Since field-deployed discharges involve no addition of N2 and no significant heat or mass transfer

at the liquid/gas interface, an acceptable Figure I -type experimental procedure should similarly

involve discharges with no significant addition of N2. In view of this, in addition to the above

criterion, two additional criteria must be satisfied if the experimental discharge is to simulate closely a field-deployed discharge: 1) the experiment must include a discharge process during which the total mass of N-, delivered from the holding vessel to the discharge vessel is small compared to the mass of

Ni in the discharge vessel immediately prior to the onset of discharge, i.e.,

require ~ << \ for t > t^^^j.

where M^-, ^ is the mass of in the discharge vessel at tguj^^\ and 2) the addition of during the

entire pressurization/discharge sequence does not lead to gas flows in the discharge vessel gas (where

velocities will be of the order of Uj^j) which are so vigorous as to lead to significant heat or mass

transfer at the liquid/gas interface. It is reasonable to expect that interface surface interactions will not be significant if Uj^j - dZ/dt never exceeds the order of a few m/s. Thus

require Uj^j - — < 5-10 mjs for t > 0 s (15)

dt

3.2.8.3 Model Input Parameters. Sets of parameters representative of those for a Figure 1- type experimental design were selected to simulate discharge of a field-deployed system with a 0.5 x

10"-^ m^ discharge vessel half-filled with liquid HCFC-22 and pressurized with to 41.4 x 10^ Pa. Parameters of the field-deployed system are included in Table 1

.

The selected parameter sets of Table 1 defined simulations of 18 difi'erent experimental systems

and the single simulation of the field-deployed system. The parameters sets were used to define and

solve the model's initial value problem. Solution results provide time-dependent histories of all model

variables and the information necessary to determine whether or not the criteria of Equations (14) and

(15) are satisfied during the simulation.

As indicated, the model simulations involve the agent HCFC-22. In solving Equations (11) and

(12) the thermodynamic properties listed in Table 2 and Table 3 were used for HCFC-22 and for N2.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 1 . Characteristics of Simulated Discharges - Model Input Parameters

51

agent

material

HCFC-22

^DV,l ^ 294 K

diameter

of dis- charge

vessel = Ddv =

0.05 m

volume of

discharge

vessel — Vdv0.5 X 10-3 m^

initial volume of

liquid

agent in discharge

vessel

= 0.25 X 10"3 m^

0.0191 m

yHT = 2.5 X 10"3 m^

(large holding tank)

2.5 X 10"^ m^

(small holding tank)

field simulation

(no holding tank)

51.7 X 10^ Pa 155 X 10^ Pa

34.5 X 10^ Pa 9.38 X 10^ Pa ^ ^sat(^dv,])

34.5 X 10^ Pa 41.4 X

10^ Pa

37.9 X

10^ Pa

44.8 X

10^ Pa

37.9 X

10^ Pa

44.8 X

10^ Pa

37.9 X

10^ Pa

44.8 X

10^ Pa

41.3 X

10^ Pa

Do =

0.5 mm,

1 mm,

and 5 mm

0.5 mm,

1 mm,

and 5 mm

0.5 mm,

1 mm,

and 5 mm

0.5 mm,

1 mm,

and 5 mm

0.5 mm,

1 mm,

and 5 mm

0.05 mm

1 mm,

and

5 mm
52 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 2. Properties of HCFC-22

Molecular Weight 86.47 kg/mole

Liquid Density [p^^ = p^^j(294 K)] 1209 kg/m^

Constant Pressure Specific Heat [Cp^Q] 57 X 10^ Jkg-^mole-^K-^

Saturation Pressure [Psat^ 6895(T/R)"^ X

[F - (T/R)]^t^ " (T/R)]/[F(T/R)]jqA - B/(T/R) + D(T/R)

where A = 29. 36, B = 3845.1, C = 7.861,

D = 0.002190, E = 305.8, F = 686.1

The following completes the information used to fully define the model equations.

The flow coefficient for the nozzle/orifice (flow of test agent liquid from the discharge vessel to the outside environment) was assigned a value of

C^^ - 0.60. (16)

The use of Equation (8) and the Oj and Oj components of Equation (11) require values of P^p for isentropic paths from the time-dependent liquid states in the discharge vessel. For the present

calculation it is assumed that throughout the discharge process P^p < P^jm ~ ^-^^ ^

spinoidal-curve type instabilities of the metastable exit liquid stream do not occur during the major

portion of the discharge-process time interval. However, as seen above in the discussion of CO2, this assumption can not be expected to hold in general, and actual estimates of P^p will have to be

included in more general applications of the mathematical model.

An expression for the flow coefficient, q, for the flow through a sharp-edged orifice (flow of

N2 from the holding tank to the discharge vessel) as a function of cross-orifice pressure ratio,

Ppyy/PnT^ is taken from Perry (1949). Data points for o(Pdv^^ht) plotted in Figure 5. In the

present calculations, values of Cp^ q are estimated by linear interpolation between these data points.

To determine the constant mass of gaseous test agent in the discharge vessel it is assumed that the process of filling the vessel included a time when pure test agent filled the entire vessel with gas and liquid volumes in thermodynamic equilibrium with each other. The mass of gaseous agent at this time is taken to be Mp^y^Q. In the present calculations it is assumed that at this time of equilibrium

the agent temperature was 294 K. This is the value taken above for the initial temperatures Tp^yj and

3.2.9 Relatively Low-Pressure Large-Volume Holding Tank. This section presents results for the

pressurization/discharge sequence using a relatively low-pressure large-volume holding tank. Note

that experimental safety considerations motivate use of relatively low-pressures in the holding tank,

especially when it is of relatively large volume.

VfPP = 2.5 X 10'-^ m-^ is chosen to be ten times larger than the initial gas volume in the

discharge vessel and Pf^pj = 51.7 x 10^ Pa is chosen to be 6.9 x 10^ Pa greater than the maximum

of the diaphragm burst pressure which can fall within the range 37.9 x 10^ Pa < Ppup^r ^^-^ ^

10^ Pa.
SECTION 3.

Table 3.

FLUID DYNAMICS OF AGENT DISCHARGE

Properties of Nitrogen

53

Molecular Weight 28.01 ;/mole

Constant Pressure Specific Heat [Cp^2] [7.440 - 3.24 X 10-3(T/K) + 6.400 X 10"^ (T/K)2 - 2.790 X 10"^(T/K)3] X 4186 J/(kgmole -K)

The objective in using the present choice of parameters is to provide for a pressurization process

which is slow enough to minimize disturbances of the liquid/gas interface, i.e., satisfy Equation (15),

but fast enough to be able to bracket clearly the time of the subsequent discharge process to a known

time interval of the order of 1 s. The 1 s time interval is required to guarantee timely triggering of

and photographic data acquisition from a high-speed camera during the discharge process. As

mentioned above, the latter is expected to occur over a time interval of the order of 10"^ s. The relatively large volume of the holding tank would allow the pressurization/discharge process

to be initiated either: 1) from a relatively high-P^,^^ state, somewhat below the minimum possible

diaphragm burst pressure and with a relatively significant initial mass of N2 in the discharge vessel; or 2) from a minimum-P^jy;; state where there is no N2 in the discharge vessel (i.e, the discharge

vessel initially contains only pure test agent at the saturation pressure Pj^yj = Psat^ where is significantly below the P^f/^^y- pressures under investigation).

3.2.9.1 Initiating the Pressurization/Discharge Sequence from the Agent's Saturation State; No N2 in the Discharge Vessel at t = 0. Simulations were carried out for experiments where the

pressurization/discharge processes were initiated at the test agent's saturation state, i.e., at P^jv;; =

PsAT^^DVi = 294 K) = 9.38 X 10^ Pa, e.g., shortly after filling the evacuated vessel with the test agent. In such experiments there is no N2 in the discharge vessel at t = 0, i.e., M^yj^jj ~ ^

Results of the calculations are plotted in Figure 6 through Figure 1 1

First the pressurization process will be considered. Figure 6 presents plots of P^y and P^jj

during vessel pressurization, 0 < t < t^m^^- Results are presented for the three orifice diameters,

Dq = 0.005 m, 0.001 m, and 0.0005 m. As seen in Figure 6, for the smallest orifice, Dq = 0.0005 m, the pressurization process, when

^DK7 - ^DV < ^BURsr^ ^ak^s 4.2 s for P^ur^ = 37.9 x 10^ Pa to 5.8 s for Pgyj^^ = 44.8 x 10^

Pa. For the intermediate-size orifice, Dq = 0.001 m, the pressurization process takes 1.0 s for Pg^,j^^ = 37.9 X 10^ Pa to 1.4 s for Pqur^ = 44.8 x 10^ Pa. For the largest orifice, = 0.005

m, the entire process takes place within 0.1 s. Uj^j values corresponding to Figure 6 are presented in Figure 7. For the smallest orifice, Dq =

0.0005 m, Uj^ starts out at approximately 4 m/s and drops to 2 - 3 m/s at tgjjj^^, the time of

diaphragm bursting. For Dq = 0.001 m, t/y^^ starts out at approximately 8 m/s and drops to approximately 4 m/s at t^^j^^. For Dq = 0.005 m, Uj^ starts out at approximately 40 m/s and

drops to as low as approximately 16 m/s for Pgu^^ = 44.8 x 10^ Pa at tguj^^-

Based on the above results it is concluded that during the pressurization process and relative to the criterion of Equation (15): 1) use of the large orifice would lead to conditions which do not satisfy
54 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0.9

0.85

0.8

J0.75

0.7

0.65

0.6

0 0.2 0.4 0.6 0.8

P^^//P DV" HI

Figure 5. Discharge coefficient, q, for compressible flow through a sharp-edged orifice as a function of cross-orifice pressure ratio, Pdv^^ht (Perry, 1949).
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 6. Pressures in the holding tank and discharge vessel for 0 < t < t^uR^ for a larg*

volume, low-pressure holding tank and no N2 in the discharge vessel at / - 0.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 7. Plots of Uj^j in the discharge vessel for 0 < r < tg^j^^ for a large-volume, lowpressure holding tank and no N2 in the discharge vessel at r 0.
3. FLUID DYNAMICS OF AGENT DISCHARGE

57

10^

CD

S- 10

0

N

N

£. 10'

-o

isi -o

If)

£ 10

UJ—3

. dZ/dP

:

UjET

Z/Zi

I

\

\

T

^ * *

*

*

10

4 I 1 L

_1 I 1-

0 0.005 0.01

t " ^URST

0.015 0.02 0.025
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

10'

10

10

0

" 10-^

10

-2

uJET

dZ/dt

Z/Z

/

'(Mn2- Mn2,b)/Mn2,B

_1 I I L. 0 0.005 0.01 0.015 0.02 0.025

^

" ^BURST

Figure 9. Ujgj, dZ/dt, Z/Zj, and (Mn2-Mis42 b)/^N2 b

for t>tBURST' large-volume, lowpressure holding tank; no N2 in vessel at t = 0; Do = 5 mm, Pburst^^-^ MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 59

CD

CD

CM

N

D

isi o

To"

E

UJ—5

10^

10"

10

-1

e 10

-2

10

-3

10

-4

0

z/z

' (Mn2 - Mn2,b)/Mn2,B

dZ/dt

UJET

\

\

0.005 0.01 0.015 0.02 0.025

t-tBURST [s]

Figure 10. Ujej, dZ/dt, Z/Zj, and (Mn2-Mn2,b)/Mn2,b for t^^BURST; large-volume, lowpressure holding tank; no N2 in vessel at t = 0; Do=0.5 mm, Pburst'^4.5 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

GO

N

4—

:o

ND

E

10'

UJET

10^ dZ/dt

100

Z/Z

10

-1

UJ—3 /

(Mn2 - Mn2,b)/Mn2,B

10'

\

\

\

_l I I I

I I I I I II I

I I I I I I [_ 0 0.005 0.01 0.015 0.02 0.025

t-tBURST [S]

Figure 11. Vj^j, dZ/dt, Z/Zj, and (Mn2-Mn2,b)/Mn2,b for t-teuRST' large-volume, lowpressure holding tank; no N2 in vessel at t = 0; D^^S mm, Pburst^'*-^ MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 61

the criterion; 2) use of the intermediate size orifice would lead to results which are barely acceptable;

and 3) use of the small orifice would lead to acceptable performance, provided that pressurization

times are consistent with timing requirements of the photographic data acquisition system.

Now the discharge behavior will be discussed. Figure 8 through Figure 1 1 present plots of

Uj^. dZ/dt, Z/Zj, and (M^2 " ^N2,b)^^N2.b during the discharge process. For the limiting burst

pressures, Pqurst = ^"^-^ ^ ^'^d 44.8 x 10^ Pa, results are presented for the small- and

large-orifice designs.

Figure 8 presents results for Pgyj^^ = 37.9 x 10^ Pa and Dq = 0.0005 m. Once the discharge

process is initiated, the liquid is seen to drop from its initial elevation {Z/Zj = 1) and to be removed

entirely from the vessel (Z/Zj 0) in approximately 0.024 s. Liquid discharge occurs at a relatively

uniform rate with dZ/dt dropping from approximately 7.0 to 4.5 m/s during the emptying process. During discharge Uj^j first rises slightly (because of relatively rapid reductions of P^y, slow

reductions of P^, and resulting increases in cross-orifice pressure difi'erence) from approximately 3 m/s, and then drops (because of the significantly increased length along the jet axis between the source of the jet, at the orifice, and the elevation at the gas/liquid interface - refer to Figure 1 and

Figure 4) to the final value of approximately 2 m/s. As can be seen from the plot of (M]^2 ' ^n2 b)^

M^2 5' during the entire emptying process the total mass addition of N2 from the holding tank to the

discharge vessel is less than 1% of the initial N2 mass in the discharge vessel. Figure 9 presents results for Pguj^^ = 37.9 x 10^ Pa and Dq = 0.005 m. Based on an approximate extrapolation from t - teuR^ ^ 0.020 s, once the discharge process is initiated the liquid

is seen to drop from its initial elevation and to be removed entirely from the vessel in approximately

0.023 s. Liquid discharge occurs at a relatively uniform rate with dZ/dt dropping from approximately

7.0 to 5.5 m/s during the emptying process. During discharge Uj^j first rises slightly from

approximately 30 m/s, and then drops to the final value of approximately 20 m/s. As can be seen from the plot of (M^2 " ^N2 b)^^N2 B' during the entire emptying process the total mass addition of

N2 from the holding tank to the discharge vessel is significant, approximately 45% of the initial N2

mass in the discharge vessel. Figure 10 presents results for Pbur^ = 44.8 x 10^ Pa and Dq = 0.0005 m. Here the

discharge is completed in 0.023 s. Liquid discharge occurs at a relatively uniform rate with dZ/dt

dropping from approximately 7.5 to 5.0 m/s during the emptying process. During discharge Uj^-p

first rises from approximately 1 to 2 m/s, and then drops back to the final value of approximately 1 m/s. During the discharge the total mass addition of N2 from the holding tank to the discharge vessel

is less than 1 % of the initial N2 mass in the discharge vessel. Figure 11 presents results for P^ur^ ^ 44.8 x 10^ Pa and Dq = 0.005 m. From this it is seen

that discharge occurs in approximately 0.021 s. Liquid discharge occurs at a relatively uniform rate with dZ/dt dropping from approximately 7.5 to 5.5 m/s during the emptying process. During

discharge Uj^j- first rises from approximately 17 to 25 m/s, and then drops to the final value of

approximately 20 m/s. During the discharge the total mass addition of N2 from the holding tank to the discharge vessel is approximately 40% of the initial N2 mass in the discharge vessel. Based on the above results it is concluded that during the discharge process and relative to the

criteria of Equations (14) and (15) : 1) use of the large orifice would lead to conditions which violate

both of the criterion; 2) use of the small orifice would lead to acceptable performance.

3.2.9.2 Initiating the Pressurization/Discharge Sequence From a Ppyj Slightly Below

^BURST- Simulations were carried out for experiments initiated from a P^y] which is slightly below

^BURsr- Since the minimum possible P^ur^^^ ^'^•^ ^ ^DV,i selected to be 34.5 x

lO^Pa.
62 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 12 presents plots of P^y and P^j during vessel pressurization. Results are presented for the three orifice diameters, Dq = 0.005 m, 0.001 m, and 0.0005 m. As seen in Figure 12, for the smallest orifice, Dq = 0.0005 m, the pressurization process, P^yj

< Pj^y < PguRsr^ takes 0.6 s for Pguj^^ = 37.9 x 10^ Pa to 1.9 s for Pguj^^ = 44.8 x 10^ Pa.

'

For the intermediate-size orifice, Dq = 0.001 m, the pressurization process takes 0.1 s for Pgu^^ =

37.9 X 10^ Pa to 0.5 s for Pguj^^ = 44.8 x 10^ Pa. For the largest orifice, Dq = 0.005 m, the

entire pressurization process takes place in less than 0.02 s.

[/y£'7' values corresponding to Figure 12 are presented in Figure 13. For the smallest orifice, Dq

= 0.0005 m, Uj^j starts out at approximately 3.5 m/s and drops to 2.0 - 2.5 m/s at the time of

diaphragm rupture. For Dq = 0.001 m, Vj^j starts out at approximately 7 m/s and drops to approximately 4.5 m/s at the time of diaphragm rupture. For Dq = 0.005 m, JJj^j- starts out at approximately 33 m/s and drops to as low as approximately 21 m/s for Pquj^^ ~ 44.8 x 10^ Pa at the time of diaphragm rupture.

Based on the above results it is concluded that during the pressurization process and relative to the criterion of Equation (15), the present results for P^y j = 34.5 x 10^ Pa are similar to the earlier

results for Pj^yj = P^at = 9.38 x 10^ Pa. In the present case where Ppyj = 34.5 x 10^ Pa, results for the discharge process are very

similar to the earlier results where P^yj - Psat = 9.38 x 10^ Pa and are qualitatively well

represented by the plots of Figure 8 through Figure 1 1 and the earlier discussion of these.

3.2.9.3 Comparing the Discharges of a Small-Orifice Test Configuration and a FieldDeployed System. The model equations were also used to simulate the discharge of a field-deployed

vessel filled with HCFC-22 and pressurized with Nj. At the onset of discharge the temperature in the

vessel was taken to be 294 K and, corresponding to the above calculations, the pressure in the vessel was taken to be 41.4 x 10^ Pa, i.e., a nominal Pquj^^ =41.4 x 10^ Pa, where simulating tests would involve experiments where 37.9 x 10^ Pa < Pguj^^ < 44.8 x 10^ Pa. In Figure 14 and

Figure 15 calculated values of P^y and Z/Zj for the deployed system simulation are compared with

the values of these variables obtained in the above test simulations with Dq = 0.0005 m. The Pj^y j = 9.38 X 10^ Pa results are presented in the figures since these compare somewhat less favorably

with the field-deployed system results than do the P^yj = 34.5 x 10^ Pa results. As can be seen in the figures, simulated field-deployed system results and simulated test results compare favorably.

3.2.9.4 Summary of Results of Simulations Involving the Relatively Low-Pressure LargeVolume Holding Tank. A Figure 1-type test configuration with a 0.25 x 10"^ m-^ holding tank and

a 0.0005 m, and possibly a 0.001 m diameter orifice will provide experimental discharges which can be expected to simulate accurately the discharge of field-deployed systems. Of the two orifice sizes,

the smaller orifice is preferable. Use of a 0.005 m diameter orifice can not be expected to adequately

simulate field-deployed system discharges under the conditions studied.

It is expected that this result can be extended to a wide range of test parameters and test agents.

3.2.10 Relatively High-Pressure Small-Volume Holding Tank. This section presents results for the

pressurization/discharge sequence using a relatively high-pressure small-volume holding tank. Vrrp —

2.5 X 10"^ m-^ is chosen to be one tenth of the initial gas volume in the discharge vessel and Pifj- j -

155.1 X 10^ Pa is chosen to be large enough to ensure that the amount of N2 stored in the holding

tank will rupture the diaphragm even when Pgyj^^ is at its maximum value, 44.8 x 10^ Pa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 12. Pressures in the holding tank and discharge vessel for 0<t<tguRST for a largevolume, low-pressure holding tank; Pdv,i = 3.4 MPa; Phti^^-^'^ ^P^' ^burst = 4.5 MPa.
64 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 13. Plots of Uj^t in the discharge vessel for 0 < t ^tguRsr for ^ large-volume, lowpressure holding tank; Pdv,i = 3.4 MPa; Pht,i=5.17 MPa; Pburst = 4.5 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 65

I I

\

I I I 1 I I 1 1

1 1 1 1 1 1 1 L 0 0.005 0.01 0.015 0.02 0.025

Figure 14. Pressure in the vessel during discharges of the field-deployed vessel and test

configuration vessel with a large-volume, low-pressure holding tank and no N2 in

- the discharge vessel at t = 0.
66 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

N

N

Figure 15. Plots of Z/Zj during discharges of the field-deployed vessel and test configuration

vessel with a large-volume, low-pressure holding tank and no N2 in the discharge

vessel at t = 0.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 67

The objective of choosing this combination of parameters is to attain a pressurization/discharge

sequence which is so rapid, well within the above-mentioned 1 s time interval, as to guarantee,

without any significant timing considerations, successful acquisition of high-speed photographic data

of the discharge process. The feature of a relatively small initial volume and mass of N2 in the

holding tank is only consistent with test sequences with relatively high initial Pdv./^^ values close to the minimum possible diaphragm burst pressure. Thus, the present test parameters would not be

expected to be useful in initiating the pressurization/discharge process from an initial state where the

vessel contained only pure test agent at its saturation pressure. Here, P^y j is again chosen to be

34.5 X 10^ Pa, slightly below 37.9 x 10^ Pa. the minimum value for Pburst3,2.10.1 Pressurization and Discharge for the Small OrilBce, Dq = 0.0005 m. Simulation

results for pressurization/discharge test using the small orifice design are presented in Figure 16

through Figure 20.

For Pgui^^ = 37.9 x 10^ Pa, plots of P^y and Pf^j are presented in Figure 16 and plots of

Uj^j, dZ/dt, Z/Zj, and (Mj^2 ' ^N2,b)^^N2.b Figure 17. Note that the latter variable is meaningful

only during discharge when t > tgyj^^. As can be seen in these figures, for this case tg^j^^ =

0.173 s, i.e., this is the time it takes for Pj^iy to rise from Ppyj = 34.5 x 10^ Pa to the diaphragm

rupture pressure. From Figure 17 it is seen that the liquid is nearly all discharged from the vessel

{i.e., Z/Zj is approaching zero) at approximately t = 0.2 s. From Figure 17 it is also seen that during discharge: the rate of liquid outflow is reduced to about half of its original value, with dZ/dt

being reduced from 7.0 m/s to 4.5 m/s; Uj^j- is reduced from an initial predischarge value of

approximately 4 m/s to approximately 2 m/s; and (Mi^2 ~ ^n2 b^^^ni b^^^^^ ^^^m 0 to approximately

0.02, i.e., at the end of the discharge the mass of N2 in the vessel is increased from its value at huRsr^y ^ factor of approximately 0.02. The latter two results indicate that the criteria of Equations

(14) and (15) are both satisfied and, in this sense, as with the relatively large holding tank, the small

orifice design can be expected to adequately reproduce field-deployed discharges.

For Pbur^ =44.8 x 10^ Pa, plots of P^y and P^j are presented in Figure 18 and plots of

Uj^j and dZ/dt are presented in Figure 19. Plots of Uj^j, dZ/dt, Z/Zj, and (Mj^2 ~ ^n2 b^^^ni b

^'^^ the times of the discharge process are highlighted in Figure 20. For this case tgjjj^^ = 0.94 s. From

Figure 18 it is seen that the has almost been reduced to P^y at the time that discharge is initiated. Indeed, if the P^jj had been much less than 155 x 10^ Pa, would not have risen to Pg^j^^ and

the discharge process would never have occurred! From Figure 19 it is seen that during the

pressurization process Uj^j- is reduced from approximately 4.0 m/s to 1.5 m/s. The liquid is nearly

all discharged from the vessel at 0.962 s. During discharge: dZ/dt is reduced from 7.5 m/s to 5.0

m/s; Uj^j rises to approximately 2.5 m/s from its initial value of 1.5 m/s and finally drops to approximately 2 m/s; and (M^2 " ^N2 b^^^N2 b

^^^^^ ^""^^ ^ approximately 0.005. The criteria of

Equations (14) and (15) are again satisfied. As noted, for both PguR^ =37.9 x 10^ Pa and 44.8 x 10^ Pa the above results indicate that both criteria of Equations (14) and (15) are satisfied with the small-holding-tank/small-orifice design

and that, in this sense, this design can be expected to adequately reproduce field-deployed discharges.

However, for the following three reasons the large-holding-tank design is significantly more robust

than the small-holding-tank design: (1) For a given maximum value for PguRsr^'^'^ discharge vessel

volume, the minimum acceptable value for Pf^j j (approximately 155 x 10^ Pa for the present test parameters) is very sensitive to the original amount of liquid in the vessel. For example, in the

present simulations the vessel was assumed to be one-half full; but if it were only one-quarter full, the minimum Pi^j j that would lead to rupture of the 44.8 x 10^ Pa diaphragm would have been

approximately 207 x 10^ Pa. (2) The small-volume-holding tank configuration does not allow for the
68 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

160. h

120,

CO Q80,

>

40,

0

t[s] tBURST=0.173S

Figure 16. Pressures in the holding tank and discharge vessel for t > 0 for a small-volume,

high-pressure holding tank; = 0.5 mm; P^vi = 3.45 MPa; Pht,i~ l^-^

MPa; Pburst =3.79 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 69

CD

CM

CVJ

N

N

ND

(f) E

UJ

—3

10

,0

10

-1

^ 10

-2

10

0

uJET dZ/dt

z/z.

(Mn2 - Mn2,b)/Mn2,B

_l I L. 0.05

J I I L. 0.1

t[s]

0.15

10" — — — - — — - — —

\

/

I

'

I

i

0.2

tBURST=0-1 ' 3s

Figure 17. Uj^. dZ/dt, Z/Z,, and (MN2-MN2,B)/Mfg2.B for small-volume, high-pressure

holding lank; D„=0.5 mm; Pdv,i=3.45 MPa; P„ti = 15.5 MPa; Pburst=3-8

MPa.
70 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 18. Pressures in the holding tank and discharge vessel for t > 0 for a small-volume,

high-pressure holding tank; = 0.5 mm; Pdv,i = 3.45 MPa; Pht,i= ^^-^ MPa; Pburst = 4-48 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 71

Figure 19. Uj£-p and dZ/dt for t > 0 for a small-volume, high-pressure holding tank; =

0.5 mm; Pdv,i = 3.45 MPa; Pht,i = 15.5 MPa; Pburst = 4.48 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

10^

10'

10-^

10

10

0.935

z/z

(Mn2 - Mn2,b)/Mn2,B

dZ/dt

Uj ET

\

\

\

\

_J I II I

I I _j ; I I I I L. 0.945 0.95 0.955 0.96

tBURST=0.940S t[s]

Figure 20. Ujej, dZ/dt, Z/Z^, and (Mn2-Mn2,b)/Mn2,b for small-volume, high-pressure

holding tank; 0^ = 0.5 mm; Pdvi = 3.45 MPa; Phti =16 MPa; Pburst""4.48

MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 73

option of initiating tests from relatively-low values of P^yj, e.g., immediately after filling the vessel

with the test agent, when P^^y = P^^p To do so would once again require highly variable and

unreasonably large values of Phjj. (3) The large-holding-tank design requires Pj^jj values only

slightly greater than the maximum Pq^j^^ value and it suffers from neither of the above shortcomings.

3.2.10.2 Pressurization and Discharge for the Large Orifice, Dq = 0.005 m. Simulation

results for pressurization/discharge test using the large-orifice design are presented in Figure 21

through Figure 24.

For PBURST ^ 37.9 x 10^ Pa, plots of P^y and P^ are presented in Figure 21 and plots of

Uj^, dZ/dt, Z/Zj, and (Mj^2 - ^n2.b)^^N2,b Figure 22. Note that the latter variable is meaningful

only during discharge when t > t^yj^^. As can be seen in these figures, for this case (burst ^

0.00173 s, i.e., this is the time it takes for P^yto rise from Pj^yj = 34.5 x 10^ Pa to the diaphragm

rupmre pressure. From Figure 21 is seen that the P^ has almost been reduced to Pj^y at t = 0.015

s, at which time approximately 40% of the original amount of liquid is still in the vessel, i.e., Z/Zj —

0.4 (Figure 22). Also, from Figure 22 it is seen that the liquid is nearly all discharged from the

vessel (i.e., Z/Zj is approaching zero) at approximately t = 0.023 s. From Figure 22 it is seen that during discharge: the rate of liquid outflow is reduced to about half of its original value, with dZ/dt

going from 8 m/s to 4 m/s; Uj^ is always dropping, being reduced from an initially large

predischarge value of approximately 40 m/s to approximately 4 m/s; and (M^2 ' ^n2 b^^^ni b

'"'^^^ from 0 to 0.35, i.e., at the end of the discharge the mass of N2 in the vessel is increased from its value at tsuRST^y ^ factor of approximately 0.35. The latter two results indicate that the criteria of

Equations (14) and (15) are not satisfied and, as when used in conjunction with a relatively large

holding tank, the large orifice design is not expected to adequately reproduce field-deployed

discharges.

For PBURST ~ ^ 10^ Pa, plots of P^y and Pfjj- are presented in Figure 23 and plots of

Ujgj-, dZ/dt, Z/Zj, and (M^2 ' b^^^ni b

Figure 24. The results are qualitatively similar to those discussed in the last paragraph for Pburst ~ ^^^ ^ P^- quantitative differences can

be easily determined from the figures. For this case t^jjj^^ = 0.0094 s. From Figure 23 is seen that the P^ has almost been reduced to P^y^i the time that discharge is initiated. Indeed, if the Pfjjj

had been much less than 155 x 10^ I^, P^y would not have risen to Pburst discharge

process would never have occurred! The liquid is nearly all discharged from the vessel at 0.03 s. During the pressurization process f/j^y- drops from approximately 40 m/s to 15 m/s. During

discharge: dZ/dt is reduced from 7 m/s to 5 m/s; Uj^ is reduced from 15 m/s to 3.5 m/s; and fM^2 " M^2 b)^^n2 b

""i^^^ ^"com 0 to 0.1. Although the criterion in Equation (14) is satisfied, that of

Equation (15) is not.

3.2.11 Summary and Conclusions. A mathematical model and associated computer program were

developed to simulate the discharge of fire extinguishment agents from N2-pressurized vessels. The

model is expected to have three applications. First, to guide the development of an experimental

procedure which both 1) simulates depressurization of a field-deployed discharge vessel and 2) allows

for acquisition of data, including high speed photography, to characterize adequately the discharge

process. Second, to evaluate the discharge characteristics of a wide range of alternativeagent/pressure-vessel configurations, thereby extending the slow and relatively costly experimental

method of making such evaluations. Finally, to predict exit flow conditions to be used to solve the

problem of agent dispersal outside of the discharge vessel.
74 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 21. Pressures in the holding tank and discharge vessel for t > 0 for a small-volume,

high-pressure holding tank; = 5 mm; Pdvi = 3.45 MPa; Pht,i = 15.5 MPa;

Prtir-^t = 3.79 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 75

Figure 22. Ujet, dZ/dt, Z/Zj, and (Mn2-Mn2,b)/Mn2,b for small-volume, high-pressure

holding tank; 0^ = 5 mm; Pdv,i=3.'45 MPa; Phti = 15.5 MPa; Pburst= 3-79

MPa.
76 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 23. Pressures in the holding tank and discharge vessel for t > 0 for a small-volume,

high-pressure holding tank; = 5 mm; i = 3.45 MPa; P^tj = 15.5 MPa;

Prttrct = 4.48 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 77

10'

CD

cJ"

CD

CM 10^

N

N

E

N

"D

uJET

,0

z/z

10" — —

dZ/dt

COE

LU

10' -1 I I 1_ 0 0.005

/

\

(Mn2 - Mn2,b)/Mn2,B

0.015 0.02 0.025 0.03

tBURST=0.0094s t [s]

Figure 24. Ujgj, dZ/dt, Z/Z^, and (Mn2-Mn2,b)/Mn2,b small-volume, high-pressure

holding tank; 0^ = 5 mm; Pdv,i = 3.'45 MPa; Phti = 16 MPa; Pburst=4.48

MPa.
78 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

The model is based on the experimental configuration depicted in Figure 1 , and the solution

method is capable of treating arbitrary choices of geometric parameters, material properties, and

initial conditions.

The model was used in example calculations which address the first of the applications, viz., the

establishment of a robust experimental design and procedure that would 1) simulate the discharge of

field-deployed fire extinguishment systems useable in aircraft and 2) meet additional experimental

constraints consistent with a program at NIST in support of aircraft fire safety issues. In the example calculations, the fixed parameters which characterized the system were a 0.0005

m-^ discharge vessel with a circular exit nozzle/orifice of diameter 0.01905 m half-filled with

HCFC-22 in the liquid phase at "room temperature," 294 K.

The deployed system was assumed to be pressurized with N2 to a total pressure of 41.4 x 10^

Pa, where the exit nozzle/orifice is closed with an explosive cap which is released "on demand." In

the experimental system, the exit nozzle/orifice is capped with a disk which ruptures at an a prioriunknown pressure, between 37.9 x 10^ Pa and 44.8 x 10^ Pa. The experimental system, is brought to the rupture/burst pressure, P^ur^^ ^om an initial pressure P£)y] by a process of pressurization by flow from the holding tank which is initially filled with N2 at some pressure, Pf^j- j, higher than Pguj^^- The holding tank communicates with the

discharge vessel through a circular orifice of diameter Dq.

The model was used to simulate the discharge of the deployed system and of the experimental

system. Experimental parameters that were varied were Pj^yj (9.38 x 10^ Pa, the saturation

pressure of HCFC-22 at 294 K, or 34.5 x 10^ Pa); P^^j j and the volume of the holding tank (51.7

X 10^ Pa and 2.5 x lO'^ m^ or 155.1 x 10^ Pa and 2.5 x 10"^ m^); and Dg (0.005 m, 0.001 m,

or 0.0005 m).

Results of the simulations are presented in Figure 6 through Figure 24. Based on the model

calculations it was determined that with Pj^y j = 9.38 x 10^ Pa or 34.5 x 10^ Pa an experimental

configuration/design using the relatively low-pressure (51.7 x 10^ Pa), large-volume (2.5 x 10-3

m^), holding tank and the smallest-diameter (0.0005 m) orifice would accurately simulate the

discharge of the field-deployed system. Furthermore, it is expected that such a design is robust in the sense that it would also simulate well the field-deployed system even when extended to a range of

parameters and agent materials well beyond the scope of the present calculations.

Model calculations also indicated that an experimental design which uses the large-diameter

orifice (0.005 m) and/or the relatively high pressure small-volume holding tank (155.1 x 10^ Pa and

2.5 X 10"-^ m^) would often not provide accurate simulations of field-deployed discharges and would

never be expected to be "robust."

3.2.12 Nomenclature

cD.N-' c

c

D.O

C

D

M

M

P,AG^ "-P.NZ V.AG' ^V,N2

O

DV.AG' ^DVAL

DV.N2' MHT.N2

^DV,N2,1^ ^DV.AL,!^

MHT.N2,1

cross-sectional area of discharge vessel area of discharge nozzle/orifice, orifice from holding tank

flow coeflftcient of discharge nozzle/orifice, orifice from holding tank

specific heat at constant pressure of agent gas, N2

specific heat at constant volume of agent gas, N2

diameter of the orifice between discharge vessel and holding tank mass of gaseous, liquid agent in discharge vessel mass of N2 in discharge vessel, holding tank

DVNZ^ Mr'"^DV.AL ^HT,N2 at t 0
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 79

N2

N2.B

AMB

^BURSr

^SAT

AG^ ^^N2 T

'AL

BURST

D

Z

^JEJ

T> lAG

f

Pal

PSAT

h

f^DV.N2^ f^HT.N2

f^DV,N2,]^ l^HT.N2.1 'AMB

TTSP

total mass of N2 in discharge vessel and holding tank, see Equation (3) mass of Nj in discharge vessel when the diaphragm ruptures

pressure

pressure of ambient environment

pressure in discharge vessel when the diaphragm ruptures

pressure in discharge vessel, holding tank

P^y, at t = 0

pressure when liquid agent is at saturation

pressure on agent spinoidal curve

gas constants for gaseous agent, N-, temperature

temperature of liquid agent

temperature of gaseous agent in discharge vessel, N7 in holding tank

T, Tfjj- at t = 0

time from beginning of experiment or from onset of discharge of field- deployed system

time when diaphragm ruptures

time at completion of liquid discharge or when P^y reaches P^^j^ during

discharge

axial velocity of orifice jet upstream of gas/liquid interface

volume of holding tank

variables of dimensionless initial value problem, i = 1, 2, 3, Equation (9)

elevation of gas/liquid interface above bottom of discharge vessel

Z at r - 0

length of discharge vessel Z^y - Z

ratio of specific heats for N2, gaseous agent

dimensionless Z, Equation (A-1)

dimensionless Z^^, Equation (A-4)

dimensionless T^^, Tj^j-, Equation (A-1)

dimensionless Tj-^yj, Tf^j- j. Equation (A-1)

density of liquid agent

density of saturated liquid agent

dimensionless parameters, i = 1 to 5, Equations (10) and (A-4)

dimensionless Mpy^2^ ^htn2' Equation (A-1)

dimensionless M^yj^j b ^ht.ni h

Equation (A-1)

dimensionless P^j^g, Equation (A-4)

dimensionless P^, Pj^y, P^jj. Equation (A-1)

dimensionless P^p, Equation (A-4)

functions of and parameters, i = 1, 2, 3, Equation (11)

dimensionless t. Equations (9) and (A-1)

3.2.13 Appendix: The Dimensionless Initial Value Problem. Define the following dimensionless

variables and their initial values:
80 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

^ = tCDNAN[2PDv,l/(MDv,AL,lADvZi)]''2

^HT - Pht'^PdvJ' ^ht^'^ = 0) = tthji - Phti^^dv,! ^ 1

ttdv = Pdv/Pdv,i; '^dv(^ = 0) = 1

A'HT.Nl = Mht,N2/^N2' MhT,N2('^ = 0) = Mj^j i^2,l/'^N2 = /^HT,N2,1 = ^ " MdV,N2,1 - ^

MDV,N2 = '^DV,N2'''^N2; MdV,N2('^ = 0) = Mi5v_jvJ2,i/Mn2 = MdV,N2,1 ^ ^ " /^HT,N2,1 - ^

Mdv,al = Mdv,al/Mdv,al,i = Mdv,al/(palAdvZi); /^dv,al(^ = 0) = 1 (A-1)

^HT " ThtRn2^N2/(^htPdv,i) = (Tht''Tht,i)(Pht,i/Pdv,i V(Mht,n2,i/'^n2)'

6^j(t = 0) = ^HT.i = (Pht,i/Pdv,i)/(Mht,n2,i/Mn2) > 1

^Dv = TDyMpY^^QRy^^Q/CZjAj^yPpYj) = Tj)vMi3v,agP-ag^al''('^dv,al,iPdv,i)'

^Dv(^ = 0) = ^DY i = Tj)Y iMQy^QR^Q/(ZiAf)yPpY i)

f = Z/Zi; f(r = 0) = 1

Define the new variable Xj and introduce X2 and X3 as new designations for fijyy^j /^dvaL'

respectively:

Xi(t) = [(1 + X3MdV,N2)/(1 + ^lMDV,N2)]7rDv(fDV " ^DV.Al) + ^2^1 " MDV,N2)^;

Xj(0) = [(1 + X3]liDYN2 + ^iMdVN2 l)](rDV " 1) + ^2X4(1 " /^DV N2 l)"^ (A-2)

^2^^) = ^DV,N2' ^2(0) ^ MdV,N2,1

x3(t) = Mdv,aL' ^3(0) ^ Mdv,al,i

Define also the following functions of the xf.

fl = [Xi - X2X4(1 - X2F](1 + X,X2)/[(1 + X3X2)(rDV " ^3)];

(A-3)

f2 = X2(l - X2)(^ - f3 = X2(l - X2)y; f4 = X2(l - X2r/fi;
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

X5f3{7[2/(7 + 1)](T + l)/(7- l)/f2}l/2. if > + l)/2][7/(7-l)]

81

f6(Xi)-

X5f3(i/f4)i/T'{27[i - {i/0y-^^'m{y - DfJ}^'^; if f4 < [(7 + i)/2][^/(^-i)i

(fi " '"'amb)^^'^' if ""amb - """sp

(fj - TTsp)^'^; if TT^MB < "^SP

where

Xj = (Mn2/Mdv,Ag)(I^N2/Rag); ^2 = ^HT.i/(MhTN2,i)^;

>^3 = (Cv,N2/Cv,ag)(Mn2/Mdv.Ag) = \(yAG " l)/(7 ' D;

X4 = [(Tag - 1)/(t - 1)][Vht/(ZiAdv)];

X5 = (Cd_o/Cd.n)(Ao/An)[(Mdv,al/Mn2)(AdvZi/Vht)/2]1^2.

fov = Zdv/Zi = constant > I; y^^ = Cpag/Cv,ag = Rag/Cv,ag + 1;

^amb ^ Pamb^Pdv.^ ^sp = Psp/Pov.i

Then the Xj can be determined from the solution of the following initial value problem

dxj/dr = - (7AG - l)fif6; dxj/dr = f^; dx^fdr = - fg

Xi(0) = [(1 + X3MdV,N2,iV(1 + XlMDV,N2,l)](rDV " 1) + ^iM^ " MdV,N2,i)^;

X2(0) = MdV,N2,1' ^3^0) ^ MdV,ALJ

and the solutions for the original dimensionless variables of Equation (A-1) can be retrieved from

(A-4)

(A-5)

""dV - '"'HT - A'HT,N2 - 1 - ^2^ MdV,N2 " ^2' /^DV,AL " f " ^3^

^HT = h'^ ^DV = (rov - X3)fl/(1 + h^2)

(A-6)

From the latter solutions the dimensional variables can be determined finally from the definitions of

Equation (A-1).

The problem formulation of Equation (A-5) and its solution are valid up to the value of t where

X3 = 0 (the last of the liquid is driven from the discharge vessel) or fj is reduced to Psat^Pdv,! (^^^
82 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

liquid in the discharge vessel is subjected to a condition where an equilibrium state would lead to flashing).

3.3 Design of Experimental Release Vessel

3.3.1 Introduction. In a conventional halon 1301 bottle, a fixed amount of agent is initially

dispensed to the bottle followed by pressurization with nitrogen to a fixed equilibrium pressure at room temperature. Depending upon the application, the amount of agent and the charged pressure can be varied accordingly. Release of the agent from the vessel is normally achieved by a squib or a solenoid valve, either manually or automatically, upon the detection of a fire. For fire protection of

dry bays, the agent has to be released from the vessel and suppress the fire in less than 50 ms. The

use of a squib or solenoid valve to study discharge dynamics in a laboratory is not appropriate

because a squib is an explosive device and the flow path through a solenoid valve is too complex to analyze. In order to obtain fundamental understanding of the discharge processes, an experimental

apparatus has to be designed such that it can eliminate the above shortcomings of using either a squib

or a solenoid valve and at the same time closely simulate the actual discharge process. There were two major design criteria that had to be considered: vessel size and release mechanism.

Since the experimental condition was at high pressure, the size of the vessel had to be small enough

to minimize weight for ease of mounting and handling of the vessel, but not so small that the size of

the vessel became unrealistic and that the characteristic emptying time of the agent was so short that it was not possible to characterize the discharge process in sufficient detail. In addition, minimization

of the amount of agents used in the discharge tests had to be considered because of the limited

availability and high cost of some of the alternative agents. An internal volume of approximately 5 x

lO""* m-' was selected. This volume was less than a factor of three smaller than the internal volume

(14 X lO^"* m-^) of the smallest vessel (type CF-2) specified in the appropriate Military Specification

(1965). In addition, this volume was large enough to hold a sufficient amount of agent to perform the

discharge experiments. An aspect ratio (length/diameter) of about 5 was used in the vessel design.

The selection of this aspect ratio was governed by the dimension of the rupture disk holder (to be

discussed).

Since the intent was not to design an actual release mechanism for practical applications, but

rather to design a release mechanism that could be used to compare the discharge dynamics of various

alternative agents, the device had to have a simple geometry, a simple flow pattern, and quick action.

Due to safety consideration in the laboratory, release mechanisms that utilized explosive cartridges to

initiate the discharge of the vessel were not considered in the design of the discharge vessel. Solenoid

valves were also not considered because of the complex flow path inside the valves (although some

valves do have a straight-through design) and a slow response time compared to the emptying time of

the vessel. Furthermore, solenoid valves are not designed to operate under extreme high and low

temperatures. A rupture disk was chosen as a quick release mechanism because the bursting of the

disk is almost instantaneous once a specified pressure has been reached and a simple straight-through

and full opening of the non-fragmented burst disk can be obtained. This technique had been

successfully used previously in a study of rapid venting of hot, high pressure liquids from a pressure

vessel (Kim-E, 1981; Kim-E and Reid, 1983).

3.3.2 Release Vessel Design. A rupture disk (FDI FA scored) with a diameter of 19.1 mm was

used as a rapid release mechanism for discharging agents from the pressurized vessel. The disk was
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 83

made of stainless steel (SS 316) with a nominal burst pressure of 4.12 MPa at room temperature

(22 °C). The acmal burst pressure varied less than + 10% of the nominal burst pressure. The vessel used in this study was made of either stainless steel or clear acrylic plastic (polymethyl methacrylate). The clear acrylic vessel was used for visualizing the internal behavior of the

agents during discharge and for obtaining emptying rates of the agents from the vessel. The stainless steel (SS 304) vessel was constructed from a tube with an internal diameter of 50.8

mm and a wall thickness of 6.4 mm. One end of the tube was welded with an end plate in which

there were four tapped access holes for mounting a piezoelectric dynamic pressure transducer (Kistler

Model 603B1), two thermocouples (K-type), and a needle valve (Whitey SS-1RS4) for dispensing the

agent into the vessel and for subsequent nitrogen pressurization. The other end of the tube was

welded to a flange with an 0-ring seal for mounting a rupture disk holder (FDI Model FF). A tapped

hole was also made on the lower side wall of the vessel so that a pressure transducer could be

mounted to monitor the pressure of the liquid during downward discharge. The internal volume of

the vessel was measured by filling the vessel with water and weighing and was found to be 4.97 x

10""* m^. The vessel was hydrostatically tested to 17.2 MPa. However, for safety reasons, the

vessel was operated at pressures less than or equal to 8.2 MPa. Figure 25 is a schematic of the

stainless steel vessel. The plastic vessel was constructed from a clear acrylic tube with a nominal internal diameter of

50.8 mm and a nominal wall thickness of 9.5 mm. A schematic of the vessel is shown in Figure 26.

The tube was sandwiched between two aluminum end plates equipped with 0-ring seals using four

threaded rods which passed through the plates. The rupture disk holder was attached to one of the

end plates. Two tapped holes were drilled for mounting a dynamic pressure transducer and a needle

valve on the other end plate. For those agents (all the HCFC's and HFC's) that were found to cause

glazing or to dissolve into the wall of the plastic vessel, a new acrylic tube was used for each run. The internal diameter of the plastic tube was found to vary from tube to tube due to the manufacturing process of extruding the tube. The average internal volume of the all plastic vessels was 5.09 x

10~^ m^ with a standard deviation of 0.15 x 10"^ m-^. Measurements were recorded for each tube

used.

3.3.3 Design Validation. The experimental set-up used in this study (to be discussed) resembles

closely the conceptual experimental configuration described in Section 3.2. In the conceptual design,

a nitrogen holding tank was proposed, whereas in the actual experimental set-up the internal volume

(3.38 x 10""^ m^) of the supply line (not shown in Figure 25 and Figure 26) between the nitrogen

cylinder and the pressure vessel acted like a holding tank. In addition, the needle valve used for regulating nitrogen flow into the vessel functioned in a similar way as an orifice plate placed between

the holding tank and the pressure vessel. A holding tank was not used in the actual experimental set- up because the combined holding tank/vessel system could not be easily assembled due to the

bulkiness of the system. The situation was more critical when changing of the orientation of the

system was required in order to study the eff'ect of orientation on the discharge (see Section 3.4). In an actual discharge vessel, the vessel is a closed system; that is, there is no continuous flow of

nitrogen into the vessel. However, the present experimental design depends on the inflow of nitrogen

to raise the internal pressure of the vessel in order to rupture the disk (see Section 3.4.2). Further- more, it was not possible to shut off the nitrogen flow to the vessel at the instant when the rupture

disk burst due to the relatively slow response time ( ~ 35 ms) of the solenoid used to connect the

vessel and nitrogen supply line. A series of experiments were performed to access the eff'ect of the continuous nitrogen inflow

during discharge. It was found that by regulating (less than a 1/4 turn; equivalent to an orifice opening of less than 0.5 mm) the needle valve to minimize the inflow of nitrogen such that the
84 SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

Nitrogen for

rupturing disc

Pressure transducer

Access holes

for thermocouples

I

I I I I

!

I I I I

\

i

I I I

I I J L

I I

I I J_L

Fast response

solenoid valve

Needle

valve

50.8 mm

240 mm

Access hole

for pressure transducer

1/

Rupture disc holder Rupture disc

Figure 25. Schematic of the stainless steel vessel.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 85

Needle

Rupture disc

Figure 26. Schematic of the acrylic vessel.
86 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

duration from the initiation of nitrogen flow to the rupture of the burst disk was greater than one

second, no appreciable diff'erences (within the experimental uncertainty due to the actual burst

pressure) were detected in the internal pressure traces during discharge. Therefore, it can be

concluded that the continuous flow of nitrogen has an insignificant effect on the discharge behavior.

Based on the above discussion, the actual experimental configuration, in essence, incorporates the

design recommendations made in Section 3.2 for a generic experimental set-up using a relatively lowpressure large-volume holding tank and a small orifice between the holding tank and the discharge

vessel.Since nitrogen is dissolved in liquid agent in an actual discharge vessel, the eff'ect of dissolved

nitrogen on the discharge process needs to be addressed. During discharge, rapid internal depressurization of the vessel could cause degassing of dissolved nitrogen from the liquid which could affect the dynamic behavior of the liquid agent inside the vessel and the flow rate through the vessel

opening. Based on observed internal pressure-time histories during discharges of halon 1301, it has

been conjectured (with no visual observations) that degassing of nitrogen occurs during the time

interval from initiation of discharge to the complete depletion of the liquid (Williamson, 1976, Elliot

er al., 1984). However, the characteristic discharge times (order of seconds) for vertically downward

release were much longer in these earlier studies than in our experiments ( ~ 30 ms). Based on the

following argument, it is hypothesized that the characteristic discharge time in our experiments is much faster than the time required for degassing of dissolved nitrogen. The time for evolution of

dissolved nitrogen in the liquid can be considered to be the sum of the waiting time for nucleation of

nitrogen bubbles and the time for nitrogen bubbles to grow (Avedisian, 1985). If the waiting time is assumed to be very short even in the case of rapid discharge, the bubble growth time is then the

dominant factor. Assuming the bubble growth is diffusion-controlled, i.e., the growth is due to the

diffusion of nitrogen into the bubble, the bubble growth velocity is of the order of 0. 1 m/s (Bankoff,

1966). This growth velocity is about two orders of magnitude slower than the velocity of the

depleting liquid level during rapid discharge (to be discussed later), implying that the discharge time

of the liquid is much faster than that of degassing of dissolved nitrogen. Hence, the presence or absence of dissolved nitrogen would not affect the behavior of the liquid inside the vessel during very

rapid discharge. In the experimental procedure used in this work (to be discussed in the next

section), the dissolved nitrogen does not play a role due to the short contact time between the liquid

and nitrogen in the vessel before initiation of discharge. The above argument is to ensure that the

experimental procedure used in this study simulates closely the actual discharge process. Although

dissolved nitrogen may not play a role in the behavior of the liquid inside the vessel during vertically

downward discharge, it may affect the flashing behavior of the liquid external to the vessel as well as the behavior of the liquid inside the vessel during vertically upward and horizontal release due to longer characteristic discharge time. However, the solubilities of nitrogen in all the liquid agents

used in this study were found to be very similar (see Section 2). Therefore, the differences in the

overall external behaviors of the liquid agents could still be considered not to be attributable to differences in the solubility of nitrogen in the liquid agents.

3.4 Vessel Discharge Rate Measurements

3.4.1 Introduction. The events occurring inside and outside of a pressurized vessel containing

alternative agent during discharge are extremely complicated due to the presence of two-phase flow. The discharge process addresses how the agent is delivered to a fire scene. The eflFectiveness of an

agent in suppressing a fire depends on how fast and how it is transported to the fire. This part of the
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 87

task is designed to address the behavior of the proposed aUernative agents inside the pressure vessel

during rapid discharge. The characteristic discharge time is on the order of tens of milliseconds.

Various parameters that could influence the discharge process will also be examined in this study.

The parameters studied are: (1) orientation of discharge, (2) orifice diameter, (3) presence of an

extension tube at the vessel exit, (4) vessel pressure at discharge, and (5) ambient temperature. The

behaviors of the agents external to the vessel during discharge are discussed in Section 3.5.

3.4.2 Experimental Method. The experimental procedure involved the following steps. The vessel was evacuated for at least 15 min. The vessel was then connected to the agent supply bottle. For

liquid agents, the vessel was filled approximately two-thirds full (by volume). For FC-116, which is a gas at room temperature, the vessel was filled until the pressure in the vessel equilibrated with the

pressure in the supply bottle, and the amount of FC-116 in the vessel corresponded to an average fill density of 110 kg/m^. This number was limited by the pressure of the agent in its supply bottle. The

total mass of an agent dispensed to the vessel was obtained by weighing on an electronic scale with an

accuracy of 1 g. The vessel was then pressurized with nitrogen to 75% of the nominal burst pressure

of the rupture disk through a solenoid valve (Skinner 7 12X1 DIB) and the needle valve (see Figure 25

and Figure 26). To initiate a discharge, the upstream nitrogen pressure in the nitrogen supply line was first raised to approximately 15% above the nominal burst pressure with the solenoid valve

closed. The solenoid valve was then opened to start a flow of nitrogen into the vessel through the

needle valve and initiate the experiment. The nitrogen flow continuously increased the internal pressure of the vessel to a point where the rupture disk could no longer sustain the rising internal pressure. Bursting and complete opening of the rupture disk followed, thus causing rapid release of

the agent from the vessel. A high-speed movie camera (Photec IV) operating at 2000 frames per second was used to document the events occurring inside the plastic vessel. Kodak Ektachrome high-speed daylight films

(ASA 400), some front lighting, and intense backlighting were used. Emptying rates were obtained

by using frame-by-frame analysis of the movies taken from the experiments. The high-speed motion

picture film acquired for each discharge was analyzed using a slide scanner (Microtek ScanMaker 35t)

to digitize the images and digital image processing software (Aldus PhotoStyler Version 1.1).

Emptying rates were extracted from the digitized images of the liquid level of the agent in the vessel

as a function of time during discharge.

The slide scanner is designed for use with 35 mm slides or film strips, both color and black and

white, but is easily adapted for use with 16 mm movie film. It uses a linear array CCD (charge

coupled device) and can digitize an image with a resolution of 3656 x 3656 dpi (dots per inch); twice

the resolution (7312 x 7312 dpi) is possible with software interpolation. Dots can be represented by

an 8-bit code indicating any one of 256 shades of gray, or by a 24-bit color code indicating any one

of over 16 million colors. For this work, two frames per image were scanned as grayscale images

with 3656 x 3656 dpi at 100% scale. The scanning module incorporated in the digital image processing software (PhotoStyler) is Twain-compliant, which allows software applications and hardware imaging devices to communicate

directly. This means that the scanner can be accessed directly from within the digital image

processing software; after the scan, the image is placed inside the imaging software, where data

processing is performed before the image is archived. The archival method used is to compress the

image in JPEG format. In this way, an image that comprises 3.8 Mbyte of disk space can be saved in a file which is 144 kbyte without loss of image quality.

The process for computing the emptying rates is straightforward. The digital image processing

software allows the measurement of the position of every pixel in the image, which is typically 2630

pixels wide by 1470 pixels high. Accurate measurements of timing and liquid level position, as seen
88 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

through the transparent vessel, allow the computation of volumetric flow rate when the vessel

diameter is known.

The edge of the film strip contains a timing mark adjacent to the sprocket holes; the spot appears on the film at a rate of 1 kHz. When the high-speed movie camera is operated at 2000 frames per

second, there is a timing spot every 2 frames (nominally 0.5 ms per frame). A more accurate framing rate can be computed by measuring the timing spot relative to the sprocket holes and using

linear interpolation. It is estimated that the imprecision in the film speed for the data in this report is ± 0.001 ms per frame, or + 0.2%.

The liquid level and a reference point are measured for each frame with a resolution of + 0.5

pixels. Pixels are related to the physical length measurement scale by means of a 25.4 mm steel ball photographed at the beginning of each film reel. The uncertainty in the length calibration is + 0.5

pixel, or 0.2 mm. Thus, the uncertainty in the liquid position measurements is typically + 0.4 mm.

Two thermocouples, one placed in the ullage and one in the liquid, were used to record the

temperature changes in both phases during discharge. The thermocouples used were K-type, fine gage (12 or 25 /xm), and unsheathed. The thermocouple were constructed by arc-welding the fine thermocouple to the two bare ends (with the original bead junction removed) of a rugged unsheathed

thermocouple (K-type) probe under a microscope. Attempts to record temperatures in the both phases

(vapor and liquid) inside the vessel proved to be difficult, although some qualitative information could

still be obtained from the temperature measurements. The diflficulty lies in the response time and the

fragility of the fine wire thermocouples. Even a fine 12 /xm unsheathed thermocouple was found to have a time constant which was too slow to follow the events occurring inside the vessel. The pressure transducer was a Kistler Model 603B1 piezoelectric dynamic pressure transducer

which was regulated by a dual mode charge amplifier (Kistler Model 5004). The outputs from the

charge amplifier and thermocouples were recorded using a high-speed data acquisition board

(Strawberry Tree FLASH- 12^^^ Model 1) at a rate of 25 kHz and stored in a personal computer for subsequent data analysis.

Unless specified, discharge experiments were conducted by discharging vertically downward at room temperature with rupture disks having a nominal burst pressure of 4.12 MPa. This condition

will be referred to as the standard discharge in the discussion. The agents studied were the eleven

potential replacement agents: HFC-236fa, FC-31-10, FC-318, HCFC-124, HFC-227ea, HFC-134a,

FC-218, HCFC-22, HFC-125, HFC-125 (40% by mass)/HFC-32 (60%) mixture, and FC-116. Halon

1301 was also included in the study as a reference agent.

The eff'ect of the size of the opening on the discharge process was assessed by mounting various

diameter orifice plates at the base of the discharge vessel to restrict the opening of the rupture disk at the exit of the disk holder. The orifice diameters used in this study were 19.1, 12.7 and 6.4 mm.

The plates were 5.08 mm thick. In order to simulate discharge through a pipe for engine nacelle fire protection applications,

experiments were performed using an extension tube 0.5 m long with an internal diameter equal to that of the 19.1 mm rupture disk. The tube was attached to the vessel exit, and the eff'ect of

orientation on the discharge dynamics with an extension tube was also studied.

The effect of the initial charge pressure on the discharge was simulated and studied by performing a series of experiments at room temperature using rupture disks with different burst pressures. Rupture disks with nominal burst pressures of 2.76, 4.12, and 5.50 MPa were selected. A few experiments were also conducted to evaluate the effect of temperature on the discharge

characteristics. The temperature effect was evaluated only at -45 °C by using burst disks with

nominal burst pressures of 2.76 MPa. Based on the measurements of thermodynamic properties of

agent/nitrogen mixture (see Section 2), this burst pressure was chosen in order to simulate the residual

pressure in an actual vessel as a result of exposing the vessel to a -45 °C environment. Two agents
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 89

(FC-31-10 and FC-218) were evaluated at this temperature. Cooling was achieved by immersing the

vessel in dry ice. Due to its large thermal inertia, the stainless steel vessel and its contents could only

be chilled to an equilibrium temperature of -45 °C. No experiments were conducted at 150 °C as originally planned because the final pressure exceeded the operating pressure of the vessel even when

the vessel was 1/3 fill (see Section 2). However, some general comments will be made in Section

3.4.3.5 with regard to the discharge characteristics at elevated temperatures.

The study on elTects of orientation on the discharge dynamics was carried out by discharging the

vessel vertically downward and upward or horizontally. Since the horizontal and vertical upwarddischarge processes were found experimentally to be slower than the downward discharge, the highspeed movie camera was operated at 500 or 1000 frames per second in these experiments.

Only three agents were used for studying the eifect of various parameters on the discharge

process. Two representative alternative agents (FC-31-10 and FC-218) and halon 1301 at room

temperamre. These two alternative agents were chosen among others because they represented the

range of boiling points encompassed by most of the eleven proposed fluids. Halon 1301 was included

as the baseline agent.

3.4.3 Results and Discussion

3.4.3.1 Standard Discharge. Visual observations of the events occurring inside the vessel

during downward discharge are described as follows. Liquid agent is propelled out of the vessel

when the rupture disk breaks. In some of the experiments, the inflow of nitrogen disturbs the

liquid/vapor interface causing a wavy motion along the liquid surface in a manner similar to that described in Section 3.2. Apparently these disturbances did not affect the agent release behavior. At

the end of the liquid discharge the vessel interior becomes foggy, possibly due to the condensation of

vapor caused by a decrease in temperature in the ullage due to the adiabatic expansion of the vapor

phase. As the discharge of the remaining vapor continues, the vessel becomes clear once again.

Figure 27 shows a photographic sequence of events during a downward discharge of FC-218

from a plastic vessel. The photographs were taken at 2000 frames per second. The frame which

corresponds to the time t = 0.5 ms is the first frame where agent is observed outside the vessel. Due

to the framing rate, there is a 0.5 ms uncertainty in the times. The dark grey horizontal shadow in the middle of the background was caused by the backlighting set-up. The first photograph shows the

plastic vessel, which is located in the middle, the vessel mount, and the liquid level. In this sequence

of photographs, the perturbation of the vapor/liquid interface caused by the inflow of nitrogen is evident from the wavy nature of the interface. The photographs also show the behavior (to be

discussed in Section 3.5) of the agent outside the vessel during discharge.

Figure 28 through Figure 39 show results for the temporal variation of the internal pressure

during downward discharges of all the agents evaluated in this study. Pressures reported in the

figures are gauge pressures. The pressure is nondimensionalized by the actual burst pressure, which is taken to be the pressure at t (time) = 0 s. The time / = 0 s is defined as the time when the

first laser (located at the immediate exit of the vessel) signal is attenuated (see Section 3.5) by the

agent coming out from the vessel. For most agents (except FC-116, HFC-125/HFC-32 mixture, and

halon 1301), there are two distinct regions, separated by an inflection point, in the pressure-time

history. The first region corresponds to the time interval during which the liquid agent is being

propelled from the vessel; the second region corresponds to the period when the remaining vapor

(mostly nitrogen) is being vented from the vessel. Based on observations from the plastic vessel

experiments, the inflection point corresponds very closely to the time at which the liquid agent has

just been completely expelled from the vessel. Since the pressure-time histories obtained from the

pressure transducer mounted on the stainless steel vessel wall in order to monitor the temporal change
90 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 27 Photographic sequence of events showing a downward discharge of

FC-218 from a plastic vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 27. (continued) Photographic sequence of events of FC-218 from a plastic vessel.

showing a downward discharge
92 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 27

100.0 ms 125.0 ms

(continued) Photographic sequence of events showing a downward discharge

of FC-218 from a plastic vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100 120 140

I (ms)

Figure 28. Temporal variations of internal pressure during downward discharges of

HFC-236fa.
94 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

0 20 40 60 80 100 120 140

t (ms)

Figure 29. Temporal variations of internal pressure during downward discharges of

FC-31-10.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 95

FC-318

0 20 40 60 80 100 120 140

t (ms)

Figure 30. Temporal variations of internal pressure during downward discharges of

FC-318.
96 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 3 1 . Temporal variations of internal pressure during downward discharges of

HCFC-124.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 32. Temporal variations of internal pressure during downward discharges of

HFC-227ea.
98 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 33. Temporal variations of internal pressure during downward discharges of HFC-134a.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 99

0 20 40 60 80 100 120 140

t (ms)

Figure 34. Temporal variations of internal pressure during downward discharges of FC-218.
100 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 35. Temporal variations of internal pressure during downward discharges of HCFC-22.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

1

\ r

I \ \ \ I \ \ i_

0 20 40 60 80 100 120 140

t (ms)

Figure 36. Temporal variations of internal pressure during downward discharges of HFC- 125.
102 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100 120

t (ms)

Figure 37. Temporal variations of internal pressure during downward discharges of the HFC125/HFC-32 mixture.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 103

FC-116

0 20 40 60 80 100

t (ms)

Figure 38. Temporal variations of internal pressure during downward discharges of FC-116.
104 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

1.2 1 r

Halon 1301

= 4.55 MPa

with extension tube

P = 4.35 MPa

20 40 60 80 100 120 140 160 180 200

t (ms)

Figure 39. Temporal variations of internal pressure during downward discharges of halon 1301.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 105

of the pressure in the liquid phase followed those obtained from the pressure transducer used to measure pressure change in the vapor phase, only the pressure-time histories from the vapor phase

measurements are presented and discussed.

The pressure-time histories obtained using plastic vessels have an oscillatory nature and more

noise. Such time traces are easily identified in the figures. This behavior is believed to have been

caused by vibration of the vessel mount during discharge. The observed oscillatory frequency of the

pressure traces corresponded closely to the natural frequency of the vessel mount, which was

estimated by simulating the four threaded rods (see Figure 26) as four springs and the vessel as a mass attached to the springs. Since a diff'erent mount was used for the stainless steel vessel, no such

oscillatory characteristic in the pressure-time histories was observed. The variations in the pressure- time histories from run to run reflect the nature of the rupture disk, that is, the actual burst pressure

of each rupture disk used in the experiments deviated somewhat from its nominal burst pressure. Therefore, difi'erent Pfs are anticipated depending on the pressure at which the disk bursts. Diff'erences in the pressure histories obtained from the stainless steel vessel and from plastic vessels

are due to the diff'erences in the internal volumes of the vessels. The visual observation of the internal behavior of liquid agent during discharge from plastic

vessels shows that during the time when the liquid level is visible (refer to Figure 26) no internal

boiling of the liquid occurs during depressurization. This observation can be explained by examining

the temporal variation of the internal pressure. From Figure 28 through Figure 36, for the nine

agents with boiling points above -45 °C, the pressures at the time when the liquid empties (at the

inflection point in the pressure-time history) is still well above the saturation vapor pressure for all of

these agents at room temperature. Although the temperature of the ullage above the liquid is very

low because of the expansion of the vapor phase, the characteristic time, t^, for the interfacial heat

transfer between the liquid and the vapor phases is much longer than the discharge time of the liquid,

typically 0(10 ms). The characteristic time for heat transfer can be estimated (Bird et al., 1960)

using

where a is the thermal diffusivity, typically O(10~^ m^/s) and 5 is the thermal penetration distance in the liquid, in our case typically O(10~^ m). Hence, is on the order of tens of seconds. Therefore,

the liquid agent remains at the initial room temperature and can be considered to undergo an

isothermal depressurization during discharge. As shown in Figure 40, this argument is supported by

the temperature-time histories in the liquid during discharge of HCFC-22. Figure 40 was obtained

using a 25 /xm fine wire thermocouple. Data were taken at a rate of 10 kHz.

For FC-1 16 (see Figure 38), the pressure-time histories show a relatively smooth decay of the

internal pressure. Since FC-1 16 is a gas at room temperature, this behavior is typical of a gas vented

from a pressurized vessel. The two regions in the temporal variation of the internal pressure

observed for the liquid agents do not exist. For HCFC-22, HFC- 125, HFC-125/HFC-32 mixture, and halon 1301, only the pressure-time

histories from the stainless vessel were recorded. Since some of these agents etch plastic and all have

relatively high vapor pressures at room temperature, the plastic vessel was not used for fear of

rupture of the vessel during its filling and handling.

The experimental results for HFC-125/HFC-32 mixture and halon 1301 deserve further

discussion. The pressure-time histories in Figure 37 and Figure 39 show that the transition from the

liquid discharge to the venting of the remaining vapor from the vessel is not distinct and abrupt. The

transition appears to be gradual. Since no visual observations were made for these agents, the events
106 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 40. Temporal variation of temperature in the liquid phase during discharge of HCFC-22.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 107

occurring inside the vessel could only be inferred from the experimental pressure-time traces. It can be seen that at some point during the release the internal pressure dropped below the saturation vapor

pressure of the agent at room temperature (refer to Section 2). Therefore, it is conjectured that boiling of the liquid occurred near the end of the liquid discharge, thus causing a gradual transition

until the remaining liquid had been discharged.

Based on visual observations of the discharges, a simple mathematical model can be formulated

to simulate the pressure-time history. As mentioned above, there are two distinct regions, separated

by an inflection point, in the pressure-time traces. The first region corresponds to the liquid

discharge while the second region corresponds to the discharge of the remaining vapor after the liquid

has been released from the vessel. The formulation of the model for liquid discharge will first be

presented, followed by a discussion on the simulation of the discharge process of the remaining

vapor.Once the rupture disk bursts, the nitrogen above the liquid agent simply acts as a piston to drive

the liquid out of the vessel. The control volume that is being considered is the ullage above the

liquid. Since the discharge process is very rapid, the process occurring in the gas phase can be

assumed to be adiabatic and reversible, i.e., isentropic. Furthermore, for simplicity the control

volume is assumed to be closed in the analysis because the inflow of nitrogen does not influence the

discharge process (see Sections 3.2 and 3.3.3). Assuming the vapor phase to be ideal, from the first law of thermodynamics for an isentropic process.

where 7 is the ratio of the constant pressure and volume heat capacities, P is the pressure, and V is the control volume. Since nitrogen is the dominant species in the vapor phase, the 7 of nitrogen with

a value of 1.4 is assumed. If a time derivative of Equation (18) is taken, then

Because of the design of the rupture disk holder, there is a short distance, L (25.4 mm), having a diameter of 19.1 mm through which the fluid has to pass before exiting from the release vessel

assembly. Because of rapid depressurization, the liquid becomes superheated as it passes through the

opening of the disk (Modell and Reid, 1983; Reid, 1976). Superheating of the liquid has been

observed in previous studies on the discharge of saturated and subcooled liquid from orifices and

nozzles (Benjamin and Miller, 1941; Burnell, 1947; Pasqua, 1953; Hesson and Peck, 1958; Henry,

1970; Prisco et al., 1911 \ Celata et al., 1983; Van den Akker et al., 1983). If we assume the flow

of metastable liquid through the disk opening (D) and the disk holder can be approximated by flow

through a short tube (0 < L/D < 3) , then the volumetric flow rate of the liquid, expressed in terms

of the volumetric expansion rate of the gas above the liquid agent, can be calculated by (Prisco et al.,

1977; Tong, 1965)

constant , (18)

dP ^ Py dV

dt

^ V dt

(19)

Pi

(20)

where A is the cross sectional area of the disk opening, is the liquid density, P^^^ is the saturation

vapor pressure of agent, and Q is the discharge coefficient for the liquid, with measured values found
108 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

to be between 0.61 to 0.64 in the literature (Frisco et al., 1977; Tong, 1965). Equations (19) and

(20) were solved numerically using a fourth order Runge-Kutta method (Carnahan et al., 1969) to obtain the pressure-time histories and the liquid level as a function of time.

Figure 41 through Figure 46 show the temporal variations of liquid level for agents obtained

from experiments using plastic vessels. Also shown in the figures are predictions obtained using

Equations (19) and (20). The liquid levels were measured from the upper edge of the lower

aluminum plate (see Figure 26), and a zero liquid level does not correspond to complete depletion of

liquid. The best curve fits were obtained by optimizing the values of Q, which varied between 0.67

to 0.80. These Q values are slightly higher than previous experimental values reported in the

literature. After complete depletion of liquid, the discharge process of the remaining vapor can be modeled

as follows. If the vapor is assumed to be an ideal gas, the process is again assumed to be isentropic,

and the pressure in the vessel is high enough that the flow can be assumed to be choked in the disk

opening (which is approximated as a round, sharp-edged orifice), it can be easily shown that the

following equation describes the rate of pressure decay in a vessel as a function of time (Kim-E,

1981).

P , CA 1/2 M 2y

2y

(21)

with

K

Y

+ 1

(22)

where P^i and T^i are the pressure and temperature in the vessel at the instant of complete depletion of

liquid agent, respectively, is the discharge coeflScient for the gas, R is the universal gas constant,

and M is the molecular weight of the vapor (assumed to be nitrogen). A choked flow is justifiable

because Pg/P, where P^ is the exit pressure (0.101 MPa), is less than the critical pressure ratio most

of the time during discharge of the remaining vapor. For an ideal gas undergoing an isentropic process, the critical pressure ratio is given by (Balzhiser et al., 1972)

( 2 ]

Y

(Y - 1) (23)

(p)critical [y + iJ

Optimum values of Q were found to be between 0.26 and 0.31 (Kim-E, 1981) during depressurization of a vessel initially filled with nitrogen.

Figure 47 through Figure 52 compare the complete pressure-time histories with predictions from

Equations (19), (20), (21), and (22). The best curve fits for the duration of vapor discharge were

obtained by optimizing the values of which varied between 0.25 to 0.35. However, if instead of

using the P^^'s from the liquid discharge calculations, one fits the experimental results of vapor

discharge by optimizing both P^i and the values of Q are found to be ~ 0.6. The average volumetric flow rates, defined as the initial volume of liquid in the vessel divided by

the emptying time of the liquid, are summarized in Table 4. Knowing the framing rate of the high-
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE

0 2 4 6 8 10 12 14 16 18 20 22 24

I (ms)

Figure 41 . Temporal variation of liquid level obtained from a HFC-236fa discharg
110 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 2 4 6 8 10 12 14 16 18 20 22 24

t (ms)

Figure 42. Te nporal variation of liquid level obtained from a FC-31-10 discharge.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 43. Temporal variation of liquid level obtained from a FC-318 discharge.
112 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 44. Temporal variation of liquid level obtained from a HCFC-124 discharge.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

Figure 45. Temporal variation of liquid level obtained from a HFC-227ea discharge.
114 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 2 4 6 8 10 12 14 16 18 20 22 24

t (ms)

Figure 46. Temporal variation of liquid level obtained from a FC-218 discharge.
SECTIONS. FLUro DYNAMICS OF AGENT DISCHARGE 115

HFC-236fa

0 20 40 60 80 100 120 140

t (ms)

Figure 47. Comparison of measured pressure-time history of a HFC-236fa discharge with

prediction.
116 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 48. Comparison of measured pressure-time history of a FC-31-10 discharge with

prediction.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 117

FC-318

0 20 40 60 80 100 120

t (ms)

Figure 49. Comparison of measured pressure-time history of a FC-318 discharge with

prediction.
118 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 119

1 r

0 20 40 60 80 100 120 140

t (ms)

Figure 5 1 . Comparison of measured pressure-time history of a HFC-227ea discharge with

prediction.
120 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

J L 0 20 40 60 80 100 120 140

t (ms)

Figure 52. Comparison of measured pressure-time history of a FC-218 discharge with

prediction.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 121

Table 4. Initial burst pressures (P,), initial agent masses (M,), and average volumetric flow rates (Q) obtained from experiments using plastic and stainless steel vessel

Agent (g) Q (mVs) X 10^

rlrC-ZJOla 4. + U. 14 468 ± 5 10.8 ± 0.5

t:/~* t 1 1 A rL-Jl-lU J.oo ± U.Uo 508 ± 18 10.3 ± 0.2

rL-Jlo 4.z(J ± U.JO 503 ± 3 10.8 ± 0.2 HLrL-1/4 4,16 + 0.18 464 + 2 10.9 ± 0.8 HrL-ZZ /ea 4.44 476 11.1 HFC- 134a 4.23 + 0.12 416 ± 8 10.7 ± 0.9

FC-218 4.31 ± 0.37 462 + 2 10.8 + 0.7 HCFC-22 4.40 ± 0.15 411 ± 14 10.2 ± 0.1 HFC-125 4.16 ± 0.09 414 ± 10 9.8 + 0.7

HFC-125/HFC-32 mixture 4.32 ± 0.04 367 ± 12 10.0 + 0.8

CFjBr'' 4.35 554 7.6

^mean ± SD

''one run

speed movie camera, the emptying time was obtained by counting the number of frames of the movies

for the second flashing of the spray (see Section 3.5) to occur at the vessel exit following the start of

the release because the time for the appearance of the second flashing corresponded closely to the

inflection point in the pressure-time histories. The scatter in the data reflects the uncertainty in identifying the exact emptying time of the liquid and variations in the initial burst pressure. Except

for halon 1301, the average volumetric flow rates of the alternative liquid agents evaluated vary less than 10% among each other. Based on Equations (19) and (20) and given the same initial conditions,

the emptying time or the volumetric flow rate of the liquid should be relatively independent of the

agents studied because the terms under the square root in Equation (20) do not diff'er significantly

among the agents. This is consistent with the experimental observations made in this work.

However, the above simple analysis is only applicable when there is no internal boiling of liquid

agent inside the vessel during discharge in a downward orientation; this may not be the case for liquid

agents with low boiling points (e.g., halon 1301).

Since FC-116 is a gas at room temperature. Equations (21) and (22) can be used to simulate the

discharge of FC-116. In this case, P^^ = and T^i = T^. The predicted and measured pressure

decays are given in Figure 53. The predicted values were calculated by assuming an average

molecular weight of 83 based on the amount of FC-1 16 and nitrogen at r = 0 and a value of 1.2

for y. There is a large discrepancy between the measured and predicted temporal variations of

internal pressure during discharge. It is conjectured that the assumption of ideal gas behavior which

underlies Equations (21) and (22) may not be applicable for FC-1 16. Support for this conjecture is obtained from studies on releases of nitrogen only (which should behave as an ideal gas). For

nitrogen-only releases from the pressurized vessel, Equations (21) and (22) predicted reasonably well

the temporal variation of the internal pressure as can be seen in Figure 54. Note that, for both cases
122 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100

t (ms)

Figure 53. Comparison of a measured pressure-time history of FC-116 discharge with

prediction.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 123

Figure 54. Comparison of a measured pressure-time history of pure nitrogen discharge with

prediction.
124 SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

FC-116 and nitrogen releases, the optimum values of are much higher than the values found in other discharge experiments wherein a liquid agent is initially present and the values in the literature (Kim-E, 1981). The low values of Q obtained by Kim-E (1981) are possibly due to obstruction of

his flow path by a pneumatic plunger. In his experimental set-up, a rupture disk was used as a

release mechanism. The plunger which was placed inside the vessel was used to break the rupture

disk in order to initiate the discharge process. However, the reason for the low and high values

obtained in this study is not known; the low values of Q were obtained when gas discharge was

preceded by liquid release, whereas the high values of were found when liquid was not present

initially in the vessel. The value for Q normally centers around 0.6 (Perry et al., 1984). It can only

be conjectured that the edge roughness of the rupture disk opening (Perry et al., 1984) and the

transient eff"ect due to the abrupt transition from the discharge of liquid to the release of the remaining

vapor may have an effect on the discharge coefficient, and that the assumption of a round, sharpedged orifice may not be valid because the flow through the disk holder resembles one through a short tube rather than through an orifice. The use of a single value of Q may also be questionable

because can be a function of the pressure ratio across the opening (see Section 3.2). Since the

pressure inside the vessel is changing with time during discharge, the pressure ratio is then a function

of time. Therefore, it is conceivable that may not be constant during the course of discharge.

3.4.3.2 Effect of an Orifice Plate. The experimental results for internal pressure obtained by

restricting the disk opening by means of an orifice plate are shown in Figure 55 through Figure 57. The agent in the vessel remains as liquid during the observation period from the initiation of

discharge to the time when the liquid level is no longer visible (see Figure 26) within the plastic

vessel. The discharge process is very similar to that without the orifice plate. The key effect of

decreasing the diameter of the orifice on the discharge process is to increase the emptying time of the

contents, as indicated by the locations of the inflection points in Figure 55 through Figure 57. For

the 12.7 mm plate, the average volumetric flow rates are 6.0 x 10"-^, 5.6 x 10"-^, and 3.2 x 10~^

rn^/s for FC-31-10, FC-218, and halon 1301, respectively. For the 6.4 mm plate, the average

volumetric flow rates are 1.5 x 10"^ 1.4 x 10~^ and 1.3 x 10"^ m^/s for FC-31-10, FC-218,

and halon 1301, respectively. It is found that increasing the diameter of the plate from 6.4 to 12.7 mm increases the average volumetric flow rates of FC-31-10 and FC-218 by approximately

fourfold, in accordance with Equation (20), i.e., dV/dt ~ /4 for a given P. The relationship between

the average volumetric flow rate (Q) and the orifice area {A) is clearly seen in Figure 58 for FC-31-

10 and FC-218 using burst disks with a nominal burst pressure of 4.12 MPa. Results discussed

earlier for standard releases (orifice diameter of 19.1 mm) are included. The symbols in the figure are measured values. The dotted and solid lines in the figure are least-square fits of the data which

are forced to pass through the origin. Figure 59 through Figure 62 show the comparisons of

measured pressure-time histories of FC-31-10 and FC-218 discharge with predictions using Equations

(19) through (22). For the experimental results obtained using a 6.4 mm restrictor plate, the model

did not simulate the discharge process well because the flow field is more complicated when a smaller

orifice plate is used to restrict the flow at the exit of the short tube section which was part of the

rupture disk holder (see Figure 25 and Figure 26). Since it has been conjectured in Section 3.4.3.1

that boiling of liquid halon 1301 inside the vessel occurs near the end of liquid discharge, the above

volumetric flow rate-area relationship is not applicable to halon 1301.

3.4.3.3 Effect of an Extension Tube. From visual observations, the presence of an extension

tube did not affect the internal behavior of the liquid during discharge. Figure 63, Figure 64, and
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 55. Temporal variations of internal pressure during discharges of FC-31-10 with

restricting orifice plate (6.4 or 12.7 mm).
126 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

1.2 1 r

a.

P_ = 4.12 MPa

6.4 mm plate

4.21 MPa

6.4 mm plate

1 r

FC-218

P = 4.21 MPa

1^.7 ram plate

P = 4.29 MPa

1^.7 mm plate

J L 100 200 300 400 500 600 700 800 900 1000

t (ms)

Figure 56. Temporal variations of internal pressure during discharges of FC-218 with a

restricting orifice plate (6.4 or 12.7 mm).
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 127

Figure 57. Temporal variations of internal pressure during discharges of halon 1301 with a

restricting orifice plate (6.4 or 12.7 mm).
128 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0.020

0.016

0.012

0.008

0.004 -

0.000

1 r

• FC-31-10

V FC-218

J L 0 50 100 150 200 250 300 350

A (mm )

Figure 58. Relationship between average volumetric flow rate and orifice area.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 59. Comparison of a measured pressure-time history of FC-31-10 discharge using

6.4 mm orifice plate with prediction.
130 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 60. Comparison of a measured pressure-time history of FC-31-10 discharge using a 12.7 mm orifice plate with prediction.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 131

Figure 61 . Comparison of a measured pressure-time history of FC-218 discharge using a 6.4 mm orifice plate with prediction.
132 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

I I I \ \ \ I 0 50 100 150 200 250 300

time (ms)

Figure 62. Comparison of a measured pressure-time history of FC-218 discharge using a 12.7 mm orifice plate with prediction.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 133

Figure 63. Temporal variations of internal pressure during discharges of FC-31-10 with a 0.5 m

long extension tube.
134 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

I

\ I I \ I 0 20 40 60 80 100 120 140 160 180 200

t (ms)

Figure 64. Temporal variations of internal pressure during discharges of FC-218 with a 0.5 m

long extension tube.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 135

Figure 39 show the internal pressure results for FC-31-10, FC-218, and halon 1301. For comparison, the results obtained with and without extension tube are shown in Figure 65, Figure 66, and

Figure 39. It appears that the presence of an extension tube does not change the temporal behavior of

the internal pressure during discharge of liquid FC-31-10 and FC-218. This is indicative of the

existence of critical flow at the vessel orifice since the downstream geometry does not aff'ect the

pressure upstream (Prisco et al., 1977). The presence of an extension tube slows down the discharge

of the remaining vapor from the vessel after liquid depletion because of friction losses in the tube, as can be seen in Figure 65, Figure 66, and Figure 39. In the case of halon 1301, the presence of an

extension tube makes the pressure-curve transition from liquid to gaseous discharge more gradual.

Such gradual transition is indicative of initiation of boiling within the vessel or extension tube (see

discussion in Section 3.4.3.1). Furthermore, when compared to the discharge behavior of halon 1301

without an extension tube, the discharge of the two-phase boiling halon 1301 in the presence of an

extension tube is much slower due to additional frictional losses in the tube.

3.4.3.4 Effect of Initial Vessel Pressure. The above discussion is based on results obtained

using rupture disks with nominal burst pressures of 4.12 MPa. In order to study the eff'ect of initial pressure on the discharge dynamics, rupture disks with different nominal burst pressures (2.75 and

5.50 MPa) were used. Figure 67 shows pressure-time histories for FC-218 for three initial burst

pressures. The oscillation seen in two of the pressure traces reflects the use of a plastic vessel. It is clear that the emptying time of the liquid decreases with increasing initial pressure in the vessel, and

that a twofold increase in the initial burst pressure only increases the average volumetric flow rate by

approximately a square root of 2. The measured average volumetric flow rates for these three runs

are 8.4 x 10~^ m^/s for = 2.60 MPa, 11.3 x 10""^ m^/s for = 4.57 MPa, and 11.9 x 10"^

m?/s for Pj = 6.18 MPa. The results are consistent with the model (Equation (20)), i.e., dV/dt -

(P)^'^ for a given A. The relationship between the average volumetric flow rate (Q) and the initial charge pressure (P,) is shown in Figure 68 for FC-218 using a 19.1 mm opening. The dotted line in the figure is a least-square fit through the origin, and the symbols represent measured values. The

calculated and measured temporal variations of internal pressure for Pj = 6.18 MPa, as an example,

are shown in Figure 69. The prediction matches the experiment quite well.

3.4.3.5 Effect of Vessel Temperature. The study of the efl'ect of temperature on the discharge

dynamics was carried out experimentally with FC-31-10 and FC-218 at -45 °C. Since similar trends were observed for FC-218, only the results for FC-31-10 are shown in Figure 70. The pressure-time

histories are very similar to those obtained using the same burst disks but at room temperature. The

internal behavior of the agent in the vessel, though not observed visually, should be similar. However, the dispersion characteristics of the agent external to the release vessel were observed to have a less-flashing nature when the agent was discharged at -45 °C (see Section 3.5.3.9).

Figure 71 shows a comparison of the measured results with predictions from Equations (19) through

(22).

Although no experiments were carried out at 150 °C, the discharge of an agent at this temperature may be characterized as a discharge of a supercritical fluid from the vessel because at this temperature, and its corresponding pressure, the agent/nitrogen mixture is in the supercritical region.

If this indeed is the case, then the discharge process should be very similar to that of a FC-1 16/nitrogen mixture.

Some interesting but important events which may occur when the vessel is rapidly discharged at moderately high temperature and pressure will be discussed. These processes are also discussed in Section 3.2.4. The following is based on the hypothesis proposed by Reid (1979) in an attempt to

(

\
136 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 65. Temporal variations of internal pressure during discharges of FC-31-10 with and

without an extension tube.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100 120 140

t (ms)

Figure 66. Temporal variations of internal pressure during discharges of FC-218 with and

without an extension tube.
138 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100 120 140

t (ms)

Figure 67. Temporal variations of internal pressure during discharges of FC-218 with different

initial burst pressures.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 139

0.016

0.000

D ,1/2 Pj (MPa)

Figure 68. Relationship between average volumetric flow rate and initial charge

pressure.
140 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

t (ms)

Figure 70. Temporal variations of internal pressure during downward discharges of FC-31-10 at -45 °C.
Figure 71. Comparison of measured pressure-time history of a FC-31-10 discharge at -45 °C

with prediction.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 143

provide a possible meclianism for pressurized-liquid lank explosions or boiling liquid expanding vapor

explosions (BLEVEs).

In order lo delineate the important features, and for simplicity and clarity, a fixed initial amount

of pure agent in the vessel without pressurization of nitrogen will be considered. For example, say the vessel has an initial fill density {ppf) of 0.5 kg of FC-31-10 per 5 x lO"'^ m_^ (approximately 2/3

full) at room temperature (22 °C). The vessel is then heated to say 7^ = 100 °C. The heating

process of the vessel is illustrated in Figure 72 (point A to point B) which is a pressure-temperature

diagram of pure FC-31-10 (Daubert and Danner, 1992). Point /4 represents the initial condition of

the vessel, that is, liquid and vapor of FC-31-10 coexist in the vessel at 7) = 22 °C with a saturation

vapor pressure of 0.24 MPa). The saturation curve in Figure 72 represents the conditions of pressure

and temperature required for the coexistence of liquid and vapor phases. To determine the final state (point B) of the heating process, that is, whether point B will lie in the single phase (gas or liquid)

region or on the saturation curve, one has to compare the saturation vapor and liquid densities at 7^ (p^.^and p^) with the initial fill density of the vessel (p^//). If p^,f

< Ppn < Pip then two phases

(vapor and liquid) still coexist in equilibrium in the final state (Smith and Van Ness, 1975). In the case of FC-31-10.

p^.f = 296 kg/m^ and Pij = 1002 kg/m^ (Brown and Mears, 1958). Given an

initial fill density of 1000 kg/m^, the final state is still two-phase. Therefore, point B lies on the

saturation curve, as shown in Figure 72. If the vessel is rapidly depressurized or discharged at the

final temperature and pressure, the liquid will be superheated, that is, at a temperature higher than the

saturation temperature at the corresponding existing pressure (Modell and Reid, 1983). The liquid is superheated as a result of crossing the saturation curve without concomitant phase change. However,

there is a pressure limit for how far the metastable liquid can be sustained. This limit is the so-called

spinodal curve (Modell and Reid, 1983). The superheated liquid thus lies in the region between the

saturation curve and the spinodal curve. Once the superheated liquid reaches the spinodal curve, spontaneous, homogeneous nucleation will occur within the bulk of the liquid. Such rapid nucleation

is expected to be accompanied by strong and damaging shock waves and huge pressure increases in the vessel. If this event happens, it is expected that substantial damage would be incurred by the

vessel, or in the worst scenario, explosion of the vessel would occur if the vessel was not designed to withstand such a pressure surge. For a pure agent, the spinodal curve can be calculated (Beegle et al., 1974) by determining

where the condition

= 0 (24)

r

is met. P is the pressure, v is the molar volume, and T is the temperature. The Peng-Robinson

equation of state (Peng and Robinson, 1976) is used for illustration (see Section 2.3.1). By applying

Equation (24) to the equation of state, the following quartic equation in v results: RT 2a(v ^b) _ Q (25)

(v - bf [v(v + b) + biv - b )]'

or

SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 145

+ {4b - —)v' + 2b{b + —)v^ + 2bH~ - 2b)v + b' - - 0 (26)

RT RT RT RT

where a and b are the parameters of the equation of state (see Section 2.3.1). For given temperatures

below the critical point, there are four real roots from Equation (26). Two of the roots, one less than

the saturated liquid molar volume and the other greater than the saturated vapor molar volume, are nonphysical and are disregarded. The smallest of the remaining two roots corresponds to the v on the

liquid spinodal curve, whereas the largest corresponds to the v on the vapor spinodal curve which is not relevant in this discussion. Once v is determined for a given T, P can be obtained from the

equation of state. The calculated liquid spinodal curve for FC-31-10 is also shown in Figure 72.

If the vessel is rapidly discharged at the condition corresponding to point B in Figure 72, the

discharge process can be approximated by that of an isothermal depressurization (see Section 3.4.3),

as depicted by the vertical line in Figure 72. If the homogenous nucleation pressure (P,,,j) at 7^, which is the pressure corresponding to the intersection of the vertical line and the spinodal curve, is greater than any instantaneous pressure during the interval when the liquid is being discharged, that is

Pit) < n„ 0 < ^ < (27)

where t^j represents the time at which complete depletion of liquid occurs, then homogeneous

nucleation would be predicted. However, if P(t) > Pf^^^ for 0 < r < t^j, then homogeneous

nucleation would not occur. The above discussion is based on a pure agent. If nitrogen is used to pressurize the vessel,

nitrogen will be dissolved in the liquid agent. The homogeneous nucleation pressure will then depend

not only on the temperature but also on the composition of the liquid mixture. The effect of dissolved

gas on the homogeneous nucleation pressure of a liquid was found to raise the homogeneous

nucleation pressure (Forest and Ward, 1977). In this case, homogeneous nucleation, if it occurs, would initiate sooner in the presence of the dissolved gas during rapid isothermal depressurization.

Although the hypothesis proposed by Reid (1979) can not be substantiated or disproved at present

(Reid, 1979, Kim-E and Reid, 1983), it is prudent to recognize such a potentially hazardous event when the vessel is rapidly discharged at high temperature and pressure, and further study is warranted.

3.4.3.6 Vertically Upward Discharge. When the vapor above the liquid is vented first during

depressurization, that is, in the case of upward discharge of the vessel, the discharge process is totally

different from that of downward discharge. Visual observation of the discharge process from a

plastic vessel reveals several stages. Once the rupture disk breaks, the vapor on top of the liquid is vented out rapidly, during which time the liquid in the vessel remains quiescent. The vapor ullage

becomes foggy (r ~ 10 ms), possibly due to the condensation of the vapor as the result of cooling by

adiabatic expansion. Then, following a delay time (r ~ 25 ms), boiling is initiated from the bottom

of the vessel, and, because of bubble growth, the liquid-gas mixture level expands all the way through

the disk opening. Frothing of the liquid continues for a period of time during which time a two-phase

flow emerges from the vessel. The liquid level then settles down {t ~ 500 ms). The interior of the

vessel becomes clear, and some liquid is observed at the bottom of the vessel. At this point, most of

the liquid in the vessel has been expelled. Pool boiling of the liquid then continues until the

remaining liquid is boiled off. The complete emptying time of the liquid agent in the vessel is on the

order of seconds. The discharge process is much slower for FC-31-10 than for FC-218 because
146 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10 has a higher boiling point. The observations, in general, are very similar to those made by

Giihler et al. (1979) and Fletcher (1982) although the discharge process in both earlier studies is much slower.

Figure 73 shows a photographic sequence of events during a vertically upward discharge of

FC-218. The photographs were taken at 1000 frames per second. The frame which corresponds to

t = 1 ms is the first frame where the agent first appears outside the vessel. Due to the framing rate, there is a 1 ms uncertainty in the time. The first photograph shows the plastic vessel, a stainless-steel tube inside the vessel for introducing nitrogen in order to burst the rupture disk, vessel mount, and

the liquid level. The behavior of the agent inside and outside of the vessel during discharge can be

seen in these photographs. The discharge characteristics of the agent outside of the vessel is discussed in Section 3.5. In Figure 74 the results of pressure measurement for upward discharges of FC-31-10 are shown.

In the upward discharge configuration, the pressure transducer measured the temporal variation of the

pressure in the liquid phase. When the stainless steel vessel was used in the experiments, the pressure

transducer mounted on the side wall of the vessel was used to monitor the pressure change in the

vapor phase. Due to its relatively high boiling point, FC-31-10 does not show distinctly in the

pressure-time histories the several stages occurring in the discharge process. Pressure traces for FC-218 and halon 1301 are shown in Figure 75 and Figure 76. The several processes discussed above are clearly depicted. Upon initiation of discharge, the internal pressure

decreases rapidly. The following small pressure rise corresponds to the expansion of the liquid.

Then, the pressure decreases gradually over a period of time. This time interval corresponds to the

continuous frothing of the liquid. Eventually, an abrupt decrease in pressure corresponds to the pool

boiling of the remaining liquid in the vessel. The above interpretation is based on comparison of the

pressure traces with high-speed movie records obtained from experiments on FC-218 using plastic

vessels. Only the stainless steel vessel was used for experiments on halon 1301 due to safety

considerations. However, it is conjectured that similar behaviors occurred for halon 1301 because

similar pressure-time histories were noted.

When an extension tube was attached to the release vessel assembly exit, the behavior of the

liquid agent inside the vessel during vertically upward discharge was very similar to that without an

extension tube, as observed in the high-speed movies and noted in Figure 74 through Figure 76.

3.4.3.7 Horizontal Discharge. Observations regarding horizontal discharge of the vessel also

show several stages. Once discharge is initiated, the vapor is vented very rapidly from the vessel. Some liquid also flows out of the vessel because the initial liquid level in the horizontally mounted

vessel covers a portion of the rupture disk. The flow of vapor causes some wavy motion along the

liquid/vapor interface. The sloshing of the liquid could also be caused by the vibration of the vessel

mount. A foggy cloud is formed (r ~ 15 ms) above the liquid for a period of time. The vessel

ullage then becomes clear (r ~ 400 ms), and pool boiling is initiated in the remaining liquid pool

which still constitutes a significant amount of the initial liquid. Boiling continues until all the liquid is depleted. The emptying time is on the order of seconds. The complete discharge of the remaining

liquid is much slower for FC-31-10 than for FC-218 because FC-218 has a lower boiling point.

Figure 77 shows a photographic sequence of events during a horizontal discharge of FC-218.

The photographs were taken at 1000 frames per second. The frame which corresponds to t = 1 ms is the first frame where the agent first appears outside the vessel. Due to the framing rate, there is a 1 ms uncertainty in the time. The first photograph shows the plastic vessel (mounted horizontally) with

its exit located to the far left and the liquid level. The photographs reveal the behavior of the agent

internal and external to the vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 147

Figure 73. Photographic sequence of events showing a vertically upward discharge

of FC-218 from a plastic vessel.
148 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 73 . (continued) Photographic sequence of events showing a vertically upward discharge of

FC-218 from a plastic vessel.
Figure 73 . (continued) Photographic sequence of events showing a vertically upward discharge of

FC-218 from a plastic vessel.
150 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

0 200 400 600 800 1000 1200

t (ms)

Figure 74. Temporal variations of internal pressure during upward discharges of FC-31-10 with

and without an extension tube.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 151

1 1 1 1 1 1 1 1 1

FC-218

upward discharge

P = 4.58 MPa

P1 = 4.36 MPa \\ P = 4.34 MPa 1 —

with extension tube \

1 \ v*f

-r 'wJ'v-^ III

0 100 200 300 400 500 600 700 800 900 1000

t (ms)

Figure 75. Temporal variations of internal pressure during upward discharges of FC-218 with and

without an extension tube.
152 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 76. Temporal variations of internal pressure during upward discharges of halon 1301 with

and without an extension tube.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 153

80 ms - 100 ms

Figure 77. Photographic sequence of events showing a horizontal discharge of

FC-218 from a plastic vessel.
154 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 77. (continued) Photographic sequence of events showing a horizontal discharge of FC-218

from a plastic vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 155

Figure 77. (continued) Photographic sequence of events showing a horizontal discharge of FC-218 from

a plastic vessel.
156 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 78 and Figure 79 show the pressure-time histories for horizontal releases of FC-31-10 and

FC-218. The results were obtained from the pressure transducer that was initially positioned at a location below the liquid level. Therefore, the measured pressure traces represent a combination of

the pressure change in two stages during discharge. In the first stage, the pressure measurements

were made when the liquid level was still above the position where the transducer was located. In the

second stage, the pressure measurements were obtained when the liquid level had dropped below the

location of the pressure transducer. In other words, the pressure in the vapor phase was measured.

In Figure 78 and Figure 79, after the initial rapid decrease in pressure, the slight rebound in pressure due to initiation of boiling and bubble growth is not apparent in the case of FC-31-10, but is noticeable in the case of FC-218. Thereafter, the pressure decreases continuously until the liquid is completely boiled off". The reason for the presence of small humps in the pressure traces of FC-31-10

at 500 ms < r < 1000 ms (Figure 78) is unclear. It is conjectured that the liquid level in the vessel

during this time interval is still slightly below the location of the pressure transducer because of the

relatively slow boiling process of FC-31-10, and that any sloshing of the liquid upon the transducer

could raise the pressure reading. The shifting of the location of the hump in the two pressure traces may be indicative of such an event. Figure 78 and Figure 79 also show results for temporal variations of pressure when an extension

tube is connected to the exit of the rupture disk holder. The behaviors are very similar to that when

no extension tube is present. In the case of halon 1301, the results obtained with and without an extension tube are shown in Figure 80. Because a stainless steel vessel was used, no visual observations were made. Therefore,

it is unclear at present why the small pressure rebound is apparent in the case of an extension tube but

is not present when the extension tube is absent.

3.4.4 Conclusions. In downward discharge, the average volumetric flow rates for all the alternative

liquid agents are very similar. No internal boiling was observed for downward discharge of

HFC-236fa, FC-31-10, FC-318, HCFC-124, HFC-227ea, HFC-134a, FC-218, HCFC-22, and

HFC-125. Based on results for FC-31-10 and FC-218, the behavior of the liquid inside the vessel

during discharge was found to be very similar whether an extension tube was present or absent at the

vessel exit. A twofold increase in the initial discharge pressure decreases the emptying time of the

liquid approximately by half. The effect of a restrictor plate at the vessel exit on the discharge

process is to increase the liquid emptying time. A twofold increase in the diameter of the orifice plate increases the average volumetric flow rate by approximately fourfold. Liquid discharged at cold

temperature behaves very similarly to that at room temperature. However, when the vessel is discharged vertically upward or horizontally, the discharge process depends on the boiling point of the

agent. Agents with higher boiling points remain in the vessel for a considerable amount time before

they can be boiled oflF.

3.4.5 Recommendations. Based on the above findings, it is recommended that in order to empty

the liquid contents from the pressure vessel within tens of milliseconds, the vessel should be

discharged in a vertically downward position. If the vessel were to be designed to increase release

rates and to withstand higher internal pressure, higher initial charge pressures are also recommended.

In addition, further study of discharge at high temperatures is needed because of concern regarding

boiling liquid expanding vapor explosion (BLEVE).
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 157

0.6

1 1 1 1 1 1

FC-31-10

horizontal discharge

?.= 4.03 MPa

with extension tube

P.

1 = 4.18 MPa \\

1 Pj = 4.33 MPa \ \

\ p_ = 4.46 MPa

\ with extension tube

1 1 1 1 1 1 0 200 400 600 800 1000 1200 1400

t (ms)

Figure 78. Temporal variations of internal pressure during horizontal discharges of

FC-31-10 with and without an extension tube.
158 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

1-2 j

1 \ 1 \ f—

FC-218

horizontal discharge

0.8 -

Oh" 0.6 - P = 4.27 MPa

i 0 200 400 600 800 1000

t (ms)

Figure 79. Temporal variations of internal pressure during horizontal discharges of FC-218 with

and without an extension tube.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 80. Temporal variations of internal pressure during horizontal discharges of halon 130i

with and without an extension tube.
160 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

3.5 Measurements of Agent Dispersion

3.5.1 Introduction. The purpose of this part of the research effort was to characterize experimentally the dispersion and evaporation behavior of the eleven potential replacement fire extinguishing

agents. The design of the experiments was based on the reasonable hypothesis that effective fire extinguishment requires both rapid dispersion of the agent within a volume as well as rapid gasification following release of the agent from a pressurized bottle. These properties are expected to be

crucial for both the dry bay and engine nacelle applications. However, they are particularly relevant

for the dry bay application due to the extremely rapid fire suppression times (on the order of tens of

milliseconds) required.

The agent must be dispersed throughout the volume to ensure that all regions of flame are suppressed. This requires that the agent not only be rapidly released from the bottle, but also that it be rapidly and effectively mixed with the surrounding gases. Rapid vaporization of the liquid is desirable for two reasons. While the liquid may, in some circumstances, be dispersed more rapidly

as a two-phase mixture, it is also true that it is unlikely to be dispersed evenly throughout the volume.

Since the fire location cannot be specified, it is necessary to have full vaporization to ensure a more

uniform distribution. The proposed halon replacement agents are not expected to be as effective as halon 1301 for these fire fighting applications since they do not have the effective chemical suppression mechanism associated with the presence of bromine atoms. Principal mechanisms for extinguishment are therefore likely to be heat extraction and dilution as well as the physical effects associated

with the rapid release of the agent (i.e., flame extinction due to stretch). Complete vaporization ensures the maximum interaction between the agent and the fire zone. With the exception of one of the chemicals (FC-1 16), all of the proposed alternatives are liquids

when stored under pressure at room temperature and are superheated when released as liquids into the

ambient atmosphere. As a result, partial rapid vaporization is to be expected when these agents are released into an ambient environment at room temperature. A review of the literature indicated that while there have been investigations of flashing liquids reported, none of these previous experiments

dealt with the rapid, short-time releases which are characteristic of the current study. Here we

summarize the literature which was reviewed during the course of this effort in order to provide an

overview of the existing understanding and to provide a basis for later discussion.

One of the earliest investigations of released superheated liquids was the study of Brown and

York (1962). These authors investigated sprays formed by the release of water, CFC-11 and water

saturated with carbon dioxide. They noted that the flashing of a superheated liquid or a pressurized

liquid with dissolved gases provides an additional mechanism for the breakup of a liquid stream (in

addition to the classical mechanisms resulting from surface tension forces (Rayleigh, 1878) and

aerodynamic forces (Weber, 1931)). They pointed out that for adiabatic conditions the flashing of a superheated liquid would occur by the extraction of latent heat from the liquid to provide the energy

necessary to evaporate the liquid. Thermodynamic equilibrium is attained when the fraction of liquid

converted to gas is suflRcient to cool the remaining liquid to it saturation pressure. Results reported included the distance required for the liquid jets to break up as well as characterization of the sprays generated. Several of their conclusions are relevant to the current

investigation. They reported that a liquid jet does not "shatter" abruptly when the liquid just becomes

superheated, but that a well defined amount of superheating is required for shattering to occur. A

dependence of the flashing behavior on the jet diameter was characterized by correlating the liquid

breakup times with the Weber number (We), defined as
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 161

P.UD (28)

la

where p„ is the density of the gas surrounding the liquid jet, is the jet velocity, D is the jet

diameter, and a is the liquid surface tension. The times required for jet breakup increased with

decreasing D. A transition in the behavior was observed for We = 12. Brown and York argued that the times required for flashing of pressurized liquid jets containing

dissolved gases should be considerably longer than those for superheated jets due to molecular

diffusivity considerations. Experimental evidence was provided which supports this conclusion.

Measurements of drop sizes generated by the flashing liquids showed that flashing provided an

effective mechanism for atomizing the liquid. The flashing of the liquids resulted in a rapid

expansion of the flow. Lienhard and coworkers (Lienhard, 1964; Lienhard and Stephenson, 1966; Lienhard, 1966;

Lienhard and Day, 1970) developed a model for predicting the time required for superheated liquids

to flash and compared the findings with experiment. The first of these papers (Lienhard, 1964)

considered the conditions for which bubbles growing in a liquid are stable to growth or collapse.

Results were based on thermodynamic arguments. Lienhard and Stephenson (1966) considered the

kinetics of bubble growth. Their analysis was based on equations for bubble growth developed earlier

by Deragarabedian (1953), Forster and Zuber (1954), and Plesset and Zwick (1954). In order for bubbles to grow spontaneously they must attain a critical size. For smaller sizes, bubbles tend to redissolve into the liquid. Nuclei bubbles therefore arise due to fluctuations in the local environment

which favor their formation. When such flucmations result in a bubble with suflftcient size to continue growth, a large bubble will form. Generally the formation of a larger bubble will "trigger"

the growth of additional bubbles. Using arguments based on the distribution of disturbances in the

liquid, Lienhard and Stephenson derived an approximate relation for the period required for a superheated liquid to flash which can be written as

^ ^' (29)

where is the delay time for bubble growth, A is the cross sectional area for the liquid flow, is the vapor pressure for the liquid at its release temperature, and P^,„^ is the ambient pressure. Note

that Py-Pamb ^ measure of the superheat based on pressure. Experimental results were presented

for the flashing of superheated water which supported the use of this equation.

In the same paper these authors derived an expression for the amount of work that an isobaric

expansion of a superheated liquid can do. For the case where the saturation temperature for the

liquid, Ti, is much greater than the superheat expressed in temperature terms, TfT^^^, the resulting

expression is

w = C ^IlI^, (30)

" 1T_^,

where w is the amount of work which can be done by the flashing and is the constant pressure

heat capacity which had been assumed to be constant. The above analysis for the flashing time was extended by Lienhard and Day (1970) [see also Day

(1969)] in order to derive a more general equation applicable to dilTerent superheated fluids. They
162 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

argued that the delay time for bubble growth should scale with the density of the saturated liquid, p/, D, a, and the superheat, Py-Pamb- Using the Buckingham Pi-Theorem, they reduced these variables

to two dimensionless variables which should correlate the experimental data. These were a dimensionless dwell time, 4>,

(P -P yf'^Dt, $ s 1^ (31)

where is the liquid density, and a dimensionless jet diameter,

Y ^ ^^^v'^aml) (32) a

By using Equation (29) and eliminating and {Py-Pamh^ ^^'^^ equations one finds that

= constant (32)

Once the constant in Equation (32) is determined experimentally, it should be possible to predict t^ for any superheated fluid. Based on measurements of delay times for jet flashing with water and

liquid nitrogen, the constant in Equation (32) was found to equal 2.12 x 10^-^. This paper also has a

detailed discussion of the conditions under which liquid jet break up occurs as the result of capillary

and aerodynamic instability. In a brief paper, Lienhard (1966) considered how the spreading of a flashing liquid jet can be

understood in terms of the amount of work which the flashing can do on the system (Equation (30)).

The analysis was done in terms of the radial velocity, U^, imparted to the spray by the flashing. The

maximum such velocity, U^^, is

Ur.max

( c Y'^ T\ sat J

Lienhard argued that all of the available work would not be used to impart radial momentum and that only a fraction, C, would be available. Based on experimental spreading angles for flashing

superheated water jets he estimated that C = 0.04.

Solomon et al. (1982) reported limited measurements for a flashing jet of CFC-11 and

superheated fuels. They found that flashing due to small amounts of superheating was very efl'ective in generating atomized liquid flows. Celata et al. (1982) have provided a very brief description of experiments on the release of

superheated water. They found that flashing took place very rapidly, and that the spreading angle of

the two-phase flow increased with increasing superheat. Using a probe, they recorded dynamic

pressures within the flow field. Unlike previous investigations, they concluded that evaporation was

occurring on the outside of the flow and that there was a narrow liquid core in the center of the flow

surrounded by a two-phase shell. Moodie and Ewan (1990) provide a nice introduction to the release behavior of liquid jets. They

discussed the behavior of flashing jets and showed a photograph of a flashing CFC-1 1 jet. The

analysis was complicated because flashing seemed to occur within the vessel orifice. They measured
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 163

the peak droplet sizes which suggested that the droplet size decreased with increasing superheating.

Time-averaged temperature measurements were also provided. The temperature profiles seem to suggest that the entire jet is flashing unlike the findings of Celata et al. (1982). The temperature

measurements also indicate that the entrainment of air plays a critical role in the ultimate evaporation

of the superheated liquid.

These authors also made mass fraction measurements of CFC-11 and air in the flow field. These

measurements also suggest the two-phase flow covered a large extent of the flow. Velocity measure- ments were also reported. These measurements showed that measured flow velocities were considerably higher than expected based on liquid flow rates at the vessel exit Radial velocity profiles in the

region near the vessel were non Gaussian, but tended to attain a Gaussian profile as the measurement

location was moved further downstream.

Superheated releases of CFC-U, chlorine, methylamine, and cyclohexane have been investigated

by Johnson and Diener (1991). In these experiments the liquid fraction of the two-phase flow was

capmred. The data indicated that there were well defined superheat points for which the liquids

flashed near the vessel exit. Under these conditions the amounts of liquid trapped decreased

significantly. Comparisons were made with a model designed to predict the behaviors. In general,

the agreement was not very good.

There are a variety of systems employed for the release of halon 1301 into dry bays and nacelles.

It was beyond the scope of this study to investigate a wide range of release configurations. For this reason, the decision was made to characterize the spreading and mixing characteristics of the agents

following release from a generic system based on the vessel developed to study the release rate of

agents (see Section 3.3).

3.5.2 Experimental Methods. As mentioned above, there were very few investigations available in the literature to guide the design of a system for characterizing the dispersion and evaporation

behavior of the agents. As a first step in the design process a high-speed video system^ was used to record a downward release of HCFC-22 from the vessel at a rate of 1000 frames/s. A burst disk

providing a nominal pressure of 4.1 MPa was employed and the 0.0005 m^ vessel was filled one-third

full. A series of sequential photographs from the high-speed video are reproduced in Figure 81.

Lighting for the camera was provided from the rear and front of the release. As can be seen from Figure 81, the mixing process is extremely complex. When the burst disk

first opens, a flashing is observed which results in a rapid expansion of the agent in both the

downward and radial directions. The lighting from behind the flow is attenuated by the flow. Shortly afterwards a vortex structure forms at the downstream edge of the release and the flow field looks very much like an impulsively started axisymmetric gas jet. Such structures have been

observed in liquid jets (Shirakashi et al., 1984) and combustion torch ignition experiments (Cattolica

and Vosen, 1987). As the flow develops behind the leading edge it has a nearly linear spreading

behavior. This is true near the orifice as well as the far field. The linear expansion seems to be

initiated immediately at the orifice of the vessel. At 13 ms into the release a second flashing starts near the vessel orifice and one ms later this flashing has resulted in a very rapid radial expansion.

Analysis of the emptying rate of the vessel for these conditions indicates that the second flashing

occurred very close to the time when the liquid was completely released from the vessel. Following

Kodak Ektapro 1000 Motion Analyzer. The authors thank Michael Howachyn of Kodak for arranging a demonstration of this instrument and for allowing us to record an initial series of highspeed images of the releases.
Figure 81. Sequence from a high-speed video (1000 frames/s) of the flow formed by a

release of 245 g (1/3 full) of HCFC-22 from the 0.0005 m^ test vessel. Time

increases to the right and from top to bottom.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 165

the second flashing, the release becomes obviously slower and within 30 ms the flow has cleared up

and the backlighting is again transmitted.

Additional high speed films indicated that the double flashing behavior identified from Figure 8

was typical for downward releases of all of the agents, with the exception of FC-116 which was not

investigated since this agent is not a liquid under the storage conditions. All of the agents resulted in nearly total attenuation of a laser beam when they were present in a two-phase state. This observation ultimately determined the type of diagnostics which could be employed to characterize the mixing

and vaporization behavior.

Our initial intention was to record the concentration of agent as a function of time using either

Rayleigh light scattering (gas concentrations only) (Pitts and Kashiwagi, 1984) or aspirated hot-film

probes (Brown and Rebollo, 1972). However, the early visualization studies indicated that these

techniques are inappropriate due to the two-phase nature of the flows and their high optical density.

A consideration of possible diagnostics revealed that there were no techniques available which would

allow quantitative concentration measurements or two-phase flow characterization over the short time

periods required.

The discussion above indicates that the two most important flow properties determining the

ability of an agent to extinguish a flame are rapid mixing and agent vaporization. A simple, but

effective, experimental system was developed to characterize the agents with regards to these two

variables. This system combined high speed flow visualimtion, laser attenuation measurements,

dynamic pressure measurements, and an aspirated hot-film concentration probe. To our knowledge,

this is the first time that such a system has been used to characterize the behavior of the short period,

two-phase flow generated by a release from a pressurized bottle. Figure 82 shows a schematic of the

overall system.

The system was used to characterize the flows formed by releases of the agents from the vessels

developed to study their release rates (see Section 3.3). Data were recorded for the various release

conditions performed during these investigations (see Section 3.4).

3.5.2.1 High-Speed Filming. The same high-speed camera which was used to record the fluid

level inside the transparent vessel (see Section 3.4.2) simultaneously recorded the release of the agent

in regions near the vessel exit. This was accomplished by placing the camera such that it recorded

the liquid behavior inside the transparent vessel as well as locations immediately downstream of the

vessel exit. The field of view imaged by the camera could be varied by adjusting the camera

placement and distance from the vessel. Most data were recorded with a field of view of roughly

0.40 m width and 0.10 m height.

Selected frames of these films were used to estimate the axial and radial velocities of the agents

immediately following release. The approach used image analysis (see Section 3.4.2) to locate the

downstream and radial edges of the release at different times and calculate the corresponding velocity.

The radial velocity was determined for the downstream location (13 mm from orifice) where the

pressure transducer (see ahead) was located. Due to the relatively small area of the near-field release

which was imaged, normally only two such measurements could be made before the agent covered an

area larger than the field-of-view. Note that frame 1 is defined as the frame where it is clear that the

release of the agent has begun. As a result, the time for frame 1 can be anywhere from 0 to 0.49 ms

(for a framing rate of 2000 frames/s) following the bursting of the disk.

3.5.2.2 Laser Attenuation Measurements. This part of the system was designed to measure

the rate at which the downstream edge of the released agent moved away from the vessel exit and to provide estimates for the period required for the agent to evaporate. Five laser beams were aligned

downstream of the vessel exit and perpendicular to the flow direction. Simple helium neon lasers
166 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

detector

detector

detector

detector

detector

nitrogen

thermocouples X

^^^^^>^ solenoid valve

pressure

transducer needle

valve

laser 1

laser 2

laser 3

laser 4

laser 5

pressure transducer hot-film probe

Figure 82. A schematic of the experimental system is shown. Helium-neon lasers and silicon

photodiodes are used for the extinction m.easurements. The piezoelectric pressure

transducers and aspirated hot-film probe are described in the text.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 167

(Melles Griot 05-LLR) were mounted on a vertical strut located 1.3 m from the vessel such that the

spacing between lasers was nominally 0.3 m. The laser closest to the vessel was positioned across the center of the orifice within 1 mm of the rupture disk holder.

After traversing the flow field the lasers were detected by a series of five photodiodes. The

photodiode amplifier circuits and mountings were based on a design by Bryner (1993). Each detector was a silicon photodiode (Hamamatsu S1337-1010BQ) having an active surface area of 100 mnr. An

operational amplifier circuit (Figure 83) was used to generate a voltage proportional to the light

intensity striking the photodiode. Each photodiode and amplifier was placed in a rigid mount with a connector for power input and signal extraction. The mount was drilled and tapped so that a rod

could be attached for positioning the photodiode. A tube having an inside diameter of 22 mm and a

length of 0.23 m was placed in front of the detector in order to shield the photodiode from ambient

light in the room. In some cases neutral density filters were used to ensure that the outputs of the

photodetectors were not saturated. Circuits housed in a single box were built to provide power for the photodiodes and signal outputs on BNC connectors for a maximum of eight detectors. For the

experiments reported here only five detectors were used.

The photodetectors were mounted on a vertical strut located on the opposite side of the room

from the lasers. Each photodiode system was aligned with one of the laser beams in such a way that the beam was centered on the active area. Figure 84 is a photograph of the photodiode detectors. The voltage outputs for the photodiodes (numbered 1 to 5 with 1 closest to the vessel) were

recorded using the data acquisition system described below. Tests showed that laser ff\ was totally

attenuated ver>' rapidly (in less than 40 /xs) when the disk burst. The attenuation of this laser beam

was used to define time 0 for the experiments. It is interesting that this laser beam was normally

found to be totally attenuated several hundred microseconds before a pressure drop was recorded by

the pressure transducers within the vessel.

3.5.2.3 Dynamic Pressure Measurements. Two piezoelectric pressure transducers were used

external to the vessel to record the dynamic pressure of the releases. The transducers were Kistler

603B1 models used with Model 5004 amplifiers. The amplifiers included 180 kHz frequency filters and were operated in the "long" integration mode to allow the pressure to be tracked over the

relatively long time (generally several hundred milliseconds) of a release. The output of the amplifier

is a voltage proportional to the pressure. Usually a range of 687 kPa/V was used.

One of the pressure transducers was located near the vessel exit at a distance 13 mm downstream

of the rupture disk holder of the vessel and 46 mm from the centerline defined by the center of the

orifice. It was oriented so that it was perpendicular to the centerline. Note that the orifice was

nominally 19.1 mm in diameter, and the pressure transducer was located 4.8 orifice radii from the

centerline.

The second pressure transducer was positioned along the centerline of the release at a downstream position just beyond the point where the fifth laser beam crossed the centerline (1.3 m

downstream of the vessel exit). It was oriented perpendicular to the flow direction.

Both pressure transducers were mounted in the center of conically shaped pieces of aluminum

designed to protect the transducers and, at the same time, minimize their interaction with the flow

field. The near-field transducer and its mount were attached to a rod which allowed positioning of

the transducer. This pressure transducer and its mount can be seen in the early frames of the highspeed films which are shown later. The downstream transducer and its mount were attached to a

plate which was capable of surviving the large dynamics pressures sometimes generated by the flow

fields at this position. Figure 85 is a photograph showing the pressure transducer mounted in the

experimental system.
168 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Incident

Light

10 picofarad

capacitor

Feedback Resistor 50 k Q

Output

Signal

Figure 83 . The circuit used to power a photodiode and generate a voltage proportional to the current generated by light striking the detector. The output is a signal proportional to the laser intensity reaching the detector.
Figure 84. A photograph of the array of five photodiodes and mounts used to record trans- mitted laser intensities across the flow. The nominal spacing between detectors is 0.3 m.
170 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 85. Photograph of the dynamic pressure transducer (left) and aspirated hot-film (right)

located 1.3 m downstream of the agent vessel. Probes are placed in metal

housings for protection and to provide a means for mounting to an aluminum

plate.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 171

An important point with regard to the use of these transducers is the expected response of the

transducer to the flow. Davis (1980) used pitot tubes to make similar measurements of dynamic

pressure in a gas-liquid mixture flow. He reported that pressure probes respond partially to the

incident flow momentum flux with pressure rises of

where is the measured dynamic pressure, e is a factor which varies between 1.0 and 1.85, p,„ is the density of the mixture, and U is the local flow velocity. Note that Equation (34) reduces to the

standard Bernoulli equation form for a pitot tube (Roberson and Crowe, 1980) when e equals 1. Davis argues that the variation in £ is the result of the relative motion of the gas and liquid in the

region of the probe tip. For the case of the piezoelectric transducers used in the current investigation,

the stagnation point is very close to the probe and such efl'ects should not be important. Equation

(34) with £ = 1 will therefore be used.

3.5.2.4 Aspirated Hot-Film Probe. The aspirated hot-film probe is a device which is designed

to record real-time concentrations in isothermal binary gas mixtures. The response time depends on a number of system parameters, but is generally less than 1 ms. The active part of the device consists

of a thin quartz rod which is coated with platinum. A small current is passed through the platinum

film and resistively heats the surface. The resistance of the device changes with temperature.

Feedback electronics are provided which maintain the resistance, and hence the temperature, constant. This is known as the constant-temperature mode. A bridge circuit provides the current necessary to balance the heat losses from the rod. By monitoring this current it is possible to characterize these

heat losses which vary with the flow velocity over the rod, the thermal conductivity of the gases

surrounding the rod, and the fluid temperature. (Bradshaw, 1971)

The most common use of hot-films is for velocity measurement in a single fluid where proper

calibration allows the measured heat losses to be related to velocity. Brown and Rebollo (1972)

showed that these devices could be utilized for accurate concentration measurements in isothermal

binary gas mixtures by placing the hot-film in the flow generated by extracting gases through a small

orifice with a sufficient backing pressure that the gas becomes sonically choked. In this case the

velocity only depends on the fluid properties. Since the flow velocity is always constant for a given

composition, it is possible to calibrate the instrument with mixtures of the two gases having a range of

concentration, and then make accurate concentration measurements in unknown mixtures.

The use of an aspirated hot-film for concentration measurements in the current investigation is complicated by the variations in temperature which occur as the result of the vaporization of the

superheated agents and the presence of two-phase flow, which dramatically alters the heat loss behavior of the fluid. For this reason, in the current investigation, these probes were not used for quantitative measurements, but rather provided qualitative indications of the presence of an agent and

whether or not a single or two phase fluid was present.

The system employed was commercially available from TSI, Inc. It consisted of a Model

1440-20 aspirating probe used in conjunction with a Model 1053B constant temperature anemometer

and Model 1051-2 power supply. The necessary vacuum source for the probe (7 1/m @ 33 kPa) was

provided by a small vacuum pump. The probe was placed in a conically shaped aluminum holder to protect it and allow easy mounting. It was positioned near the centerline just downstream of the fifth laser beam by mounting on a plate placed in the flow. It can be seen in the photograph in Figure 85.

The voltage output of the anemometer circuit was recorded in the manner described below.
172 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

3.5.2.5 Data Acquisition System. The data acquisition requirements for this experiment were

demanding. Counting the sensors employed within the vessel, a maximum of eleven sensors were

used for a single run. The processes of interest occur at submillisecond time scales. A data

acquisition rate of 25 kHz per channel was ultimately used. Note that this corresponds to a total maximum data rate of 275 kHz. Total data collection times varied up to a maximum of 2 s. The design of the agent release vessel presented an additional difficulty. Due to variations in the

pressure required to rupture the disk and the need to flow the pressurization nitrogen into the vessel at a reasonably slow rate, it was not possible to know exactly when the experiment would be initiated. This meant that it was necessary to "trigger" data acquisition using some property of the experiment

itself. It was also desirable to record some data prior to the initiation of the experiment in order to provide baselines. This was done by continuously recording the data until a trigger was received and

then saving a specified number of data points before and after the trigger signal. This type of data

acquisition is referred to as "pretriggering."

The system chosen for the experiments was the Flash 12-Model 1 digitizer from Strawberry

Tree. This system provides 16 channels of 12 bit data acquisition at a maximum total data rate of 1 MHz. Voltage ranges can be specified for each channel, and provisions are available for making

thermocouple measurements. A daughter board was attached which provided a total high-speed

memory of 1 Mbyte. The board was mounted in a 486 personal computer, and data acquisition

functions were controlled using Workbench 4.0 software.

The system was operated in the pretrigger mode with the trigger signal provided by the near-total attenuation of laser #1 which was located immediately downstream of the vessel orifice. All of the

transducer outputs described above as well as those internal to the vessel (see Section 3.4.2.) were

connected to the digitizer, and the voltage ranges adjusted to match the expected signals. Following

the conclusion of an experiment, the data was stored for later analysis using SigmaPlot 5.0 software.

3.5.2.6 Test Conditions. These experiments were performed in conjunction with the release

experiments summarized in Section 3.4. The majority of these experiments were downward releases

at room temperature using nominally 4.1 MPa burst disks and a vessel orifice of 19.1 mm.

However, additional runs were performed with a cooled vessel exhausting into a room temperature

environment, using burst disks set for higher and lower pressure, and smaller orifices. In a few

cases, a half-meter extension tube was attached to the vessel. In this case, the near-field pressure

measurements were not made, but the transmitted intensities of the three downstream lasers, the

downstream dynamic pressure, and the aspirated hot-film response were recorded. Some measure- ments were also made for upward releases. In this case laser #5 was placed just above the vessel

orifice and provided the timing for the experiment.

3.5.3 Experimental Results

3.5.3.1 High-Speed Filming. The high-speed films provide excellent qualitative insights into the behavior of the agents upon release from the pressure vessel. Figure 86 and Figure 87 show time

series taken from films recorded at 2000 frames/s for releases of FC-31-10 and halon 1301. These

films were recorded primarily to monitor the behavior of the released agent. The region imaged is 0.19 m (downstream) x 0.33 m (radial). Remarkable diff'erences are observed in the mixing

behaviors of the two agents.

The behavior of the FC-31-10 (Figure 86) is summarized as follows. At time 0 the disk bursts

and the agent appears as a white plume. During the next 3 ms the agent spreads rapidly in both the

radial (maximum of 0.14 m) and axial directions (0.17 m). The flow then develops a mushroom

shape which is characteristic of an impulsively started axisymmetric jet. As the flow further
Figure 86. A sequence from a high-speed film (2000 frames/s) of the flow formed by a release

of 386 g of FC-31-10 from the vessel. The times indicate the period since the

initiation of the release.
174 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 175

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTION

il.O ms

3. FLUE) DYMAl./nCS OF AGENT DISCHARGE

11.5 ms

Figure 86. (continued) A sequence from a higii-speed fiirn (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 177

14.0 ms . 14.5 ms

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
178 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

17.0 ms 17.5 ms

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 179

20.0 ms 20.5 ms

Figure 86. f continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
180 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 181

24.0 ms 24.5 ms

26.0 ms ' 26.5 ms

j

i Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

! by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

\ smce the initiation of the release.

\
182 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

29.0 ms 29.5 ms

Figure 86. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 386 g of FC-31-10 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 183

Figure 87. A sequence from a high-speed film (2000 frames/s) of the flow formed by a release

of 270 g of halon 1301 from the vessel. The times indicate the period since the

initiation of the release.
184 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 87. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 185

Figure 87

8 0 ms 8.5 ms

(continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
186 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 87 (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3 FLUID DYNAMICS OF AGENT DISCHARGE

Figure 87

14.0 ms 14.5 ms

(continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
188 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 87

17.0 ms 17.5 ms

(continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 189

20.0 ms 20.5 ms

Figure 87. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
190 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 87 (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 191

Figure 87. (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
Figure 87 (continued) A sequence from a high-speed film (2000 frames/s) of the flow formed

by a release of 270 g of halon 1301 from the vessel. The times indicate the period

since the initiation of the release.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 193

develops, the neck of the flow grows thinner until at roughly 6 ms following release the agent is exiting the vessel as a liquid stream which exists over the entire observed downstream length. At

approximately 10 ms partial flashing of the liquid jet occurs near the orifice and subsequently grows

in time and space. Even after 15 ms the two-phase flow is compact and has not spread substantially over the observed downstream distance. At later times the jet begins to spread slowly until a very

strong expansion occurs 21 ms after release. This point corresponds closely to the emptying time

estimated from pressure measurements recorded internal to the vessel. This strong flashing behavior

lasts a few milliseconds before only condensed vapor in the nitrogen flow is observed to exit the

vessel.The halon 1301 (Figure 87) is released in a very different way. The initial release of the agent

from the bottle is much more violent than observed for the FC-31-10. The flow develops fingers

which look almost like an explosion. By 3.5 ms the agent covers a wider radial extent (approximately 0.29 m) than it has moved downstream. The flow near the nozzle seems to be assuming a conical

shape. The jet continues to spread in the radial direction as the flow field moves beyond the farthest observed downstream position. For the next 8 ms the flow seems to settle down into a flow condition

for which the flow leaves the nozzle and spreads rapidly with downstream distance. The spreading

angle is roughly 60°. During this period the material ejected to large radial positions at earlier times seems to have evaporated. This behavior can be contrasted to that for FC-31-10 where the released

agent appears to form a liquid stream during this phase of the release. Fourteen milliseconds after the

release a very strong flashing is observed which disperses material over extremely wide radial distances for short downstream distances. Again, internal pressure measurements indicate that this is roughly the time when all of the liquid has been emptied from the vessel. Films were recorded for all of the agents over the smaller 0.40 m x 0.10 m view area. Table 5

lists measured values of radial and axial velocities calculated from the films for the ten proposed

alternative agents which are liquid at room temperature and halon 1301. Due to the limited number

of measurements the data do not show definite trends. However, it is clear that the axial near-field

velocities increase with decreasing boiling point (i.e., the amount of superheating).

3.5.3.2 Near-Field Dynamic Pressure Measurements. These measurements were intended to

test two aspects of the released agent behaviors. The first was whether or not there were shock

waves generated by either the bursting of the disk or the sudden flashing of the liquid. The second was to detect the presence of the agent at the transducer and, if possible, estimate the degree of

vaporization.

The high-speed films showed that during an agent release there were two flashings which might

expel liquid to a distance where it could strike the pressure transducer. Time records of pressure

traces all show signals corresponding to the second flashing near the end of the release. The lower

boiling-point liquids also showed pressure rises shortly after the disk burst (recall that time 0 is defined as the time when laser ffl is strongly attenuated). Interestingly, no pressure changes were

generally detected for the higher boiling-point liquids at short times.

Figure 88 shows a time plot of pressure from the near-field pressure transducer for two room

temperature releases of HFC-236fa from a vessel which is 2/3 full of the liquid agent. The two sets of data are in good agreement. No pressure increases are observed until near the time when the fluid

is emptied from the vessel. Apparently, the flashing observed in the high-speed films when the disk

burst did not generate shock waves or impart sufficient momentum to the HFC-236fa to reach the

pressure transducer which was 46 mm from the centerline. In contrast, the strong flashing observed

near the ends of the releases generated a very large dynamic pressure at the transducer. Table 6 lists the maximum pressures observed for the two releases.
194 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

600

0)^ 500

400

300

200

100

0

HFC-236fa

1 1

-

1 1 1 1

/ \\w

B

— ^l||^JHBjUUI||l^|Lu||jL| 1

_ A 1

1

1 1 r 1 1 0 25 50 75 1 00 1 25 150

Time (ms)

Figure 88. Two time traces of near-field dynamic pressure for releases of HFC-236fa. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 464 g of agent, release

pressure of 4.30 Mpa; B, 471 g

agent, release pressure of 4.51 MPa.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 195

Table 5. Values of Velocity Determined from High Speed Films of Liquid Agents Released from

Pressurized Vessels

AGENT

Radial Velocity

Frame 2-3

(m/s)

Radial Velocity

Frame 3-4 (m/s)

Axial Velocity

Frame 2-3 (m/s)

Axial Velocity

Frame 3-4 (m/s)

HFC-236fa 9

a 42 57

FC-31-10 12 2.4 49 a

FC-318 14 0 67 a

HCFC-124 10 0 57

a

HFC-227ea a 2 55 57

HFC- 134a 10 5 69 63

FC-218 6 11 86

a

HCFC-22 a a 58.1 49.3

HFC- 125 2.1 18 76

a

HFC-125/HFC32 mixture17 9 80

a

halon 1301 11 1.5 > 80

a

^ Measurements could not be made for these conditions due to poor contrast in the images

A number of releases with various vessel fill levels have been made for FC-31-10. Figure 89

shows time plots of pressures from the near-field pressure transducer for three runs having initial vessel fill levels of 1/3, 1/2, and 2/3. The observed behaviors are very similar to those found for HFC-236fa. Table 6 summarizes the pressure increases observed during the second flashings for a number of experiments. There are no clear trends in the maximum pressure with burst pressure or

the mass of agent in the bottle. A range of values are observed which are likely to be characteristic

of run-to-run variations.

Figure 90 shows two pressure time traces for FC-318. The results are similar to those for HFC236fa and FC-31-10 except that one of the pressure traces has a slight jump with a maximum of 41

kPa starting 3.48 ms after the agent release. This pressure rise indicates that for this release the

initial flashing behavior was strong enough to eject material as far as the pressure transducer. A total of six experiments were run with this agent. Only one recorded the initial flashing behavior. Pressure increases associated with the second flashing are listed in Table 6. A fairly wide range of

pressures is observed.

Two time traces for the near-field pressure behavior are shown in Figure 91 for HCFC-124. No

signals were recorded immediately following the opening of the vessel orifice. The relevant data for observed dynamic pressures as a result of the second flashing are summarized in Table 6.
196 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 6. Maximum Dynamic Pressures Observed by the Near-Field Transducer During the

Second Flashings of Agent Releases

A (jpnt Tnfprnjil PrpQQiirf=» at Release (MPa)

IVldAllIlUIIl LfylldllllL Pressure (kPa)

HFC-236fa 464 4.30 499

471 4.51 480

FC-31-10 261 4.50 314

262 4.45 293

263 3.98 364

266 4.33 314

386 4.21 448

387 4.52 397

495 3.93 557

515 4.20 312

520 3.85 315

FC-318 257 3.90 487

270 4.25 469

404 3.92 516

501 4.43 647

505 3.99 561

521 4.34 629

HCFC-124 465 4.30 708

462 4.05 668

HFC-227ea 476 4.45 900

489 4.41 792

HFC- 134a 419 4.21 1176

418 4.13 1101

423 4.38 1317
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 197

Table 6. (continued) Maximum Dynamic Pressures Observed by the Near-Field Transducer

During the Second Flashings of Agent Releases.

Agent Mass of Agent (g) Internal Pressure at Release (MPa)

Maximum Dynamic

Pressure (kPa)

FC-218 468 4.28 559

463 4.06 421

460 4.58 579

HCFC-22 412 2.81 572

421 4.48 1402

417 4.52 1271

HFC- 125 423 4.24 1047

403 4.19 367

416 4.07 566

HFC-125/HFC-32 375 4.36 1508

mixture

358 4.31 1626

halon 1301 270 4.43 602

554 4.36 113

Figure 92 and Table 6 contain similar results for HFC-227ea. One of the releases showed a

quite substantial pressure increase starting 1.12 ms after the initial release of the agent. The pressure

reached a maximum of 304 kPa.

Results for HFC- 134a are given in Figure 93 and Table 6. Three sets of data were recorded for

this agent. Two of these showed pressure increases (169, 102 kPa) shortly after (1.80 ms and

1.52 ms) the initiation of the experiments. The experiment which isn't shown in Figure 93 did not have a pressure increase at the short delay time.

Three sets of data were also recorded for FC-218. Figure 94 and Table 6 summarize the

findings. As found for the HFC- 134a, two of the releases resulted in small pressure increases (28

and 31 kPa) shortly following the release time (1.92 and 1.60 ms). As can be seen by comparing

entries in Table 6, the pressure increases as a result of the second flashing observed for the FC-218

are considerably smaller than those found for the HFC-134a.

HCFC-22 generated pressure traces which were quite complicated and variable. Examples of

two of the traces are shown in Figure 95. The pressure time histories are quite different than those

for the agents discussed up to this point. Not only are pressure increases observed near the beginnings and ends of the releases, there also indications of some flashing behavior at intermediate times.

Table 6 summarizes the findings for three releases of HCFC-22. The dynamic pressures observed for the releases using nominal 4.14 burst disks were the highest recorded during the study. The pressure

increase was found to be considerably smaller for the one case using a lower-pressure burst disk (see

Table 6).
198 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

J L 0 25 50 75 100 125 150

Time (ms)

Figure 89. Three time traces of near-field dynamic pressure for releases of FC-31-10.

Traces B and C are offset by 50 kPa. Conditions: A, 515 g, release pressure

4.20 MPa; B, 386 g, release pressure 4.21 MPa; C, 263 g, release pressure of

3.98 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE

FC-318

700 r-n -n \ \—

0 25 50 75 100 125 150

Time (ms)

Figure 90. Two time traces of near-field dynamic pressure for releases of FC-318. Trace B

is oifset by 100 kPa. Conditions: A, 501 g, release pressure 4.43 MPa; B,

270 g, release pressure 4.25 MPa.
200 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HCFC-124

900 r-n \

i I

Time (ms)

Figure 91 Two time traces of near-field dynamic pressure for releases of HCFC-124. Trace

B is offset by 100 kPa. Conditions: A, 465 g, release pressure 4.30 MPa; B,

462 g, release pressure 4.05 MPa.
SECTION 3. FLUTD DYNAMICS OF AGENT DISCHARGE

HFC-227ea

1000 r-1 , ^ ^

0 25 50 75 100 125 150

Time (ms)

Figure 92. Two time traces of near-field dynamic pressure for releases of HFC-227ea. Trace

B is offset by 100 kPa. Conditions: A, 476 g, release pressure 4.45 MPa; B,

489 g, release pressure 4.41 MPa.
202 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HCFC-134a

0 25 50 75 100 125 150

Time (ms)

Figure 93. Two time traces of near-field dynamic pressure for releases of HFC- 134a. Trace

B is offset by 100 kPa. Conditions: A, 423 g, release pressure 4.38 MPa; B,

418 g, release pressure 4.13 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 203

FC-218

700 r—} \ \ r—

0 25 50 75 100 125 150

Time (ms)

Figure 94. Two time traces of near-field dynamic pressure for releases of FC-218. Trace B is offset by 100 kPa. Conditions: A, 463 g, release pressure 4.06 MPa; B, 468 g, release pressure 4.28 MPa.
204 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HCFC-22

1600 r-n —n \ \

0 25 50 75 100 125 150

Time (ms)

Figure 95. Two time traces of near-field dynamic pressure for releases of HCFC-22. Trace B

is offset by 100 kPa. Conditions: A, 412 g, release pressure 4.52 MPa; B, 417 g, release pressure 4.48 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 205

HFC- 125 also generated pressure traces which varied from release to release. Figure 96 shows

two examples. Both plots indicate that flashing occurred at the beginning and end of the releases, but

there were indications that flashing also took place at intermediate times. The third set of data

summarized in Table 6 was similar. Each release has a pressure rise at early times. The maximum

pressures observed were 20, 81, and 122 kPa.

The HFC-125/HFC-32 mixture has the lowest boiling point of all of the potential replacement

agents except FC-116. Time traces for two releases of this agent (see Figure 97) indicate that the

liquid tends to flash strongly during most of time the liquid is exiting the vessel. Pressure increases

(249 and 293 kPa) first appear immediately (1.24 and 1.12 ms) following the release. These are followed by a short quiet period before the dynamic pressure begins to increase rapidly. The

maximum pressure readings occurred near the end of the releases, reaching values greater than 1500

kPa. Table 6 summarizes the findings.

For comparison purposes, releases of halon 1301 were also performed. Figure 98 and Table 6

summarize the findings. Note that trace B was recorded for a release in which the vessel was only

1/3 full of liquid. In both cases, the pressure traces indicate that flashing occurred immediately

following (1.16 and 1.84 ms) the opening of the orifice. Early pressure increases of 91 and 43 kPa

were observed. Trace A indicates that the liquid flashed near the orifice during most of the release. The flashing associated with the emptying of the vessel is also obvious. Despite the fact that the

boiling point of halon 1301 is 20 K lower than for the HFC-125/HFC-32 mixture, the dynamic

pressures measured are considerably lower for the halon 1301. The maximum pressures observed due

to the second flashing are quite different for the two experiments shown in Figure 98.

These measurements will be discussed further in Section 3.5.4. However, a few general

conclusions can be drawn from the results summarized in Table 6 and the figures. The data for the

various agents have been presented in order of decreasing boiling point. Only for the lowest boilingpoint agents are dynamic pressures generated by the flashing observed immediately following the

opening of the burst disk. For agents having boiling points higher than HCFC-22 (232 K), flashing is only detected near the beginning and end of the releases. Pressure increases observed for HCFC-22

and agents having lower boiling points suggest that strong flashing is taking place at intermediate

times as well. With the exceptions of FC-218, HFC- 125, and halon 1301, the pressure increases

observed near the time when the vessel empties increase with decreasing boiling point.

3.5.3.3 Laser Attenuation Measurements. Laser attenuation measurements were made at the

five downstream positions listed in Table 7 for each of the agents investigated. Figure 99 through

Figure 109 show examples of the laser signals recorded for releases of all of the alternative agents,

with the exception of FC-116, and halon 1301. In each case the vessel was 2/3 full of the liquid

agent and a nominal 4.14 MPa burst disk was used.

For the higher boiling-point agents, the arrival of the agent at a laser beam fully extinguishes the

beam nearly instantaneously (within 40 /zs). This is clearly seen in Figure 99 through Figure 104

where sharp drops are seen in the laser intensities at successively later times as the agent moves away

from the vessel orifice. Recall that time 0 corresponds to the extinguishment of laser #1. The light

extinguishment is the result of scattering by the two-phase flows generated by the releases. For the

lower boiling-point agents, the fall ofl" in light intensity with time is not as sharp. This is most clear

for lasers #3, #4, and #5 for the halon 1301 release (see Figure 109). This observation is evidence

that these liquid agents are evaporating at a sufficient rate on the leading edges of the releases that the

two-phase flow is not optically dense enough to fully extinguish the laser beams. Even for the low

boiling-point liquids, the laser beams are eventually fully attenuated as the flow develops fiirther. At varying times following the complete attenuation of the laser beams, the optical density of the

flows decreases and the laser beams are partially transmitted. The exact nature of the optical density
206 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 96. Two time traces of near-field dynamic pressure for releases of HFC- 125. Trace B is

offset by 100 kPa. Conditions: A, 416 g, release pressure 4.07 MPa; B, 423 g, release pressure 4.24 MPa.
SECTIONS. FLUro DYNAMICS OF AGENT DISCHARGE 207

HFC-125/HFC-32 Mixtiire

Time (ms)

Figure 97. Two time traces of near-field dynamic pressure for releases of the

HFC-125/HFC-32 mixture. Trace B is offset by 100 l<Pa. Conditions: A,

375 g, release pressure 4.36 MPa; B, 358 g, release pressure 4.31 MPa.
208 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 98. Two time traces of near-field dynamic pressure for releases of halon 1301.

Trace B is ofi'set by 100 kPa. Conditions: A, 554 g, release pressure

4.36 MPa; B, 270 g, release pressure 4.43 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 209

Table 7. Downstream Distances from the Orifice of the Release Vessel for the Five Laser

Beams. Distances Between the Laser Beams are also Indicated.

Downstream Position (m) Separation Distance (m)

Laser #1 0.0

0.320

Laser #2 0.320

0.335

Laser #3 0.655

0.305

Laser #A 0.960

Laser #5

0.340

1.300

of a flow is expected to be quite complex. It should be a function of the total mass concentration of

the agent, the degree of vaporization of the agent, the size of droplets in the liquid phase, and the

degree to which the agent has cooled and condensed water vapor from the surrounding air entrained

by the flow. Even so, there are clear trends in the data. The total attenuation of laser #1 lasts for much longer than the agent release time (30-40 ms). The explanation for this behavior is available

from the high-speed films of the behavior of the agents inside the vessel (see Section 3.4.1). These

show that agent condensation occurs in the nitrogen above the liquid in the vessel during a release. This is attributed to the cooling of the nitrogen and gaseous agent by the adiabatic expansion of the

gas (see Section 3.4). Due to the gas used to pressurize the vessel, there is a slow flow of nitrogen even after a release is complete and eventually the condensed agent is swept from the vessel. The time behaviors of laser §2 transmission for releases of difl'erent agents provide insight in the

evaporation characteristics of the various agents. For all of the agents except FC-218, the time of

total laser beam attenuation is on the order of 60 ms or roughly 50% longer than the total release time

of the liquid. The total attenuation time for FC-218 is roughly the same as its release time. The total attenuation times of the beam for laser #3 tend to be somewhat shorter than for laser #2. The

recovery of light transmission tends to be quicker for the lower boiling-point agents.

The periods of total light attenuation for lasers #4 and §5 are longer and the recovery of light

transmission takes place more slowly. It is likely that this observation is the result of the condensation of water vapor which should require a longer period to evaporate than the agents.

As can be seen in Figure 99 through Figure 109, the arrival time of an agent release at a laser beam is easily detected as a rapid, near-total attenuation of the transmitted light. It is a simple

exercise to accurately determine these arrival times. For example, for the data shown in Figure 99

the arrival times are 0, 5.88, 10.55, 15.31, and 20.01 ms for lasers #1-^5, respectively. Since the

distances between the laser beams are known, four average velocities can be calculated from these

results. Such calculations have been made for the series of releases from 2/3 filled vessels. In the

figures which follow, the average velocities are plotted as a function of downstream distance defined

as the midpoints of the laser beam locations.

Figure 110 through Figure 120 show results for the ten agents which are liquids at room

temperature and halon 1301. There are dramatic diff"erences among the results. HFC-236fa,

FC-31-10, FC-318, HCFC-124, and HFC-227ea (Figure 110 through Figure 114) are similar. Average velocities measured at the first measuring station fall in a range of 50 to 60 m/s. As the
210 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 99. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7

following a release of HFC-236fa at 0 s. Conditions: 464 g

agent, 4.30 MPa

release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 211

Figure 100. Time behaviors for transmission of laser beams #l-#5 spaced as shown in

Table 7 following a release of FC-31-10 at 0 s. Conditions: 520 g

agent,

3.85 MPa release pressure.
212 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 101. Time behaviors for transmission of laser beams #l-#5 spaced as shown in

Table 7 following a release of FC-318 at 0 s. Conditions: 501 g agent,

4.43 MPa release pressure.
SECTION 3. FLUID DYNAlvlICS OF AGENT DISCHARGE 213

Figure 102. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7 following a release of HCFC-124 at 0 s. Conditions: 462 g

agent, 4.05 MPa release pressure.
214 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HFC-227ea

-| 1 1 r

80 100 120 140

Time (ms)

Figure 103. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7 following a release of HFC-227ea at 0 s. Conditions: 476 g

agent, 4.45 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 215

HFC- 134a

T

1 1 1 1 1 1 T

Time (ms)

Figure 104. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7 following a release of HFC- 134a at 0 s. Conditions: 418 g agent,

4.13 MPa release pressure.
216 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 105. Time behaviors for transmission of laser beams #l-#5 spaced as shown in

Table 7 following a release of FC-218 at 0 s. Conditions: 463 g

agent,

4.06 MPa release pressure.
SECTION 3. FLUID DYN./^CS OF AGENT DISCHARGE 217

Figure 106. Time behaviors for transmission of laser beams #l-#5 spaced as shown in

Table 7 following a release of HCFC-22 at 0 s. Conditions: 417 g agent,

4.52 MPa release pressure.
218 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 107. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7 following a release of HFC- 125 at 0 s. Conditions: 423 g

agent,

4.24 MPa release pressure.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 108. Time behaviors for transmission of laser beams #l-#5 spaced as shown

Table 7 following a release of the HFC-125/HFC-32 mixture at 0 s. Conditions: 375 g agent, 4.36 MPa release pressure.
220 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

halon 1301

T

1 r \ 1 T r

0 20 40 60 80 100 120 140 Time (ms)

Figure 109. Time behaviors for transmission of laser beams #l-#5 spaced as shown in Table 7 following a release of halon 1301 at 0 s. Conditions: 554 g

agent, 4.36 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 221

Figure 110. Average velocities for releases of HFC-236fa determined by the time of

arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
222 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

\0

FC=31-10

1

4.21 MPa • 4.52 MPa

I I I I I 0 20 40 60 80 100 120

Average Downstream Position (cm)

Figure 111. Average velocities for releases of FC-31-10 determined by the time of

arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 223

80

FC-318

70 -

60

50

o

1 3.90

1 1 MPa

1 1

• 3.92 MPa

V 3.99 MPa

T 4.25

4 37

MPa

MPa

4.43 MPa p

„ V

/ »

^^^^

B

/'

//'''

/ / ^

y

/ -/

/ _-

V

/ • ^

1

d

1

1 1 1 0 20 40 60 80 100

Average Downstream Distance (cnn)

120

Figure 112. Average velocities for releases of FC-318 determined by the time of arrival

at laser beams #l-#5 plotted as a function of average downstream distance. Pressures are those inside the vessel at disk bursting.
224 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 113. Average velocities for releases of HCFC-124 determined by the time of

arrival at laser beams ff\-tt5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 225

HFC-227ea

80

50

0 20 40 60 80 100 120

Average Downstream Distance (cnn)

Figure 114. Average velocities for releases of HFC-227ea determined by the time of

arrival at laser beams M\-M5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
226 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HFC- 134a

1 r~ 1

V 4.38 MPa

0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 115. Average velocities for releases of HFC- 1 34a determined by the time of

arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 227

0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 116. Average velocities for releases of FC-218 determined by the time of arrival

at laser beams #l-#5 plotted as a function of average downstream distance. Pressures are those inside the vessel at disk bursting.
228 SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

HCFC-22

70

60

50

40

30

J(

s'x V\\ \

1 1

- -

\

o 3.63 MPa .,,\\\\\ • 3.92 MPa V \0

V 4.24 MPa \ V

T 4.34 MPa \\

4.48 MPa

4.49 MPa

A 4.52 MPa

A 4.56 MPa

1 1 1

i 1 n

0 20 40 60 80 100

Average Downstream Position (cm)

120

Figure 117. Average velocities for releases of HCFC-22 determined by the time of

arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 229

Figure 118. Average velocities for releases of HFC- 125 determined by the time of

arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
230 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 119. Average velocities for releases of the HFC-125/HFC-32 mixture determined by the time of arrival at laser beams #l-#5 plotted as a function of

average downstream distance. Pressures are those inside the vessel at disk

bursting.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

Figure 120. Average velocities for releases of halon 1301 determined by the time of

arrival at laser beams U\-U5 plotted as a function of average downstream

distance. Pressures are those inside the vessel at disk bursting.
232 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

agent moves further away from the vessel, the average velocities actually increase, remaining roughly

constant at approximately 70 m/s.

HFC-134a (Figure 115) has a different velocity dependence on downstream distance. Two of the

releases had high velocities (65-70 m/s) near the vessel which stayed roughly constant over the two

nearest measurement locations. At the third and fourth locations the velocities actually decreased to values of roughly 50 m/s. Note that these results are nearly exactly the opposite of those for the

higher boiling-point agents discussed above. A third release of HFC- 134a looks similar to the higher

boiling-point agents {i.e., the measured velocity near the nozzle is roughly 56 ms/s and increases to roughly 70 m/s) except that the velocity again decreases at the furthest measuring station. The velocity results for FC-218 (Figure 116) are similar to those for HFC-134a. Velocities for the two locations near the nozzle are on the order of 80 ms/s. As the agent moves further from the

exit the velocity begins to fall off, decreasing to approximately 40 m/s for the last station. The

results for HCFC-22 (Figure 117) are similar, but there is substantial scatter in the results. For HFC-125 (Figure 118) the velocities near the vessel are again roughly 70 m/s. However,

these velocities decrease faster with downstream distance becoming less than 20 m/s at the most

distant location. The HFC-125/HFC-32 mixture (Figure 119) displays a similar, but enhanced

behavior. Velocities near the vessel exit are very high, approaching 100 m/s. By the second

measuring station the velocity has dropped rapidly and eventually falls to less than 30 m/s. Interestingly, the results for halon 1301 (Figure 120) are similar to those for FC-218 and HFC-125.

These measurements have been made for a range of pressures in the release vessel due to variations in burst disks. Close examination shows that the measured velocities depend somewhat on

pressure, with higher pressures yielding higher velocities. However, for the range of pressures

shown the dependence is relatively weak and on the order of experiment-to-experiment variation.

3.5.3.4 Far-Field Dynamic Pressure Measurements. These measurements were made on the

centerline of the releases at a distance 1.3 m from the vessel exit. The pressure transducer was

oriented such that it faced the release direction. Figure 121 through Figure 131 show examples of the

dynamic pressures recorded as a function of time for the ten room-temperature liquid replacement

agents as well as halon 1301.

Figure 121 shows time plots of downstream dynamic pressure for two releases of HFC-236fa. In both cases a rapid pressure rise is observed at approximately 20 ms following release. This agrees

closely with the arrival of the agent at laser #5 as determined by strong beam attenuation (18.8 ms

and 20.1 ms). The pressure jumps roughly 200 to 300 kPa immediately following the arrival of the

agent at the transducer. The pressure then continues to increase over the next 25 ms. Maximum

pressures observed in the two cases were 1,100 kPa and 1,420 kPa. As we will be discussed shortly,

such pressure increases are indicative of a two-phase flow striking the transducer.

Similar pressure traces for FC-31-10 are shown in Figure 122. The behaviors are similar to those for HFC-236fa with maximum pressures reaching 820 and 1029 kPa. The pressure traces for FC-318 (Figure 123) also have a similar behavior except the initial pressure increase is somewhat

higher (500 to 700 kPa) with maximum observed values of 1000 and 1200 kPa. Results for HCFC-124 are shown in Figure 124. This agent generates strong pressure increases (500 to 700 kPa)

immediately upon arrival at the transducer which then grow to very high levels (1200 to 1500 kPa).

The results for HFC-227ea (Figure 125) are very similar to those for HCFC-124. The pressure

traces for HFC-134a (Figure 126) have very rapid and large increases (> 1000 kPa) upon arrival at the transducer and reach the highest values (1900-2100 kPa) of any of the high boiling-point agents.

The dynamic pressure measurements for FC-218 (Figure 127) have a very different behavior than

those discussed thus far. A rapid pressure increase is observed at roughly 25 ms corresponding to the

time when the agent reaches the transducer. A maximum value is quickly reached. This maximum
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 233

Time (ms)

Figure 121. Dynamic pressure measured 1.3 m from the vessel for HFC-236fa. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 471 g

agent,

release pressure of 4.51 MPa ; B, 464 g of agent, release pressure of

4.30 MPa.
234 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

Time (ms)

Figure 122. Dynamic pressure measured 1.3 m from the vessel for FC-31-10. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 495 g agent,

release pressure of 3.93 MPa ; B, 515 g of agent, release pressure of

4.20 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 235

FC-318

0 20 40 60 80 100 120 140

Time (ms)

Figure 123. Dynamic pressure measured 1.3 m from the vessel for FC-318. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 521 g agent,

release pressure of 4.34 MPa ; B, 505 g of agent, release pressure of

3.99 MPa.
236 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 25 50 75 100 125 150

Time (ms)

Figure 124. Dynamic pressure measured 1.3 m from the vessel for HCFC-124. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 462 g agent,

release pressure of 4.05 MPa ; B, 465 g of agent, release pressure of

4.30 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 237

HFC-227ea

0 25 50 75 100 125 150

Time (ms)

Figure 125. Dynamic pressure measured 1.3 m from the vessel for HFC-227ea. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 489 g

agent,

release pressure of 4.41 MPa ; B, 476 g of agent, release pressure of

4.45 MPa.
238 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HFC- 134a

D_

CO

CO

0)

CL

O• o

c

O

(D

CO

c

OQ 25 50 75 100

Time (ms)

125 150

Figure 126. Dynamic pressure measured 1.3 m from the vessel for HFC-134a. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 418 g agent,

release pressure of 4.13 MPa ; B, 419 g of agent, release pressure of

4.21 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 239

O

CL

300

FC-218

0) 250 -

CO

CO

CD

o

o

c

O

0)

-I-' CO

c

o

200 -

150 -

100 -

50 -

0

-'

25 50 75 100

Time (ms)

125 150

Figure 127. Dynamic pressure measured 1.3 m from the vessel for FC-218. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 460 g agent,

release pressure of 4.58 MPa ; B, 468 g of agent, release pressure of

4.28 MPa.
240 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HCFC-22

Time (ms)

Figure 128. Dynamic pressure measured 1.3 m from the vessel for HCFC-22. For

clarity, trace B is offset by 100 kPa. Conditions are: A, 421 g agent,

release pressure of 4.48 MPa ; B, 417 g of agent, release pressure of

4.52 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 241

HFC- 125

O 175 I—I 1 1 r CL

Time (ms)

Figure 129. Dynamic pressure measured 1.3 m from the vessel for HFC-125. For

clarity, trace B is offset by 50 kPa. Conditions are: A, 403 g agent,

release pressure of 4.19 MPa ; B, 423 g of agent, release pressure of

4.24 MPa.
242 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HFC-125/HFC-32

0 25 50 75 100 125 150

Time (ms)

Figure 130. Dynamic pressure measured 1.3 m from the vessel for the HFC-125/HFC32 mixture. For clarity, trace B is offset by 100 kPa. Conditions are: A,

358 g agent, release pressure of 4.31 MPa ; B, 375 g of agent, release

pressure of 4.36 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 243

Figure 131. Dynamic pressure measured 1.3 m from the vessel for halon 1301. For

clarity, trace B is offset by 75 kPa. Conditions are: A, 1/3 full, 270 g

agent, release pressure of 4.43 MPa ; B, 554 g of agent, release pressure

of 4.36 MPa.
244 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

value is considerably smaller (100-150 kPa) than observed for any of the higher boiling-point agents.

The pressure time traces for HFC-125 (Figure 129) are similar except for the fact that the maximum

observed pressures are even lower (on the order of 50 kPa). Dynamic pressures observed for halon

1301 (Figure 131) were also quite low, except near the end of one of the releases where a rapid

increase to 375 kPa was observed.

While the pressure increases for the low boiling-point agents FC-218, HFC-125, and halon 1301

tended to be smaller than observed for higher boiling-point cases, the results were quite different for HCFC-22 (Figure 128) and the HFC-125/HFC-32 mixture (Figure 130). Both of these agents showed

substantial variation between experiments. For HCFC-22 the time required for the pressure to reach

its maximum value differed somewhat between releases, but the maximum values reached were

greater, being on the order of 2100 kPa for both cases shown. The data for the mixture are also

variable. In one case the maximum pressure was in excess of 2200 kPa, while in the other the

maximum only reached 1200 kPa.

To summarize, the results for the high boiling-point liquids seem to be consistent and show that very high dynamic pressures are generated at a distance of 1.3 m from the vessel by the release of

this type of agent. Two distinct types of pressure curves are found for the lower boiling-point

liquids. HCFC-22 and the HFC-125/HFC-32 mixture generate maximum pressures which are similar

to those observed for the higher boiling-point agents even though they have more complex time

histories, and there are larger variations among experiments. FC-218, HFC-125, and halon 1301

releases generate maximum pressures which are considerably reduced from those observed for all of

the other tested liquids.

3.5.3.5 Aspirated Hot-Film Measurements. The aspirated hot-film has only been used to characterize qualitatively the nature of the flows reaching the 1.3 m downstream position. It should

be noted that the these probes can be calibrated to accurately measure concentration in isothermal

binary mixtures of gases. If the temperature is changing, or if two phases are present the response of

these probes is more complex. An ongoing project at NIST is evaluating the utility of aspirated hotfilm probes for concentration measurements under these complex conditions.

Since the results are qualitative, only a few representative examples are reported here. Even so,

it will become clear that the hot-film response provides useful information for detecting the present of an agent and for determining whether it is a single or two-phase flow. Figure 1 32 shows an example of the aspirated hot-film output observed during the release of

HFC-236fa. A nearly constant signal (roughly 6.3 V) is observed until 20 ms following release at which time an abrupt increase in voltage is observed. Figure 99 and Figure 121 show that 20 ms is the time when the agent first reaches the probe. Thus the constant signal is that representative of room temperature air. The signal induced by the arrival of the agent is very "spikey" and reaches

values of nearly 10 V. Such voltages are much higher than would be expected for a 100% concentration of the gaseous agent (Pitts and McCaffrey, 1986). The large heat transfer from the hot-film

probe which these voltages represent is attributed to the presence of two-phase flow in the probe.

Evaporative cooling of the droplets requires large amounts of heat. The large voltage spikes are observed for roughly 40 ms. This period is on the order of the period required for release of the

agent. Afterwards, much smaller voltages are observed which decay slowly toward that expected for

air. This is taken as evidence that the mixture is becoming more gaseous due to evaporation of the

agent as it mixes with the air. This is consistent with the transmission observed for laser #5

(Figure 99) which indicates that the flow becomes partially transmitting at roughly 100 ms following

the release. The hot-film responses for FC-31-10, FC-318, HCFC-124, HFC-227ea, and HFC-134a are very

similar to those for HFC-236fa. During the period when the releases are passing the probe there are
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 245

HFC-236fa

6.0

^ ^ ^ ^ ^ ^ 0 25 50 75 100 125 150

Time (ms)

Figure 132. The voltage output of the aspirated hot-film located 1.3 m from the vessel

is plotted as a function of time for a release of HFC-236fa. Conditions:

464 g of agent, release pressure of 4.30 MPa.
246 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

voltage signals which must be associated with two-phase flow. Afterwards, the voltages begin to fall and approach those for room-temperature air. The hot-film response for FC-218 releases was quite different. Figure 133 shows an example of

the hot-film voltage plotted as a function of time. The signal begins to rise at the time when laser beam #5 is strongly attenuated (Figure 105) and the downstream pressure transducer (Figure 127, curve A) detects the arrival of the agent. The signal changes rapidly with time indicating the

presence of highly turbulent flow, but during the initial part of the signal following the arrival of the

agent the voltage increases are relatively small and consistent with those expected for a binary gas

mixture. At roughly 50 ms following the release, there is a rise in the hot-film voltage and spiking

appears which suggests the presence of a two-phase flow. Interestingly, the downstream pressure

transducer recorded a small rise in dynamic pressure at the same time. These results suggest that near the end of the release a small amount of two-phase fluid reached the 1.3 m downstream position,

but the responses of the detectors suggest that the FC-218 was much more completely vaporized than

the higher boiling-point agents.

Figure 134 shows simultaneous plots of downstream dynamic pressure and hot-film output as functions of time for a release of HFC- 125. The flow of this agent results in relatively small

increases in both dynamic pressure and hot-film voltage. Roughly 20 ms after the agent arrives at the

measurement location there are sharp increases in both signals. It is difficult to determine whether

this is the result of the arrival of two-phase flow or a higher concentration (i.e., density) or velocity

of gaseous agent. In either case, it is clear that this agent arrives nearly completely vaporized.

Figure 135 shows a similar plot for halon 1301. Here there is a significant increase in dynamic

pressure associated with the end of the release which appears to have two-phase behavior. However,

it is clear from the traces that during most of the release primarily gas-phase agent is reaching this downstream position.

Recall that releases of HCFC-22 and the HFC-125/HFC-32 mixture resulted in large downstream

dynamic pressures. Figure 136 and Figure 137 show time plots for the hot-film outputs for these two

agents. In both cases there is evidence for two-phase flow near the beginning and ends of the

releases. This conclusion is consistent with the observed dynamic pressure behaviors (see Figure 128

and Figure 130).

3.5.3.6 Effects of Orifice Diameter. Experiments were done for FC-31-10, FC-218, and halon

1301 using plates with orifices of 12.7 and 6.4 mm. The same flow characterizations recorded for the standard release conditions were used.

For the FC-31-10 the near-field pressure transducer recorded a pressure rise of roughly 450 kPa

when the 12.7 mm orifice was used and no measurable increase when the 6.4 mm orifice was used.

These results can be compared with the results in Table 6 for the 19.1 mm opening where maximum

dynamic pressures were between 300-400 kPa. The pressure behavior for a release of FC-218

through the 12.7 mm orifice is shown in Figure 138. It can be seen that the transducer responded

shortly after the start of the release and that the dynamic pressure remained high over the remaining

release period of approximately 75 ms. Comparison with Figure 94 shows that the maximum

pressures reached are lower for this orifice diameter than for the 19.1 mm opening, but that the

pressures are sustained over a much longer period. For the 12.7 mm orifice, the fluid seems to be

flashing continuously. For the smallest orifice no pressure increase was detected. The dynamic

pressure time behaviors recorded for releases of halon 1301 had very similar behaviors to those for the FC-218.

The size of the orifice has a strong efl^ect on flashing behavior. For the higher boiling-point

FC-31-10 no flashing is observed for the smaller orifices. For the lower boiling-point liquids, the

maximum dynamic pressure observed decreases when the 19.1 mm orifice is reduced to 12.7 mm, but
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 247

Figure 133. The voltage output of the aspirated hot-film located 1.3 m from the vessel

is plotted as a function of time following a release of FC-218. Conditions:

460 g of agent, release pressure of 4.58 MPa.
248 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

HFC- 125

50 75 1 00

Time (ms)

80

- 60

- 40

- 20

0

- -20

- -40

- -60

-80

150

o

CL

(D

L_

Z3CO

CO

0)

CL

O

D

C

Q

O

CO

100 ^

Figure 134. The voltage output of the aspirated hot-film and the dynamic pressure for probes located 1.3 m from the vessel are plotted as a function of time

following a release of HFC- 125. Conditions: 403 g of agent, release

pressure of 4.19 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 249

Figure 135. The voltage output of the aspirated hot-film and the dynamic pressure for probes located 1.3 m from the vessel are plotted as a function of time

following a release of halon 1301. Conditions: 554 g of agent, release

pressure of 4.36 MPa.
250 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 136. The voltage output of the aspirated hot-film located 1.3 m from the vessel

is plotted as a function of time following a release of HCFC-22. Conditions: 421 g of agent, release pressure of 4.48 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 251

HFC-125/HFC-32

75 100 125 150

Time (ids)

Figure 137. The voltage output of the aspirated hot-film located 1.3 m from the vessel

is plotted as a function of time following a release of the HFC-125/HFC- 3 mixture. Conditions; 375 g of agent, release pressure of 4.36 MPa.

It
252 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

0 25 50 75 100 125 150

Time (ms)

Figure 138. Time trace of near-field dynamic pressure for a release of FC-218 through

a 12.7 mm orifice. Conditions: 457 g of agent, 4.30 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 253

the period of flashing is substantially extended. No increases in dynamic pressure are observed for the smallest orifice. Figure 139 through Figure 141 compare the average velocity variations with average downstream

position for releases of FC-31-10, FC-218, and halon 1301 through the three orifices. For FC-31-10

releases, smaller orifices result in significant changes in the velocity development. With the 19.1 mm

opening (Figure 111 and Figure 139) the velocity first increases and then levels off at values near 70 m/s. For the two smaller orifices the initial velocity is on this order, but it then decreases with

increasing downstream distance. The smallest velocity is observed at the farthest position using the

6.4 mm orifice. Velocities near the vessel exit are roughly the same for each orifice size when FC-218 is considered (Figure 140). The velocities drop ofl" in each case with increasing downstream distance.

Once again the smallest orifice results in the lowest velocity at the last measuring position.

The results for halon 1301 (Figure 141) are similar to those for FC-218 except that the velocities nearest the orifice vary somewhat with the smaller orifices producing higher velocities. Overall, the

smaller orifices generate higher velocities near the vessel and lower velocities further downstream.

The evaporation characteristics of the flows for releases of the three agents are indicated by

Figure 142 through Figure 145 where the time behaviors for the far-field pressure transducer and hotfilm aspirated film voltages are plotted as functions of time. The responses for the release of

FC-31-10 from the 12.7 mm orifice (Figure 142) indicate that this agent is a two-phase mixture 1.3 m

from the vessel. This is clear from the high dynamic pressures and spikey nature of the hot-film

signal. The period required to empty the bottle is roughly 65 ms. This is also roughly the length of

the two-phase flow period at the downstream position. The plots for the release from the 6.4 mm

orifice indicate that the agent first arrives in the gas phase and remains so throughout the release

period.

The data for the release of FC-218 through the 12.7 mm orifice (Figure 143) suggest that this agent is fully vaporized when it reaches the sampling position, except, perhaps, at the very end of the

release. Both the low dynamic pressures and small increases in hot-film voltages support this conclusion. Similar results were obtained for the 6.4 mm orifice. Halon 1301 has a behavior which is very much like that for FC-218 except that a very strong

two-phase component is observed near the end of the release when the 12.7 mm orifice is used

(Figure 144). Both the dynamic pressure and hot-film signals indicate the presence of two-phases. It

is possible that this change in flow composition is associated with the strong flashing which occurred

near the vessel as the liquid emptied from the vessel. Figure 145 shows results for a release of halon

1301 through the 6.4 mm orifice. Here the signals from both transducers are small indicating that only gaseous agent is present 1.3 m from the orifice. A small pressure increase was recorded at the

end of the release which may be associated with the presence of a small amount of liquid agent or a higher concentration or velocity of the gaseous agent. Note the long time period required for release

through this small opening.

3.5.3.7 Effect of Vessel Pressurization Level. The efl"ect of the vessel pressurization level on

the mixing and dispersion behavior of an agent was investigated for releases of FC-218 using burst

disks designed to break at nominal pressures of 2.76, 4.14, and 5.52 MPa. Figure 146 and

Figure 147 show examples of the results for near-field dynamic pressure as a function of time for vessels having high and low internal pressures, respectively. Figure 94 shows two examples of

results for releases having intermediate levels of pressurization. None of the releases indicated a pressure rise associated with the first flashing following the release of the FC-218. Interesting!)/, -he highest dynamic pressure associated with the second flashing is observed for the release with an
254 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 139. Average velocities determined by the time of arrival at laser beams #l-#5

plotted as a function of average downstream distance for releases of

FC-31-10 through 19.1, 12.7, and 6.4 mm orifices.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 255

FC-218

0 20 40 60 80 100 120

Average Downstream Distance (cnn)

Figure 140. Average velocities determined by the time of arrival at laser beams #1-^5

plotted as a function of average downstream distance for releases of

FC-218 through 19.1, 12.7, and 6.4 mm orifices.
256 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

halon 1301

0 20 40 60 80 100 120

Average Downstream Distance (cnn)

Figure 141. Average velocities determined by the time of arrival at laser beams #l-#5

plotted as a function of average downstream distance for releases of halon

1301 through 19.1, 12.7, and 6.4 mm orifices.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 257

FC-31-10

1500 O

CL

- 1000

- 500

- 0

- -500

- -1000

-1500

25 50 75 100 125 150 175 200

0)

D

W

CO0

Q_

o

D

c

Q

E

D

(D

COc

o

Time (ms)

Figure 142. Voltage output of the aspirated hot-film and dynamic pressure for probes

placed 1.3 m from the vessel as a function of time following a release of

FC-31-10 through a 12.7 mm orifice. Conditions: 521 g of agent, release — pressure of 4.48 MPa.
258 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Z5

o

8.0

FC-218

> 7.5 -

3

Q_

7.0 -

O 6.5 -

6.0

Figure 143. Voltage output of the aspirated hot-film and dynamic pressure for probes

placed 1.3 m from the vessel as a function of time following a release of

FC-218 through a 12.7 mm orifice. Conditions: 463 g of agent, release

pressure of 4.23 MPa.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 259

halon 1301

Time (ms)

Figure 144. Voltage output of the aspirated hot-film and dynamic pressure for probes

placed 1.3 m from the vessel as a function of time for a release of halon

1301 through a 12.7 mm orifice. Conditions: 537 g of agent, release

pressure of 4.54 MPa.
260 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

halon 1301

Time (ms)

Figure 145. Voltage output of the aspirated hot-film and dynamic pressure for probes

placed 1.3 m from the vessel as a function of time for a release of halon

1301 through a 6.4 mm orifice. Conditions: 546 g of agent, release

pressure of 4.62 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 146. Two time traces of near-field dynamic pressure for releases of FC-218

using nominal 5.52 MPa burst disks. Trace B is off'set by 100 kPa.

Conditions: A, 455 g, release pressure 6.21 MPa; B, 457 g, release

pressure 5.97 MPa.
262 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

0 25 50 75 100 125 150

Time (ms)

Figure 147. Time trace of the near-field dynamic pressure for a release of FC-218

using a nominal 2.76 MPa burst disk. Conditions: 436 g, release pressure

2.61 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 263

internal pressure of 4.28 MPa. The releases at higher internal pressures resulted in somewhat lower

dynamic pressures. The smallest pressure rise was observed when the vessel pressure was 2.61 MPa.

Measurements discussed in Section 3.4.3.4 showed that the agent was expelled from the vessel at higher rates as the internal pressure was increased. Figure 148 shows the average velocities measured

for three releases of FC-218 subject to a range of vessel pressures. Consistent with the emptying rate measurements, the velocities increase with internal pressure. The largest difference is observed when

reducing the internal vessel pressure from 4.28 MPa to 2.61 MPa.

3.5.3.8 Effect of Attaching a 0.5 m Extension Tube to the Vessel. A 0.5 m tube having a 19.1 mm inside diameter was attached to the orifice of the vessel. The near-field pressure transducer was not employed in these experiments. The tube blocked lasers #1 and §2. As a result, average

velocity measurements are only available for the two measurement locations farthest from the vessel. Note that the exit for the tube was located 0.8 m from the downstream position of laser tf5. Measurements were made for FC-31-10, FC-218, and halon 1301.

Figure 149 through Figure 151 show measurements of average velocities for the three agents

with the tube in place. For comparison purposes the results for releases without the tube are also

shown. In each case the velocities observed with the tube are considerably higher than when the tube

is absent. For the higher boiling-point FC-31-10 the increase is only 20 to 30%. However, for the

two lower boiling-point agents the increase is more than a factor of two. The downstream dynamic pressures and hot-film responses recorded for releases of the three

agents are shown in Figure 152 through Figure 154. In each case very high dynamic pressures are measured suggesting that the agents are arriving as two-phase flows. The spikey nature of the hotfilm signals support this conclusion. For the two low boiling-point agents, FC-218 and halon 1301,

the hot-film detects the presence of the agents before a large pressure increases is observed. This

suggest that the initial agent reaching the detectors is primarily vapor. However, after only a few

milliseconds, the flow displays two-phase character. Comparison of Figure 154 with Figure 152 and Figure 153 shows that the two-phase flow period

is considerably longer for the halon 1301 (~ 100 ms) than for the other two agents. This, in turn,

implies that the time required for the release of the agent from the vessel is substantially longer. In the absence of the extension tube (see Section 3.4.3.1) the liquid release period for the three agents is on the order of 40 ms. The presence of the extension tube therefore more than doubles the liquid

release period for halon 1301. The same conclusion was reached in Section 3.4.3.3 based on internal

vessel pressure measurements. The longer liquid release period is attributed to the generation of a two-phase flow of halon 1301 in the tube.

3.5.3.9 Effect of Cooling the Agent Prior to Release. Two releases of FC-31-10 were studied

in which the agent inside the vessel was cooled to 228 K, i.e., a temperature far below its boiling

point of 271 K. As would be expected, the high-speed films indicated that there was very little flashing of the liquid as it exited the vessel. No near-field dynamic pressure measurements were made, but no pressure increase would be expected due to the absence of flashing behavior.

Figure 155 is a plot of the measured average velocity as a function of distance from the vessel. These results can be compared with similar plots shown in Figure 111 for releases of FC-31-10 with

the vessel at room temperature. The velocities near the vessel exit are much lower for the cooled

agent as expected for the lower release pressure. There are small increases in these velocities with

downstream distance, but the increases are much smaller than observed for the room temperature

releases (roughly 10 m/s versus 15 mis).
264 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 148. Average velocities as a function of downstream distance determined by time

of arrival at laser beams #l-#5 for releases of FC-218 with burst disks

covering a range of nominal pressures. Pressures are those inside the vessel

at disk bursting.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 265

FC-31-10

CO

O

o

>

(D

o

>

<

110

100

90

o 80

70

60

50

O 4.47 MPa, no tube

• 4.27 MPa, tube

V 4.14 MPa, tube

_

0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 149. Average velocities determined by the time of arrival at laser beams #l-#5

as a function of downstream distance for releases of FC-31-10 with and

without an extension tube. Pressures are those inside vessel at disk bursting.
266 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

J L 0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 150. Average velocities determined by the time of arrival at laser beams #l-#5

as a function of downstream distance for releases of FC-218 with and

without an extension tube. Pressures are those inside vessel at disk bursting.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 267

E

o

o

>

X

<

Qi

D

>

<

200

Halon 1301

175 h

150

125

100 [-

75

50

25

T

O 4.36 MPa, no tube • 4.56 MPa, tube

O

0 20 40 60 80 100 120

Average Downstream Position (cnn)

Figure 151. Average velocities determined by the time of arrival at laser beams #l-#5

as a function of downstream distance for releases of halon 1301 with and

without an extension tube. Pressures are those inside vessel at disk bursting.
268 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10

1000 ^

25 50 75 100

Time (ms)

125

800

600

400

200

- 0

- -200

- -400

- -600

- -800

-1000

150

i_

Z5

CO

CO0

1^

CL

o

E

D

c

o

o

-M

COc

oQ

Figure 152. Aspirated hot-film voltage and dynamic pressure for probes located 1.3 m

from the vessel with an attached 0.5 m tube plotted as functions of time

following release of FC-31-10. Conditions: 513 g of agent, release

pressure of 4.14 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 269

FC-218

Time (ms)

Figure 153. Aspirated hot-film voltage and dynamic pressure for probes located 1.3 m

from the vessel with an attached 0.5 m tube plotted as functions of time

following release of FC-218. Conditions: 465 g of agent, release pressure

of 4.48 MPa.
270 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

halon 1301

50 75 100

Time (ms)

o

Q_

Z3

(/) U)

(D

CL

O

Q

D

(D

CDC

o

Figure 154. Aspirated hot-film voltage and dynamic pressure for probes located 1.3 m

from the vessel with an attached 0.5 m tube as functions of time following

release of halon 1301. Conditions: 523 g of agent, release pressure of

4.56 MPa.
SECTIONS. FLUro DYNAMICS OF AGENT DISCHARGE

FC-31-10

• 3.32 MPa, 228 K

0 20 40 60 80 100 120

Average Downstream Distance (cnn)

Figure 155. Average velocities for releases of cooled FC-31-10 determined by the time

of arrival at laser beams #l-#5 plotted as a function of average downstream

distance. Vessel internal temperatures and pressures are indicated.
272 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

The downstream dynamic pressure and aspirated hot-film voltage time plots shown in Figure 156

indicate that when the agent arrives at the 1.3 m sampling position it has a great deal of two-phase

character. This is consistent with a slow evaporation due to its initial cold temperature.

A single release of FC-218 cooled to 228 K was run. At times during releases of agents which

reached the downstream detectors as mostly liquid, a strong interference was recorded on all of the

data channels. This was the case for the release of cold FC-218. For this reason, plots of the data

are not shown. However, the results are discussed because they provide important insights as to the

importance of superheating on the agent behavior. The release pressure was 3.34 MPa. Due to the

positioning of the vessel, laser #1 was located 60 mm from the vessel orifice. The initial pressure

drop inside the vessel was noted 1 ms before laser #1 detected any attenuation. This suggests agent

began flowing immediately and achieved a near-vessel velocity on the order of 60 m/s. This estimate

is consistent with measurements of the average velocity determined from laser beam attenuation

measurements shown in Figure 157. For comparison, the results for a room temperature release are also included in the figure. The most remarkable finding from Figure 157 is the nearly constant flow

velocity at positions removed from the vessel. A steady stream of fluid is detected with a velocity of

57 m/s.

Strong interferences in the signals were detected when the release reached the 1.3 m sampling

position. The aspirated hot-wire signal was very large and saturated the digitizer. The dynamic

pressure measurement was also quite noisy, however, it is possible to estimate that the dynamic

pressure rose quickly and reached a level of roughly 1 .4 MPa that persisted over the period of the

release. Using the velocity of 57 m/s from the laser measurements, this dynamic pressure corre- sponds to a density of 860 kg/vn?. The room temperature density of the liquid is 1321 kg/n?

suggesting that very little of the liquid was evaporated over the 1.3 m length of the flow. Recall that for room temperature releases of FC-218, the degree of vaporization was quite high,

approaching 100%. These results demonstrate the dramatic effect that the degree of superheating can

have on the vaporization of these agents.

3.5.3.10 Effect of Orientation—Upward Releases. Releases of the agents were made with the

vessel orifice oriented upward and horizontally. Measurements of the mixing behavior have only

been recorded for the vessel oriented upwards. High-speed films, near-field dynamic pressure, laser attenuation, downstream dynamic pressure, and aspirated hot-film measurements have all been made.

The experimental system was modified by moving the vessel to a location just below laser #5 and the

downstream pressure transducer and aspirated hot-film to a location just upstream of laser #1. Keep

in mind for the figures which follow that laser #5 is closest to the vessel exit. Releases of FC-31-10,

FC-218, and halon 1301 were studied.

The high-speed films demonstrated that releases of the agents upward were quite different from

downward releases. The behaviors of the agents within the vessel are described in Section 3.4.3.6

There it is noted that for an upward release the nitrogen first exited the vessel. During this period some gaseous agent was condensed. After a fairly substantial delay the agent in the bottle began to boil and froth, filled the vessel, and generated a two-phase flow exiting the vessel. Approximately 20

ms was required for the agent to begin to boil and a longer period to fill the vessel and generate the

two-phase flow. The periods required to expel the agent completely were substantially longer for the

upward releases than when the agents were released downward.

The films show related behaviors for the released agents near the vessel orifice {e.g., see Figure 73). A plume is easily observed when the disk first opens. After a short period of time the

plume seems to disappear, only to reappear when the fluid inside the vessel begins to boil.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 273

0 25 50 75 100 125 150

Time (ms)

Figure 156. Aspirated hot-film voltage and dynamic pressure for probes located 1.3 m

from the vessel plotted as functions of time following release of FC-31-10

cooled to 228 K. Conditions: 517 g of agent, release pressure of

3.34 MPa.
274 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

80

FC-218

70

60

o

o 50

>

40 -

30 -

20

O room temperature, 4.06 MPa

• cold, 3.34 MPa

0 20 40 60 80 100 120

Average Downstream Distance (cm)

140

Figure 157. Average velocities for releases of room temperature and cooled FC-218

determined by the lime of arrival at laser beams #l-#5 plotted as a function

of average downstream distance. Vessel pressures are indicated.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 275

None of the time plots of near-field dynamic pressure showed an increase following the bursting

of the disk. This indicates that the disk bursting does not generate a shock wave, and that the strong

flashing behaviors which are characteristic of downward releases do not occur for the upward cases. Figure 158 shows the laser attenuation signals over the first 200 ms for a release of FC-31-10.

Laser #5 is totally attenuated very rapidly when the disk bursts. The flow then allows light to pass

for a very short period before extinguishing the beam once again. The high-speed film shows that the

second attenuation occurs as the result of an increase in optical density in the gas within the vessel

which is subsequently transported through the orifice. Presumably this is agent vapor which

condensed as the nitrogen cooled during expansion. At roughly 18 ms the flow becomes partially

transmitting for a brief period. The film shows the same behavior. It appears as if the nitrogen flow

has ceased before the agent begins to boil in the vessel. At roughly 24 ms following release, the

liquid begins to boil at the bottom of the vessel and vapor is once again forced through the orifice leading to total attenuation of laser #5. At approximately 54 ms after release the boiling liquid fills the vessel and the flow from the orifice begins to spread much more rapidly, Since laser ,^^5 is already totally attenuated, no eff'ect is observed on the laser signal.

Further downstream the lasers are attenuated in different ways. Laser #4 is nearly totally

attenuated very shortly after the disk opens, but then is partially transmitted for a period before the

two-phase flow arrives. The degree of laser beam attenuation due to the initial flow of nitrogen

decreases with increasing distance from the nozzle, but is still observable at laser #1. The furthest

laser detects the flow very shortly (approximately 5 ms) after the release indicating the flow velocity

is quite high. Figure 159 shows the average velocities determined for the initial gas flow as a function of distance from the orifice. At progressively later times, lasers #4-#l show a strong

extinguishment which is attributed to the arrival of condensed agent from the vessel. Estimates of the

arrival times for this flow have been made and used to calculate average velocities which are also

included in Figure 159. Lasers #3-#l become partially transmitting after the initial arrival of the twophase flow. The flow continues to transmit light for periods of tens of milliseconds before all three

lasers are again totally attenuated. The later attenuations are attributed to the arrival of the two-phase

flow generated by the boiling liquid filling the vessel. The arrival times of this third flow at lasers #3-#l have also been estimated and used to calculate average velocities. Results are shown in Figure 159.

Three points should be noted concerning Figure 159. First, the velocities of the two-phase flows

are considerably less than the velocities associated with the nitrogen release. Second, the average

two-phase flow velocities are much lower than observed for comparable downward releases of

FC-31-10 (see Figure 111). Three, the velocities of the two two-phase flows are similar. Figure 160 shows a similar time plot for the laser signals over a much longer period of time. It

is now clear that an optically dense flow is present until roughly 800 ms after the disk opens and the

period required for full release of the agent approaches 1 s. This result contrasts dramatically with

the observed emptying times when this agent is released downward which are on the order of 40 ms

for a two-thirds full vessel. The signals recorded by the downstream pressure transducer and aspirated hot-film are consistent

with the above conclusions. Results over short and longer time periods are shown in Figure 161 and

Figure 162. The arrival of the released nitrogen at the position 1.3 m from the vessel is clearly

detectable in the hot-film signal even though it causes only a very small dynamic pressure rise (Figure 161) despite the high velocity. The response of the hot-film requires that this initial gas flow

either contains some of the agent, or that it is substantially cooled by the nitrogen. On the other

hand, when the first flow identified as consisting of two phases arrives at 75 ms following release,

there is a substantial increase in both the hot-film and pressure signals consistent with the presence of

a two-phase flow. Note that the velocity of the two-phase flow is considerably lower than the flow
276 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 158. Short time behaviors for transmission of laser beams #5-#l spaced as shown in Table 7 following an upward release of FC-31-10 at 0 s. Conditions: 505 g

agent, 4.40 MPa release pressure.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 277

CO

>^

• o

_o

>

a

0)

>

<

200

175

150

125

100

75

50

25

0

FC-31-10, Upwards

1 1 1 1 1

O

-

-

o

-

o o Nitrogen

• First Two-Phase Flow

V Second Two—Phase Flow

•

1 1 V

•

1 1 1 0 20 40 60 80 100 120

Average Downstream Distance (cm)

Figure 159. Average velocities for upward releases of FC-31-10 determined by the time

of arrival at laser beams #5-#l plotted as a function of average downstream

distance. Results for arrival of nitrogen and two two-phase flows are indicated (see text).
278 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10. Upwards

0 200 4O0 600 800 1000 1200 1400

Time (ms)

Figure 160. Long time behaviors for transmission of laser beams #5-#l spaced as shown in Table 7 following an upward release of FC-31-10 at 0 s. Conditions: 505 g agent, 4.40 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 279

Time (ms)

Figure 161. Voltage output of the aspirated hot-film and the dynamic pressure for probes located 1.3 m from the vessel are plotted for a short time period

following an upward release of FC-31-10. Conditions: 505 g of agent,

release pressure of 4.40 MPa.
280 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-31-10, Upwards

1 1 1 1 r 800 a_

- 600

- 0

- -200

-600

200 400 600 800 1000 1200 1400

Time (ms)

CO 400 CO

(D

CL

200 o

Q

C

Q

E

a

CD - -400 ^COc

oQ

Figure 162. Voltage output of the aspirated hot-film and the dynamic pressure for probes located 1.3 m from the vessel are plotted for a long time period

following an upward release of FC-31-10. Conditions: 505 g of agent,

release pressure of 4.40 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 281

resulting from the nitrogen release (see Figure 159). Following the first arrival of two-phase flow,

both signals fall consistent with laser #1 being transmitted during this period. The arrival of the

second two-phase flow is accompanied by a large dynamic pressure jump and much stronger

indications of two-phase flow in the aspirated hot-film signal. Figure 162 shows that the later twophase flow persists until roughly 800 ms, again consistent with the attenuation behavior of laser beam #1.

Figure 163 and Figure 164 show the laser intensity behaviors observed for upward releases of

FC-218. For short times (Figure 163), the results are similar to those for the FC-31-10. When the

disk bursts, laser #5 is immediately fully attenuated. Some transmission is then observed shortly after the disk opens before the laser is totally attenuated a few milliseconds later due to condensation of the

agent vapor in the nitrogen. Transmission is observed for a period from roughly 12 to 20 ms

following the nitrogen release. At this point, boiling in the vessel begins from the bottom and

expulsion of gas from the vessel again extinguishes the laser beam. Note that the boiling begins at a somewhat shorter time than observed for the FC-31-10. High-speed films indicate the vessel is filled with boiling liquid at 39 ms following release and the more vigorous two-phase flow should form at this time.

The attenuation signals for lasers #4-#5 do not show the distinct arrival of two separate types of

two-phase flow. However, it is possible to estimate roughly when the flow generated by the twophase flow formed when the vessel is completely filled with boiling liquid reaches the lasers. For the

position of laser #5 we have used the high-speed film to provide the starting time for the flow. The

arrival times have been used to calculate the average velocities for the downstream edge of the twophase flow shown in Figure 165. Velocities for the nitrogen release are also included in this figure.

The results are very similar to those for FC-31-10 shown in Figure 159.

Figure 164 shows that the total attenuation of the laser beams lasts until roughly 250 ms after the

release and that the laser beams are nearly completely transmitted at 500 ms. These results suggest

that the upward release of FC-218 from the vessel requires only 1/3 as long as a similar release of

FC-31-10.

Time plots of the downstream dynamic pressure and aspirated hot-film voltage for the FC-218

release are shown in Figure 166 and Figure 167. Figure 166 indicates that the two-phase flow

reaches the sampling position roughly 55 ms following release. This is attributed to the two-phase

flow which extinguishes laser #5 at 20 ms following release. A much higher pressure rise appears

starting at roughly 80 ms. This time is consistent with the attenuation behavior seen for laser #1.

The hot-film signal is consistent with the arrival of a strong two-phase flow. The plots of these

signals over longer time periods (Figure 167) show that the two-phase flow is present until roughly

250 ms following release, consistent with the conclusion based on laser measurements.

The plastic vessel was not used for releases of halon 1301 so the internal behavior cannot be

described. However, high-speed films of releases outside of the vessel indicated that the behaviors

for halon 1301 were similar to those already described. The first two-phase flow appeared at 16 ms

following release and the two-phase flow from the filled vessel appeared at approximately 30 ms.

Figure 168 through Figure 171 show measurement results for a release of halon 1301. Short and

long time plots of laser intensities are given in Figure 168 and Figure 169. For the short time data

(Figure 168), the behaviors observed are similar to those described for FC-31-10 and FC-218. It is

difficult to estimate arrival times for the two-phase flow at the positions nearer the orifice so

velocities will not be reported. However, it is clear that two diff'erent two-phase flows reach laser #5

at roughly 30 and 50 ms. The eff'ects of these flows are also evident in the hot-film and dynamic

pressure signals shown in Figure 170. As found for the other agents, the second flow generates much

higher dynamic pressures and hot-film response. This is attributed to the arrival of a two-phase flow

generated by boiling when the vessel is totally filled with a two-phase mixture. Figure 169 and
282 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218, Upwards

0 20 40 60 80 100

Time (ms)

Figure 163. Short time behaviors for transmission of laser beams #5-#l spaced as shown

in Table 7 following an upward release of FC-218 at 0 s. Conditions: 461 g

agent, 4.35 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 283

Figure 164. Long time behaviors for transmission of laser beams #5-#l spaced as shown

in Table 7 following an upward release of FC-218 at 0 s. Conditions: 461

g agent, 4.35 MPa release pressure.
284 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

CO \

o

_o

>

O

(D

>

<

200

175

150

125

100

75

50

25

0

0

FC-218, Upwards

O

O

O

O

O Nitrogen

• Two-Phase

20 40 60 80 1 00 1 20

Average Downstream Distance (cm)

Figure 165. Average velocities for upward releases of FC-218 determined by the time of

arrival at laser beams #5-#l plotted as a function of average downstream

distance. Results for arrival of nitrogen and the two-phase flow are indicated

(see text).
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 285

i

FC-218, Upwards

0 20 40 60 80 100 Oi

Time (ms)

Figure 166. Voltage output of the aspirated hot-film and the dynamic pressure for probes

located 1.3 m from the vessel are plotted for a short time period following an upward release of FC-218. Conditions: 461 g of agent, release pressure of 4.35

MPa.
286 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Time (ms)

Figure 167. Voltage output of the aspirated hot-film and the dynamic pressure for probes located 1.3m from the vessel are plotted for a long time period

following an upward release of FC-218. Conditions: 461 g of agent,

release pressure of 4.35 MPa.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 287

Figure 168. Short time behaviors for transmission of laser beams #5-#l spaced as shown in Table 7 following an upward release of halon 1301 at 0 s. Conditions: 545 g

agent, 4.46 MPa release pressure.
288 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 169. Long time behaviors for transmission of laser beams #5-#l spaced as shown in Table 3 following an upward release of halon 1301 at 0 s. Conditions: 545 g

agent, 4.46 MPa release pressure.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 289

>

13

Cl

Ll_

14

halon 1301, Upwards

1 r

12 -

3 10

O

8

-

6

-

0 10 20 30 40 50 60 70

Time (ms)

o

CL

0

13

CO

CO

CD

CL

O

c

Q

D

0

CO

c

o

Figure 170. Voltage output of aspirated hot-film and dynamic pressure for probes

located 1.3 m from the vessel are plotted for a short time period following an upward release of halon 1301. Conditions: 545 g of agent, release

pressure of 4.46 MPa.
290 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0 50 100 150 200 250 300 350 400 450

Time (ms)

Figure 171. Voltage output of aspirated hot-film and dynamic pressure for probes

located 1.3 m from the vessel are plotted for a long time period following an upward release of halon 1301. Conditions: 545 g of agent, release

pressure of 4.46 MPa.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 291

Figure 171 show that the two-phase flow from the vessel lasts until roughly 200 ms following the

start of the release.

3.5.4 Discussion

3.5.4.1 Downward Release of Pressurized Superheated Liquids. The experimental findings

provide a great deal of insight into the mechanisms responsible for the dispersion and vaporization of

superheated liquids from pressurized bottles. To our knowledge, this investigation represents the first systematic investigation of these important physical processes which are expected to be crucial in the

ability of an agent to extinguish a fire in a compartment. These findings are particularly relevant for the aircraft dry bay application. In the following we will summarize the important conclusions of the

study and attempt to provide physical explanations for the observations.

The following discussion will be clearer if one considers the following figures which show

calculated values of Jakob number (Figure 172) for each agent as a function of agent boiling point and

the dynamic pressures (Figure 173 and Figure 174) expected for a given velocity and a range of

densities chosen to span pure air to those typical of the liquid agents.

The Jakob number {Ja) is a dimensionless number which estimates the fraction of a superheated

liquid which can be vaporized adiabatically. In other words, it is the ratio of the sensible heat within

the liquid, which can be extracted from the liquid by cooling to its boiling point for the ambient

pressure, and the heat of vaporization, aH^^^, for the liquid. Mathematically, it is defined as

Ja - (35)

where Ah is the change in enthalpy of the liquid on cooling from the ambient temperature (T^,,,^) to the boiling point of the liquid (7^^,) for the ambient pressure,

C I is the constant pressure heat capacity for the liquid which varies slightly with temperature.

Values of Ja calculated for each of the potential replacement agents, with the exception of FC-116,

and halon 1301 have been calculated. The results are plotted in Figure 172 and listed in Table 8. In order to calculate the specific heats at a constant pressure of one atmosphere for temperatures

ranging from room temperature to the liquid boiling point, the specific heat was assumed to be

independent of pressure. The specific heats were fit to a second- or third-order polynomial and

integrated from the boiling point to room temperature (293 K) to yield the change in internal energy. The specific heat was also fit by a spline fit technique. Results for the two approaches are included in Table 8. The two sets of calculations are in excellent agreement.

The required specific heats of HFC-227ea, HCFC-22, HFC-134a, HCFC-124, HFC-125,

FC-218, FC-318, and halon 1301 were obtained from the PROZPER program (Ely and Huber,

1993). The values for FC-31-10 were estimated by the method of Tarakad, Danner, and Lee-Kesler

(Daubert and Danner, 1992). The specific heat of the HFC-125/HFC-32 mixture was calculated

using Rowlinson's corresponding states method (Reid et al., 1987) coupled with the Genetron

equation (Allied Signal, 1991) for the ideal heat capacity. Values for HFC-236fa were calculated

using Rowlingson's corresponding states method and the method of Joback (Reid et al., 1987). Heats

of vaporization are those listed in Table 1 of Section 2.1.

amb (36)
292 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

0.7 r—

0.6 A FC-218

CD

E

15

_Q

O

O

0.5

0.4

0.3

0.2

^HFC-125

halon 1301

HFC-125/HFC-32

O

HCFC-22

O

HFC- 134a A

HFC-227

FC-318

HCFC-124 V

FC-31-10

FC-236

0.1

200 225 250 275

Agent Boiling Point (K)

Figure 172. Values of Jakob number are plotted as a function of boiling point for each

of the potential replacement agents, except FC-116, and halon 1301.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE

0 25 50 75 100 125 150 175 200

Velocity (m/s)

Figure 173. Calculated dynamic pressures for: A) liquid halon 1301, B) liquid HFC125/HFC-32 mixture, C) 50% liquid/50% vapor mixture FC-218, D) 25%

liquid/75% mixture FC-218, E) 10% liquid/90% gas mixture FC-218, and

F) room-temperature air.
294 SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE

O I I I I I I I I I

I I 0 10 20 30 40 50 60 70 80 90 100

Velocity (m/s)

Figure 174. Calculated dynamic pressures for: A) 10% liquid/90% mixture FC-218, B)

5% liquid/95% mixture FC-218, C) 1% liquid/99% mixture FC-218, and

D) room-temperature air.
SECTION 3 . FLUID DYNAMICS OF AGENT DISCHARGE 295

Table 8. Jakob Numbers for Room Temperature (293 K) Releases of Ten Potential Replacement Agents and Halon 1301. Heat Capacity Temperature Dependencies have been

Fit Using Polynomial and Spline Fits.

AUtJN 1 Ja (Polynomial Fit) Ja (bphne hit)

HFC-236fa 0.165 0.165

FC-31-10 0.236

FC-318 0.255 0.255

HCFC-124 0.228 0.228

HFC-227ea 0.306 0.306

HFC- 134a . 0.283 0.283

FC-218 0.587 0.578

HCFC-22 0.313 0.313

HFC- 125 0.520 0.520

HFC-125/HFC-32 0.391 0.391 •

halon 1301 0.509 0.509

Note that three of the agents, HFC- 125, FC-218, and halon 1301, have higher Ja than the

remaining agents. The value for the HFC-125/HFC-32 mixture is intermediate. All of the other

agents have values which fall below 0.35.

During the course of this investigation dynamic pressures were recorded near the vessel exit and on the flow centerline 1.3 m from the vessel orifice. Since actual velocities or densities are not

available, it is not possible to derive density or velocity information from these measurements using

Equation (34). However, the magnitudes of the dynamic pressures do provide insights into the nature of the flows. This becomes clear when one considers Figure 173 and Figure 174.

Figure 173 shows the calculated dynamic pressure as a function of velocity for densities corre- sponding to those for liquid halon 1301 (1551 kgW), liquid HFC-125/HFC-32 mixture (1040

kg/m^), a 50% liquid/50% vapor mixture of FC-218 (665 kg/m-^) at its boiling point temperature, a 25% liquid/75% gas mixture of FC-218 (338 kg/m^), a 10% liquid/90% gas mixture of FC-218 (141

kg/m^), and room-temperature air (1.20 kg/m^). The dependencies of the calculated dynamic

pressures on density and velocity are clear. Note that for high velocities and densities, very large

dynamic pressures are attained, while for air the pressures are relatively low for the entire velocity

range.Figure 174 is a similar plot of dynamic pressure, but the velocity and density ranges have been

reduced. Results are for a 10% liquid/90% gas mixture of FC-218 (141 kg/m^), a 5% liquid/95%

gas mixture of FC-218 (75.3 kg/m^), a 1% liquid/99% gas mixture of FC-218 (22.8 kg/m^), and

room-temperature air (1.20 kg/m^). The efl'ect of even a small fraction of liquid phase agent on

measured dynamic pressures is obvious from this figure.
296 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

The two most dramatic events that were identified in the high-speed films of the releases were

the violent mixing of agents with their surroundings observed at the initiation of a release and at times

close to when the final liquid agent exits the containment vessel. These events were termed

"flashings" even though it will become clear that term is most likely inappropriate for the later of

these events. The initial rapid spreading of the agent as the burst disk opens is fairly easy to explain qualitatively by considering the events when the disk opens, even though the actual behaviors are exceedingly complex to describe quantitatively. Immediately following the rupture of the disk, the liquid agent

in proximity to the orifice experiences a rapid depressurization to the outside ambient pressure. At

this point the liquid is not in motion. Pressure measurements in the liquid within the vessel indicate

that this pressure drop is not transmitted into the vessel which means the drop must occur within the

orifice. The liquid at the interface should undergo a rapid flashing. This flashing may be induced by

local perturbations associated with the bursting of the disk or may even take place as a miniature

boiling liquid expanding vapor explosion (BLEVE). Such processes have been discussed in Sections

3.2.4 and 3.4.3.5.

The first detectable pressure drops inside the vessel within the nitrogen occur 0.36 to 0.76 ms

following the near-total attenuation of laser beam #1, with the majority requiring roughly 0.4 ms.

Since pressure changes are expected to be transmitted at the speed of sound, less than 1 ms is required for pressure fluctuations to travel from the orifice to the internal pressure transducer at the

top of the vessel. The observed delay is therefore attributable to the period required for the liquid to begin to accelerate through the orifice and thus drop the nitrogen pressure due to a volume increase of

the ullage. The internal pressure fits discussed in Section 3.4.3.1 indicate that the liquid flow reaches

a choked condition in a very short period of time.

Even though the initial flashing behavior is evident in all of the high-speed movies of the

releases, it was only detected by the near-field pressure transducer for the lower boiling-point liquids.

This demonstrates that shock waves are not generated by the disk bursting, and that any signals

recorded by the transducer must either result from shock waves generated by liquid flashing or direct

impingement of a two-phase flow on the transducer. The periods required for the near-field external

pressure transducer to sense the releases varied from 1.28 to 1.84 ms. Such time delays are consistent with estimates for radial velocities listed in Table 5 based on high-speed-films measurements.Once the pressure at the orifice is released, the liquid in the vessel begins to accelerate and flow

through the orifice. The high-speed films indicate that the flows quickly form a liquid stream. Based on measurements given in Section 3.4.3.1, the flow velocities are on the order of 60 m/s.

The "flashing" near the conclusion of the release of the liquid agent is at first glance more

difficult to understand. The films show a very vigorous event which appears to occur as the result of

very rapid boiling of the agent. For all of the agents the near-field pressure transducer recorded

sharp increases in dynamic pressure which varied from several hundred to 1600 kPa (see Figure 88

through Figure 98).

Reference to Figure 173 and Figure 174 shows that the fluid striking the pressure transducer

must have high velocities and densities. Such a behavior is difficult to attribute simply to flashing of

the liquid. The key is to recognize that a release of a vessel's contents is not complete once the liquid

is expelled. A significant pressure of nitrogen and vaporized agent remains. These gases will now

exit the vessel. The question is: How are the gases released from the vessel? The answer has

already been provided in Section 3.4.3.1. The flow becomes that of a sonically choked gas flow

through the orifice.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 297

By assuming that there is only ideal-gas nitrogen in the vessel, it is straightforward to estimate a flow velocity for the gas. Assuming an initial pressure of 4.14 MPa and that the liquid fills two

thirds of the vessel, the expression for the adiabatic expansion of a gas,

P,V] = PfVj, (37)

allows the initial temperature for the gas release to be calculated. Here P and V are pressure and

volume, the subscripts / and / refer to initial and final conditions, and 7 is the ratio of constant pressure and constant volume heat capacities. Using the ideal gas law this becomes

T = r.(_/) Y , (38)

/ p

where T is the temperature. For V, = 0.000333 m-^ and = 0.000500 m-^, these equations yield

Pjr = 2.40 MPa and 7^ = 25 IK. With these values, it is now possible to calculate the sonic velocity

using

where R is the gas constant and M is the molecular weight for the gas. Substituting in Equation (39)

gives U^^^i^ = 323 m/s.

It is now clear that the flow leaving the vessel will accelerate dramatically once the liquid has

been expelled. This rapidly expanding gas quickly catches up with the slower moving, but much

denser, liquid or two-phase jet. To the gas stream, the liquid will appear much as a solid body due to

its much higher density. As a result, the gas should pass around the agent flow which explains the

very rapid expansion of the jet observed in the films. The nitrogen flow is also expected to interact

with the liquid or two-phase flow, both fragmenting it and mixing with it. As a result of this mixing,

a two-phase mixture will be generated. It is presumably such mixtures which are detected by the

dynamic pressure transducer. These flows can be both high speed and relatively high density due to their two-phase character, thus explaining the very high dynamic pressures which are observed.

To our knowledge, this is the first report of such an aerodynamic mechanism for rapidly

dispersing releases of liquid agents. It is unclear how significantly this process will aff'ect the overall dynamics for the mixing of an agent with its surroundings. However, it seems likely that for shortperiod releases, such as those investigated here, the eff'ects can be substantial, and that the release of

high pressure nitrogen can substantially augment the mixing of a released liquid agent as the result of

its own velocity and flashing behavior.

The velocities measured from the extinguishment of the laser beams provide insights into the

flashing behaviors of the various agents. The measured velocities nearest the vessel exit were

between 50 and 60 m/s for HFC-236fa, FC-31-10, FC-318, and HCFC-124 (see Figure 110 through

Figure 113). The results for HFC-227ea (Figure 114) are only slightly higher. These velocities are consistent with the flow velocity of the liquid stream from the vessel suggesting that for these agents

the liquid stream extends at least 0.32 m downstream of the orifice. These observations are also

consistent with the observation of a liquid stream in the high-speed films {e.g., see Figure 86).

For each of these high boiling-point agents the velocity was higher at the second measurement

position indicating that flashing had begun and the flow was accelerating due to the volume expansion

associated with the flashing. Assuming an average liquid flow velocity of 60 m/s, this means that the

period required for these agents to flash lies somewhere between 5 and 10 ms.
298 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

As noted earlier, the measured velocities for HFC- 134a (Figure 115) suggested that this agent was an intermediate case. In two experiments the highest velocities were observed between lasers ^\

and #2 suggesting that flashing of the liquid jet occurred somewhere in between. One case showed

the flashing taking place between lasers #2 and #3. For the two cases where flashing occurred near

the vessel, the velocity then fell ofi" with increasing distance for distances beyond laser #3. This

suggests that flashing is complete and the resulting two-phase flow is slowing down as it mixes with

surrounding air as would be expected for a gas jet.

The remaining low boiling-point liquids, FC-218, HCFC-22, HFC-125, the HFC-125/HFC-32

mixture, and halon 1301 all have high velocities near the vessel exit, and then slow down significantly

with increasing distance. Again, this is taken to mean that the agents rapidly expand near the exit due

to rapid boiling and then spread more slowly as two-phase flows entraining air. The fact that all have

increased velocities near the vessel exit indicates that the time required for flashing must be significantly less than 5 ms.

Additional evidence for this conclusion is found in the near-field pressure traces for FC-218,

HCFC-22, HFC-125, the HFC-125/HFC-32 mixture, and halon 1301 (Figure 94 through Figure 98).

These traces show that dynamic pressures are recorded not only at the start and end of a release, but

during the release. This can only be the case if the fluid leaving the vessel is flashing over a very

short flow distance. This flashing generates the two-phase flows necessary for the observed pressure

increases. The pressure increases for the intermediate times are strongest for the HFC-125/HFC-32

mixture and halon 1301 consistent with their low boiling points. A very rapid expansion of halon

1301 is also consistent with the rapid spreading of this flow which can be seen in Figure 87.

The general trends observed for the flashing times are consistent with those discussed by Day

(1969) and Lienhard and Day (1970). The times decrease with increasing superheat which is directly related to the difl'erence in pressure between the ambient and saturated values. Using Equations (31)-

(32), it is possible to derive the following expression for the flashing time.

This equation has been tested for those agents for which surface tension value estimates were

available from Daubert and Danner (1992). Other values for the variables in Equation (40) are taken

from Table 1 of Section 2. Table 9 lists the values of surface tension used along with calculated

values of t^.

It is clear that Equation (40) vastly underestimates the times required for the liquid streams to

flash. At the present time we are unable to provide an explanation for this failure. On the basis of

this finding, it is concluded that the prediction of flashing lengths for these agents is currently not

possible.

The discussion thus far has shown how flashing of the agents leads to enhanced mixing of the

two-phase flow with its surroundings. It is clear that the amount of superheating is crucial to the

speed and violence of this flashing. As noted earlier, it is also necessary for the agents to be

vaporized in order to be effective. The downstream dynamic pressure measurements and hot-film

response provide insights here. All of the high boiling-point agents show similar behaviors. When the agent first reaches the

pressure transducer there is a rapid jump in the pressure. Following this jump the pressure continues

to increase for a period roughly equal to the time required for release of the liquid agent from the

containment vessel. This observation suggests that velocity measured from the propagation of the

downstream edge of the release should provide a good estimate for the entire release period.

For HFC-236fa, FC-31-10, FC-318, HCFC-124, and HFC-227ea these velocities were on the

order of 70 m/s. The maximum dynamic pressures observed for these agents were in range of 800 to

(40)
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 299

Table 9. Values of Time Required for Flashing of Superheated Liquid Streams Calculated

Using Equation (40)

Agent Surfece Tension (N/m) Calculated (s)

FC-31-10 0.0153 4.0 X 10"^

FC-318 0.0117 8.5 X 10"^ HFC- 134a 0.0230 2.7 X 10"^

FC-218 0.0159 1.3 X 10-^ HCFC-22 0.0184 1.1 X 10"^

HFC-125 0.0171 3.3 X 10"^

halon 1301 0.0156 2.1 X 10-9

1600 kPa with a typical value being 1200 kPa. Using Equation (34), the density necessary to generate such a pressure rise is on the order of 430 kg/m^. The densities of these liquids range from

1356 to 1499 kg/m^. This suggests that roughly 30% of the mass for these agents reaching the 1.3 m

downstream position on the centerline of the flow is not vaporized. As discussed in the experimental

section, the aspirated hot-film responses were consistent with the presence of two-phase flows at this downstream location for all of the releases. The HFC-134a results are somewhat difl'erent. The laser measurements (Figure 115) indicate the

velocity for this flow at the 1.3 m position is on the order of 50 m/s, but the dynamic pressures

observed were on the order of 2000 kPa yielding a density of roughly 1600 kg/vn?. This estimate is higher than the original density of the liquid which is 1209 kg/m^. This suggests that this agent is only slightly evaporated over the flow distance and its velocity must be somewhat higher than

estimated. It is difficult to understand this finding given the behavior of the other agents. However,

the experimental results were repeatable.

The results for the low boiling-point liquids varied dramatically from agent to agent. When

FC-218 reached the pressure transducer there was a fairly rapid increase in dynamic pressure to values on the order of 100 to 200 kPa (Figure 127). Figure 116 indicates the flow velocities are on

the order of 40 m/s. This suggests that the local density is roughly 125-250 kg/n?. Since the liquid

density is 1321 kg/m?, this suggests that 10-20% of the agent is present as liquid (also see Figure 173

and Figure 174).

For HFC-125 the dynamic pressures recorded when the agent first reached the pressure

transducer (Figure 129) were ver>' low, being less than 10 kPa. Measured velocities (Figure 118) were on the order of 18 m/s. This indicates that the flow density was less than 60 kg/m?. Based on

the liquid density for HFC-125 (1190 kg/m^), this suggests that much less than 5% of the flow is in the liquid state. The results for halon 1301 were variable. One of the releases (B in Figure 131) gave no

measurable pressure increase when the agent first arrived. The pressure eventually began to rise slowly after 20 ms and then increased at the end of the release to almost 400 kPa. The hot-film

response (Figure 135) indicates that two-phase flow is associated with the large pressure rise. Flow

velocities are on the order of 30 m/s (Figure 120). These observations suggest that the flow initially
300 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

arrives nearly fully vaporized, but that near the end of the release the flow density increased to roughly 900 kg/vn?. This observation is difficult to understand until one realizes that the large

dynamic pressure increase is likely associated with the flow generated by the release of nitrogen from

the vessel. Flow velocities are expected to be much higher for this case. It seems likely that the high

pressure results from a high-speed flow of nitrogen, which has entrained some two-phase halon 1301,

striking the pressure transducer. Recall it was concluded that a two-phase mixture of halon 1301 is released from the vessel near the end of the liquid release due to flashing within the nozzle or vessel

(see Section 3.4.3.1). The high-speed nitrogen flow would be likely to entrain the resulting twophase flow. The results for FC-218, HFC-125, and halon 1301 indicate that most of the liquid agent is vaporized over the 1.3 m flow length. This is in sharp contrast to HCFC-22 where very high

dynamic pressures are generated during most of the time the agent flow is striking the transducer

(Figure 128). Maximum dynamic pressures were on the order of 2000 kPa. Using an average

velocity of 40 m/s (Figure 117), these pressures yield calculated densities on the order of 2500 kg/m-^

which can be compared to a liquid density of 1192 kg/m-^. This is another case where the flow

velocity estimate must be low. However, it seems clear that a large fraction of the agent arrives in the liquid state at the 1.3 m location.

The HFC-125/HFC-32 mixture also generated complicated dynamic pressure time dependencies

(Figure 130). When the agent first reached the transducer there were pressure increases on the order

of several hundred kPa. Using an estimate for the flow velocity of 30 m/s from Figure 119, this yields densities on the order of 700 kg/m? or roughly 60% liquid. The aspirated hot-film response

(Figure 137) for this agent also indicated the presence of a strong two-phase flow. Near the end of

the releases there were strong pressure increases which are likely associated with an increase in flow

velocity, perhaps as a result of the nitrogen release. The conclusions from the above discussion can be summarized as: • The high boiling-point agents generate flows which consist of roughly 30% liquid on the centerline at 1.3 m from the vessel. • FC-218, HFC-125, and halon 1301 produce flows which have most of the agent (> 90%) in the gas state. • HCFC-22 and the HFC-125/HFC-32 mixture yield pressure traces which suggest that a large

fraction of their mass is still liquid at the measurement location.

It is impossible to provide a detailed description of the physical processes which are occurring

during the mixing following a high pressure release of a superheated liquid due to the complex and

interactive nature of the physical processes which are taking place. Flashing generates a two-phase

flow at some unknown position which has unknown velocity, temperature, and droplet size distributions. This two-phase flow then begins to entrain room-temperature air. For temperatures near 295 K, the entrained air will also introduce heat into the flow which can evaporate additional liquid.

The short duration of the flows further complicates attempts to understand the behavior since flow

development is likely to last for a large fraction of the total flow time. As an example, the impulsively started flow should immediately interact directly with room temperature air. However, as the flow

develops, the flow volume will contain more and more agent and air will enter only from outer regions of the flow. As a result, the amount of agent which is vaporized as a function of downstream

distance is expected to vary with time. This may explain some of the time dependencies observed in the downstream dynamic pressure measurements.

The lack of detailed measurements also makes a complete understanding impossible. Most of the

downstream measurements are only valid for the centerline of the flow and the fraction of two-phase

flow is only estimated for two locations. No information is available concerning the radial structure of the flow.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 301

Despite the limitations of the measurements, it is possible to draw some general conclusions

concerning the evaporation of the agents. The measurements suggest that the high boiling-point

agents are only roughly 70% vaporized at the 1.3 m downstream position. Referring to Figure 172 it can be seen that these six agents, including HFC- 134a, are clumped together at the lower right-hand

side of the plot of Ja versus agent boiling point. The Ja numbers are all 0.3 or less indicating that the superheating of these agents is sufficient to vaporize only a small fraction of the agent. In order

for the remainder of the agent to be vaporized, additional heat must be absorbed from the surroundings. The fact that these agents reach the pressure transducer as two-phase flows indicates that

insufficient heat is available to fully vaporize the agents. Since it was estimated that roughly 30% of

the agent flows was liquid at the transducer, this suggests that sufficient air was entrained to vaporize

half of the remaining liquid.

It should be pointed out that the actual dynamic pressure which is measured will be a complicated function of flashing behavior and the degree of vaporization. This is due to the diff'erent dependencies of the dynamic pressure on density and velocity (see Equation (34)). Flashing of an

agent increases its velocity and decreases it density due to the accompanying volume expansion. As a

result, it is possible to either increase or decrease the measured dynamic pressure. This efi'ect may

explain the high densities calculated for some of the flows such HFC-134a.

Based on this discussion, it is now possible to gain some insight into the vaporization behaviors

of the low boiling-point liquids. The three liquids which seemed to be nearly fully vaporized were FC-218, HFC-125, and halon 1301. Figure 172 shows that these three liquids have the highest Ja

numbers of any of the agents tested. In each case, the agent can provide over 50% of the heat

required for full evaporation. They also have quite high superheats, and, as a result, they tend to flash strongly near the vessel exit. The resulting two-phase flow has a relatively long distance over which to entrain air. The wide dispersion of the agent due to flashing also aids in the evaporation

process. Apparently, sufficient air is entrained into these flows to provide the heat necessary to complete the evaporation for these agents.

HCFC-22 is an interesting case. Even though its boiling point is lower than that for FC-218, its Ja is only slightly more than 0.3. As a result, flashing vaporizes considerably less of the fluid than

for the low boiling-point liquids just discussed. Another peculiarity of this agent is that for constant

fill level in the vessel (generally 3.33 x 10"* m^), significantly higher number of moles of this agent

are released during the same liquid flow period (4.6 moles versus 2.3, 3.3, and 3.6 for FC-218,

HFC-125, and halon 1301, respectively). This means the amount of expansion and acceleration of

the liquid will be higher for a given percentage of vaporization. This expansion will tend to shield

the liquid from entrained air and may well contribute to a slower rate of agent evaporation. The

combination of low Ja and high number of moles most likely accounts for the very high downstream

dynamic pressures observed for this agent (see Figure 128).

The same arguments provide an explanation for the two-phase character of the HFC-125/HFC-32

mixture which was observed. Its Ja number is intermediate between that of halon 1301 and

HCFC-22. The liquid itself only provides 40% of the heat necessary to vaporize all of the liquid. In

addition, this agent provides, by far, more moles of agent (5.2 moles) than any other for a release

from the 2/3 full vessel. Again, the degree of expansion and acceleration associated with this large

number of moles is likely to limit the entrainment of air and complete vaporization of this agent.

On the basis of the mixing results, three parameters have been identified which control the

degree of vaporization of an agent. These are the amount of superheat, Jakob number, and the

number of moles contained in the liquid release. Efficient evaporation is favored by high superheat,

high Ja, and a low number of moles of fluid in the vessel.
302 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

3.5.4.2 Effects of Orifice Diameter on Mixing and Evaporation. The orifice size affects the

dynamic pressure measurements in the near field, the dispersion velocities, and the degree of

evaporation at the far stream position. For FC-31-10 the maximum dynamic pressure associated with

the release of the nitrogen using the 12.7 mm opening was roughly the same as for the 19.1 mm

opening, while the 6.4 mm opening generated no pressure increase. For FC-218 and halon 1301 the

intermediate size orifice yielded lower maximum overpressures, but also provided indications that flashing was taking place over much of the liquid release periods. Such an indication of continual

flashing for the largest nozzle was not observed for FC-218. For both of the low boiling-point liquids no near-field pressure increases were observed when the smallest orifice was used.

These dynamic pressure results can be understood in terms of the expected flashing behavior of

the agents and the effects of the released nitrogen. Due to the low superheat of the FC-31-10, this agent never flashes in the near field. The pressure rise detected for the 12.7 mm orifice is the result of high-speed nitrogen striking the slower moving liquid. For the smaller orifice, no pressure rise is detected due to the reduced mass flow of both the liquid and the nitrogen gas. The flow simply does

not expand sufficiently to reach the pressure transducer.

The most interesting observation for the FC-218 was the flashing of the liquid for flow through

the 12.7 mm opening, which was not observed when the 19.1 mm opening was used. This is likely

due to a reduction in the flashing time for the smaller orifice which caused the flashing to occur

nearer the nozzle with the result that two-phase flow struck the transducer. Such a behavior is consistent with the dependence of flashing time on liquid stream diameter reported by Lienhard and

Day (1970). The reduction of the maximum dynamic pressure associated with the release of nitrogen on going from the 19.1 to 12.7 mm orifices and the absence of a pressure rise for the 6.4 mm orifice are explained by the reduction in mass flows of the liquid and nitrogen as discussed above for FC-31-10. The near-field pressure behaviors for halon 1301 are consistent with the conclusions for FC-218.

The average velocity measurements with downstream position shown in Figure 139 through

Figure 141 indicate that, in general, the velocities nearest the vessel increase as the orifice size is decreased and that those at the farthest downstream position decrease. These higher velocities are

consistent with the liquid flashing earlier and more strongly near the vessel when the liquid stream

diameter is reduced as predicted by Lienhard and Day (1970). Due to the smaller diameters and

lower mass flows of the liquid jet through the smaller orifices, the two-phase flows will mix with air and spread faster over shorter distances in a manner similar to turbulent axisymmetric jets (see

Richards and Pitts (1993) and references therein). As a result of air entrainment they slow down as observed experimentally.

The measurements of downstream dynamic pressure and aspirated hot-film response confirm the

above conclusions. Due to the low superheat of FC-31-10 it reached the 1.3 m downstream position

as a two-phase flow for releases through both the 19.1 mm and 12.7 mm orifices. However, due to the shorter flow distance required for the entrainment of a suflftcient mass of air to evaporate the

agent, the 6.4 mm orifice release reached these detectors as a gas-phase flow. The FC-218 seems to be fully vaporized for each of the orifice sizes. Interestingly, halon 1301 showed (Figure 144) two- phase character near the end of the release when the intermediate orifice size was used. This was

also the case for the 19.1 mm orifice. The smallest orifice yields single-phase flow due to the more

efficient entrainment of air as can be seen in Figure 145.

3.5.4.3 Effects of Vessel Pressurization Level on Mixing and Evaporation. The effects of

vessel pressurization are only available for FC-218. The near-field dynamic pressure traces are

qualitatively similar when using nominal 5.51 MPa, 4.14 MPa, and 2.76 MPa burst disks. No

pressure rise is observed until very near the time when liquid is completely expelled from the vessel.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 303

In each case the pressure rise is attributed to the collision of high-speed nitrogen with the liquid stream. The actual magnitudes of the dynamic pressure increases are difficult to interpret physically

due to the complex interactions involved.

The average flow velocities observed for the FC-218 as a function of downstream distance were

very similar for overpressures of 6.21 MPa and 4.28 MPa (Figure 148). This suggests that the

flashing behavior is similar for these two flows and that the velocities resulting from the flashing are dominating the flow behavior. On the other hand, the velocities were lower at each station for a

release at the lowest pressure (2.61 MPa). It would appear that the significantly lower liquid

velocities near the vessel exit in this case resulted in significantly slower mixing.

3.5.4.4 Effects of an Extension Tube on Mixing and Evaporation. The addition of a 0.5 m

extension tube on the vessel had very little eff'ect on the release behavior of the agents, but it significantly modified the mixing behavior. Figure 149 through Figure 151 show that measured

velocities at the two locations furthest from the vessel exit were much higher when the tube was in place than when no tube was used for FC-31-10, FC-218, and halon 1301. These observation can be

easily understood if one makes the reasonable assumption that each of these liquids flashed somewhere within the tube due to a reduction in pressure. The flashing liquids must expand. For a free expansion, the gas is free to expand in all directions. However, for the liquids in a tube the

expansion is constrained to occur only along the tube direction. Much as for a bullet constrained

within the barrel of a rifle, the mass of the fluid is highly accelerated down the tube. For flashing

liquids with high Ja number, the accelerations are quite dramatic indeed. Velocities approaching

200 m/s were observed.

The dynamic pressures measured 1.3 m from the vessel opening were extremely high for each of

the fluids (Figure 152 through Figure 154), and, in the case of halon 1301, approached 3 MPa. Even

assuming very high velocities, this suggests that a significant amount of the liquid was unevaporated.

The measured responses of the aspirated hot-films support this conclusion. To our knowledge, the

extremely high velocities and overpressures generated by using a tube to dispense superheated fire- fighting agents have not been previously documented.

3.5.4.5 Effects of Agent Cooling on Mixing and Evaporation. Cooling of the agents has a very large eff'ect on the dispersion and evaporation of the released liquid. FC-31-10 has a boiling

point close to room temperature and, as a result, does not flash or evaporate rapidly. At the 1.3 downstream measurement position, the flow generated by a room temperature release of this agent has

a large percentage of the agent in the liquid phase. Cooling will slow down the mixing and evaporation and increase the liquid fraction of the two-phase flow. In fact, very nearly 100% of the agent is expected to arrive at the 1.3 m position as liquid.

The effects of cooling FC-218 are considerably more dramatic than for FC-31-10. A room

temperature release of FC-2 1 8 flashes very close to the vessel exit rapidly evaporating and dispersing

the agent. By the time the agent reaches the 1.3 m downstream position the flow has entrained

sufficient air to essentially evaporate all of the released agent. Upon cooling to a temperature of

228 K, which is below its boiling point of 236 K, these advantages of FC-218 for mixing disappear.

Flashing no longer takes place and the agent reaches the downstream location mostly as a liquid flow.

It should be noted that for these tests the agent entered room temperature air. If the ambient

environment was also cooled, the liquid would not be expected to vaporize at all, and the release

would be a liquid stream which would eventually pool somewhere within the space.

3.5.4.6 Effects of Release Orientation on Mixing and Evaporation The efl'ects of releasing

the agents upwards instead of downward are dramatic. The observations are easily understood in
304 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

terms of the changes in release behavior which were observed inside the vessel (see Section 3.4.3.6).

It was noted that for downward releases the nitrogen acts on the liquid inside the vessel much as a pressure applied to a piston. It pushes the liquid rapidly from the vessel where it subsequently flashes

if the liquid at the vessel temperature is superheated. The behavior for upward releases is completely

difi'erent. Since the nitrogen is now at the top of the vessel next to the orifice, it is immediately

released from the vessel upon the opening of the burst disk. The pressure within the vessel drops

very rapidly to close to that of the ambient surroundings. There is no force to drive the liquid agent

from the vessel. If the pressure in the vessel falls below the ambient saturation pressure of the agent,

the agent will begin to boil, ultimately filling the vessel with a two-phase mixture. As the pressure

rises due to expanding bubbles, this two-phase mixture is expelled from the vessel. The measurements of nitrogen velocity shown in Figure 159 and Figure 165 show that downstream velocities are on the order of 175 m/s. These high velocities are due to the very large

velocities (estimated to be on the order of 300 m/s) expected for choked flows of nitrogen through the

19.1 mm orifice. There is very little agent vapor in the nitrogen and the fraction of the total agent mass released initially is quite small. The very small laser attenuations (Figure 158, Figure 163, and

Figure 168) observed downstream of the vessel when the nitrogen flows arrived are consistent with

this conclusion.

Once the nitrogen was expelled from the vessel, delay times were observed before the liquid

agent began to flow. For FC-31-10 this delay time was 24 ms and for FC-218 it was 19 ms. It was

not possible to measure a time accurately for the halon 1301 release, but the behavior of laser #4 in Figure 168 suggests a value on the order of 16 ms. These initial releases of agent are attributed to the expulsion of condensed vapor above the liquid as the superheated liquids in the vessel begin to boil following depressurization. The decreasing delay times observed with increasing superheat are consistent with the behaviors expected for flashing of superheated liquids (see the discussion in Section 3.5.1). It is interesting that the observed delay times are significantly longer than those

estimated for streams of liquid outside of the vessel, which were much less than 5 ms for the two low

boiling-point liquids.

The initiation of boiling does result in the release of some agent from the vessel. However, the

laser beam attenuation, downstream dynamic pressure, and aspirated hot-film measurements all indicated the mass of the agents released during these release periods were fairly low. Only when the

two-phase mixture fully filled the vessel did significant agent begin to leave the bottle. These times were estimated as 54, 39, and 30 ms for FC-31-10, FC-218, and halon 1301, respectively. For

FC-31-10 the time before significant mass of agent was released during an upward release was longer

than required for the complete downward release of this agent. Even for the lower boiling-point

agents the times required for release of a significant mass of the agents were quite long for upward

releases as compared to downward releases. The change in release mechanism for upward releases as compared to downward releases also

has significant eff'ects on the mixing and evaporation behavior of the agents. The results shown in Figure 159 and Figure 165 show that both of the two-phase flows have velocities which are significantly lower than observed for downward releases. As a result, much longer times are required for an agent to reach the 1.3 m downstream distance. Even though the two-phase flows are moving more

slowly and therefore provide longer times for evaporation, the very high dynamic pressures and

aspirated hot-film responses measured indicate that much of the agent mass is still liquid. Estimates

of densities for FC-31-10 and FC-218 based on the measured velocities and dynamic pressures are

significantly higher than the densities of the pure liquid agents. This suggests that velocities later during the flow period are somewhat higher than estimated based on the leading edge measurements.

This does not contradict the conclusion that the flows must consist of large fractions of unevaporated

liquid. For the low boiling-point liquids this result is completely the opposite to that for downward
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 305

releases, where it was observed the FC-218 and halon 1301 were nearly fully vaporized over this distance.

The change in evaporation behavior can be understood in terms of the release mechanism. The

pressure necessary to expel the liquid from the vessel is generated by boiling of the liquid at a pressure close to atmospheric. As a result, the temperature of the liquid drops to levels close to it boiling point. The gas-liquid mixture leaving the vessel is expected to have a similar temperature.

No additional flashing will occur to evaporate liquid. The only way that additional vaporization can

occur is through entrainment of room temperature air into the flow with subsequent heat transfer to the agent mixture. The downward releases demonstrated that insufficient air is entrained to vaporize

a large fraction of the agent.

The times required for complete release of the agents estimated from the laser beam attenuation

measurements, 800 ms for FC-31-10, 500 ms for FC-218, and 250 ms for halon 1301, are consistent

with estimates based on measurements within the vessel, (see Section 3.4.3.6).

These results demonstrate how detrimental the eff'ects of changing the release direction of a pressurized agent from a downward to upward orientation can be on the time required for the start of

the release, the release rate, the efi'ectiveness of agent dispersion, and the degree of evaporation. It should be noted that while no tests were performed using cooled agents, it is extremely likely that essentially no agent would be expelled from the vessel by an upward release of one of these agents

cooled below it boiling point.

3.5.5 Summary and Agent Ranking. The primary purpose of the work described in this section was to allow the proposed halon-alternative agents to be ranked with regard to their dispersion

behaviors and evaporation rates following release from pressurized bottles. It has been argued that

fire extinguishment capability is enhanced by efficient dispersion and mixing and rapid evaporation.

The measurement techniques developed during the study provide sufficient data to allow the proposed

alternative agents to be ranked based on these criteria.

It is significant that the current fire extinguishing agent, halon 1301, was found to be efficiently

dispersed following release from a pressurized vessel at room temperature as the result of the rapid

and strong flashing of the superheated liquid. Rapid vaporization was also observed due to its high

Jakob number and relatively low number of moles per unit volume of the liquid. During the course

of this investigation a workshop was held with a selected group of representatives of manufacturers

(Walter Kidde Aerospace, Pacific Scientific, and Systron Donner) of fire-extinguishment systems for

aircraft in an eff'ort to obtain practitioners' viewpoints on desirable dispersion, mixing, and evaporation behaviors for alternative agents. Their conclusion was that the replacement agent's physical

properties should resemble as nearly as possible those for halon 1301. This conclusion is consistent

with our own, which is based on our understanding of the fire extinguishment process. The experimental findings show that agents which are relatively high boiling~FC-236fa,

FC-31-10, FC-318, HCFC-124, HFC-227ea, and HFC-134a-have dispersion, mixing, and evaporation properties following downward release which are not well suited for rapid fire extinguishment.

This results from their relatively high-boiling points and low Jakob numbers. In the case of the lower

boiling alternative agents~FC-218, HCFC-22, HFC-125, and the HFC-125/HFC-32 mixtureimportant diff'erences were observed. HCFC-22 and the HFC-125/HFC-32 mixture were shown to have strong flashing behaviors due to their low boiling points, but also were found to vaporize

relatively slowly due their low Jakob numbers and relatively high number of moles per unit volume of

liquid. As a result, their mixing and vaporization behaviors are quite diff'erent than observed for halon 1301. On the other hand, FC-218 and HFC-125 were found to mix very efficiently by flashing

and to vaporize fapidly since they have higher Jakob numbers and fewer moles per unit volume of the

liquid than HCFC-22 and the mixture.
306 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 10. Relative ranking of ten alternative agents based on dispersion, mixing, and evaporation

behavior. Note that FC-1 16 has not been included since it is a critical fluid at room

temperature.

Agent Relative Ranking

FC-218 1 HFC- 125 1 HFC-125/HFC-32 Mixture 3

HCFC-22 4

HFC- 134a 5

HFC-227ea 6

HCFC-124 7

FC-318 8

FC-31-10 9

FC-236fa 10

On the basis of the above discussion it is possible to rank the alternative agents based on their

dispersion, mixing, and evaporation behaviors as shown in Table 10. Boiling point alone has been

used to rate agents having boiling points which are higher than FC-218. Note that FC-1 16 is not included in the table. This is because this agent is not a liquid at room temperature.

This investigation has yielded a number of findings which provide an improved understanding of

the physical processes which are responsible for the dispersion, mixing, and evaporation behaviors of

the agents and may be useful for the engineering design of practical fire extinguishment systems.

The importance of the boiling point, Jakob number, and number of moles of agent per unit

volume of liquid for determining the behaviors during downward releases have already been

discussed. To our knowledge, this is the first time that these parameters have been identified as the

controlling physical properties for short-period releases of pressurized agents. Also for the first time,

it has been shown that the release of the nitrogen pressurization gas following expulsion of an liquid

agent from the vessel interacts strongly with the agent flow and provides an additional mechanism for dispersion and mixing of the agent. This mechanism is expected to be important for releases having

very short periods.

The eff'ects of variations in a variety of system parameters on the mixing behavior have been

characterized. Modifications in dispersion and evaporation behaviors with changes in the pressurization level of the vessel and diameter of the vessel orifice for downward releases can be understood in terms of the understanding of these processes obtained during the current investigation. Higher vessel

pressurization of the agents results in more rapid discharges of the liquids from the containment vessel

and, therefore, more rapid mixing. The higher pressures of nitrogen present result in more intense

mixing when the high-speed nitrogen exits the vessel following depletion of the liquid.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 307

Changes in the mixing and evaporation behavior with orifice diameter result from a number of

dependencies. Use of smaller diameters significantly reduces the mass flow rate of the liquids. As a

result, the mixing times are increased. Reduction of the orifice diameter also reduces the time

required for the liquid flow to flash. In some cases, this eff'ect improves mixing even though the total release time has been increased. The known mixing behavior for gaseous axisymmetric jets indicates

that downstream mixing behavior scales with jet diameter. For this reason, flows through smaller

orifices mix over much shorter flow distances than flows from larger orifices. Keep in mind that this apparent advantage is offset by the reduced agent mass-flow rate. The effects of adding an extension tube to the vessel are dramatic. When flashing occurs within

the tube very high exit velocities are observed. The resulting two-phase flows are capable of

generating quite high dynamic pressures. Even for the low boiling liquids, mixing and evaporation

are inhibited in this configuration.

The system parameter which has been found to have the most profound effect on the mixing and

evaporation behavior is the liquid temperature. This is a direct result of the dominant role of the

degree of superheating on the flashing and evaporation. It is interesting that the role of agent

temperamre has been characterized as being underappreciated with regard to testing of halon 1301

dispersion in nacelles (Chamberlain, 1970). The current investigation has considered the effects of

cooling the agent. Particularly for FC-218, the dispersion and evaporation were drastically degraded

by cooling below the boiling point. It is likely that cooling drastically reduces the effectiveness of an

agent for suppression of nacelle or dry bay fires. As an example of the possible effects which might result from cooling. Figure 175 shows the

calculated Jakob number for superheated FC-218 liquid for a range of liquid temperatures at atmospheric pressure. It is clear that Ja decreases rapidly with temperature. Most measurements for

this investigation were for 294 K for which Ja = 0.6. A decrease in temperature of less than 30 K

results in a decrease of the Ja number to 0.3. This value is similar to those the high boiling liquids at room temperature which were found to vaporize very slowly. Clearly, relatively minor cooling will

result is dramatic variations in flashing and evaporation behavior.

Effects of heating the liquid agent were not considered during this work. However, dramatic

changes, which might aid or hinder fire-fighting effectiveness are to be anticipated.

The significance of these findings for designing fire extinguishment systems for aircraft are discussed further in Section 3.7. Another parameter which significantly changes the mixing and evaporation behavior is the vessel

orientation. In general, highly effective dispersion and evaporation is only observed for an orientation

where the nitrogen pressurization gas is located such that the agent is expelled from the vessel as a

liquid driven by the gas pressure. If an orientation allows the pressurization gas to exit the vessel

before the liquid, the release rate is substantially reduced. Due to the slower release rates, mixing is less effective. Furthermore, in these cases the agent is generally expelled as a two-phase flow which

does not flash on exiting the vessel. A significant mechanism for mixing and rapid evaporation is thus eliminated.

3.6 Transient Spray Computations

3.6.1 Introduction. Over the past three decades, an increasing acceptance of and reliance upon

numerical solutions for transient fluid flow problems have taken place. In many cases, experimental

studies are prohibitively expensive, whereas high-speed computers are comparatively economical and

allow a wide range of parameters to be examined in a short time. As a result, computational fluid
308 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

FC-218

I I I

1 1 1 1 r

230 240 250 260 270 280 290 300 310 320

Liquid Temperature (K)

Figure 175. Calculated values of Jakob number for releases of FC-218 as a function of

liquid temperature over a range of 310 to 235 K.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 309

dynamics (CFD) has become a powerful tool for both basic and design studies in a broad range of

physics and engineering problems. CFD refers mainly to the use of computers for numerically

solving the nonlinear equations governing fluid motion. Pioneering work in this field first employed

hand computations to demonstrate that nonlinear equations of fluid motion could be solved numerically. Computers have automated, improved, and dramatically sped up what was first done by hand.

The basic CFD equations express the balances of mass, momentum and energy of a fluid, and

can be solved analytically for only a limited number of simplified situations. The success of CFD is often based on innovative numerical algorithms which discretize either the full set of equations or some simplified set derived from them, forming a large but finite number of algebraic or ordinary

differential equations which can be solved by computers (Oran, 1988). The fundamental idea in CFD

is that spatial dimensions are divided into discrete contiguous cells, usually called finite volumes or

finite elements, and time is discretized in short intervals called timesteps. This discretization is forced

by conventional computers having finite-sized memories segmented into floating-point words of data. The numerically determined values of the fluid variables in each cell are advanced from one timestep

to the next using the nearby fluid variables determined at the previous timestep. While such marching

techniques are developed with time-dependent problems in mind, they are also used to find steadystate solutions when appropriate.

The number of possible CFD algorithms is enormous and the correct choice depends on many

properties of the problem being solved and the computer resources available. In all cases, however,

only a finite number of discrete values comprise the representation, and each value is only specified

to finite precision. This means that information is inevitably lost in the computational solution

relative to the continuous problem being approximated. The result is uncertainty and errors in the

computation arising from the discretization. The source of the uncertainty is the missing information

about the detailed solution structure within the discrete spatial cells and timesteps. All approaches

which use a finite number of values to represent a continuous profile have this problem. Nowadays,

computers with hundreds of millions of high-precision floating-point calculations per second are readily available, and this translates into computations which can simulate some real fluid flows more

accurately than can be validated by modern experimental diagnostic techniques. Three-dimensional

simulations with several million cells are possible giving spatial resolution on the order of 1 % of the

size of the physical system. Two-dimensional solutions, where they are appropriate, can provide even

higher spatial resolution. In this section preliminary eff'orts to extend the use of existing CFD codes for the computation of

the flows generated by the release of halon 1301 and alternative agents are summarized. Also

included is an approach for predicting the initial conditions for the two-phase flow generated by the

release of superheated liquids into an ambient environment.

3.6.2 Numerical Simulation of Discharge Sprays

3.6.2.1 Models Considered and General Capabilities. The fundamental CFD problem for an

extinguishing agent discharge and dispersion is to develop the capability of qualitatively and quantitatively predicting the time evolution of the agent concentration field. It inherently involves a multidimensional analysis of a one- or two-phase, multi-component, compressible medium where

thermodynamic efl'ects and transport phenomena are important. This, in turn, imposes even more

stringent requirements on the numerical algorithms used and computational times.

The dispersion of sprays of halon and halon-alternative fire extinguishing agents throughout a protected space is a very complex process involving many physical phenomena occurring simultaneously and afl'ecting each other. The agent stored in the liquid state is discharged from nitrogenpressurized vessels through a simple, nozzle/orifice-like exit. The systems under consideration
310 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

involve vessel volumes of the order of 10'^ m^, with required discharge times of the order of 10"^ s. To achieve such rapid discharge times, the initial pressurization levels must typically be of the order

of several megapascals. As the liquid agent exits from the vessel, thermodynamic and fluid-dynamic

instabilities lead to flashing and break the agent into a two-phase droplet/gaseous jet mixture. This

occurs in a short transition region which starts at the vessel exit and ends where thermodynamic and

fluid-dynamic equilibrium has been achieved. If the temperature of the discharged liquid agent is far above its normal boiling point, the liquid very quickly evaporates and creates a cloud of agent vapors

surrounding the moving spray with remaining liquid having a temperature equal to its boiling point. In the transition region the spray consists of a flowing mixture of liquid-agent fragments of various

shapes and sizes and gaseous agent. This flow then entrains initially-quiescent air from the surrounding environment.

Downstream of the transition region the flow begins to develop as a mixed, two-phase, agent/air

jet where thermodynamic equilibrium is maintained and where droplet collision and agglomeration do

not play an important role in the ensuing jet dynamics and in the dispersal of the agent throughout the

protected space. The entrainment, mixing and evaporation continue as the spray is dispersed throughout the space. For advanced multi-dimensional mathematical models describing complex transient flow problems, numerical methods solving sets of nonlinear partial difl'erential equations as well as relevant

computer codes have been developed at the Los Alamos National Laboratory. Two of these codes--

CONCHAS-SPRAY (Cloutman et ai, 1982) and KIVA-II (Amsden et al., 1985a,b,c; 1987; 1989)--

have been considered during the present study. The reader is referred to the cited references for complete details concerning these models.

CONCHAS-SPRAY was the first computer code developed at Los Alamos National Laboratory

to solve the equations of transient, multi-component, dissipative, chemically reactive fluid dynamics,

together with those for the dynamics of an evaporating liquid spray. It was a descendant of an earlier

experimental code called CONCHAS (Butler et al., 1979), which did not include the spray modeling

capability. The code is quite general, but has been designed with applications to internal combustion

engines specifically in mind. The formulation is spatially two-dimensional, which requires that the

dependent variables depend on only two of the three spatial coordinates because of symmetry. The

option is provided to select either rectangular or cylindrical coordinates, corresponding to linear or

axial symmetry, respectively. All the transport coefficients such as dynamic viscosity, heat conductivity and mass diflusivity as well as heat capacities are assumed to be temperature dependent. The

efi'ects of turbulence are represented by an optional subgrid scale (SGS) turbulence model (Deardorff",

1971). This model was found to be very useful for parabolic flows which occur in, as an example,

jet flows {e.g., Gmurczyk et al., 1992). The code allows a Lagrangian, Eulerian, or mixed

description, and is particularly useful for representing curved or moving boundary surfaces.

Evaporating liquid sprays are represented by a discrete-particle technique (Dukowicz, 1980), in which each computational particle represents a number of similar physical particles. The radius and

other attributes of a particle are statistically assigned using a Monte Carlo sampling technique. The

evaporation rate is obtained from a quasi-steady model (Dukowicz, 1979). The particles and fluid interact by exchanging mass, momentum and energy. These interactions are treated by implicit

coupling procedures to avoid the prohibitively small time steps that would otherwise often be

necessary. The spray is assumed to be thin, as defined by O'Rourke (1981).

KIVA-II (a descendent of KIVA) is one of the latest in a sequence of multidimensional numerical

models (Amsden et al., 1985a,b,c; 1987; 1989) designed to solve the dynamics of evaporating fuel sprays interacting with flowing multicomponent gases undergoing mixing and heat transfer in a manner similar to CONCHAS-SPRAY, but extended into full three-dimensional situations, including

droplet break-up/coalescence phenomena. Turbulence modeling is more sophisticated than in
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 311

CONCHAS-SPRAY. A standard model (Amsden et ai, 1989; Launder and Spaulding, 1972)

extended to variable density flows is used to simulate turbulence effects in the calculations. Transport

equations for the turbulence kinetic energy and turbulence dissipation rate are solved with appropriate source terms for the spray interaction.

The spray model is fundamentally based and, as such, has broad applicability The dynamics of

the atomized fuel spray are handled by a Monte Carlo based discrete-particle technique (Dukowicz,

1980). The spray is considered to be composed of discrete computational particles, each of which

represents a group of droplets of similar physical properties. The distribution functions for droplet

size, velocity, and spray pattern produced by the fuel injector are statistically sampled and the

resulting Lagrangian parcels are followed as they locally interact and exchange mass, momentum, and

energy with the surrounding gas. The model accounts for turbulence effects and interactions between

droplets (O'Rourke, 1981) but other dense spray effects are neglected. When a computational parcel

is selected to be injected into the flow field, its size is determined by sampling from a chi-squared

distribution function (Bracco, 1985). The masses of the injected parcels are constant, thereby

permitting each parcel to represent a different number of actual droplets depending on its size. The

influence of droplet collisions on the spatial variation of the droplet mean size is taken into account by

a sampling procedure. Collision frequency is then calculated and used to determine the probability

that a droplet in a given parcel undergoes a collision with a droplet in a nearby parcel. All droplets

in the given parcel behave in the same manner; either they do or do not collide and no new computational parcels are created. However, if a collision occurs and the resulting parcel has more than twice

the original injection mass of the parcel, it is broken into two identical parcels with half the number

of droplets as the original. Trajectories of these two parcels subsequently diverge because of their

different interactions with the local mrbulence field.

3.6.2.2 Calculational Approach and Code Modification. As a first step in the development of

codes capable of modeling releases of halon 1301 and alternative agents, it was decided to focus on

the use of CONCHAS-SPRAY. The two-dimensional axisymmetric variant of the generalized model

was adapted as a subset of the KIVA-II code to accommodate the needs of the specific problem under

consideration. The set of time-dependent equations is expressed in a two-dimensional form since the

experimental discharges were designed to be axisymmetric. It should be noted, however, that the

experiments discussed earlier in Section 3 were in a large rectangular room, while those which were modeled are for much smaller closed volumes. The choice to model a smaller closed volume was

dictated by practical computational time limitations. CONCHAS-SPRAY consists of a set of 26 primary subroutines controlled by a short main

program. In addition to the primary subroutines there are thirteen supporting subroutines that perform tasks for the primaries. For various applications, all required geometrical specifications,

initial and boundary conditions may be specified using input data alone. In other cases where the

input data is inadequate to define the problem of interest, the required specifications must be inserted

directly into the appropriate subroutines. The modular structure of the program enables such

modifications.

The CONCHAS-SPRAY code was used almost as received from Los Alamos. However, the

evaporation routine was judged to be inadequate. As a result, the following expression for the agent

equilibrium vapor pressure, PJTi), at liquid temperature 7) suggested by Daubert and Danner (1992),

PJJ) = exp(A+-+C]iiTj+DTi\ (41)

was incorporated into the expression for the mass fraction of agent vapor, 1^, at the droplet surface,
312 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

(42)

where M^,y and are the molecular weights for air and agent, respectively, and P is the total pressure in the gas. A, B, C, D, and E are empirical constants chosen to fit the temperature variation

of a given agent. When the vapor pressure data of HCFC-22 were modeled, CONCHAS-SPRAY

was unable to compute beyond the first timestep because the liquid vapor pressure exceeded 101 kPa

for a temperature corresponding to the boiling point of HCFC-22. A different set of coefficients was

chosen in subsequent calculations to maintain a vapor pressure below the ambient. Rather than a

flashing spray, the results are more applicable to an evaporating spray. Special attention to this issue

will be required to properly model the flashing process which was experienced by many of the agents

in the experimental discharges.

Several other minor modifications to the code were necessary to ensure numerical stability. For

instance, strong temperature and pressure gradients were generated by the rapid evaporation of the

liquid that prevented convergence of the calculations. This type of numerical instability was removed

by the introduction of a viscous pressure (Thompson, 1972), which was added to individual pressures

within a computational cell and also to the pressure-volume work term appearing in the energy

equation. Following these modifications, it was found that stable calculations could be carried out. The situation investigated was the discharge of a hypothetical liquid agent through an orifice located on the centerline of a 0.42 m diameter cylindrically shaped volume. Agent inflow boundary

conditions were derived from the results of the discharge vessel experiments (see Section 3.4), in which the agent mass flow rate was measured. The initial pressure was 101 kPa, and the liquid was assumed to be a mono-sized spray at a temperature of -41 °C. Computational parcels representing

the droplets were introduced at the injector with an axial velocity of 61.4 m/s directed at an angle of

0° relative to the symmetry axis. The corresponding mass flow rate was 21.2 kg/s. A zero azimuthal component of velocity was assumed for the droplets at the injector. All of the gas and

liquid parameters were taken as uniform across the orifice exit. The air in the compartment was

specified to be quiescent and uniform in temperature at time equals zero. The surrounding walls were

treated as solid, free-slip boundaries maintained at a constant temperature of 21 °C. The subsequent

penetration and shape of the spray resulted solely from the interactions of the liquid droplets with the

ambient air. A zero-order model of the transitional region between the nozzle exit and a fully-developed spray

is described in Section 3.6.4. Because the details of the actual fluid flow were not measured and are

essentially unknown, it has been assumed that a spray of uniform droplet size is already established at the discharge orifice exit and is injected into the area at a prescribed rate. It should be noted that,

due to these assumptions, the model does not reproduce many critical aspects of the discharge

process, and cannot be directly compared to the experiments. The value of the calculations has been

in the identification of the limitations of CONCHAS-SPRAY and the clarification of the experimental

parameters which must be measured or controlled before a realistic representation of the discharge

process can be expected. Six cases were chosen for the numerical simulations: a chamber length of

1.0 and 2.5 m, an initial air temperature of 21 and -30 °C, and an initial droplet diameter of 200 or 600 ixm.

3.6.2.3 Sample Calculations and Results. All the computations were performed on a Convex

supercomputer with a UNIX operating system. The wall clock computational time for one case (around 15 ms of the real process) required three hours on the average. The data were collected in

PM . -M.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 313

the form of large ASCII files (typically 12 Mb/file/case) which were converted subsequently into small graphics subfiles. The results were then plotted as contours of constant velocity, density, pressure, temperature, gas species mass fraction, vorticity, kinematic viscosity, mass flux, and liquid

phase mass fraction. In order to demonstrate the types of eff"ects which were observed, results for the

vapor phase mass fraction of agent into the one meter long volume are described.

Figure 176 shows the time behavior of the agent vapor mass fraction distribution for four times

following release. Assuming that the mass fraction of agent vapor follows the two-phase jet flow

(likely to be an excellent assumption), it can seen that the downstream edge of the flow field moves

away from the orifice with increasing time and that the jet spreads with increasing downstream

distance. By 14.3 ms the flow has moved roughly 0.9 m from the source. An interesting feature is the "head" which grows on the downstream edge of the flow field. Recall that a similar structure was

observed in high speed films of the experimental agent releases. Calculated velocity fields indicate

that the head forms as the result of a vortical motion developed by the flow field. As a result, this is a region of efficienl heat transfer from the air to the flow which evaporates the liquid droplets. The

resulting agent evaporation causes a volume expansion which further enlarges the head.

Figure 177 shows the results of an identical calculation except that the diameter of the injected

liquid droplets is increased from 200 /xm to 600 ^im. The downstream penetration of the two-phase

flow is little changed from that found when smaller droplets are injected. However, the spreading

rate of the jet is decreased and the head at the downstream edge of the flow is much narrower. Larger droplets are expected to result in lower evaporation of the liquid due to their smaller surface

area. The reduction in spreading rates for Figure 177 is likely due to a combination of a smaller

volume expansion and the lesser entrainment of air associated with fewer large rather than many small

droplets.

The eff'ect of the ambient temperature on the mixing behavior can be seen by comparing

Figure 176 and Figure 178. The overall shapes of the mass fraction contours are very similar even though the air temperature has been decreased over 50 °C. At first glance this appears to be counter

to what one would expect; i.e, a lower ambient temperature should diminish the vaporized agent mass

fraction. The explanation is that the vapor pressure curve used in the calculation ensures that the

agent pressure is less than the ambient pressure even for the higher temperature computation. For an

agent like HCFC-22, an increase in temperature has a much more dramatic impact on the vapor

pressure, which most likely would lead to a significantly higher vapor fraction when the ambient

temperature is increased as much as 50 °C. This could not be verified, however, because of

numerical instabilities which occurred under more realistic flashing conditions.

As noted above, the calculations generate two-dimensional predictions for a wide range of

variables including velocity, pressure, temperamre, density, species mass fraction and spray contours. A few additional characteristics of the results are summarized here to give an indication of what

physical processes may be important in these systems. At certain regions in the flows very high

velocities were calculated due to the rapid evaporation of the liquid droplets. These velocities were

sufficient at times to generate shock waves which subsequently interacted with the two-phase flow and

modified its mixing and evaporation behavior. It should be noted that shock waves were not detected

in the experiments, but this may have been the result of diff'erences in geometry between the

experiment and calculation. Since the vessel was closed, major modifications were found when the

vessel length was increased from 1.0 m to 2.5 m. This is to be expected due to the higher pressures

and stronger shock waves which occur in the smaller volume.
314 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 176. Agent vapor mass fraction contours at 3.5, 8.3, 10.7, 14.3 ms following spray release at 0 ms. Maximum mass fraction = 0. 15, contour inter- val =0.0188. spray mean Sauter diameter = 200 ^m, ambient T = 294 K,

P=101 kPa, compartment length=l m.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 315

Figure 177. Agent vapor mass fraction contours at 3.5, 8.3, 10.7, 14.3 ms following

spray release at 0 ms. Maximum mass fraction =0.14, contour inter- val =0.0175, spray mean Sauter diameter = 600 jim, ambient T = 294 K,

P=101 kPa, compartment length=l m.
316 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 178. Agent vapor mass fraction contours at 3.5, 8.3, 10.7, 14.3 ms following

spray release at 0 ms. Maximum mass fraction = 0.209, contour inter- val = 0.0261, spray mean Sauter diameter = 200 fxm, ambient T = 243 K,

P=101 kPa, compartment length =1 m.
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 317

3.6.3 Establishing Initial/Boundary Conditions from the Vessel Discharge Simulations

3.6.3.1 Overview. A critical element in exercising CONCHAS-SPRAY, KIVA, or any other

computational fluid mechanics computer code, for the purpose of providing agent dispersion

simulations, is the establishment of a set of initial/boundary conditions. Especially important and

problematic is a determination of the characteristics of the agent jet at a location on the jet axis, near

the exit section of the discharge vessel, as it enters the protected space. These initial/boundary

conditions should be determined from results of a simulation of pressure vessel agent discharge. The

purpose of this section is to present a method of estimating the required time-dependent

initial/boundary conditions by expanding on the vessel discharge model described in Section 3.2. Although not used in the dispersion computations reported above, this new methodology is expected

to be used and evaluated in future CONCHAS-SPRAY and KIVA simulations.

3.6.3.2 Modeling the Early Development of the Jet: From the Exit Section of the Discharge

Vessel to a Nearby Section in a State of Thermodynamic Equilibrium and Completed Droplet

Formation. It is assumed that within the discharge vessel and upstream of the exit nozzle/orifice the

velocity of the liquid agent is so small that its kinetic energy can be neglected in defining its thermodynamic state. Therefore, the thermodynamic state upstream of the nozzle/orifice position can

be estimated by the pressure in the discharge vessel, P^y, and the temperature there of the liquid

agent, Tj^yAL- '^^^^ ^^^^ most applications it is expected that T^yjj^ is well approximated by

^AMB' the ambient temperature outside the vessel. A point representing the state of the liquid agent in the vessel is indicated in the pressure-enthalpy {P,h) diagram of Figure 179.

Because of the relatively short nozzle/orifice design under consideration, the time interval for the

liquid to approach it, pass through it, and enter the outside environment is very small. For conditions

of this study this time is of the order of 10"-^

s. During this time interval the pressure of the

traversing liquid is reduced from the high P^y values inside the vessel, of the order of several MPa,

to pressures of the order of P^if^g, 101 kPa.

As discharging liquid enters and traverses the region of the nozzle/orifice, it will enter and move

along superheated metastable thermodynamic states. While dependent on its changing thermodynamic

state in the vessel, {P£,y, T^vaD^ '^^ particular thermodynamic properties, as the agent

material penetrates the ambient environment it approaches, and likely achieves the P^mb pressure

while still in its metastable liquid state. Thus, downstream of a vena contracta the liquid agent

develops into a near-uniform-radius liquid jet, r^j, which can be predicted from traditional incompressible fluid-dynamic considerations. See Figure 180. For the relatively large-vapor-pressure agent

materials of interest here, a combination of fluid-dynamic and thermodynamic instabilities will then

lead to breakup into small droplets and flashing (i.e., rapid evaporation to a two-phase equilibrium

thermodynamic state) of the metastable liquid jet (Lienhard and Day, 1970).

Prior to the onset of breakup and/or flashing the metastable liquid is expected to move through

thermodynamic states on near-isentropic paths. However, as discussed in Sections 3.2 and 3.4.3.4,

for a given agent and initial-state condition, (Pj^y, Tj^y^j), it is possible that the pressure along such

a path will not reach P^^^fB prior to an intersection with the agent's spinoidal curve. Moreover, if such an intersection does occur, then at that instant spontaneous nucleation in the liquid is to be

expected (Reid, 1978; Modell and Reid, 1983). It is conjectured that this would lead to near- explosive breakup and flashing of the jet, all of this being initiated within a jet penetration depth (into

the outside environment) of the order of a single nozzle/orifice diameter.

Whatever the thermodynamic-state path of the discharging metastable liquid, it is reasonable to assume that a relatively short distance downstream of the nozzle/orifice a thermodynamic equilibrium
318 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Figure 179. Path in the enthalpy - pressure plane of the thermodynamic state of the

initially liquid agent as it flows from inside to outside the pressure vessel

and achieves a stable two-phase state at the initial section.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 319

\
320 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

State and an end to the droplet-formation phenomenon will be achieved. It is also reasonable to assume that, as the agent flows further downstream, developing as a mixed, two-phase, agent/air jet,

thermodynamic equilibrium will be maintained, and droplet collision and agglomeration will not play an important role in jet dynamics. These latter conditions are required if CONCHAS-SPRAY, KIVA,

or similar CFD simulations are to be applicable.

The position along the jet axis where both approximate thermodynamic equilibrium and

completion of droplet formation is first achieved will be denoted here as the initial position of the jet.

Associated with this position is the plane section normal to the jet axis, called the initial section of the

jet. Similarly, the thermodynamic state and properties of the jet at the initial position will be denoted

as its initial state and initial properties. The region of the flow between the nozzle/orifice discharge

section and the initial section will be referred to as the transition region of the jet. The initial section

is a critical portion of the surface that bounds the protected space throughout which the agent is being

dispersed.

A detailed description of the initial state and properties of the jet would be extremely complicated. This would include the variation across the initial section of air/gaseous-agent concentrations,

temperature, and velocity, and of the size, velocity, and temperature distributions of the liquid agent

droplets. Note also that in the present application all of these variables of the initial jet flow field would be time-dependent. In general, an accurate detailed description of the initial state of the jet,

whether determined by theoretical or experimental means, is either impractical or beyond the current

state of technology.

Although it is not possible to provide a detailed description of the initial state of the jet, it is reasonable to expect that an achievable approximate description, suitable for use in establishing

boundary conditions for in the formulation the dispersing-agent-flow-field problem, would lead to good approximations of agent dispersal phenomena. It is the purpose of this work to provide such an

approximate description of the initial state. The dynamic processes that initiate breakup and flashing of the metastable liquid jet and bring it to its initial state are not completely understood. At one extreme, it is possible that the abovementioned instabilities lead to violent breakup and flashing of the liquid jet immediately upon leaving

the exit nozzle/orifice. Such behavior was observed, for example, in some experiments of Brown and

York (1962). This was also observed in the current experiments described in Section 3.5 for lowboiling agents.

The metastable state can also persist even as the liquid jet penetrates relatively deeply into the

ambient environment. For example. Brown and York (1962) described a liquid water jet of diameter

D = 7.6 X lO""^ m, initially at 9.03 x 10^ Pa and 415 K, which penetrated up to 0.0254 m, corre- sponding to L/D = 32, before initiation of violent breakup/flashing. Similar behaviors were

observed for high-boiling agents in the current study. In this regard it is of interest to note that in the

literature of the discharge of flashing liquids from high pressure vessels there is a degree of conventional wisdom that analyses of the two-phase flow phenomena, based on thermodynamic equilibrium

states, tend to yield good results when the nozzle/piping from the inside to the outside of a high

pressure discharge vessel is at least 0.1 m (Lueng and Nazario, 1990).

The initial section will be some incremental distance downstream of the penetration distance of

the metastable liquid jet. The studies of Brown and York (1962) and the recent experiments at NIST

suggest that this incremental distance will typically not exceed the order of the penetration distance

itself. Whatever the distance, it is assumed here that the total length of the transition region is small

enough, compared to the characteristic length of the overall agent dispersal problem of interest, that the processes within the transition region can be treated as quasi-steady. In the transition region it is assumed that momentum transfer and any net work or heat transfer interactions between the penetrating agent jet and the ambient air environment are negligible. Thus,
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 321

consistent with the analysis of Epstein et al. (1983) for steady, two-phase, air/evaporating-liquid jets,

it is assumed that the jet at the initial section can be approximated as consisting only of a two-phase

mixture of agent, with no entrained air. Finally, in the transition section it is assumed that increases

in the energy of the liquid due to mechanical work by surface tension forces to form the initial-section

droplets are negligible compared to the energy transfers involved in phase-change processes. The above assumptions are used with the following additional considerations to determine the

initial state. The initial section of the jet is defined by a characteristic initial radius Vj. The region r <

consists of a pure mixture of saturated-agent vapor and liquid (droplets) in thermodynamic equilibrium at ambient pressure, Pj = P^/^^^ ^^'^ saturation temperature, Tj - T^^^f^^MS^- The liquid-agent

component of the jet consists of droplets of uniform-radius r^^,, and these are uniformly dispersed

over the region r < Vj. The agent droplets and gas have the same uniform axial jet velocity, Vj ^. There is no agent in r > r^, and in the initial section, this outer region is assumed to consist of a uniform ambient-temperature/pressure environment (P^mb^ '^amb^ quiescent air. Conservation of mass and energy is invoked by assuming a quasi-steady state between the initial section and a section in the discharge vessel somewhat upstream of the exit nozzle/orifice. Conservation of mass yields

dMDVAL

dt

' (43)

where

(1-X,)"

(44)

'ML

and pj ^, pj and Pj are the average density of the agent, the density of the liquid agent

(droplets), and the density of the agent gas at the initial section, respectively; Xj is the quality of the

two-phase flow at the initial section, i.e., the fraction of a mass of agent which is saturated gas;

^DVAL instantaneous mass of liquid agent in the discharge vessel; and the right side of

Equation (43), determined previously from, e.g., the discharge model of Section 3.2, is assumed to be specified.

When invoking conservation of energy, the kinetic energy at the section in the discharge vessel is assumed negligible compared to that at the initial section. The resulting expressions are

hovAL - Kal^^^alg^TJ^^ (45)

and

Klg^Tsat) - ^AG-hiAL' (46)

where h^y^^ and /z/^^ are the specific enthalpy of the agent liquid in the discharge vessel and at the

initial section, respectively; hj^Q is the specific enthalpy of the agent gas at the initial section; and

^ALG ^^^^ of vaporization of the agent at the temperature T^^-p. In Equation (46),
322 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Klg = Klg^T)^ T<T^r (47)

is a specified function of temperature, T, where Tcan not exceed the critical temperature, Tf^j^. A point representing the state of the agent at the initial section is presented in the (P, h) diagram

of Figure 179. Consistent with Equation (45), this reflects the fact that some of the original enthalpy

of the liquid agent in its high pressure "near-rest" state, hj^y, was exchanged for its now-potentiallysignificant kinetic energy. The sketch of the transition region, consistent with the all the above assumptions/

approximations, is presented in Figure 180.

The initial droplet radius is assumed to be specified. According to Brown and York (1962) it is expected that this will be in the range 10 ^m - 100 pim. The sensitivity of the downstream agent

concentrations to variations of this parameter should be evaluated from future dispersion computations.The initial jet radius is assumed to be specified. Based on photographic data in Brown and York

(1962), r/r^ seems to be of the order of ten to twenty. Thus, in (what is reasonable to construe

visually in high speed photographs as) a region of breakup and flashing of the liquid jet to an initial equilibrium state, the radius of the jet increases by a factor between ten and 20. Again, the

sensitivity of the downstream agent concentrations to variations of this ratio should be evaluated from

future dispersion computations. One would hope that such sensitivity is not great. In describing the thermodynamic properties of the agent, the liquid state is modeled everywhere

as being incompressible, and the gaseous state at the initial section is modeled as a perfect gas. For

problems of interest here, agent liquid temperatures are expected to be small enough compared to Tj-^

and vary in a small enough range to permit accurate modeling of the liquid as being incompressible,

with density p^^, and as having a constant specific heat, C^^. Thus,

(48)

The values of p^^ and C^^ will be taken as

P

- PAiiT,,); C,, = C,,iT,,y, T,, = (Ts^r^T^y^L)/^, (49)

where 7^^ is taken to be the average temperature of the liquid in the discharge vessel and at the initial

state, and where p^i(T) and C^i(T} are specified functions of T. From the assumption of incompressibility it follows that

h -h = C (T -T (50)

'^DVAL "lAL ^AL<^DVAL ^ SAT' ' ^

'

The agent gas is modeled as a perfect gas.

Pal

P P _ ^AMB - (51) lAG 1? T R T

^AG^I ^AG^SAT

where R^q is the gas constant. As discussed above, it is expected that in most instances the liquid jet will achieve P^^B prior to

its breakup/flashing. Then, Bernoulli's equation leads to
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 323

dMDVAL 2 dt

'AL

V2 (52)

where rjj is determined from the radius of the exit nozzle/orifice, r^, and its orifice discharge

coefficient, C^,

' u

V )

(53)

Substitutirg Equations (46) and (50) into Equation (45) leads to

^ C (T^ _T \ . ^ DV AMB^ _ I, (J' \ ^ ^AL^^DVAL SAT ^rALG<^ SAT'' (54)

Substituting Equations (44) and (52) into Equation (43) leads to

2

/ r \ U (^DK ^AMB^

\2

9al I e j

(55)

where

'IAG

'AL

Comparing Equations (55) and (54) leads to

(56)

X2+A2^-(A.j + l) = 0, (57)

where

X = -(1-e) — Xe

(1-e)

(58)

and
324 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Pal^alg^sat

(1-6)

P^ -P DV AMB^

I r \2 (60)

The solution to Equation (57) is

2)

(61)

With Equation (58) and other previous definitions. Equation (61) provides the solution for Xj. This

would be used, in turn, to obtain V}^ from Equation (54) or (55), thereby completing the description

of the initial state.

IP -p ^DK AMh

'AL

(62)

A sketch of the velocity distribution and the mass flux distributions of gaseous and liquid agent at the initial section are presented in Figure 181.

3.6.3.3 Algorithm for Determining the Initial State. The above solution can be implemented

with the following algorithm:

1. Specify agent and its thermodynamic properties, including T^^j(P) and h^^Q(T), p^i(T),

and C^i(T).

2. Specify r/rj^ (expected range 10 - 20).

3. Specify P^^q, T^j^q, and T^^yAL- (I" most cases expect to specify T^y^^ = T^mb-)

4. Specify = T^^B r > rj. Compute T^at(^amb) ^^^^ specify 7} = for 0 <

r < rj.

5. Compute h^Q(T^^) from Equation (47).

6. Compute T^^ according to Equation (49) and then p^^ and according to Equation

(49) and p^(TJ and C^l(T), respectively.

7. Compute Pj^q according to Equation (51) and then £ according to Equation (56).

8. Use a component-1 model, e.g., that of Section 3.2, to find the time-dependent values of

Pj~fy for a particular Figure 1 -configuration of interest.
SECTION 3, FLUID DYNAMICS OF AGENT DISCHARGE 325

AT z=Zj :

V

I,

V

-(dMj^Q/dt)/(7tr^)

MASS FLUX OF GASEOUS AGENT

MASS FLUX OF LIQUID AGENT

Figure 181. Sketch of the velocity and mass fluxes of agent gas and agent liquid

droplets at the initial section, z = Zj.
326 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

9. For a particular value of P£,y compute and X2 from Equations (59) and (60), and then

X from Equation (61).

10. Find Xj from Equation (58) and Vj ^ from Equation (62).

3.6.3.4 Example Applications of the Initial State Algorithm. The initial-state algorithm was

applied in four example simulations. These involve the potential replacement agent HCFC-22 and

halon 1301 and two different rj values, 0.074 m and 0.148 m, corresponding to rj - lOr^ and 20r^

and to a value rjj of 0.0074 m which corresponds, in turn, to a discharge vessel nozzle radius of

0.0095 m and a discharge coefficient of 0.6. The examples involve time-dependent discharges of

room-temperature agent from a Figure 1-type configuration with no holding tank. The geometric

parameters, orifice coefficient, and initial conditions assumed in the simulated example discharge

calculations are presented in Table 1 1 . These are the same parameters used in the example calculations of Section 3.2. During the discharge, the time-dependent value of Pj~,y was computed using the

component- 1-type model of Section 3.2. The material property functions for HCFC-22 and halon 1301 which were used in the example

calculations are taken from the prescriptions given in Table 12 (Daubert and Banner, 1992).

'^SAT obtained as the root of P(T^^j-) = P^ji^g. Then T^l was calculated using Equation

(49). Finally, ^^^^(T^^j-), P^iiT^i), and C^^iT^ were calculated from the property functions of

Table 12. Results are presented in Table 13. The results of the example calculations are tabulated in Table 14 (HCFC-22, = lOr^ =

lOC^^^V^ = 0.0738 m). Table 15 (halon 1301, Kj = lOr^ = lOC^^'^r^ = 0.0738 m).

Table 16 (HCFC-22, rj = 20r^ = lOC^^'^r^ = 0.148 m), and Table 17 (halon 1301, = 20r^ =

20C^^^^r^ = 0.148 m). Included in the tables are computed values of the variable

M 7

r = ^'^'^^= ^

(63)

^DVM^t-Q) Z(r=0)

which is the dimensionless mass of liquid agent, Mj^y^^^, or dimensionless elevation of the liquid

agent in the discharge vessel, Z. The tabulated results of Table 14 for the discharge of HCFC-22 {rj = lOr^ = 0.0738 m) are plotted in Figure 182 through Figure 185. These include plots of the time-dependent values of Pj^y

(Figure 182), Vj^ (Figure 183), f and Xj (Figure 184), and dMj^^/dt and dM^^^/dr (Figure 185).

A study of the results leads to the following observations:

1. For both HCFC-22 and halon 1301, the times to complete the discharge of liquid agent

from the pressure vessel are almost identical at 0.024 + 0.001 s (Table 14 through

Table 17).

2. At the initial section, the quality of the two-phase agent flow, Xj, is significantly different

for the two agents considered, but for a given agent Xj is relatively independent of t and

rj, for the two initial-section radii values considered (r^ = 0.0738 m and 0.148 m).

Thus, for the cases considered, Xj = 0.31 ± 0.01 for HCFC-22 (Table 14 and Table 16)

and = 0.54 + 0.02 for halon 1301 (Table 15 and Table 17).
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 327

Table 11. Geometric parameters, orifice coefficient, and initial conditions for example discharges

from a Figure 1-type configuration with no holding tank

Length of discharge vessel (Z^,^) 0.255 m

Cross-sectional area of discharge vessel (A^y) 0.196 X 10-2

Volume of discharge vessel (V^y) 0.500 X 10"3

/\rcd Ui CaIL IIUZZIC/UI lULC k^q) nw.zoj VA 1lun"3 tn2 in

Kaaius or exii nozzie/ormce vq) yj.yjj X lU m

Discharge coefficient of exit nozzle/orifice (C^) 0.6

Initial volume of agent liquid in discharge vessel 0.250 X 10"^ m^

Initial temperature of gas in the discharge vessel

(Tdvj)

21 °C

Initial pressure in discharge vessel (P^y j) 41.4 X 10^ Pa

Temperature of agent liquid in discharge vessel 21 °C

3. At the initial section, for a given agent, and for the two initial-section radii values

considered, the mass flow rates of liquid or gaseous agent are almost identical at any

particular time during the discharge. (For example: for HCFC-22, approximately half- way through the discharge at t = 0.0125 s, values of dMj ^Jdt for Kj = 0.0738 m and rj = 0.148 m are 8.59 kg/s (Table 14) and 8.49 kg/s (Table 16), respectively; and for halon

1301, approximately half-way through the discharge at t = 0.0122 s, values for dMj ^Jdt

for r; = 0.0738 m and = 0.148 m are 6.95 kg/s (Table 15) and 8.59 kg/s (Table' 17),

respectively.) However, over the course of the discharge, these mass flow rates are reduced to approximately two thirds of their initial value. (For example: for HCFC-22

and rj = 0.0738 m, the values of dMj ^Jdt at t = 0 and t = 0.233 s (near the end of the

liquid discharge) are 11.78 kg/s and 7.34 kg/s, respectively (Table 14 and Table 16); and

for halon 1301, and rj = 0.0738 m, the values of dM;^^/dr at r = 0 and t = 0.244 s (near the end of the liquid discharge) are 9.26 kg/s and 5.88 kg/s, respectively

(Table 15)).
328

Table 12.

SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Material Property Functions for HCFC-22 and halon 1301

HCFC-22 halon 1301

Molecular Weight (kg/mole) 86.47 148.9

P{Tsat}I^ = exp[^ + BliTs^jJR) + aniTs^jJR) + D{Ts^jJR)\ ^1 - ^SAT - ^2

A 9.78 X 10^ 6.28 X lOi

B -4.77 X 10"^ -3.38 X 10^

C -1.23 X 10^ -6.75

D 2.42 X 10"^ 1.44 X 10-^

Ti 116 K 105 K

369 K 340 K

^^LG^wV(J/kg-mole) = ^(1 - ^W^C/?)^' ^1 - ^SAT - ^2

A 2.96 X 10^ 2.51 X lO"^ B 3.86 X 10-1 3.57 X 10-1

^CR 369 K 340 K

Ti 116 K 105 K

369 K 340 K

p^^fr)/(kg-mole/m3) = Al^\ F = {1 + [1 - {T/R)/Q^}; Tj < T<T2

A 1.60 1.34

B 2.66 X 10-1 2.70 X 10-1 C 3.69 X 10^ 3.40 X 10^

D 2.81 X 10-1 2.80 X 10-1

Ti 116 K 105 K

369 K 340 K

C4^a)/[J/(kg-mole-K)] = A + B{TIR) + C{T/R)^ + D{TIR)^\ TI < T < T2

A 9.47 X 10"^ 1.14 X 10^

B 3.93 X lOi -1.32 X 10^

C -8.17 X 10-1 6.12 X 10-1 D 2.79 X 10-^ 0

Ti 116 K 193 K

T2 369 K 298 K
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 13. Constant properties of the initial state for the example calculations

329

HCFC-22 halon 1301

^SAT -41 °C -58 °C

Tal -10 °c -18.5 °C

0.234 X 10^ J/kg 0.118 X 10^ J/kg

Pal(^ai) 0.132 X 10"^ kg/m^ 0.180 X 10"^ kg/m^

Cal(Tai) 0.115 X 10^J/(kg-K) 0.809 X 10^ J/(kg-K)

3.6.3.5 Nomenclature

LJn LlidillCLCi Ui a iiC[LiiLl JCL h specific enthalpy

h of agent liquid in the discharge vessel

L length of liquid jet

P pressure

^AMB pressure of the ambient environment

pressure in the discharge vessel

Pi pressure at the initial section

P-AG gas constant for the agent gas

r radius

n radius of two-phase agent jet at the initial section

radius of agent droplets at initial section

ru radius of the liquid jet outside the discharge vessel

T absolute temperature

^AMB temperature of the ambient environment

^DVAL temperature of liquid agent in the discharge vessel

Ti temperature at the initial section

^SAT saturation temperature of the agent

t time

velocity at the initial section

^lA Vj of agent

quality of the two-phase agent flow at the initial section Z distance along the jet axis from exit nozzle/orifice z at the initial section

3.6.4 Concluding Remarks on Modeling the Discharge Spray. Realistic numerical modeling of

the spray discharge process requires an accurate knowledge of the initial conditions surrounding the

release of the agent. The conditions prior to bursting the rupture disc in the agent storage vessel are

fully established, but a mathematical description of the disintegration of the disc and the motion of the
330 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 14. HCFC-22 Jet Behavior for r^^ 10r^=0.148 m; p^^= 1320 kg/m^ P/ag=3.58

kg/m^ 7^^=232 K

tis) (Pa) f

(m/s) (kg/s)

dMj (^^J^X

(kg/s)

0.00000 0 4135 X 10^ 1.0000 32.21 0 3172 5 36 11.53

0.00179 0 3663 X 10^ 0.9030 30.13 0 3158 5 01 10.86

0.00358 0 3299 X 10' 0.8117 28.45 0 3148 4 73 10.30

0.00537 0 3008 X

. ^7 10^ 0.7248 27.05 0 3139 4 50 9.83

0.00716 0 2769 X 10^ 0.6417 25.86 0 3132 4 30 9.43

0.00895 0 2570 X 0.5620 24.83 0 3126 4 13 9.08

0.01073 0 2401 X 10^ 0.4851 23.92 0 3121 3 98 8.77

0.01252 0 2255 X 10^ 0.4108 23.12 0 3117 3 85 8.49

0.01431 0 2128 X 10^ 0.3389 22.40 0 3113 3 73 8.24

0.01610 0 2016 X 10^ 0.2690 21.75 0 3109 3 62 8.02

0.01789 0 1917 X 10^ 0.2010 21.16 0 3107 3 52 7.81

0.01968 0 1829 X 10^ 0.1347 20.62 0 3104 3 43 7.62

0.02147 0 1749 X 10"^ 0.0701 20.12 0 3102 3 35 7.44

0.02326 0 1677 X 10^ 0.0069 19.67 0 3099 3 27 7.28

liquid agent through the exit plane and for several orifice diameters downstream is far beyond the

state of the art for even the simple geometry chosen for the discharge experiments. What is known

are the size of the orifice and the emptying rate of the storage vessel, both from the experimental

measurements and the model of the emptying process described in Section 3.2. The first-order

analysis in the previous section provides a starting point for the droplet enthalpy, but because

prediction and/or measurement of droplet size have not yet been developed for these conditions, the

droplet size distribution can only be hypothesized.

The general approach for numerically simulating a developing, two-phase, multi-component,

compressible, evaporating spray has been presented, and two difi'erent Los Alamos computer codes

have been described. KIVA-II is the most versatile because it can deal with fully three dimensional

flow and boundary conditions, it has a more sophisticated description of the turbulent interactions

between the droplets and the gas phase, and it accounts for droplet breakup. There is a severe computational penalty paid for this increased versatility, however. Because the discharge experiments were close to axisymmetric, and because the initial droplet diameter was unknown anyway, the long

computational times required for KIVA-II convinced us to chose the more efficient CONCHASSPRAY code for our initial parametric study. It was felt that at this stage of our knowledge of the
SECTIONS. FLUID DYNAMICS OF AGENT DISCHARGE 331

Table 15. HCFC-22 Jet Behavior for r^= 10r^=0.074 m; p^= 1320 kgW; P/^g=3.58

kg/m^; r^ ^ = 232 K

r(s) (Pa) r

(m/s) (kg/s)

dMj y^j/dt

(kg/s)

0 0 4135 X 10' 1.000 122.73 0 3022 5 10 11.78

0.00179 0 3663 X 10^ 0.9031 115.46 0 3025 4 80 11.07

0.00358 0 3299 X 10^ 0.8117 109.49 0 3028 4 55 10.48

0.00537 0 3008 X 10^ 0.7248 104.46 0 3030 4 34 9.99

0.00716 0 2769 X 10^ 0.6417 100.14 0 3032 4 16 9.57

0.00895 0 2570 X

- ^7 10' 0.5620 96.37 0 3033 4 01 9.20

0.01073 0 2401 X 10^ 0.4851 93.05 0 3034 3 87 8.88

0.01252 0 2255 X 10^ 0.4108 90.08 0 3035 3 75 8.59

0.01431 0 2128 X 10^ 0.3389 87.41 0 3036 3 63 8.33

0.01610 0 2016 X 10^ 0.2690 84.99 0 3037 3 53 8.10

0.01789 0 1917 X 10"^ 0.2010 82.78 0 3038 3 44 7.89

0.01968 0 1829 X 10^ 0.1347 80.76 0 3039 3 36 7.69

0.02147 0 1749 X 10^ 0.0701 78.88 0 3039 3 28 7.51

0.02326 0 1677 X 10^ 0.0069 77.15 0 3040 3 21 7.34

phenomena, the penalty for using a less sophisticated axisymmetric code would be offset by the

increased number of parameters which could be varied.

Sample computations were carried out to evaluate the potential of CONCHAS-SPRAY to provide

qualitative understanding of agent spray dispersion. The input data for the calculations were based on

the agent discharge experiments described earlier. The geometrical dimensions of the computational

domain were smaller than the dimensions of the actual compartment in which the discharges were

performed to ensure enough spatial resolution to resolve the controlling parameters.

The calculations have demonstrated that the Los Alamos computer code CONCHAS-SPRAY can be used to analyze the dynamics of halon-alternative agent dispersion and mixing in a space designed

for fire protection. Within its physical limitations, it can perform the necessary computations of the

discharge process quickly and relatively cheaply, aiding one's understanding and indicating directions

for future experiments.

The numerical analysis based on the full fluid dynamics equations as well as a description of the

heat and mass transfer leads to the time dependance of the two-dimensional field of all parameters

characterizing the agent dispersion phenomena, including velocity, pressure, temperature, density,

species mass fraction, mass flux and spray contour. Much of these data are difficult or impossible to measure in experiments.
332 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Table 16. Halon 1301 Jet Behavior for r^- 10r^= 0.074 m; p^= 1800 kgW; p^^^=6.17 kg/m^

7}.^=215K

t (s) U V (Pa) f l,nU dMr AT /dt

(m/s) (kg/s) (kg/s)

0 0.4135 X 10^ 1.0000 142 83 0 5181 9.95 9.26

0.00204 0.3687 X 10^ 0.9030 135 28 0 5205 9.43 8.69

0.00407 0.3338 X lO'^ 0.8111 128 99 0 5224 8.99 8.22

0.00611 0.3057 X 10^ 0.7236 123 64 0 5239 8.62 7.83

0.00814 0.2826 X 10^ 0.6397 118 99 0 5252 8.29 7.50

0.01018 0.2631 X 10^ 0.5590 114 91 0 5263 8.01 7.21

0.01222 0.2465 X 10^ 0.4812 111 28 0 5273 7.76 6.95

0.01425 0.2322 X 10^ 0.4058 108 03 0 5281 7.53 6.73

0.01629 0.2197 X 10^ 0.3327 105 08 0 5289 7.32 6.52

0.01832 0.2086 X 10^ 0.2616 102 40 0 5295 7.14 6.34

0.02036 0.1988 X 10^ 0.1924 99.94 0 5301 7.97 6.17

0.02240 0.1899 X 10^ 0.1248 97.67 0 5307 6.81 6.02

0.02443 0.1820 X 10^ 0.0588 95.58 0 5311 6.66 5.88

A comparison between KIVA-II and CONCHAS-SPRAY for the same set of conditions is needed

to determine the impact of the turbulence and droplet breakup model on the results. If the outcome is independent of the code, additional analyses with CONCHAS-SPRAY could be conducted. If the

turbulence and droplet models influence the results, KIVA-II (or better yet, KIVA-III, (Amsden et

al., 1993)) should be used for future calculations. KIVA has the additional advantage that fully three

dimensional spaces can be handled. Extended resources on the NIST Cray computer would be

required for these investigations. The spectrum of parameters influencing the spray dynamics needs

to be broadened to include additional compartment geometries, discharge orifice geometry, thermodynamic and gas dynamic parameters of the agent and ambient air, agent injection and atomization

parameters as well as consideration of the transition region of the spray.

3.7 Summary and Recommendations

The proposed alternative agents are ranked based on their release, mixing and dispersion, and

evaporation behavior. Measurements of the release behaviors for the alternative agents have

demonstrated that when the liquid agents are initially at room temperature, all of the agents are released as superheated liquids. Theory predicts that only very small difl'erences in release rates are
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 333

Table 17. Halon 1301 Jet Behavior for r^= 10r^=0.148 m; p^^=1800 kgW; P/^g=6.17

kgW; 7)^=215 K

t r

0.00000 0.4135 X 10^ 1.0000 38 48 0 5583 10.73 8.49

0.00204 0.3687 X 10^ 0.9030 36 17 0 5566 10.08 8.03

0.00407 0.3338 X 10^ 0.8111 34 28 0 5552 9.56 7.66

0.00611 0.3057 X 10^ 0.7236 32 69 0 5541 9.11 7.33

0.00814 0.2826 X 10^ 0.6397 31 33 0 5532 8.73 7.05

0.01018 0.2631 X 10^ 0.5590 30 15 0 5524 8.41 6.81

0.01222 0.2465 X 10^ 0.4812 29 11 0 5518 8.12 6.59

0.01425 0.2322 X 10^ 0.4058 28 19 0 5512 7.86 6.40

0.01629 0.2197 X 10^ 0.3327 27 36 0 5507 7.63 6.22

0.01832 0.2086 X 10^ 0.2616 26 60 0 5503 7.42 6.06

0.02036 0.1988 X 10^ 0.1924 25 92 0 5499 7.23 5.91

0.02240 0.1899 X 10^ 0.1248 25 29 0 5496 7.05 5.78

0.02443 0.1820 X 10^ 0.0588 24 71 0 5492 6.89 5.65

to be expected between agents when this is the case. In fact, experiments indicate that all of the

agents have comparable volume release rates. The only indication of a different behavior was

observed for halon 1301 for which the liquid apparently flashed just upstream of or in the orifice. As

a result of the formation of a two-phase flow, the flow rate of halon 1301 was substantially reduced.

Since the room-temperature release rates are essentially identical for all of the proposed

alternative agents, this parameter does not provide a means to differentiate between the agents. As a

result the ranking is based solely on their behaviors once released. The results summarized in Section 3.5.5 thus provide the final ranking of the agents based on release behavior, mixing and

dispersion, and evaporation rate. The final ratings on these criteria are summarized in Table 10. Note once again that FC-116 has not been included since it is not a liquid at room temperature.

Table 10 presents the result which represent the primary goal of the investigation to rank the

alternative agents with regard to release rate, dispersion and mixing, and evaporation rate. However,

during the course of the research program a number of observations and secondary conclusions have

been obtained which have important implications for the design and engineering of systems for suppression of nacelle and dry bay fires using the proposed alternative agents. During the discussion

it should be kept in mind that the experimental system developed for this investigation was designed

solely to allow a ranking of the proposed alternative agents. Actual commercial systems designed for suppression of fires on aircraft are likely to be significantly different. On the other hand, many of the

findings and conclusions of this work should be applicable to more complicated vessels and flow
334 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

>

Q_

50(10^)

40(10^)

30(10^)

20(10^)

10(10^)

0

0 0.005 0.01 0.015 0.02 0.025

t[s]

Figure 182. Plot of Table 14 results for P^v as a function of time during vessel discharge of HCFC-22.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 335

Figure 183. Plot of Table 14 results for ^ as a function of time during vessel discharge of HCFC-22.
Figure 184. Plots of Table 14 results for f and Xj as functions of time during vessel discharge of HCFC-22.
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 337

Figure 185. Plots of Table 14 results for dMj ^b/dt and dMi^g/dt as functions of time

during vessel discharge of HCFC-22.
338 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

geometries. Where appropriate, suggestions are provided for additional studies which should be

considered.

Downward releases of agents will be considered first. Results discussed in Sections 4.2 and 4.4

demonstrate that when agents are released from pressurized vessels as liquids, it is possible to predict

release rates and pressure time histories within the vessel quite accurately. The effects of changing a

vessel orifice size or degree of pressurization are easily predictable. The only case where premature

flashing within the vessel or orifice was found to result in a reduction in release rate was for halon

1301. This agent has the lowest boiling and therefore the highest superheat of the liquids tested. The

potential replacement agents may be exposed to ambient temperatures higher than standard room

temperature. These higher temperatures result in higher overheats which may ultimately lead to flashing which disrupts the liquid flow. As a result, the release of an agent may require a longer

period than predicted based on the assumption of a liquid flow. If this occurs, the agent's fire- fighting eff'ectiveness could be reduced. It would be worthwhile to determine more carefully the

conditions for which flashing of a superheated liquid can reduce the release rate. Recall also that it has been hypothesized that a superheated liquid which is rapidly depressurized might reach its spinodial condition and undergo a boiling liquid expanding vapor explosion. Such an event has not been characterized experimentally, but due to the dire consequences which could result, it would be

prudent to investigate the possibility further. The primary eff'ect on the release rate of cooling a vessel containing an agent below room

temperature is to reduce the pressure of the driving gas within the vessel. It should be possible to predict these eff"ects using the models developed during the course of this project.

The eff'ects of various system parameters on the dispersion and mixing and evaporation behavior

following release of the liquid agents from a pressurized vessel are complex. During the course of

this eff'ort several major influencing parameters have been identified. The parameter which has the

largest efl'ect on the initial mixing behavior following a room-temperature release of an agent is the

degree of superheating. The higher the superheat, the more likely the agent is to flash near the vessel

exit. Flashing also appears to occur more rapidly with higher superheats. As a result of flashing the

agent is accelerated in multiple directions and is mixed with it surroundings. The velocity measure- ments also demonstrate that the agent spreading is accelerated by flashing leading to enhanced and more rapid dispersion and better mixing. If the ambient temperature is raised, a liquid will become

more superheated and it dispersion and mixing behavior should be enhanced. Cooling of the vessel

should have an opposite effect. Measurements with FC-218 demonstrated that the latter was indeed

the case. For the experimental system investigated here only the vessel was cooled. If an agent release

occurred into an environment which was also cooled it would be expected that the dispersion and

mixing would be hindered even more. It would be useful to investigate release of agents into cooled

environments in order to fully characterize these elTects. It would also be useful to run flame

extinguishment tests such as the Wright-Patterson dry bay test facility under cold conditions to assess

the effects on flame-extinguishment effectiveness.

The effects of changes in the degree of vessel pressurization and orifice size on the dispersion

and mixing and vaporization behavior are understandable in terms of the flashing and vaporization

behavior of the agents once discharged.

Orientation of the release vessel is a parameter which strongly influences the liquid-agent release

mechanism and subsequent mixing and agent evaporation. If the bottle is oriented such that the

pressurization gas is located next to the release orifice it is released first and release of the liquid

agent occurs by boiling and subsequent release of a two-phase flow. Experiments show that not only

are the release times for an agent substantially increased, but the dispersion and evaporation behavior

are significantly degraded as compared to a downward release. If cooled sufficiently, release of agent
SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE 339

may not occur at all. Designers of extinguisher systems should assess the probability of a pressure

vessel being oriented incorrectly as a result of aircraft flight profiles. If such events are probable it may be necessary to design pressure vessels in such a way that nitrogen will always expel liquid from

the vessel.

Initial efforts have been described which are aimed at modeling the dispersion and mixing and

evaporation of agents following their release as superheated liquids from pressurized bottles. Considerable work remains to be done, but this initial work suggests that simulations sufficiently accurate for engineering design work can be developed. It should be possible to calculate behaviors

for releases of agents into complex three-dimensional compartments such as dry bays and engine

nacelles. Similar models are widely employed in a wide range of engineering disciplines to reduce

design times and limit expensive testing. Efl'orts to develop efi'ective models for agent releases should

be pursued.

The experimental diagnostics which have been used to investigate the release behavior of agents

following discharge of the potential fire-fighting agents have provided sufficient characterization to allow a ranking. However, there are a number of characteristics of these releases which have not been addressed. Many of these are expected to be crucial for efi'ective fire fighting. Such properties

as the flow fields, the concentration fields, temperature fields, droplet sizes, entrainment rates, and

degree of vaporization in the space are required before true models of flame extinction can be developed. Such experimental measurements should be performed to provide a data base for validating

models of mixing behavior and flame extinction. Such information would also be very useful for fumre eff'orts to develop new and improved flame suppression processes.

3.8 References

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Allied Signal, Inc., Genetron Products Technical Bulletin: R-32/125 Azeotrope, Allied Signal

Inc., Genetron Products, Morristown, NJ, July, 1991

Amsden, A. A., Ramshaw, J.D., O'Rourke, P.J. , and Dukowicz, J.K., KTVA: A Computer

Program for Two- and Three-Dimensional Fluid Flows with Chemical Reactions and Fuel Sprays, Los

Alamos National Laboratory Report No. LA-10245-MS, 1985a.

Amsden, A. A., Butler, T.D., O'Rourke, P.J., and Ramshaw, J.D., KTVA: A Comprehensive

Model for 2D and 3D Engine Simulations, SAE Technical Paper No. 850554 (1985b).

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Transactions (1987).

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Mesh for Complex Geometries, Los Alamos National Laboratory Report No. LA-12503-MS, 1993.

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14, 695 (1985).

Balzhiser, R.E., Samuels, M.R., and Eliassen, J.D., Chemical Engineering Thermodynamics,

The Study of Energy, Entropy, and Equilibrium, Prentice-Hall, Englewood Cliff's, NJ, 1972.
340 SECTION 3. FLUID DYNAMICS OF AGENT DISCHARGE

Bankoff, S.G., " Diffusion-Controlled Bubble Growth," in Advances in Chemical Engineering,

Vol. 6, Drew, T.B., Hoopes, Jr., J. W., Vermeulen, T., and Cokelel, G. R., Eds., pp. 1-60,

Academic Press, New York, 1966.

Beegle, B.L., Modell, M., and Reid, R.C., "Thermodynamic Stability Criterion for Pure

Substances and Mixtures," AIChE Journal 20, 1200 (1974).

Benjamin, M.W., and Miller, J.G., "The Flow of Saturated Water through Throttling Orifices,"

Transactions ASME 63, 419 (1941).

Bird, R.B., Stewart, W.E., and Lightfoot, E.N., Transport Phenomena, John Wiley & Sons,

New York, 1960.

Bracco, F.V., Modeling of Engine Sprays, SAE Technical Paper No. 850394 (1985).

Bradshaw, P., An Introduction to Turbulence and Its Measurement, Pergamon, New York, 1971.

Brown, J. A., and Mears, W.H., "Physical Properties of n-Perfluorobutane, " J. Phys. Chem. 62,

960 (1958).

Brown, G.L., and Rebollo, M.R., "A Small, Fast-Response Probe to Measure Composition of a Binary Gas Mixture," AIAA J. 10, 649 (1972).

Brown, R., and York, J.L., "Sprays Formed by Flashing Liquid Jets," A.I.Ch.E. Journal 8, 149

(1962).

Bryner, N.P, personal communication, 1993.

Burnell, J.G., "Flow of Boiling Water through Nozzles, Orifices, and Pipes," Engineering 164,

572 (1947).

Butler, T.D., Cloutman L.D., Dukowicz J.K., and Ramshaw J.D., CONCHAS: An Arbitrary

Lagrangian-Eulerian Computer Code for Multicomponent Chemically Reactive Fluid Flow at All

Speeds, Los Alamos Scientific Laboratory Report No. LA-8129-MS, 1979.

Carnahan, B., Luther, H.A., and Wilkes, J.O., Applied Numerical Methods, John Wiley & Sons,

New York, 1969.

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345

4. FLAME SUPPRESSION EFFECTIVENESS

Anthony Hamins, Grzegorz Gmurczyk, William Grosshandler,

R. Gregory Rehwoldt, Isaura Vazquez, and Thomas Cleary

Building and Fire Research Laboratory

Gary Presser Chemical Sciences and Technology Laboratory

and Kalyanasundaram Seshadri

AMES Department, University of California at San Diego, La Jolla, CA

4.1 Overview

A flame will be extinguished when the time required for the chain reaction which sustains

combustion exceeds the time it takes to replenish the necessary heat and reactants. A characteristic

time for reaction, t^^^^^, can be estimated from the inverse of a global kinetic rate coefficient, k,

expressed in Arrhenius form as

tchem ~k-'=B -' exp(-^, {seconds) (D

where B is a molecular collision frequency factor, Eq is a global activation energy, R is the ideal gas

constant, and T is the gas temperature. Assuming reactant species and heat are transported at about

the same rate {i.e., unity Lewis number), a characteristic time, r^^^^, for replenishing both can be

estimated from a convective flow velocity, U, and a length scale, L, by

(5ecoAufe). (2)

The Damkohler number is the ratio of the characteristic flow time to the characteristic chemical

reaction time, i.e., = ^y^ou/^c/iew Linan (1974) showed that as the maximum flame temperature

and the fuel burning rate decrease, a critical value Damkohler number is reached, D^*, where the

flame abruptly extinguishes. The Damkohler number criteria suggests a number of strategies for extinguishing fires that include increasing the strain rate or flame stretch (to decrease t^i^^^,), or

cooling, reactant removal, and chemical inhibition (to increase t^f,g^). A large number of studies have been conducted on fire suppression. Williams (1980) summarizes the fundamental processes of flame extinction. The suppression efl'ectiveness of a number of inert,

halogenated, and metal agents been studied phenomenologically in counterflow flames (Milne et al.,

1970; Kent and Williams, 1974; Seshadri and Williams, 1975; Mitani, 1983) and in co-flowing cup

burner flames (Simmons and Wolfhard, 1956; Creitz, 1961, 1972; Bajpai, 1974a; Hirst and Booth,

1977; Tucker et al., 1981; and Sheinson et al., 1989). A number of investigators have studied the

structure of inhibited flames in order to learn about the mechanisms of flame inhibition. A global measure of the chemical effectiveness of an agent has been quantified through heat capacity considerations (Tucker et al, 1981; Sheinson et al, 1989). Pitts et al. (1990) outline current understanding
346 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

of fire suppression, incorporating a comprehensive review of the relevant literature, in addition to a discussion of agent effectiveness, test methods, and the role of different suppression mechanisms.

The objective of this part of the study was to rank experimentally the effectiveness of various

chemicals for suppressing different laboratory-scale flames, and to project the relative effectiveness to the performance one would expect in the Wright Laboratory simulated full-scale fire suppression

facilities.

4.1.1 Comparison of Experimental Configurations. Fundamental studies have shown that the

character of the fluid mechanical flow field is a key parameter in controlling flame extinction. A

discussion of the importance of the flow field strain rate or flame stretch is given by Williams (1980).

As the strain rate increases, r^^^^ in Equation 2 decreases and a flame becomes weaker and easier to extinguish by agent addition. In order to understand the chemical effectiveness of an agent, other

parameters such as fluid mechanics which influence flame extinction must be well-controlled. For the

applications considered here, an engine nacelle and a dry bay, the fluid mechanics and thus the strain

rate is not well defined due principally to variability in the flowfield geometry. Observation in the

Wright Laboratory simulated fire suppression facilities, however, indicate that the fires of interest are turbulent and thereby characterized by a range of strain rates. The configurations listed in Table 1 were designed to improve our understanding of the suppression effectiveness of agents under a variety

of possible combustion scenarios.

Four experimental configurations were used to rank the relative effectiveness of suppression

agents. They are listed in Table 1 and include the laminar opposed flow difliision flame burner

(OFDF), the cup burner, the turbulent spray flame burner, and the deflagration tube. Table 1 also

lists differences in the parameters which characterize combustion in each of the test configurations

including the nature of the flow field and the pressure regime. In a typical fire the fuel and air are nonpremixed, as represented by the cup burner and the OFDF burner.

Opposed flow and coflow systems provide two fundamental approaches for studying difliision flame extinction. Past experience has shown the OFDF configuration to be a convenient geometry for detailed fundamental studies of the structure and mechanisms of extinction of difliision flames

(Hamins, 1985; Hamins and Seshadri, 1984a, 1984b, 1986, 1987; Hamins etai, 1985; Seshadri and

Williams, 1975, 1978a). The critical conditions of flame extinction measured in the opposed flow

configuration, can be readily interpreted using currently available asymptotic theories of flame

extinction (Fendell, 1965; Peters, 1982, 1984; Williams, 1985), and the results can also be extended

to turbulent nonpremixed combustion (Peters, 1982, 1984). Interpretation of experimental results of

flame extinction using asymptotic theories requires knowledge of the characteristic flow time in the

flame, t^^^. In the opposed flow configuration, the value of t^g^ is related to the velocity of the

reactant streams or the flow field strain rate (Seshadri and Williams, 1978b), which is an independently controlled experimental parameter.

This is not the case for a coflow system like the cup burner. In the coflowing configuration, the

shape of the flame is approximately conical and the value of t^^^ cannot be related easily to the flow

velocities of the reactant streams. It has been recently shown that in the coflowing configuration the

value of tjig^ at the flame varies with height above the exit of the burner (Seshadri et ai, 1990).

Therefore, interpretation of extinction data obtained in the coflowing configuration is not straightforward. Also, in the coflowing configuration premixing of the fuel and oxidizer can occur near the

vicinity of the surface of the liquid fuel and the exit of the oxidizer duct where the fuel vapors and the

oxidizer initially come into contact. This premixing can influence the mechanisms of flame extinction. However, in the opposed flow configuration, premixing, other than that resulting directly from

diffusion of reactants through the flame, is eliminated completely. A significant advantage of the

opposed flow configuration is that there are numerous computer programs available for predicting the
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 1. Experimental Configurations

347

Configuration Type of Combustion and Flow Field

Pressure Agents Fuels

Opposed Flow

Diffusion Flame

laminar non- premixedambient 13 gaseous agents & NaHC03 heptane, JP-8

Cup Burner quasi-laminar non- premixedambient 14 gaseous agents, NaHCO-,,

& 18 additional agents

JP-5, JP-8, 5606, 83282,

heptane, propane

Spray Burner turbulent spray ambient 14 gaseous agents &

NaHCOj

83282, JP-8

Deflagration Tube

turbulent premixed variable 14 gaseous agents ethene

Structure and critical conditions of flame extinction (Bui and Seshadri, 1992; Chelliah, 1992; Smooke

et ai, 1986). Although, such programs have also been developed for predicting the structure of

flames stabilized in the coflowing configuration, these programs require considerably more computational time. Careful interpretation of the data, however, show that there exists a correspondence

between the results obtained in the opposed flow and coflow configurations.

Consideration of in-flight fire scenarios in engine nacelles and dry bays prompted the development of two unique test configurations, the deflagration tube and the spray flame burner. Although it

is not possible to estimate the value of tjj^^^, for the spray burner or the deflagration tube, extinction

experiments in these configurations are of interest because they mimic combustion behavior in engine

nacelle and dry bay fires, where r^^^^, is also not well characterized.

In contrast to all of the other configurations listed in Table 1, combustion is premixed in the

deflagration tube. In the spray burner, although the fuel and oxidizer are initially nonpremixed,

combustion occurs in the form of a turbulent two-phase spray which is neither purely premixed nor nonpremixed, but can be thought of as an ensemble of partially premixed flames. Whereas the flames

are laminar in the OFDF and are nearly laminar in the cup burner, they are highly turbulent in the

other two configurations. Atmospheric combustion occurs in all of the configurations except the

deflagration tube, where final pressures can be much larger than atmospheric.

The diff'erences in the structure of the flames in each of the test configurations is likely to have a

large impact on the relative suppression eff'ectiveness of the agents. The introduction of an inert or chemical inhibitor to a flame impacts the stability of the flame by influencing the flame structure through changes in the temperature and species concentration fields. At the same time, the structures of premixed and nonpremixed flames are distinct. The mechanism for suppression may be distinct in detonations and diffusion flames because the chemistry may take place at different pressures, and

chemical rate constants are pressure sensitive which could enhance or diminish the importance of

particular chemical reactions. In non-premixed flames, combustion can be limited by the rate at which the reactants are brought together, whereas in premixed flames, combustion is limited by the

chemical kinetics. Spray flames represent a complicated combustion situation where fuel droplets

vaporize and react with the oxidizer, which can be composed of either pure air, air mixed with

combustion products or air mixed with fuel vapor. Because of the inherent differences in the
348 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Structure of the combustion zone in each of the configurations, it would be very surprising if the

agents displayed identical agent suppression effectiveness in the different test configurations.

4.1.2 Fuels and Agents Tested. In the four configurations, a number of different fuels were tested. Table 1 lists these fuels. They consisted of gaseous fuels (propane and ethene), multi-component

liquid fuels (JP-5 and JP-8 jet fuels, and 5606 and 83282 hydraulic fluids) and a single component

liquid fuel (heptane). The jet fuels and hydraulic fluids contain many components and can vary in composition from batch to batch. Thus, the heats of combustion (Gascoyne Laboratories, 1993) and

elemental composition (Galbraith Laboratories, 1993) for the fuel batch used in the experiments

reported here were determined as listed in Table 2.

4.1.3 Agents. Thirteen gaseous agents were tested in all configurations. The agents included

nitrogen; four perfluorinated compounds, CjF^, C^F^, C4F10, and cyclo-C4Fg; four hydrogen/fluorine

compounds, CF3HFCF3, CF3H2CF3, C2HF5, and CFH2CF3; the mixture 60% CH2F2/40% C2HF5;

the chlorinated agents CHF2CI and CHFCICF3; and CF3Br. In addition, CF3I and sodium bicarbonate powder (NaHC03) were tested in all configurations except the OFDF burner and the deflagration

tube, respectively.

Eighteen additional agents were tested in the cup burner including inert gases, other halocarbons,

a silicon compound, and a nitrogen compound. In addition, several liquid agents were tested. A

detailed explanation for the selection of additional agents for cup burner testing is given in Section 5. In addition to the gaseous and liquid agents described above, sodium bicarbonate, a solid

powder, was also tested in the first three configurations listed in Table 1. The objective of these tests was to quantitatively compare the suppression effectiveness of sodium bicarbonate to CF3Br and the

other gaseous agents.

Sodium bicarbonate powder is currently a popular fire suppressant and its eflTectiveness is well

known. Because of its zero ozone depletion potential, it deserves consideration for a wider range of

applications. Few flame extinction measurements have been conducted with powders in test devices

where the effectiveness of the powder could be directly compared with gaseous agents, and where

flow field effects were minimized.

Some previous studies indicated that particle size is an important parameter in flame extinction

using dry powders. Milne et al., (1970) studied the extinction of methane/air opposed flow diffusion

flames using NaHC03 powders with different diameters and found that smaller powders were more

effective flame suppression agents. Measurements were reported by Dodding et al., (1977) for extinction of methane/air opposed flow diffusion flames with 10 to 20 jum sodium bicarbonate

particles added to the oxidizer. However, a key experimental parameter, the strain rate, was not

reported. Others have tested the effectiveness of powders in less fundamental geometries. Ewing et

al., (1989) reported that approximately 4% (by mass) of NaHC03 ( ~ 10 ^m particles) was needed to extinguish heptane pan fires. In order to study the impact of particle size, the powder was milled and then silica (Si02) was

added to the powder to improve flow behavior and to prevent agglomeration. The particles were then

sized using a centrifuge by VORTEC Industries into several fractions, hereafter referred to as nominally 0-10, 10-20, 20-30, 30-40, 40-50, and 50+ jum. The powder was stored in sealed plastic

bags until use. Particle size distribution was measured using optical microscopy and an average (equivalent

sphere) diameter is listed in Table 3. Table 3 also lists the density of the NaHC03/Si02 mixtures

which was measured using helium pycnometry (Lum, 1993) and the (mass based) percentage of Si02

in each size fraction as measured by X-ray fluorescence (Dow-Corning, 1993).
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 2. Fuel Properties

349

Fuel Type Heat of Combustion

(MJ/kg)

Elemental Analysis

Mass Percent H Ma*;*; Pprrent P

JP-5 46.3 13.3 85.4 < 0.5

JP-8 46.5 13.9 86.2 < 0.5

5606 44.0 12.6 80.5 2.03

83282 45.1 13.2 79.5 4.78

ethene 50.3 14.3 85.7 0

propane 50.3 18.2 81.8 0

heptane 48.1 16.0 84,0 0

Table 3. Particle Characterization

Nominal Particle Size

(/xm)

Equivalent Particle Diameter

(/xm)

SiOj %

(kg/kg)

Measured

Density (kg/m?)

0 - 10 4.5 + 0.5 1.34 2220

10 - 20 7.0 ± 0.5 0.67 2220

20 - 30

a 0.46 2220

Not measured

The remainder of this Section is broken into four parts which include detailed descriptions of the

experimental apparatus listed in Table 1. For each configuration, experimental methods, results, and

discussion are included. Key findings of the work are compiled in Section 4.6 and reference are

listed in Section 4.7.

4.2 Opposed Flow Diffusion Flames

4.2.1 Introduction. An experimental study was conducted to characterize the relative influence of

various inhibiting agents in extinguishing difiiision flames burning the liquid hydrocarbon fuels heptane and JP-8. Thirteen gaseous inhibiting agents and NaHC03 were tested, and their eff'ectiveness

in extinguishing diff\ision flames burning liquid fuels was compared to CF3Br. The experimental

configuration employed in the study was a diflFusion flame stabilized in the mixing layer produced by

directing an oxidizing gas stream downward on the burning surface of a liquid fuel. This is referred

to in the literature as the OFDF configuration and is a convenient geometry for detailed fundamental

studies of the structure and mechanisms of extinction of difiiision flames. In the OFDF configuration.
350 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

shown in Figure 1 , there exists a stagnation plane, and the stoichiometry of the overall combustion

processes is such that the flame (defined as the position where the temperature attains a maximum

value) is stabilized on the oxidizer side of the stagnation plane. The oxidizing gas used in the

experiments was a mixture of air and the inhibiting agent. To test the influence of the initial temperature of the oxidizing gas stream on the critical conditions of flame extinction, the experiments were performed with the oxidizing stream at temperatures of 25 °C and 150 °C.

Milne et ai, (1970) employed the opposed flow technique to evaluate the influence of gaseous

inhibiting agents, including CF3Br and powdered agents, on the critical conditions of extinction of

methane-air, propane-air and n-butane-air diffusion flames. These experiments employed the "Tsuji"

type opposed flow configuration, where the diff\ision flame is stabilized in the vicinity of the forward

stagnation line of a porous cylinder from which the gaseous fuel is introduced into the reaction zone. From the experimental measurements it was concluded that on a mass basis some powder agents are approximately ten times more efl'ective than CF3Br in extinguishing dilTusion flames. Kent and

Williams (1974) and Seshadri and Williams (1975) studied the inhibiting efl'ect of CF3Br on opposed

flow difliision flames stabilized over the surface of pools of heptane. The velocity and composition

of the oxidizer stream at extinction were measured (Kent and Williams, 1974 and Seshadri and

Williams, 1975). The concentration of CF3Br in the oxidizing stream at flame extinction was compared to the amount of the inert diluent N2 which must be added to the air stream to extinguish

the flame. At selected conditions the flame temperature near extinction was also measured. As

mentioned in Section 4.1, interpretation of experimental results of flame extinction using asymptotic

theories requires knowledge of the characteristic residence time in the flame, t^Q^. At a fixed value

of the flow velocity of the oxidizer stream, or at a fixed value of tfl^^, the concentration of CF3Br in the air stream at extinction was found to be considerably lower than the concentration of N2 added to the air stream at extinction. In addition, the value of the measured temperature in the vicinity of

extinction was higher or the value of the characteristic chemical-reaction time tchem lower when

CF3Br was added to the oxidizer stream than for nitrogen dilution. Since nitrogen is an inert compound, these results imply that inhibition of chemical reactions by CF3Br may be responsible for extinguishing the flame at a lower value of tchem- These previously developed techniques (Kent and

Williams, 1974; Seshadri and Williams, 1975) are used to interpret the experimental results obtained

in the present study.

4.2.2 Experimental Apparatus. Figure 2 shows a schematic illustration of the OFDF burner. The

burner consisted of a fuel cup which had a diameter of 50 mm and a depth of 18 mm, and an oxidizer

duct through which the oxidizing stream containing the gaseous inhibitor or NaHC03 powder was

introduced into the flame. The fuel cup was cooled by water at the bottom in order to prevent the

fuel from boiling, but was not cooled at the rim as that would establish a radial temperature gradient

within the liquid fuel. A narrow annulus surrounded the wall of the cup near the rim to catch any

overflow of the fuel and prevent it from entering into the exhaust duct. Since, the flame sheet was

very sensitive to surface movements, the fuel height in the cup was accurately controlled by a device

similar to that designed by Bajpai (1974b). A fine pointer projecting up through the liquid pool was

used as an aid in adjusting the height of the fuel surface. Satisfactory repeatability had been obtained

by setting the pool at the level at which the pointer just forms a discernible dimple on the fuel surface. The oxidizer duct had an inner diameter of 50. 1 mm. For experiments with gaseous

inhibitors approximately five layers of fine wire screens (80 mesh/cm) with each layer of the screen

separated by a 1 mm split ring were placed in the duct to reduce the scale of turbulence and to ensure

a flat velocity profile at the exit of the duct. For experiments with NaHC03 powder, only two layers

of coarse wire screen (8 mesh/cm) were employed to prevent accumulation of the powder on screens. Also, in the experiments with NaHC03 powder, it was necessary to replace the screens frequently.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 351

FLAME

POOL OF LIQUID FUEL

Figure 1 . Schematic illustration of the counterflow configuration.
352 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

METERED OXIDIZER

BAFFLE

FUEL

FUEL LEVEL INDICATOR

WATER OUT

WATER JACKET FOR COOLING FUEL CUP

WATER IN FUEL IN FUEL OVERFLOW DRAIN

Figure 2. Schematic illustration of the OFDF burner.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 353

The distance between the oxidizer duct and the fuel cup was adjustable and the oxidizer duct could be

swivelled to allow access to the fuel cup. Experiments were performed with the separation distance,

L, between the surface of the liquid pool and the exit of the oxidizer duct equal to 10 mm. Mild

suction was used to pull the products of combustion from the flame into a heat exchanger surrounding

the fuel-overflow rim. The cooled gases then flowed into the exhaust. The suction minimized the

influence of the ambient air currents on the flame, and prevented after-burning in the heat exchanger.

A steady flame like that shown in Figure 3 could be indefinitely stabilized in this apparatus.

Measurements were made with the burner enclosed within a plexiglass box with an outlet directly

connected to the exhaust to minimize operator exposure to the products of combustion.

The volumetric flow of air, nitrogen and the gaseous inhibitors were measured to ±3% by use of

variable area flowmeters. Pressure gauges were arranged in parallel with the flowmeters to insure

that the volumetric flow of the gas stream was not aff'ected by fluctuations in the output pressure from

compressed gas cylinders or from variations in stagnation pressure caused by adjusting the volumetric

gas flows. The flowmeters were calibrated using wet-test meters. The uncertainty in the concentration of agent at extinction are estimated as approximately 10%. ^ Figure 4 shows a schematic illustration of the device employed to introduce NaHC03 powder

into the flame. This device, originally designed by Williams and Peters (1982), was made of

plexiglass and consisted of a bed of powder resting on a porous disk. Air was introduced from the

bottom fluidizing tube and passed through a bed of glass beads which distributed the flow in the

region between the glass beads and the porous disk. The air passed through the porous disk, lifted the particles and formed a fluidized bed in the region above the bed of powder. The air mixed with

powder was removed from the exit tube and introduced into the oxidizing duct of the OFDF burner.

The jet ring shown in Figure 4 was used to prevent agglomeration of the particles. The jet ring

consisted of a thin mbe closed on one end and bent to form a ring. A number of holes were punched

along the periphery of the ring and air was introduced from the open end of the tube to raise the

particles out of the fluidized bed. The mass fraction of the powder in the mixture of air and the

particles flowing out of the exit tube was altered by adjusting the volumetric flows of air in the

fluidizing tube and the jet ring.

Particle analysis was done by isokinetic sampling and weighing of the samples. After completion

of the experiments, the portion of the OFDF burner containing the fuel cup was removed and a probe

with a 10 mm diameter was placed under the oxidizer duct at nearly the same position as that of the

fuel cup. A filter was placed inside the probe to collect the particles. The volumetric flows of air in

the fluidizing tube and the jet ring were adjusted such that they were identical to those used in the

extinction experiments. A wet-test meter was placed downstream to measure the volumetric flow rate of air in the probe. Mild suction was used such that the velocity of the mixture of air and the

particles entering into the probe was approximately equal to that at the exit of the oxidizer duct. The

flow into the probe was maintained until a measurable amount of powder was collected in the filter. The probe together with the powder was then removed and its weight was measured accurately. The

weight of the powder in the probe was calculated from the difference between the weight of the probe

with and without the powder. The total weight of the air flowing through the probe was calculated

from the measured total volume of air recorded by the wet-test meter. The mass fraction of the

particles in the air stream was then calculated from the ratio of the weight of the powder collected in

the probe to the sum of the total weight of air and the weight of the particles. An average of three

such runs was made at each measured extinction point to determine the mass fraction of powder in the

oxidizing stream. The uncertainties in the amount of powder determined using this method was

estimated to be approximately 15%.
354 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 355

Exit Tube

(air + particles out)

Bed of Powder

Jet Ring

Porous Disk

Glass Beads

25.4mm

I 1 Fluidizing Tube

(air in)

Figure 4. Schematic illustration of the powder dispensing device.
356 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

4.2.3 Experimental Procedure. For experiments with gaseous agents the volumetric flow of air through the oxidizer duct was maintained at some predetermined value and the liquid pool was

ignited. The inhibitor was then added gradually to the oxidizer stream until the flame extinguished.

Care was taken to ensure that the flame attained a steady state after each incremental addition of the

inhibitor. The volumetric flows of air and inhibitor at extinction were recorded. The experiments were repeated with a difi'erent air flow. The velocity of the oxidizer stream at the exit of the duct, U, was presumed to be equal to the ratio of the total volumetric flow of air and the inhibitor to the crosssectional area of the oxidizer duct. Following the analysis of Seshadri and Williams (1978b), the

strain rate, a, is equal to the ratio of U to L, where L is the separation distance between the surface

of the liquid pool and the exit of the oxidizer duct. For experiments with NaHC03 powder, the uniformity of the concentration of powder in the

oxidizing stream was monitored using the optical arrangement shown in Figure 5. For a given level of powder in the oxidizing gas stream, the attenuation of the He-Ne laser radiation was recorded by

the oscilloscope and the strip-chart recorder. For experiments with NaHC03 powder, two separate

air streams flowed into the oxidizer duct. One stream of air, termed the primary air, flowed directly

from the flowmeters into the oxidizer duct, and the other stream of air flowed into the fluidizing and

jet-ring tube of the powder dispensing device shown in Figure 4.

Initially the volumetric flow of primary air was maintained at some predetermined value and the

liquid pool was ignited. Secondary air was then introduced through the fluidizing tube of the powder

dispensing device until a sufficient amount of powder was observed to flow through the oxidizer tube. The uniformity of powder concentration in the oxidizer stream was monitored by recording the output

of the strip-chart recorder. The flow of air was slowly increased until the flame was extinguished.

Care was taken to ensure that the flame attained a steady state after each incremental addition of

powder by monitoring the output of the strip-chart recorder. The flowmeter readings for the primary

and secondary air as well as the fractions of the secondary air flowing through the fluidizing tube and

the jet-ring were recorded. The amount of powder in the oxidizing stream was then determined as described in the previous section. The experiment was then repeated at a different volumetric flow of

the primary air.

4.2.4 Experimental Results. Figure 6 shows the mass fraction of the agents as a function of the

strain rate, a, at extinction. The fuel was heptane and the initial temperature of the oxidizer stream was 25 °C. For a given agent, the region below the curve represents flammable mixtures. For any

value of the strain rate the value of the mass fraction of CF3Br required to extinguish the flame is lower than the mass fraction of all other agents. Therefore on a mass basis CF3Br is considerably more eff'ective in extinguishing the flame than all the other agents tested. Figure 6 shows that the

relative efl'ectiveness of some of the agents in extinguishing the flame varies with the strain rate. This

becomes more clear by examining the results shown in Figure 7, where the data are replotted with an expanded vertical scale. At low strain rates HFC-236 is most eff'ective and FC-116 least efi'ective in extinguishing the flame, whereas at high strain rates HCFC-22 is most eff'ective and FC-318 least

eff'ective in extinguishing the flame. In fact at high strain rates only CF3Br and HCFC-22 are more

eff'ective than N2 in extinguishing the flame. However, the results plotted in Figures 6 and 7 show

that with the exception of CF3Br, the diff'erences in the mass-based eff'ectiveness of the various agent

are not large.

Figure 8 is a plot of the same data shown in terms of the mole fraction of the agents. The fuel was heptane and the initial temperature of the oxidizer stream was 25 °C. For any given value of the

strain rate the value of the mole fraction of CF3Br required to extinguish the flame was lower than the

mole fraction of all other agents. On a mole basis, at low and high strain rates FC-31-10 was most

eff'ective and HFC-32/125 least eff'ective in extinguishing the ffame. Note that on a mole basis, all
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

357

Strip Chart

Recorder

Oscilloscope

Signal

Amplifier

Tube Attenuator

n

Photo Diode

Oxidizer Duct

He-Ne Loser

Liquid Fuel

Figure 5. Schematic of the optical arrangement employed for monitoring powder

concentration.
Figure 6. Mass fraction of various agents as a function of the strain rate for heptane flames at extinction, with an oxidizer temperature of 25 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 359

Figure 7. Mass fraction of various agents as a function of strain rate for heptane flames at extinction (expanded vertical scale), with an oxidizer temperature of 25 °C.
360 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

the agents tested were more effective than N2 in extinguishing the flame, which is the inverse of the

results when plotted versus mass fraction.

4.2.4.1 Impact of Fuel Type. Figures 9 and 10 show, respectively, the mass and mole

fractions of the agents as a function of the strain rate at extinction for JP-8 flames. The oxidizer

stream temperature was 25 °C. A comparison of Figures 8 and 10 shows that the JP-8 flames were somewhat easier to extinguish than the heptane flames. The average agent extinction concentration

for the JP-8 flames was 65 % to 93 % of that required to extinguish the heptane flames over the range

of strain rates, with the differences being greatest at the highest values of strain rate. The reasons for the variation of this value with strain rate is not clear.

4.2.4.2 Impact of Air Temperature. To clarify the influence of the initial temperature of the

oxidizer on the critical conditions of flame extinction experiments were performed with the initial temperature of the air equal to 150 °C. Prior to performing these experiments, radial temperature

profiles were measured in the apparatus without igniting the flame using thermocouples in the region

between the exit of the gas duct and the surface of the liquid pool. The results are plotted in Figure 11, where the quantity DT refers to the difference between 150 °C and the measured value of

the temperature, and r refers to the radial distance in millimeters from the axis of symmetry. Four

sets of measurements were made: at the exit of the duct, 2 mm below the exit of the duct, and at approximately the position where the flame was stabilized above the surface of the liquid pool. The

latter measurements were made with and without the pool covered with a thin metal sheet. All

measurements were made with liquid heptane in the cup. The profiles near the exit of the duct show

that the temperature is roughly 18 °C lower than 150 °C in the region near the walls of the duct.

However, over a radius of approximately 15 mm from the axis of symmetry, the temperature was

nearly uniform and differed from 150 °C by amounts less than 5 °C. The measurements made at the

approximate location where the flame would stabilize showed that the radial temperature profile was

relatively uniform. The values of the temperature measured with the pool covered were found to be

lower than those measured with the uncovered pool, which was attributed to heat losses to the metal

sheet. The temperature measurements clearly show the profile to be nearly uniform at the approximate position of the flame, and the value to differ from 150 °C by an amount less than 5 °C.

The relative mass-based effectiveness of the various agents differed slightly for different oxidizer

stream temperatures. Figures 12 and 13 show the mass fraction of agent as a function of strain rate for heptane and JP-8, respectively, with an oxidizer temperature of 150 °C. A comparison of the

results in Figures 12 and 13 with Figures 6 and 9, shows that an enhanced oxidizer temperature leads

to flames which require somewhat more agent to extinguish. This result is consistent with theories of

flame extinction. As more enthalpy is added to a flame, it is more difficult to extinguish. Figures 14

and 15 replot the data from Figures 12 and 13 in terms of agent mole fraction.

4.2.4.3 Powder Agent: Impact of Size. Figure 16 shows the mass fraction of NaHC03

powder with a nominal particle size of 10-20 ^m in the oxidizer stream as a function of the strain rate

at extinction. The fuel was heptane and the oxidizer stream temperature was 25 °C. For comparison

the critical conditions of extinction measured with CFgBr and N2 added to the air stream are also

shown. For measurements with NaHC03 powder, the experimental uncertainties are expected to be

less than 30 % . For any given value of a, the value of the mass fraction of NaHC03 required to extinguish the flame is lower than the mass fraction of CF3Br and N2. Therefore on a mass basis

NaHC03 was more effective in extinguishing the flame than all the agents tested including CF3Br.

Figure 16 shows that at all values of the strain rate tested, the mass of NaHC03 powder required to extinguish the flame was a factor of three lower than that of CF3Br. Figure 16 shows the mass
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 8. Mole fraction of various agents as a function of strain rate for heptane flames at extinction, with an oxidizer temperature of 25 °C.
362 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 9. Mass fraction of various agents as a function of strain rate for JP-8 flames at extinction witii an oxidizer temperature of 25 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Strain rate at extinction [s* ]

Figure 10. Mole fraction of various agents as a function of strain rate for JP-8 flames at extinction with an oxidizer temperature of 25 °C.
364 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

—I 1 T I [ I O AT 2rTm under tip [K]

AT @tip [K]

AT @ tlasm pos. w/ covered pcxDl • AT @ flame pos. w/ uncovered pool

-| 1 I r T

o

25

20

1 5

o

1 0

o

o o

o

o

[3

• o • o»o#o»o

_J 1 1 1 I . 1 I 1 I I 1 1_

- 5

0

0.00 5.00 10.00 15.00 00

r [mm]

20.00 25.00 30

Figure 11. Radial temperature profiles measured without the flame at various locations at the

region between the exit of the gas duct and the surface of the liquid pool for an oxidizer stream temperature of 150 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 365

Figure 12. Mass fraction of various agents as a function of strain rate at extinction for heptane

flames with an oxidizer temperature of 150 °C.
366 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

0.00 I—'—'

'

i ''—'—'—i ' ' i

'

'——— ———

'—

0 50 100 150 200 250 300

strain rate at extinction [s'^]

Figure 13. Mass fraction of various agents as a function of strain rate at extinction for JP-8

flames with an oxidizer temperature of 150 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 367

0.40

0.00 I—'—'

—'—^—'—'—'—i

•—>—'——^'—'—•—^——.—.——^—''—'—-J

0 50 100 150 200 250 300

strain rate at extinction [s* ]

Figure 14. Mole fraction of various agents as a function of strain rate at extinction for heptane

flames with an oxidizer temperature of 150 °C.
368 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 15. Mole fraction of various agents as a function of the strain rate at extinction for JP-8

flames with an oxidizer temperature of 150 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 369

fraction of NaHC03 powder (10-20 /xm) in the oxidizer stream required to extinguish diffusion

flames burning heptane and JP-8 as a function of the strain rate. The oxidizer stream temperature was 25 °C. Consistent with extinction data measured using gaseous agents, Figure 17 shows that at any value of the strain rate, the mass of NaHC03 required to extinguish JP-8 flames was less than

that required to extinguish a heptane flame.

Figure 18 shows the influence of initial particle size of NaHC03 in extinguishing heptane flames.

The initial temperature of the oxidizer stream was 25 °C and the nominal particles sizes tested were

0-10, 10-20, and 20-30 ^m. The results show that particles between 0-10 /xm were more eflFective than those between 10-20 and 20-30 fim. However, the results show that 20-30 fxm particles were more effective than those 10-20 //m. Further studies are required to explain this observation.

4.2.5 Interpretation of the Experimental Results. As mentioned at the begirming of Section 4, in fundamental theories of flame extinction where the gas-phase chemical reaction can be approximated

as a one-step process, changes in the value of the Damkohler number, D^, can be used to qualitatively

explain the results of the OFDF suppression experiments. For a given chemical system, flame

extinction is predicted to occur at a well defined and experimentally reproducible value of the critical Damkohler number D^* (Fendell, 1965; Linan, 1974; Peters, 1982, 1984). The value of t^^^em (and,

hence, D^) will depend on the relative concentrations of the various reactants including that of the

agent and the local gas temperature, and the value of t^^^ will depend inversely on the strain rate a. In the absence of temperature measurements at conditions close to flame extinction, the calculated

adiabatic flame temperature Tf, can be used in Equation (1) to provide a rough measure of the

characteristic chemical time. For flames near extinction, the value of t^,[^gf^ decreases with increasing

values of Tf for a fixed global activation energy. The concentration of fuel, air, and agent at extinction in a heptane flame were used (from

Figure 6) to calculate the adiabatic flame temperature, assuming complete combustion to stoichiometric amounts of CO2 and H2O, and frozen levels of inert agent. Figure 19 is a plot of Tf as a function of strain rate for each of the gaseous agents. The temperatures at extinction concentrations

of CF3Br are higher than the value of Tf calculated using the measured mass fractions of all other

agents. Therefore for any value of t^^, the critical conditions of extinction measured with CF3Br

yield values of t^,f,g^ which are smaller than those for the other agents. Hence the values of D^* with

CF3Br is higher than the corresponding values of D^* from the other agents. The values of Tf for N2

suppression are 100 to 300 K lower than for CF3Br. The temperatures for agents without chlorine

are below nitrogen, and the two containing chlorine have slightly higher calculated temperatures.

Since the nitrogen data are representative of a purely thermally acting agent, Tf can be used as a

qualitative measure of the thermal versus chemical influence of the agent on the flame.

A global activation energy can be estimated from the theoretical value of D^* found by performing an asymptotic analysis (Linan, 1974; Peters, 1982, 1984; Krishnamurthy et al., 1976) of the

flame. Knowing the value of at extinction from the analysis, and the flame temperature and strain

rate from Figure 19, Equation (1) can be used to estimate the global rate coeflficient k. Figure 20 is a plot of the natural log of a function (Krishnamurthy et al., 1976) which is directly proportional to k

(see Equation 1), versus 1/Tf. (The value of Tf was calculated assuming that the agents were

chemically inert.) The value of the Eg for each of the agents can be estimated from the slope of the

respective curves. For N2, Eg is 167 kJ/mole. Deviations of Eq from this value for the other agents

is a rough measure of the chemical influence of these agents. The value of Eq calculated using

extinction data with CF3Br is equal to 894 kJ/mole, an unrealistically high value. Clearly the
370 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 16. Mass fraction of NaHC03 (10 - 20 ^m) powder in air as a function of the strain rate

at extinction for heptane flames with an oxidizer temperature of 25 °C.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 17. Mass fraction of NaHC03 (10 - 20 ixm) powder in air as a function of the strain rate

at extinction for heptane and JP-8 flames with oxidizer temperatures of 25 °C.
372 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

O

o

0.050

0.040

0.030 -

0.020

0.010

0.0

— (O-IOti) ^NaHC03 — (20-30n) - ^NaHC03 —

(10-20^) '^'NaHCOS

50 100 150 200 250 300

strain rate at eztinction [s'^]

Figure 18. Mass fraction of NaHC03 (0-10, 10-20, and 20-30 jum) powder in air as a function

of the strain rate at extinction for heptane flames with an oxidizer temperature of

25 °C.
Figure 19. Calculated adiabatic flame temperatures at extinction assuming complete combustion

of fuel and frozen composition of agent.
374 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

-7.50

Figure 20. Arrhenius plots obtained using the extinction data shown in Figure 6. The agents were considered to be chemically inert.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 375

one-step approximation is not valid, and CFgBr appears to be having a significant influence on the

flame chemistry.

Figure 2 1 shows the adiabatic flame temperatures at diff'erent strain rates calculated assuming that the agents and stoichiometric fijel/air mixture react completely to CO2, HjO, HP, HBr, and HCl.

The halogenated agents tend to have higher calculated temperatures in Figure 21 as compared to Figure 19 due to the enthalpy associated with conversion to the relatively stable acid gases, which

implies that at least part of the chemical behavior exhibited by the agents may be related to simple

conversion to equilibrium combustion products, HF and HCl. It is difficult to distinguish between the

chemical and thermal influences of these agents on the flame because of possible diff'erential transport

eff'ects. These efl'ects are not accounted for in the current analysis.

4.2.6 Conclusions from OFDF Study. The opposed flow configuration was employed to determine

the relative efl'ectiveness of various inhibiting agents in extinguishing diffusion flames burning liquid

heptane and JP-8. Eleven gaseous inhibiting agents and NaHC03 powder were tested. The oxidizing

gas used in the experiments was a mixture of air and the inhibiting agent at initial temperatures of

25 °C and 150 °C. The following remarks summarize the major conclusions of the study:

1 . The OFDF burner can be used to measure accurately the amount of a powdered agent required to suppress a variable strain- rate, laminar flame.

2. The strain rate, agent, fuel type and oxidizer temperature (in decreasing order of significance) all

aff"ect the amount of agent necessary to extinguish the opposed-flow difiPusion flame.

3. The ranking of relative efl'ectiveness of the various gaseous agents in quenching difl\ision flames on a mole basis diff'ers substantially from ranking on a mass basis, due to variation in molecular

weight.

4. The rankings of relative mass- and mole-based efi'ectiveness of the various agents are aff'ected somewhat by the strain rate at which the tests are run, but only slightly by the fuel type (JP-8 or

heptane) and the oxidizer temperature (25 °C and 150 °C).

5. On a mass or mole basis CF3Br is considerably more eff"ective in extinguishing the flame than all of the other gaseous agents tested.

6. On a mass basis at low strain rate HFC-236fa is most eff'ective and FC-116 least eff"ective of the

gaseous alternatives, whereas at high strain rates HCFC-22 is most eff'ective and FC-318 least

effective.

7. On a mole basis, at low and high strain rates FC-31-10 is most efl'ective and HFC-32/125 least

efi'ective of the gaseous alternatives.

8. Interpretation of the gaseous agent results based on one-step activation energy asymptotic theories

shows that CF3Br has a significant influence on the flame chemistry, and that the one-step

approximation is not realistic. The results also show that all the agents except CF3Br have a

significant thermal influence on the flame. However, it is difficult to distinguish between the two

influences accurately because preferential difliision has not been considered. Further studies

using detailed and reduced mechanisms are required to clarify the influence of the gaseous agents on the flame.
376 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 21. Calculated adiabatic flame temperatures as a function of extinction concentrations

and strain rates assuming complete conversion of agents to CO2, H2O, HF, HCl,

and HBr.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 377

9. The effectiveness of NaHC03 powder on a mass basis is greater than that of CF^Br; powder with

0-10 particles is more effective than the 20-30 ixm, which in turn is more effective than the

10-20 fj.m. The mechanisms of flame extinction by powders are not well understood and merit

further investigation.

4.3 Coflowing Nonpremixed Flames

4.3.1 Apparatus. An experimental study was conducted to rank the relative suppression effective- ness of various inhibiting agents on nonpremixed flames stabilized in a cup burner. The cup burner is a very common test apparatus for screening of suppressants. It is not, however, a standardized test method. A number of different size apparatus exist and various fuels and different air flows have

been used (Gann et ai. 1990). Nevertheless, the experimental results from the different investigators

have shown general agreement. The cup burner was selected as a test apparatus not to determine

absolute extinction concentration data, but to rank the relative suppression effectiveness of the various

agents. As such, it is an adequate test apparatus.

The concentration of gaseous agent in the oxidizer stream needed to obtain extinction was measured in the cup burner shown in Figure 22. The oxidizer, a mixture composed of dry air and

agent, flowed through a 96 mm diameter chimney around a fuel cup. The cup had a 45° ground

inner edge (28 mm diameter) and was used for both liquid and gaseous fuels. The ground inner edge

minimized heat losses from the flame to the Pyrex cup. A fine wire mesh screen (40 mesh/cm) and

small glass beads ( = 0.5 mm diameter) were placed inside the fuel cup for the gaseous fuel to flatten the velocity profile. The liquid fuel was gravity fed from a reservoir using a previously developed

technique which facilitated an adjustable and constant level in the cup (Bajpai, 1974b).

Figure 23 is a schematic of the oxidizer flow control system. Compressed breathing air was dried

by passing it through a two-stage filtering system composed of a silica gel water vapor trap followed

by a particle trap. The flows of air and agent were monitored by rotameters which were calibrated

with a dry test meter. Rotometer float and pressure readings were both measured. When suflRcient agent was available, direct calibration was performed with that agent. Otherwise, an indirect

calibration method was performed with dry air. Correction for density and viscosity differences were

then applied to determine the flow. A series of tests were conducted with HCFC-22 to compare the

precision of the two calibration methods. On average, the indirect method, calibrating with dry air, was within 3% of the direct calibration method. The values of gas viscosity (at 23 °C) for a number

of test agents are listed in Table 4 (Reid et al., 1977). For gases where data were not available

(HFCs -125, -134a and -227, and HCFC-124) Reichenberg's corresponding states method was

utilized. This method is described by Reid et ai, (1977). The method requires a number of physical

parameters including the critical temperature, molecular weight and group contributions. For the

mixture (HFC-32/HFC-125), Wilke's approximation was used. The critical temperatures are listed in Table 1 of Section 2.

4.3.2 Experimental Method

4.3.2.1 Gaseous Agents. A detailed experimental protocol was developed based on previous

studies (Garm et ai, 1990). The fuel level was carefully set such that the ground lip of the pyrex cup was completely covered with fuel. The air flow was initiated and the flame was ignited with a torch.
378 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 22. Schematic drawing of the cup burner.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 379

SYKCOLS 0 PFESSUFE GAUGE

X

VALVE TO CUP BURNER

A

'

3-WAY VALVE 0

DRY

TEST

METER

TO

ATMOSPHERE

-O ' <H

0-

ROTAMETER

0-

t

ROTAMETER

PARTICLE FILTER

SILICA GEL FILTER

DRYING

SYSTEM

BREATHING

AIR

CYLINDER

Figure 23. Schematic drawing of the oxidizer flow control system.
380 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 4. Absolute Viscosity of Some Agents

Agent Molecular Weight,

g/mole

Absolute Viscosity @ 23 °C, N-s/m^

HFC-32 (CH2F2) 52.1 1.20 X

10-5

HFC-32/HFC-125 (CH2F2/C2HF5) 67.3 1.32

HFC-227 (C3HF7) 170.0 1.59 HCFC-22 (CHF2CI) 86.5 1.29 HFC- 134a (C2H2F4) 102.0 1.30

FC-116(C2F6) 138.0 1.44 HCFC-124 (C2HCIF4) 136.5 1.20 HFC- 125 (C2HF5) 120.0 1.39

FC-128 (C3F8) 188.0 1.35

FC-31-10 (C4F10) 238.0 1.25

1301 (CF3Br) 148.9 1.55

1211 (CF2ClBr) 165.4 1.36

FC-318 (C-C4F8) 200.3 1.18

CO2 44.0 1.50

N2 28.0 1.76

Ar 40.0 2.23

He 4.0 1.96

C3F6 150.0 1.36

HFC-236 (CF3H2CF3) 152.0 1.22

SiF4 104.1 1.60

C2BrF3 160.9 1.36

C2F3CI 116.4 1.28

CF4 88 1.72

Tests showed that the jet fuels and hydraulic fluids did not readily ignite without the use of a microtorch. Once ignition occurred, the torch was immediately removed to prevent excess preheating of

the Pyrex cup. The chimney was placed about the fuel cup. Figure 24 is a photograph of the flame

in the cup burner. A 100 second warm up period was used prior to agent addition to minimize

variation in apparatus heat-up during ignition. With the flame established, the agent concentration in the oxidizer stream was incrementally increased until the flame was extinguished. Agent was added to the oxidizer stream in small amounts (< 1% of flow needed to attain flame extinction), with a waiting a period from five to ten seconds between increases in the agent flow. This allowed time for the

changing oxidizer flow to reach the burner and to impact flame stability. The flame was observed
II

SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 381

Figure 24. Photograph of cup burner and liquid fuel supply.
382 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

during this waiting period. After flame extinction was achieved, the used fuel was drained and the

test was repeated several times (with fresh fuel each time) to insure experimental precision.

Extinction tests were conducted on six fuels including one gas (propane) and five liquid fuels, which

were heptane, two jet fuels (JP-5 and JP-8) and two hydraulic fluids (5606 and 83282) as shown in Table 1.

4.3.2.2 Liquid Agents. A liquid agent was introduced into the oxidizer stream using the

delivery system shown in Figure 25. A regulated nitrogen stream flowed into a beaker containing the

liquid agent, displacing the agent from the beaker into the heated oxidizer stream. The beaker

holding the agent was situated in an ice-water bath. The flow of nitrogen was monitored using a

calibrated flowmeter and a pressure gauge above the agent was observed. The liquid density at the

bath temperature was needed to calculate the mass of agent addition to the oxidizer stream. The

liquid agents CF2Br2 and CH2BrCF3 had densities of approximately 2.46 and 1.79 g/ml, respectively.

A variable voltage regulator allowed control of the temperature of the heated oxidizer lines. A

thermocouple above the glass beads inside of the chimney monitored the oxidizer temperature which

was typically 10 to 15 °C above ambient. A test, described in Section 4.3.3, was conducted to substantiate the viability of the agent delivery method.

4.3.2.3 Sodium Bicarbonate Powder. The cup burner shown in Figure 22 was modified to foster a controlled, homogeneous flow of powder and to facilitate repeatable extinction measurements

using powder agents. Figure 26 is a schematic diagram of the modified cup burner system, the

system had five components: the cup burner, the air drying and control system, the powder feed

system, the powder monitoring system, and the iso-kinetic sampling system used for calibration of the

powder concentration.

A series of wire mesh screens (10 mesh/cm) functioned as flow straighteners. They were placed

across the entire cross sectional area of the chimney, below the fuel duct, after cutting the chimney

into several sections. Without these screens, non-homogeneous powder flow and recirculation zones were observed near the fuel cup. Flow visualization under non-combusting conditions showed that uniform upward flow was achieved for the powder/oxidizer dispersion for velocities greater than = 9

cm/s.

The powder was introduced into the air stream by turning a 9 mm drill bit through a closed

container of powder. The drill bit carried the powder through a 12.7 mm tube. The powder was fed

into the oxidizer stream via a tee located just below the base of the cup burner where it was swept

into the oxidizer flow. The delivery rate was controlled by a variable speed motor and a voltage

regulator. Flow visualization showed that the 0-10 and 10-20 fxm powder traveled through the

chimney as a uniform bulk flow, whereas the 20-30 fxm size fraction did not, for flows achievable in our laboratory. Thus, no powder testing was conducted with the larger size fraction. For the powder

experiments, the air stream flow was maintained at 40 1/min (in the 96 mm diameter chimney).

The concentration of powder in the flow field at the time of extinction for locations near the

flame was quantified by a laser transmission measurement in combination with iso-kinetic vacuum

sampling and a wet chemistry technique - sodium ion analysis. The method assumes a uniformly

seeded flow field. A 10 mW helium-neon laser beam (632.8 nm) was directed across a =7 cm chord

of the pyrex chimney which traversed the flow field. The beam was positioned = 2 cm below the

fuel duct to avoid light attenuation by soot particles or other combustion products. The laser signal was monitored by a silicon photo-detector with a 10 nm bandwidth line filter and connected to a data

acquisition system (4 Hz sampling rate). This allowed monitoring of the line-of-sight integrated

NaHC03 concentration as a function of time. Increased powder concentration in the flow field was

indicated by increased laser attenuation as represented by the following equation:
I!

SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 383

j ICE BATH METERED

TEMPERATURE AIR FLOW

I BATH

Figure 25. Schematic drawing of the cup burner modified for delivery of liquid agents.

1
384 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 26. Schematic drawing of the cup burner modified for deHvery of a solid powder.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 385

- = exp(-K-[NaHCO^]D (3)

where 1^ is the unattenuated laser signal, I is the attenuated signal, [NaHCOj] is the powder concentration, L is the path length, and K is an absorption/scattering coefficient which is dependent on

particle diameter, and shape as well as a number of other parameters. At the moment of flame

extinction, a signal was sent to the data acquisition system, the laser transmission was monitored and

the flow system was immediately shut off". This prevented further powder agglomeration on the

chimney walls and allowed determination of the laser signal (Ig), unattenuated by powder in the flow

field. Laser attenuation was related to mass percent of powder in the oxidizer by iso-kinetic vacuum

sampling in combination with a sodium ion electrode/analysis system.

For calibration, small holes (2 mm) were drilled in the chimney for passage of the laser beam,

eliminating the possibility of laser attenuation due to powder adsorption on the chimney walls. Flow

visualization showed that a negligible amount of powder escaped through the orifices. During

calibration, the top section of the chimney was removed and the probe/collection system was

positioned at the plane defined by the rim of the fuel cup, where the flame was typically anchored.

This avoided possible losses of powder to the chimney walls, which would have occurred if the

sample was extracted at the top of the chimney.

Calibration of the NaHC03 mass flow in the oxidizer stream was determined by iso-kinetically

sampling a portion of the NaHC03/air dispersion into a vacuum collection system while simultaneously monitoring the laser signal (I) for diflFerent powder loadings as shown in Figure 27. The absorption/scattering coefficient K was assumed to be invariant during the extinction and calibration

measurements. The powder/oxidizer dispersion was iso-kinetically collected through a 20 mm

diameter tube connected to a 20 1/min vacuum pump. The sample passed through = 25 cm of

polyvinyl tubing into a 200 ml glass impinger filled with deionized water which captured the water

soluble powder particles. After approximately a 30 second sampling time, the sample probe and lines were washed with deionized water, combined with the impinger contents, and diluted with deionized

water to 500 ml. A calibrated sodium ion-selective electrode was then used to determine the ion

concentration in the powder/water solution. In this manner, a calibration curve was developed

covering an appropriate range. Typically, 0.5 to 2 g of powder were collected over the sampling

period, which was completely salable in the 500 ml water solution.

4.3.3 Assessment of Operating Parameters. A series of preliminary measurements tested the eff'ect of a number of parameters on the agent concentration in the oxidizer stream at extinction. The

parameters tested include the fuel cup diameter, the oxidizer flow, the chimney diameter, and the

preburn time. The oxidizer flow was found to have a negligible eff'ect on the concentration of agent

at extinction, once a minimum flow was achieved, as seen in Figures 28 and 29 for N2 and HCFC-22

suppressing JP-5, JP-8, 5605, and 83282 flames. The efi'ect of varying chimney diameter (96 and

113 mm diameters) was also tested. The chimney size was found to have a negligible eff'ect once a minimum flow velocity was achieved as shown in Figure 30 for HCFC-22 extinguishing 5606 and

JP-8 flames. Two fuel cup diameters (12 and 28 mm) were tested as shown in Figure 31 and found

to have a negligible impact on the agent concentration (HCFC-22) at extinction of JP-8 flames.

Because fuel consumption in the two cups was expected to be difl'erent, this implies that the fuel burning rate had a minimal impact on the measured extinction concentration. In order to directly test

this, flames burning a gaseous fuel (propane), where the fuel flow was an independently controlled

experimental parameter, were used. The results presented in Figure 32 show only a small variation in
386 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

t

TO VACUUM PUMP

< VALVE

FILTER

< ISOKINETIC SAMPLING SYSTEM

METERED

AIR FLOW

Figure 27. System used to calibrate powder flow in cup burner.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 387

35

30 -

c

<DO

03 2 5

Q.

(D

I 20

c

0)O)

<

15 -

1 0

1

-

I I I

J t T I T

1

1 1 1

' 1

' 1 1 1

1

1 1 1 1

1

1 1 1 1 -

-

o o o o

D

a

° 5606 ; 83282 - Agent = o JP-5 ;

_ 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1

JP-8 -

1 1 1 1 1 i. —1

0 1 0 20 30 40

Air Flow (l/min)

50 60

Figure 28. N2 concentration at extinction as a function of air flow.
388 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

•| g I—1—I—I—I—

I—I—I—I—I—

I—I—I—I—I—I—I—I—I—I—I—I—I—I—I—

I—I—I—I—I—

I—I—I—I—I—

I—I—I—I—I

S 10

o

o

Q.

E

O

>

c 5

<

O <> o

Agent = R-22

5606

83282

o JP-5

JP-8

—I 1 1 1 1—I—I 1 1—I—I—L—I 1 1 1—I 1—I 1—I—I 1 1—I 1 1 1 1—I 1 1—I 1—I—I 1—I 1—

0 10 20 30 40 50 60 70 80

Air Flow (l/min)

Figure 29. CHF2CI concentration at extinction as a function of air flow.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 389

j 5 r*""—'—'—'—'—'—'—'—'—I—'—'—'—'

-|—I—I—I—I—I—r—I—I—r—

o 10 -

o o

• o • ° o

5 ^

-J I I I I I I I u

Agent = R-22

o large chimney (D=11.3 cm); fuel = JP-8

• small chimney (D=9.6 cm); fuel = JP-8

large chimney (D=11.3 cm); fuel = 5606

small chimney (D=9.6 cm); fuel = 5606

-J I I I u J I I L 0 5 10

Air Velocity (cm/s)

15

Figure 30. CHF2CI extinction concentration as a function of air velocity for different chimney

sizes.
390 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

1 5 I—'—'—'—'—I—'—'—'—'—

r

T—I—I—

r

T—

I—I—I—I—I

C

CDo 10

(D

O O

A

5

-

Agent = R-22

Fuel = JP-8

o large cup

A small cup

0

J 1 1 1 1 1 1 1 1 1 1 u J 1 1 1 u 10 20 30 40

Air Flow (l/min.)

50 60

Figure 31. CHF2CI extinction concentration as a function of air flow for difi'erent fuel cup

sizes.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 391

the N2 and HCFC-22 extinction concentration with the propane flow. In order to test if the global

equivalence ratio was a factor in these results, a particular fuel flow (0. 1 1 1/min) was used and the

eff'ect of increasing air flow on the agent extinction concentration was tested and found to be

negligible. Thus, the fuel flow had only a small eff"ect on the agent concentration at extinction. This

eff'ecl may be related to flame liftoff" and air/agent premixing with the fuel near the lip of the fuel cup. Further study is needed to investigate the mechanism responsible for these eff'ects. A number of the fuels contained multiple components. Thus, preferential distillation effects might have led to differences in the agent extinction concentration as a function of time after ignition.

This eff"ect was tested and found to be negligible for preburn times from 50 to 400 s. However,

because the fuel burning rate was found to be unsteady at times less than 100 s, the experimental

protocol required a 100 s preburn time after ignition and before agent was added. In addition, all extinction experiments were completed in less than 300 seconds. For all measurements, the 96 mm

chimney was used. For the gaseous and liquid agents, an air velocity in the chimney of approximately 5 cm/s was used. For the extmction measurements with NaHC03, an air velocity of approximately 10 cm/s was used to insure that the powder travelled uniformly through the chimney.

Measurement variance and a propagation of error calculation considering the air and agent flows,

yielded an estimate of the combined standard uncertainty in the gaseous agent concentration at extinction as ± 10%. The combined standard uncertainty in the measurement of the NaHC03

concentration at extinction was found to be approximately 25 % on average and was always less than

40%. This value was dominated by measurement variance. In order to test the impact of heating the liquid agent delivery system, the concentration of N2 - needed for extinction of a propane flame was tested with and without a heated (to 35 °C) oxidizer stream. The extinction concentration showed negligible difl'erence. As agents were added to the oxidizer stream the color of the flame changed. With the addition of an inert non-halogenated agent such as He or N2, the flame lifted off" the burner, became narrower and blue in color. With the addition of halogenated agents, the flame lifted off" the burner, and

became more intensely yellow in color for all of the fuels studied. At large agent concentrations,

copious amounts of soot were produced, covering the pyrex chimney. A greenish luminosity was

emitted from the flame when a chlorinated agent was used.

4.3.4 Experimental Results

4.3.4.1 Core Agents. Figure 33 shows the cup burner extinction results for N2, CF3Br and the core gaseous agents on a mole basis for flames burning propane, heptane, JP-8, JP-5 and the

hydraulic fluids 5606 and 83282. The propane flames were the most difficult to extinguish, followed

by heptane, the jet fuels, and the hydraulic fluids. The measurements compare favorably to previous

results as shown in Table 5 for heptane flames. Consistent with our results. Hirst and Booth (1977)

also found propane more difficult to extinguish than heptane. Cup burner extinction measurements

using the jet fuels and the hydraulic fluids have not been previously reported.

The different agents in Figure 33 had dissimilar suppression effectiveness. For all fuels, CF3Br

was a highly effective suppressant as compared to the other gaseous agents, being approximately an

order of magnitude more effective than N2 in extinguishing the cup burner flames. Other agents had

intermediate effectiveness as compared to halon 1301 and N2. On a mole basis, the larger molecular

weight FC, HFC, and HCFC compounds had a higher suppression effectiveness. The most efficient agent was C4F10, followed by cyclo-C4Fg, C3Fg, and C3HF7. The chlorinated agents (CHF2CI and

CHFCICF3) did not exhibit flame suppression behavior superior to the non-chlorinated agents.

Figure 34 replots the mole based results shown in Figure 33 on a mass basis. The most efficient gaseous agent was halon 1301 requiring approximately 15% by mass in the oxidizer stream to
392 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

40

30

c

0)o

(D QO

I

20

o

>

c

CD <

10

T T r— I I

1 r T 1——

r

_, L

• agent = N2

o agent = R-22

I I I I I I I

0.05 0.1 0.15

Propane Flow (l/min.)

0.2 0.25

Figure 32. N2 and CHF2CI extinction concentration as a function of the propane fuel flow rate.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 393
394 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 34. Agent (mass percent) concentration at extinction.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 395

Table 5. Comparison of (mole based) Extinction Measurements with Previous Work. Fuel is Heptane.

Agent Halon 1301 CF3I C2F6 N2

this study 3.27c 3.2% 8.1% 32%

Hirst and Booth, 1977 3.5% a a 30.2%

Harrison, 1992 3.5% a a a

Bajpai, 1974a 3.3% a a a

Sheinson et al., 1978 3.1% 3.2% 7.9% 30%

^ Not measured

extinguish the cup burner flames. The perfluorinated agents all had approximately the same efl'ective- ness as on a mass basis. This is directly related to the agent heat capacity, as listed in Table 6,

which shows that all of the perfluorinated agents have approximately the same heat capacities at flame

temperatures.

4.3.4.2 Sodium Bicarbonate Powder. The color of the flame became bright yellow with the

addition of even minute amounts of NaHCOj into the flame. Equilibrium calculations (Gordon and

McBride, 1970) indicate that at flame temperatures, the sodium containing species in the flame with

the highest concentrations are Na"^ and NaOH. The intense yellow luminosity was likely due to high

temperature Na"^ ion emission in the flame. As the powder concentration in the oxidizer stream was

increased, the flame became thinner and less luminous. Near extinction only a small luminous region was apparent.

For suppression of the liquid fuel flames, some powder was observed to accumulate inside of the

fuel cup and submerge. The effect of NaHC03 powder on the extinction resistance of the fuels was

tested by measuring the extinction concentration of N2 in the oxidizer stream for flames burning fuels with and without powder sprinkled on the fuel surface. No diff'erences were determined for any of

the five liquid fuels tested. Experiments showed that flame extinction using Si02 was not possible for powder loadings

achievable with our delivery system. With as much silica as could be delivered by our apparatus, the

cup burner flames appeared unaltered. The Si02 concentration attainable by the delivery system was

not quantified.

Figure 34 shows that the concentration of NaHC03 required to extinguish the cup burner flames

for two powder size fractions. Approximately 4% by mass of NaHC03 in the oxidizer stream was

necessary to extinguish the heptane cup burner flames. A comparison shows that in general, NaHC03

was a factor of three to five times more efficient than halon 1301 on a mass basis depending on fuel type and powder size. Unlike the gaseous agents, propane flames were less difficult to extinguish than

heptane for both powder sizes. The reason for this warrants further study.

There were some inconsistencies between the 0-10 and the 10-20 /^m powder results. Ewing

et al., (1992) observed enhanced extinction eff'ectiveness with decreasing powder size in heptane pool
396 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 6. Agent Heat Capacity in the Oxidizer Stream and Calculated Adiabatic Temperature near

Extinction for Heptane Cup Burner Flames

Agent

CJ1627 °C1

(J/g-°C) Tf (°C)

CFjBr (1301) 0.0406 1981 CO2 0.0779 1567

N2 0,0731 1586

C3HF7 (HFC-227) 0.0822 1614

CHF2CI (HCFC-22) 0.0678 1689 CH2FCF3 (HFC- 134a) 0.0946 1567

CjFf, (FC-116) 0.0726 1628

CHFCICF3 (HCFC-124) 0.0717 1706

CHF2CF3 (HFC-125) 0.0836 1615

C3Fg (FC-218) 0.0764 1597

C4F,o (FC-31-10) 0.0764 1569

cyclo-C4Fg (FC-318) 0.0798 1560

Ar 0.0296 1607 He 0.296 1747

CF3H 0.0831 1685

CF4 0.0683 1522

fires, until a critical powder size ( = 30 jum for NaHC03), when the effectiveness was nearly a constant. The results reported here are not consistent with these findings. Figure 34 shows that the

larger diameter particles were more eflficient than the small size fraction. It is possible that the 0-10

^m size fraction was exposed to heat or water vapor under preparation or storage, causing the

NaHC03 to react and form Na2C03, a reaction which is known to begin at 50 °C. This possibility is being currently investigated. In any case, NaHC03 is three to five times more efficient than halon

1301 on a mass basis.

4.3.4.3 Additional Agents. The mole and mass-based heptane extinction results for all gaseous

agents tested are shown in Table 7, grouped by molecular type (inert, a silicon containing compound,

a sodium containing compound, perfluorinated agents, and chlorine, bromine, and iodine containing

compounds). On a mass basis, the most efficient agent was the solid powder, NaHC03. The most

efficient gaseous agent was He, followed by CHCI2F, CF3Br, and CF3I. The most efficient gaseous

agent on a mole basis was the liquid agent CF2Br2, followed by CF3Br, CF3I, CH2BrCF3, and

CHCI2R

A number of agents, especially those with a large percentage of hydrogen atoms, proved to be

flammable. As CF2H2, for example, was added to the oxidizer stream, an independent premixed
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 397

flame was established in the pyrex chimney. The location of the premixed flame in the chimney

depended on the oxidizer flow velocity which was related to the flame speed of the CF9H2-air

mixture.

One agent, NF3, behaved as an oxidizer, creating an intensely luminous flame with increasing

agent concentration. A simple thermodynamic analysis shows that stoichiometric combustion of NF3

in conjunction with a hydrocarbon fuel and air yields rather than demands O atoms. Thus, NF3 should

act as a hydrocarbon oxidizer.

A detailed discussion of the results for the additional agents can be found in Section 5 of this publication.

4.3.5 Discussion of Results. The heat capacities of the agents are approximately equal on a mass

basis as shown in Table 6. Table 7 and Figure 34 show that on a mass basis, CF3Br and CF3I were

very efl'ective compared to the other gaseous agents tested. The remainder of the agents exhibited

essentially the same suppression performance, which was not significantly more efficient than N-,. Some investigators have found that interpretation of the results are facilitated by flame temperature

calculations (Tucker et al, 1981; Sheinson et al., 1989).

Table 7 shows the results of such calculations assuming (a) that the agents behave as an inert, (b) that the combustion process is stoichiometric, with only H2O and COj as products, (c) that there is no

dissociation, and (d) that the agent and O2 diffrise into the flame at the same rate. Both of the inert non-halogenated agents, CO2 and N2, had adiabatic flame temperatures approximately 1587 °C. All

of the fluorinated agents had calculated adiabatic flame temperatures near 1590 °C, actually between

1557 and 1617 °C. Halon 1301 had a very high calculated temperature, suggesting an amount of

chemical suppression by this compound. The chlorinated agents, HCFC-22 and HCFC-124, also had

relatively high calculated temperatures compared to the inert agents and to the perfluorinated agents,

indicating some measure of chemical activity. From an energy balance it is possible to quantify the

percentage of physical and chemical action associated with each of these agents.

This interpretation of the data, however, is problematic when considering results for the inert monatomic gases (helium and argon). The high calculated flame temperature associated with helium

in Table 6 is likely due to large preferential diffiision or Lewis Number efl'ects. An interpretation of

Lewis Number efi'ects on coflow extinction studies has not been addressed. Understanding preferential diffiision effects could be an important parameter for predicting the effectiveness of a replacement

agent and understanding the relative importance of an agent's chemical activity to its transport

properties.

Figure 35 compares the mole fraction of agent needed for extinction of heptane flames burning in the cup (from Figure 33) and OFDF burners (from Figure 8). The relative agent suppression ranking

and even the absolute values of agent concentration at extinction show excellent agreement between

the low strain rate OFDF results and the cup burner results. It is conjectured that the strain rate near

extinction in the cup burner flame is characterized by a low strain rate. Table 8 shows the results of a statistical analysis considering all of the agents tested which

indicates that propane flames were the most difficult to extinguish, followed by heptane, the jet fuels,

and the hydraulic fluids.

4.3.6 Summary and Reconunendations. An apparatus and methodology were developed to test the

suppression eff'ectiveness of gaseous, liquid and solid powder agents in a simple, coflowing flame

burning gaseous and liquid fuels. The major findings were:
398

Table 7.

SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Agent Fraction in the Oxidizer Stream at Extinction of Heptane Cup Burner Flames

Agent Type Agent Mass Percent Volume Percent N2 31 32 CO2 32 23

inert He 6.0 31 Ar 38 41

nitrogen containing NF3 a a

silicon containing SiF4 36 13

sodium containing NaHCOj (10-20 fxm) 3.0

b

CF3H 25 12

c c

CF2H2/C2HF5 30 15 HFC CH2FCF3 29 10 CHF2CF3 29 8.7 CF3CH2CF3 27 6.5 C3HF7 28 6.2 CF4 37 16 C2F6 30 8.1 C3F6 29 7.3 C3F8 30 6.3 FC

C-C4F8 32 6.3

C4F1Q 32 5.3 CHF2CI 28 12 CHCUF 32 11 CH3CF2CI c c

CF2 = CHC1 c c

chlorine containing CF2 = CFC1 31 10 CHFCICF3 26 7.0 CF3Br 14 3.1

CFjBrz (!) 16 2.6 CHjBrCFj (!) 17 3.5 bromine containing CH2 = CHBr

c c

CF2 = CFBr 27 6.3 CF2 = CHBr 24

' 6.0

iodine containing CF3I 18 3,2

acted as an oxidizer, promoted flame stability solid powder not expressed in volume percent agent observed to be flammable
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 399

a>N

'xO

35

30 -

25 -

c

o) 20

<

c 15^

a>u

a.E

O

>

1 0 -

5 -

H Cup Burner

Countertlow (a = 25 s"^ ) B Countertlow (a = 180s'^)

T- O

O -r- m ^

CD6

00

C\J

CO O O

U_ LL

CM6

u.X

CO

CVJ 6

U-X

CM6

oX

in

CM6

LLX

o

in

C\J

C\J CO

I O

CM

Figure 35. Comparison between the volume percent of agent required to suppress heptane

flames in the OFDF and the cup burner.
400

Table 8.

SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Relative Difficulty in Extinguishing Different Fuels in the Cup Burner

Fuel Resistance to Extinction

Propane 1.00

Heptane 0.87

JP-8 0.84

JP-5 0.82

5606 0.76

83282 0.74

1 . The measurements showed that a mass-based agent ranking of agent suppression effectiveness

significantly differed from a mole based ranking. On a mass basis the EC, HFC, and HCFC

compounds had essentially the same suppression effectiveness, whereas on a mole basis the

higher molecular weight compounds were more effective.

2. Propane flames were most difficult to extinguish (required highest agent concentration), followed

by heptane, the jet fuels, and the hydraulic fluids.

3. The relative extinction effectiveness of the agents was essentially unchanged for all fuels tested.

4. NaHC03 powder was the most effective of all agents tested on a mass basis.

5. Several agents (CF2H2, CH3CE2CI, CE2 = CHC1, and CH2 = CHBr) behaved like fuels, as characterized by the appearance of an independently stabilized premixed flame inside the burner.

Another agent (NF3) behaved like an oxidizer.

6. CE3I and the bromine containing compounds were generally very effective on both a mole and a mass basis.

7. The relative rankings of the effectiveness of various gaseous agents in extinguishing the opposed

flow diffusion flame burner at low strain rates agree with the cup-burner studies. In addition, at low strain rates, the absolute agent concentration at extinction was found to agree well with

measurements made in the cup-burner. Although it is very likely that flame stretch in the cup

burner is similar to the low strain rate generated in the OEDE, a systematic study to show this correspondence should be undertaken.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 401

4.4 Turbulent Spray Flames

4.4.1 Background. A fuel spray provides a unique opportunity for a fire. Small droplets quickly

evaporate and the momentum from the spray efficiently entrains the air necessary for combustion. A

nipmred high pressure fuel, lubricant or hydraulic fluid line can supply a steady flow of fuel for a

fire stabilized behind obstacles in the engine nacelle, or create a cloud of droplets from a punctured

fuel tank adjacent to a dry bay. Extinguishment of the burning spray will occur when the critical level of agent is mixed with the air that gets entrained into the primary reaction zone. The process is affected by the velocity of the flow, the rate of mixing, the system temperature, and the agent and

fuel concentrations and properties.

Neither the cup burner nor the OFDF can simulate the combined droplet evaporation and

turbulent mixing which occur in a realistic spray flame. The objective of this portion of the

investigation was to develop a laboratory facility which could capture some of the additional critical elements that could exist within an engine nacelle or dry bay fire. The apparatus is useful as a screening device only if the flame and agent injection process can be precisely controlled and operated

in a reliable and repeatable manner. In addition to the flexibility offered by a reduced scale facility,

the laboratory spray burner provides an opportunity to separate the effects of agent transport and

vaporization from flame suppression eff'ectiveness, and to control independently the rate at which the

agent enters the flame.

No standard laboratory setup for evaluating the effectiveness of an agent for extinguishing an

aircraft in-flight spray fire existed when this study was initiated. The design described in the next

section was based upon the desire to maintain flame stability over a reasonable dynamic range, and to accommodate the application of gaseous and powdered agents. A coaxial jet with a spray nozzle

typical of a simple oil furnace or gas turbine combustor was chosen for the spray flame apparatus.

4.4.2 Experimental Facility. The spray burner facility consisted of an air delivery system, a fuel delivery system, an agent injection system, and a combustion zone, as laid out in Figure 36. Figure

37 shows a cross-sectional view through the combustion zone. Air at atmospheric pressure co-flowed

around a 6 mm diameter fuel tube within a 0.5 m long, 50 mm diameter stainless steel pipe. A pyrex

tube with a 65 mm inner diameter, supported on a Teflon ring, contained the flame for 75 mm

beyond the outer steel casing. The fuel was injected along the centerline through a pressure-jet nozzle

(Delevan model 0.5-45-B) that formed a 45° solid-cone spray. The exit of the nozzle was flush with

the open end of the surrounding pipe. The flame was stabilized on a 35 mm diameter steel disk

attached to the nozzle body.

The air was supplied by a high capacity compressor at 800 kPa, and its flow was monitored with

a sonic-orifice. Temperatures from ambient to in excess of 150 °C, measured with a type-K thermocouple 1.1m upstream of the fuel outlet, could be obtained using a 60 kW electric air heater. Average velocities across the burner cross-sections reached 50 m/s. The fuel (JP-8 or hydraulic fluid MIL-H-83282C) was stored in an 18 liter tank and delivered to the burner with an electric pump at pressures controllable to 1.0 MPa-g. The nozzle was designed to deliver fuel oil at a rate of 0.5 ml/s

when the fuel line gauge pressure is 687 kPa. The flow was calibrated at the actual operating

temperatures and pressures for each of the fuels. The gaseous agents were injected impulsively into the air 0.54 m upstream of the nozzle.

Uniform dispersion across the air stream was enhanced by injecting the gas in a radial direction into a reduced diameter (25 mm) section of the air pipe through two 6 mm diameter tubes. Screens with

50% open area were placed 40 and 80 mm downstream of the injection point to ensure complete

mixing between the air and agent prior to encountering the flame zone.
402 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

AIR INLET

i

AIR

COMPRESSOR AIR

HEATER

AGENT

INLET

METERING

VALVE

T

^

ORIEICE

METER

AGENT

NJECTOR

SOLENOID

VALVE

COMPUTER

Figure 36. Block diagram of turbulent spray burner facility.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 403

AGENT

MIXING

SCREEN

GLASS

TUBE

NOZZLE

I FUEL

THERMOCOUPLE

FLAME

Figure 37. Cross-section of combustion zone of spray burner facility.
404 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

A photograph of the overall facility is shown in Figure 38. The injection mechanism, shown in Figure 39 (a), consisted of the agent supply bottle connected

to a stainless steel storage vessel through a metering valve, and to the burner through a computer

controlled solenoid valve. The storage volume, including the one liter pressure vessel and associated

plumbing, was 1040 + 10 ml and the agent pressure was adjustable up to 687 kPa-g. The agent

temperature and pressure in the storage vessel were measured with a type-K thermocouple and a pressure transducer located upstream of the solenoid valve. The amount of injected agent was

controlled by varying the initial pressure and the time that the valve was open. The controlling

computer acquired the digital signal from the pressure transducer at 700 Hz. The actual mass

delivered was determined from the difference between the initial and final pressures in the storage

vessel using the Redlich-Kwong equation of state (Van Wylen and Sonntag, 1978).

Figure 39 (b) shows how the gaseous injection system was modified to accommodate powders.

The powder was loaded into two nylon tees downstream of the computer-controlled solenoid valve. Straight-through ball valves isolated the powder from the burner to minimize back-flow and powder

loss during loading. Aluminum caps were machined to reduce any excess volume within the storage

tees. Compressed air stored in the agent vessel was used to propel the powder into the burner.

4.4.3 Experimental Parameters and Procedures. The independent parameters which were

controlled in the spray burner facility were the air flow and temperature, the fuel flow and type, the

agent composition, and the mode of injection (fixed time or fixed pressure). For powder agents there were two additional independent parameters: the particle size distribution and the transport gas

pressure. The primary dependent parameters were the mass and the rate of injection required for suppression. Table 9 lists the range of independent parameters investigated, and other relevant

operating conditions derived therefrom.

The air and JP-8 fuel flows were varied to ascertain how flame stability was afl'ected by the

operating conditions. The fuel pressure was fixed at 687 + 10 kPa-g (corresponding to a mass flow

rate of around 0.42 g/s). The spray was ignited without the air flowing using a propane torch. The

flame extended well beyond the exit of the pyrex tube and was highly luminous under these conditions. As the air flow was increased, the flame attached itself to the stabilizing disc and the plume

length decreased until the flame was stationed within the horizontal tube passage. At high air flows

little soot radiation was observed beyond the exit plane, although the flame itself maintained some

luminosity. A moderate amount of soot formed on the nozzle face in a matter of minutes. A stable flame was sustainable until the air flow rate exceeded 73 g/s. The average gas velocity across the air duct was about 33 m/s at the above-mentioned mass flow, which translated to an estimated residence

time in the recirculation zone behind the stabilizing disc of 5 ms. Blowout experiments were repeated

for fuel nozzle pressures of 515 and 858 kPa-g. At the reduced pressure the fuel flow rate decreased

by about 14%, which resulted in an equivalent decrease in the amount of air necessary to extinguish

the flame. The higher nozzle pressure had no appreciable afi'ect on the blowout limit. The operating

conditions chosen for baseline measurements were a mass flow of JP-8 equal to 0.42 g/s and an

ambient temperature and pressure air flow of 33 g/s, resulting in an average inlet air velocity of

14 m/s. This produced an 18 kW flame with an overall equivalence ratio of about 0.17.

4.4.3.1 Gaseous Agents. The injection system under idealized conditions (incompressible flow,

massless valves, no pressure losses) was designed to deliver a square-wave pulse of agent to the

burner for the amount of time programmed by the computer controller. The actual flow deviated

substantially from this scenario. There was a 15 ms delay between when the solenoid was triggered

and the flow of agent actually began. When the valve started to close, pressure waves were created

which reverberated in the injection system at the acoustic velocity, causing the flow rate to modulate.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 405

Figure 38. Photograph of turbulent spray burner facility.
406 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

METERING VALVE T ^

PRESSURE ^

o-r^BA,..= TRANSDUCER

STORAGE . VESSEL S(

AGENT

VESSEL

THERMOCOUPLE

VACUUM GAUGE

VACUUM

PUMP

VALVE

(a)

^ BURNER

SOLENOID

VALVE

COMPUTER

CONTROL

COMPRESSED

METERING VALVE

POWDER

STORAGE TEES

PRESSURE

TRANSDUCER

STORAGE . VESSEL i(

BURNER

THERMOCOUPLE

(b)

BALL VALVES

SOLENOID

VALVE

COMPUTER

CONTROL

Figure 39. Schematic of agent injection mechanism for (a) gases, and (b) powders.
SECTION 4. FL.\ME SUPPRESSION EFFECTIVENESS

Table 9. Operating conditions in turbulent spray burner

407

Fuel

Parameter JP-8 MIL-H-83282C

Air Temperature 19 - 25 °C 146 - 153 °C 118 - 123 "C

Air Pressure ambient ambient ambient

Air Flow ' 22 to 73 g/s 33 + 1 g/s 30 + 1 g/s

Nominal Air Velocity 10 to 33 m/s 20 m/s 18 m/s

Fuel Temperature 19 - 25 °C < 150 °C < 120 °C

Fuel Pressure 515 to 858 kPa-g 1034±10 kPa-g 896 ±10 kPa-g

Fuel Flow 0.43 to 0.55 ml/s 0.46±.05 ml/s 0.65 ±.05 ml/s

Thermal Input 15 to 20 kW 16±2 kW a

Equivalence Ratio 0.14 to 0.19 0.15

a

Gaseous Agents core 11, N2. CFjBr. CF3I core 11. N2, CFjBr core 1 1 (except FC-31-10), N,. CFjBr

Powder NaHCOj + silica (0-10 ixm. +50 ^m)none none

Agent Temperature 20 - 25 "C 20 - 25 °C 20 - 25 °C

Agent Pressure 20 - 687 kPa-g 20 - 60 kPa-g 20 - 60 kPa-g

Injection Interval 20 to 900 ms 75 + 10 ms 75 ±10 ms

^ molecular composition and enthalpy of formation not provided

An example of these pressure modulations and the injection timing sequence is seen in Figure 40.

The curve labeled P2 is the voltage from the pressure transducer between the solenoid valve and the

metering orifice and PI relates to the pressure upstream of the orifice. The mass of agent, m. contained in the vessel of volume V, at pressure P and temperature T, was

determined from the Redlich-Kwong equation of state,

m = [

^-^^ -] RT l-bm/V l+bm/V

M is the molecular weight of the gaseous agent. R is the gas constant, and a and b are constants dependent upon the critical properties of the agent (Van Wylen and Sonntag, 1978). The initial temperature was measured, and the final temperature was determined by assuming that the expansion

occurred isentropically following the relation
408 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

(5)

Gamma is the specific heat ratio for the gas. Thus, by measuring the change in pressure, Equation (4) can be used in an iterative fashion to determine the total amount of mass injected into the

burner. Equation (5) assumes the gas is ideal. (From Equation (4) the deviation from ideal gas

behavior was found to be a maximum of 7% for the gaseous agents.) The pressure data were

collected at a rate of 700 Hz, with the initial and final conditions found from the average of at least 100 points measured one-half second prior to the release of the agent, and one second after the

solenoid valve closed, respectively. Figure 41 presents an example under minimum suppressing

conditions of the calculated mass delivered as a function of time using Equations (4) and (5) for six

different trials at the same computer setting.

A figure of merit for extinguishing the flame, B, can be defined in terms of the mass flows of

agent, w,, and air, m^-/.

A small value for fi is desirable. The mass of agent added to the flame is determined from Equation

(4), and the actual time interval of agent injection into the burner is estimated from the linear portion

of the slope in the mass-time curve, {e.g., see Figure 40).

The storage vessel was pressurized with air to 687 kPa-g, and the solenoid valve was opened for successively longer periods of time, up to a maximum of 910 ms. Figure 42 presents the cumulative

increase in mass delivery with set time. The results indicate that the mass of air (solid symbols)

increased almost linearly with time in the range between about 25 and 250 ms. For short time

settings, less mass was delivered because the valve did not have sufficient time to fully open. The

deviation from linear behavior when the valve was open for a long period resulted from depletion of

the entire mass ( = 10 g) contained in the injector storage vessel. Figure 42 also presents the time

over which the pressure decreased in the vessel (open symbols). There was a minimum time for the

solenoid valve to respond due to inertia, which explains the greater measured time for settings below

25 ms.A number of experiments were carried out with the burner operating at baseline conditions and

with air as the extinguishing agent to ensure that the flame could not be suppressed simply by blowing

it out. When air was injected into the burner, the flame was observed to fluctuate momentarily, but was never extinguished even when the storage pressure and injection period were at their maximum

values (viz., 687 kPa-g and 910 ms, respectively).

The protocol used in the experiments with the gaseous agents was to ignite the fuel spray and set the air flow to the desired level. The flame was allowed to burn for several minutes to ensure steady

operation. If the air was at an elevated temperature, it was necessary to wait until the temperature of

the air at the burner stabilized, which took as long as 20 minutes in some cases. The storage vessel was evacuated and then flushed several times with the agent under investigation to purge contaminating gases from the system. The pressure in the vessel was adjusted with the solenoid valve closed

using the inlet metering valve. Initially a pressure was chosen which was expected to be insuflftcient to extinguish the flame. The computer control/data acquisition was begun and the response of the

flame to the injection process was observed. If the flame was not extinguished, the pressure in the

agent vessel was increased and the experiment was repeated immediately. Eventually a pressure was

m.

(6)

m+mAir
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 409

Figure 40. Pressure during injection of nitrogen; 6.4 mm metering orifice, time - 23
410 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Time, seconds

Figure 41. Mass delivered to burner as a function of time; HFC-236fa/iiydraulic fluid flame.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 411

1 000

750 -

500

250

0

-1—I—I—I—I—I—I—I—I—I—I—I—1—I—I—I—I—

I I I I I

1 T

' r " r I I 1 I I T"

-'O

.-O

0

/O

_i I I I 1 I I I I I J I I I I I J I i_ 0

o

• mass

O time

. I I 1 I I I I I t I 250 500

Set Time, ms

750 1000

Figure 42. Injection period (open circles) and mass of air delivered (solid circles) as a function

of set time. Initial pressure is 687 kPa-g.
412 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

found which was sufficient to suppress the flame. This procedure was normally repeated four times

for each agent.

4.4.3.2 Powder Agents. Two powder samples and two different pressures were used to determine the effects of particle size and pressure on suppression performance. NaHC03 POwder

was mixed with silica gel and classified by size as described earlier in Section 4. The two distributions used were 0-10 ixm and -1-50 ixm. Two air pressures were selected for powder delivery, 172

and 241 kPa-g. Lower pressures were avoided because of poor delivery performance in preliminary

testing.

The storage vessel was filled to the desired air pressure with the solenoid valve closed using the

inlet metering valve. With both downstream ball valves closed, a preweighed amount of powder,

expected to be insufficient to extinguish the flame, was loaded into the storage tees (see Figure 39

(b)). Once loading was complete, both ball valves were opened, the computer control/data acquisition was begun, and the flame response to the injection process was observed. The nominal injection

period was set for 65 ms in all the experiments presented here. If the flame was not extinguished,

any powder accumulation within the injection mechanism and the burner was removed. Total

removal was accomplished by using a minimum of five high pressure air injections ( > 240 kPa-g)

and by increasing the burner air flow to at least 60 g/s for approximately 20 minutes. Preliminary

testing showed that this procedure was necessary to properly clean the injection mechanism. The

storage vessel was again pressurized to the same injection pressure, a larger preweighed amount of

powder was loaded into the storage tees, and the experiment was repeated. Eventually an amount of

powder was found which was sufficient to suppress the flame. This procedure was normally repeated

three times for each powder sample and pressure. Once the powder suppression mass was determined the delivery tubes were disconnected from

the burner and a series of tests were run to determine powder injection efficiency. This was done by

loading a preweighed amount of powder into the storage tees and injecting it into a one liter plastic

bag. The injection was performed according to the previously discussed experimental procedure.

The plastic bag was weighed before and after the injection to determine the weight gain. Five air injections at 240 kPa-g were also performed in a different plastic bag to determine the amount of

residue left in the injection mechanism after the first injection. This procedure was repeated a minimum of five times to determine reproducibility. The percentage of powder injected was

determine by dividing the plastic bag weight gain by the preweighed amount of powder loaded into the injection mechanism. The percentage of powder residue was calculated in the same manner, using

the plastic weight gain from the five high pressure injections. The results are summarized in Table

10. Powder losses within the burner itself were not quantified; however, inspection of the burner

walls and mixing screens showed almost no powder accumulation beyond a very thin layer of dust.

4.4.4 Experimental Results

4.4.4.1 Characterization of Facility. The influence of air velocity, injection period, and

injection pressure on the amount of N2 required to extinguish the JP-8 spray flame and on the value

of B was investigated as a means to validate the operation of the experimental facility. The storage

vessel was pressurized with nitrogen to 113 kPa-g and the turbulent burner set to baseline conditions

(ambient temperature, air flow equal to 33 g/s). The fuel flow was kept constant at 0.42 g/s. The

injection interval was increased one millisecond at a time until the flame was extinguished. Flame

extinguishment occurred between 23 and 26 ms for five different runs, delivering an average of 0.33

+ 0.03 gnitrogen at a mean flow rate of 11.2 ± 0.5 g/s. From Equation (6), the figure of merit
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 413

0.8

0.6

c

o

0.4

CO

COD

0.2

1 1

-

—11 1 r 7 1

1

JP8/N2

I I I T

1

1

1 1 1 1 1 1 1 1 1 1 1

1 1

n

(6.4 mm orifice)

-

Extinction

No Extinction

1 1 i 1 1 1 1 1

1 1 1 1 1 1 1 1 1 1 1 1 1 1 J —1

0 50 100 150 200

Initial Pressure, kPa

250

Figure 43. EflFect of vessel pressure on extinguishment of JP-8 spray flame by nitrogen. Solid

symbols indicate extinction; open circles indicate no extinction.
414 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 10. Powder injection efficiency

Particle

Size (jum)

Powder Injected (%) Powder Residue (%) Air Injection Pressure

(kPa)

0-10 81.2 ± 3.0 17.8 + 3.4 172

0-10 88.8 + 1.2 10.7 + 1.7 240

+ 50 82.3 + 3.4 17.9 + 4.4 172 . + 50 93.8 + 1.5 7.5 + 0.1 240

varied between 0.24 and 0.26. This value compares to a figure of merit for nitrogen of 0.28

measured in the cup burner apparatus with JP-8 as the fuel

.

Additional experiments were carried out for air flows of 44 g/s and 22 g/s. The high velocity

(19 m/s) required an average total nitrogen mass of 0.29 g at a rate of 11.5 g/s. The amount of

nitrogen required to extinguish the lower air velocity flame was 0.32 g, with an average flow of

10.7 g/s. In this case, doubling the air flow reduced the mass of nitrogen required by 10%. The

high air flow case yielded fi = 0.21 and the low air flow case yielded B = 0.33.

A series of experiments was carried out with the air and fuel flow rates at baseline conditions,

and the time interval was fixed as the nitrogen pressure was increased. Figure 43 is a plot of the mass of nitrogen delivered to the burner as a function of initial vessel pressure for an injection

interval set to 65 ms. The open squares indicate that the flame remained lit; the solid squares

correspond to successful extinguishments. The minimum vessel pressure necessary to extinguish the

flame 100% of the time was 167 kPa-g. The amount of nitrogen delivered at this pressure and time

interval was 0.58 + .03 g. The injection time interval had an eff'ect on the minimum amount of nitrogen required to extinguish the flame. The open squares plotted in Figure 44 illustrate this effect. For these experiments, the pressure was fixed and the injection time interval was gradually increased until extinction

occurred. The minimum mass of nitrogen was about 0.32 g, for a set injection period of 23 ms. Reducing the set time to 6 ms had no impact on the amount of nitrogen required to quench the flame

because, as shown in Figure 42, the actual period of injection did not change appreciably. Injection

times longer than 23 ms lead to delivering greater amounts of N2, with more than three times as much N2 required when the injection time was set to 260 ms. A limit was reached at long time

intervals where the transient mass addition was insufficient to extinguish the flame.

Nitrogen was allowed to flow continuously in one experiment, with the rate increasing until the

flame was extinguished. The figure of merit was found from Equation (6) to be 0.11, and is indicated by the continuous flow arrow in Figure 44. This compares to a value of 0.28 found in the

cup burner with the same fuel/agent combination. Less nitrogen was required for extinguishment of

the spray flame because of the greater turbulence levels and reduced time available for the combustion

to occur. The solid squares also plotted in the figure are values of 6 that correspond to the different

injection time intervals. As the time was shortened, 13 increased, reaching a limiting value of about

0.28. (It is a coincidence that this value is identical to the value of B measured in the cup burner.)
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 415

0.4

0.3

0)

o

to

COO

0.2

0.1

0.0

1 1 1

'

1 T I r -T 1 1—1—

r

Beta

1 ' ' ' .

- - Mass

J

-

-

-

;

Continuous

1 1 1(1111

Flow ^-

-

1 1 1 1 1

Linnit '.

1 ...

10 100

Tinne, ms

Figure 44. EfTect of injection time interval on mass of nitrogen (open squares) required to

extinguish JP-8 spray flame; solid squares: /3.
416 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

If the value of B were the sole criterion for evaluating an extinguishing strategy, one would chose

to inject the agent over an extended period of time. However, as seen in Figure 44, this has the

undesirable effect of greatly increasing the amount of agent required to put out the flame. For an

agent that is to be used in a transient manner, the total mass must also be considered. This is distinct

from the quasi-steady state measurements taken with the cup burner apparatus, for which B is a reasonable measure of performance for a total flooding agent.

4.4.4.2 Gaseous Agents/JP-8/Ambient Air Test Series. A fixed injection time of 65 ms was

chosen to compare the performance of the alternative agents. This value is intermediate between the

estimated residence time in the spray flame (5 ms) and a typical time interval for injection in an actual

engine nacelle (500 ms). At this setting, an initial pressure of 167 kPa-g was necessary to extinguish

the flame with nitrogen. Nitrogen is considered chemically inert during suppression, and has a much

lower molecular weight than halon 1301 and the alternative gaseous agents that were evaluated. As a

result, the initial pressure in the vessel required for flame extinction was much less for the alternative

agents. The 6.4 mm flow-metering orifice was therefore replaced with a 4.8 mm orifice to increase

the required pressure and reduce the per cent uncertainty in the pressure change measurement.

Extinguishment experiments were performed using halon 1301 to establish a performance reference

(Grosshandler et al., 1993). An average over five experiments led to a required initial CFgBr

pressure of 24 kPa-g to extinguish the flame, which translates to a mass of 0.44 + .04 g and a (8 of

0.15.

The repeatability of the experimental technique is demonstrated in Figure 41, which is a plot of

the mass of agent delivered to the burner as a function of time for six diff'erent trials where the mass

of the agent, HFC-236fa, was just above the extinction threshold. The average mass injected was

0.78 g with a range of ±0.02 g. The initial pressure needed to cause extinction was 52 + 2 kPa-g,

and the calculated injection interval was 80 + 8 ms, as compared to the set time of 65 ms. The

difl'erence between the average injection interval and the computer setting was slightly larger than that found using nitrogen. This was attributed to the lower initial pressure used in the HFC-236fa

experiments, which was less effective in causing the solenoid valve to close. The results of the ambient temperature air/JP-8 flame suppression experiments for all of the

gaseous chemicals examined are summarized in Table 11. The CF3I required the least mass (0.54 g)

and had a figure of merit of |3 = 0.16, close to that of CF3Br. Nitrogen was only slightly less

effective. Of the core agents, HCFC-22 required the least (0.65 g) and FC-31-10 required the most

(1.00 g) material to extinguish the flame. FC-31-10 had the poorest figure of merit, 0.27.

4.4.4.3 Effect of Air Temperature and Fuel Type. The air was preheated to 150 °C ( + 3/

-5 °C) and the experiments were repeated with all of the gaseous agents except for CF3I (Vazquez et

al., 1994). The increase in temperature affected the flame stability in several ways. First, since the mass flow of air remained fixed, the average velocity across the air duct increased about 50% because

of the drop in density. The JP-8 in the line was also heated as it flowed through the heated air annulus. The higher temperatures and lower fuel density required the fuel pressure to be increased to deliver the same amount. However, even at a maximum fuel line pressure of 1.03 MPa-g, the mass

of fuel was only 90% of the ambient temperature condition, resulting in a slightly leaner flame. The

increase in air velocity and decrease in equivalence ratio destabilized the flame; but this was counteracted by the increase in enthalpy of the reactants due to the higher initial temperature.

As seen in Table 1 1 , increasing the temperature, on average, increased the amount of alternative

agent required to suppress the JP-8 spray flame by 0.04 g. The halon 1301 remained the most

effective, but in relative and absolute terms, required the largest increase in mass of all the chemicals

investigated. The nitrogen remained better than the other halogenated compounds, and FC-31-10
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 1 1 . Amount of agent required to suppress turbulent spray flames

417

JP-8,

Tair = 20 °C

JP-8,

Tair=150 °C

Hydraulic Fluid

Tair-120 °C

Agent mass /3 mass mass

llalUll 1 D\J i 0.44 0.15 0.53 0.19 0.44 0.16

0.54 0.16

a a a a

111 Li U^Cli 0.58 0.18 0.63 0.19 0.42 0.16

HrFr-97 0.65 0.20 0.70 0.23 0.70 0.24

T-IFr-195 0.73 0.22 0.77 0.24 0.78 0.26

0.74 0.22 0.75 0.22 0.70 0.23

FP-I 1 6 0.75 0.22 0.74 0.23 0.73 0.23

0.76 0.24 0.78 0.23 0.79 0.26

0.78 0.23 0.84 0.25 0.78 0.24

HFC-227 0.80 0.24 0.81 0.24 0.82 0.26

HFC-32/125 0.81 0.24 0.89 0.26 0.82 0.25

FC-218 0.89 0.24 0.87 0.25 0.86 0.28

FC-318 0.97 0.25 0.99 0.26 1.08 0.30

FC-31-10 1.00 0.27 1.02 0.25 a a

^ not measured

remained the least eff'ective. It can be speculated that the relatively poorer behavior of the halon 1301

is attributable to the decreased residence time of the agent in the flame, such that the bromine has less opportunity to scavenge the chain carrying radicals. A flame could not be stabilized using Mil-H-83282C hydraulic fluid and ambient temperature air. By increasing the fuel volume flow by 27% over the JP-8 flame and the air temperature to 120 °C

(±3 °C), sufficient stability could be maintained. A bluish appearance of the hydraulic fluid spray

flame suggested that less soot was being formed.

There was little diflTerence in the amount of halogenated agent necessary to suppress the hydraulic

fluid spray when compared to the JP-8 flames. (See Table 11, and note that neither FC-31-10 nor CF3I was tested with the hydraulic fluid.) In particular, the amount of halon 1301 was identical to the unheated jet fuel experiments. About 10% more FC-318 was used to suppress the hydraulic fluid. By contrast, 28% less nitrogen extinguished the hydraulic fluid flame. An explanation for this unique

behavior is not available.
418 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

4.4.4.4 Sodium Bicarbonate Powder Experiments. There was a definite particle size effect on

the efficiency of the powder as a fire suppression agent. Table 12 summarizes the results. The

smallest particle size powder was 2.4 to 3.0 times more effective in its fire suppression capability than

the large particle size powder. There was no significant difference in performance created by

changing injection pressure in the small size range. On the other hand, increasing pressure increased

the effectiveness of the large particle size by more than 20%. The (3 values indicate a very effective

agent for small particles, and a rather average agent for the large particle size. High speed movies of injection of the 0-10 fxm particles showed what appears to have been a uniform powder cloud passing through the burner in about 80 ms. This compared to an injection

interval equal to 75 ms based upon the recorded air pressure in the storage vessel. The photographs

also showed that flame extinction happened within the first 50 ms from the time the powder reached

the flame. This time was independent of injection pressure, and was close to the 40 ms estimated

from high speed photographs of HCFC-22 suppressing a similar flame.

4.4.5 Discussion of Results. If the agents are ranked according to the mass required to inhibit the

flame, the order does not change by more than plus or minus one position for the three series of

experiments. The exception is HFC- 125, which drops two positions in both of the high temperature

tests. Expressing the results in terms of the flame suppression number (FSN) is a convenient way to compare the performance of the different agents in the three series of experiments. The FSN is defined as the mass of agent relative to the mass of halon 1301 used to suppress the equivalent flame.

Figure 45 summarizes the results in this fashion, using (a) ambient temperature air and JP-8 (open

bars), (b) 150 °C air and JP-8 (cross-hatch bars), and (c) 120 °C air and hydraulic fluid (solid bars).

The superior performance of the 0-10 /xm NaHC03 powder is evident in the figure.

A flame suppression number also can be defined in terms of 13, the mass fraction of agent in the

flowing mixture, with /3 for halon 1301 used to normalize the results. The FSN computed on a mass

flow basis can be seen in Figure 46. For the ambient temperature JP-8 flame, the cup burner mass

flow FSN is about 2.0 ( + 0.15) for the various agents, which is about 0.5 greater than the FSNs

plotted in Figure 46. The implication is that it takes a significantly lower mass fraction of agent to extinguish the turbulent spray flame than the cup burner flame. The cup burner was not operated at elevated temperatures, but one series of tests were run using ambient temperature hydraulic fluid. The same trend was observed when the cup and turbulent spray extinguishing mass fractions were compared: a measurable decrease in agent concentration was required for the turbulent spray flame

(in spite of the fact that the higher reactant temperatures were expected to stabilize the spray flame,

everything else being equal).

Minimizing the storage volume on board the aircraft is as critical as minimizing the mass of

agent. There are a number of densities which one could use to convert the required mass of agent to the required volume, including the dispersed gas density and the density of the saturated liquid agent

at ambient temperature. The former scales with molecular weight. The latter density also has

practical significance. It provides a logical conversion from the mass required to the storage volume

because the saturated liquid condition at ambient temperature is close to the condition maintained

when the fire bottle is filled (assuming negligible solubility of the pressurizing gas). The volume

factor, VF, is defined as the volume of the agent, computed with either of the above densities,

normalized by the equivalent volume of halon 1301. Figure 47 compares the dispersed volumes.

Figure 48 displays the volume factors under storage conditions. Nitrogen requires a storage volume

36 times that of halon 1301, and is off the scale in Figure 48 because it does not liquify under typical

bottle conditions. The FC-116 also rates poorly on volume factor because its critical temperature is less than the ambient. In the figure, the density of FC-116 was calculated at 20 °C and 4.1 MPa, a

typical bottle pressure. The remainder of the agents have storage volume factors between 1.5 and
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Table 12. Sodium bicarbonate powder results

419

Particle Air Injection Powder Mass, g /3 Size, Pressure, kPa

0-10 172 0.20 0.08

0-10 240 0.19 0.07

+ 50 172 0.59 0.20

+ 50 240 0.46 0.16

2.5, depending on the agent, fuel and temperature. Of these, the HFC-32/125 mixture has the highest

volume factor, and HFC-227, HCFC-124 and HCFC-22 have the lowest. The powdered agent was

not compared on a volume basis since the volume depends upon the pressurizing method as much as on the volume of the powder.

4.4.6 Conclusions from Turbulent Spray Burner Study. The turbulent spray burner has been

found to be suitable for comparing the performance of gaseous and fine powder extinguishing agents

in transient operation. The facility is not overly sensitive to the air or fuel flows, and the agent

delivery system is able to control accurately the injection period between 20 and 910 ms. The following conclusions can be made regarding the ability of difi'erent agents to extinguish the spray flame:

1 . The mass fraction, total mass, and minimum volume of agent required to extinguish a given

flame must all be considered when ranking the performance of diff'erent fire fighting agents.

2. Of the chemicals evaluated in the turbulent spray burner, NaHC03 was the only compound more

effective than halon 1301. CF3I required the least mass and volume of the gaseous agents to extinguish the flame, followed by nitrogen (FSN only) and the chlorinated agents, HCFC-22 and

124.

3. The larger perfluorocarbons, FC-31-10 and 318, required the greatest mass. On a volume basis,

nitrogen was the poorest performer. The mixture of HFC-32 and HFC- 125 had the largest

volume factor of the condensible agents.

4. No statistically significant difl'erence in agent performance was found between the room temperature JP-8 and hydraulic fluid flame testing, indicating little fuel effect.

5. The majority of the agents required slightly more mass to extinguish the higher temperature JP-8

flame, indicating a small temperature effect. This trend is not completely unexpected since a higher enthalpy flame is normally more stable. However, the temperature eff'ect did not alter the

ranking on agent performance, with the exception of HFC- 125.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

3.0

2.5

2.0

1 .5

1 .0

0.5

JP8, 20 C

JP8, 150 C

Hydr. Fluid

UJo

o

0.0 1.

(O

o

o

<

X

in

<N

CM

ro

I o

in

CM

I o

oo

CM

I o

Ll.

CM

CM

I o

Li_ (J

m

o

O

Li_ X

Li_ O

o

CM

CM

I u

CM

I O

Li_ o

oo

ro

I O

ro

I n O

toro

CM

I O

roO

X

D

Agents, in increasing boiling point order

Figure 45. Flame suppression number (FSN) computed on a relative mass basis.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 421

3.0

2.5

2.0

1 .5

1 .0

0.5

0.0

jpa, 20 c

JP8, 150 C

Hydr. Fluid

CO

I o

c

CPO

m

o

in

CNI

CM

I o

in

CMT

oo

CNJ

I o

CM

CM

I o

O

X

T

o

D

CM

CM

I o

00m

I o

Li_ O

CM

I O

CJ

m

I o

COro

CM

I O

m

o

o

Agents, in increasing boiling point order

Figure 46. Agent flame suppression number (FSN) based upon relative rate of mass injected.
422 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Agents, in increasing boiling point order

Figure 47. Dispersed gas volume factor (based upon 101 kPa and 22 °C air).
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 423

5.0

4.5

4.0

3.5 [-

•n 3.0

2.5

2.0

1 .5

1 .0

0.5

0.0

1 1 1 r —Nitrogen

CD

o

c

oX

in

CN

CN

I o

X

u

Li_ X

CN

(N

I O

|j_ o

X

CO

I o

ro

I O

Li_ O

JP8. 20 C

JP8. 150 C

Hydr. Fluid

o

0)

CM

I O

CNO

Li_ O

00 o

i 7

o

COtn

CN

I o

Agents, in increasing boiling point order

Figure 48. Storage volume factor of different agents based upon saturated liquid conditions at 20 °C.
424 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

6. The core gaseous agents all performed better in the turbulent spray burner relative to halon 1301

than was predicted from cup burner measurements. Generally speaking, about twice as much

mass and volume were needed to suppress the spray flame using the core alternative agents when

compared to halon 1301.

4.5 High Speed Premixed Flames and Quasi-detonations

4.5.1 Introduction. An anti-aircraft device entering a dry bay could lead to a situation in which a vaporizing fuel spray produces a combustible mixture which is then ignited by a glowing fragment. If the space were confined, the pressure would increase behind the reaction front, accelerating the

flame. A transition to turbulence would likely occur as the flame encounters clutter in the dry bay.

If the ventilation is insufficient to relieve the pressure build up, the possibility of a supersonic

detonation would exist, leading to destructive over-pressures in the dry bay.

A supersonic flame is distinct from the flames simulated in the cup, OFDF and spray burners.

As a result, the efl'ectiveness of an agent in preventing a detonation depends upon diff'erent chemical

and physical mechanisms. A shock wave precedes the supersonic flame. Obstructions in the flow

promote intense mixing of the fresh reactants with the combustion products and cause the pressure waves to interact with the mixing region. Given enough distance, the flame can accelerate dramatically, increasing the temperature of the reaction zone behind the shock and further adding to the heat

release rate. Depending upon the geometric details, the wave can approach its theoretical ChapmanJouguet velocity and accompanying high pressure ratio. Even a slight variation in composition of the

reactants near the limit of detonation can cause a dramatic change in the wave velocity and cause

destructive pressures to be attained.

There exists an extensive literature describing the kinetics and dynamics of flame/shock wave

systems formed within classical detonation tubes (e.g., Lefebvre et ai, 1992; Nettleton, 1987; Lee,

1984; Baker et ai, 1983; Westbrook, 1982). Chapman and Wheeler (1926) were the first to note

that a methane/air flame could be accelerated to a terminal velocity in a shorter distance within a

circular tube by placing obstacles into the flow. Lee et al., (1984) built on this observation to study

quasi-detonations in hydrogen/air and hydrocarbon/air mixtures. A quasi-detonation propagates more

slowly than a true detonation due to pressure losses in the flow, but its structure is more complex

than a true detonation, and the mechanism of its propagation is not fully understood. Although an

obstructed flow is more difficult to analyze than the flow in a smooth-walled tube, the complex

arrangement was chosen for the current study because it more closely simulates a potentially

damaging condition in the dry bay. The desire to rapidly suppress a flame and the associated

pressure build up in such a situation is the primary motivation behind this study.

4.5.2 Experimental Facility. The effectiveness of a fire fighting agent in suppressing a high speed,

premixed flame or quasi-detonation can be rated by the extent to which it decelerates the propagating wave and simultaneously attenuates the hazardous shock which is always ahead of the ffame. Because

the fire extinguishant is unlikely to be released prior to the establishment of a turbulent flame, the

traditional experiment in which the flame inhibitor is premixed with the fuel and air prior to ignition

{e.g.. Das, 1986) does not replicate the chemistry critical to the actual situation. Each dry bay on an

aircraft has a different geometry, and the release of the agent once a fire is detected is highly

variable. Heinonen et al., (1991) injected suppressant into a chamber shortly following the ignition

of a fuel spray in air, but had difficulty controlling the mixing and in duplicating the process. The
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 425

complexities and biases associated with the fluid dynamics of release can be avoided by premixing the

agent with the fuel and air in a portion of the mbe distinct from where the flame is initiated. The two-section, deflagration/detonation tube shown in Figures 49 and 50 was designed to produce the desired environment for both the flame initiation and flame suppression regimes. A

repeatable, uninhibited mrbulent flame was fully established in the driver section, the design of which

was based directly upon the work of Peraldi et al., (1986). They found that a 50 mm inner diameter

mbe with a blockage ratio of 0.43 could be used to create repeatable, high-speed flames and quasidetonations within the first several meters of an 18m tube. By varying the equivalence ratio of

ethene/air mixmres from 0.5 to 2.1, they were able to produce flame velocities between about 600

and 1300 m/s.

The driver section was 5 m long and was equipped at the closed end with a spark plug. This

section was filled with the combustible mixture of ethene and air. The gas handling system consisted

of a vacuum pumping network; pressurized gas cylinders for the fuel, oxidizer and agent; and a dual

circulating pump. The ignition energy was delivered in a microexplosion of a tin droplet shortcircuiting the tips of nichrome electrodes connected to an 80 V power supply. Spiral-shaped obstructions made of 6.4 mm stainless steel rods with a pitch equal to the inner diameter of the tube were

inserted into the tube, to produce an area blockage ratio of 44%, close to the value shown by Lee

et al. (1984) to promote a quasi-detonation in their facility. The second section of the deflagration/detonation tube contained the gaseous agent along with the same fuel/air mixture used in the

driver section. The diameter was the same and its length was 2.5 m. An identical spiral insert was

used to maintain a high level of mixing. The two sections were separated from each other by a 50

mm inner diameter, stainless steel, high vacuum gate valve, which remained closed until just before

ignition. Pressure transducers and photodiodes were located along the test section to monitor the strength

and speed of the combustion wave. Their output was recorded on a multi-channel, digital storage

oscilloscope. The pressure transducer mounting is depicted in Figure 51 and a schematic of the

photodiode mounting is shown in Figure 52. A logic diagram of the photodiode sequencing system is shown in Figure 53.

4.5.3 Operating Procedure. The whole system was evacuated to 10'^ Pa before filling the two

sections separately with the desired mixtures, which were attained through the method of static partial pressures. The fuel/air ratio and total pressures were held constant across the gate valve. The initial temperamre was the ambient value, 22 + 3 °C. The oxidizer used in all experiments was breathing

grade air. Ethene (CP grade, 99.5% volume purity) was chosen as the fuel because it had been

demonstrated (Lee et al., 1984) that subsonic flames, quasi-detonations, and full detonations all could

be obtained in a tube of this geometry simply by varying the stoichiometry. After filling, the gases were homogenized independently using a double, spark-free circulating pump, recirculating the entire

tube volume a total of 20 times. The mixtures were left for five minutes to become quiescent. About

ten seconds prior to ignition, the gate valve was opened manually.

After ignition, the flame propagated into the driver section and accelerated quickly due to the

intense mrbulence created by the interactions of the flow with the obstacles, generating a shock wave

ahead of it. After passing through the open gate valve the flame/shock system encountered the same

combustible mixture and a certain amount of agent in the test section. A rendering of the flame/shock

system passing through the gate valve is shown in Figure 54. Depending on the concentration of the

agent, the flame was or was not extinguished and the pressure wave attenuated.

Experiments using 5 % ethene in air mixtures were run under a variety of conditions to assess the

operation of the system. Figure 55 is an example of the pressure trace when the N2 partial pressure

is (a) insufficient to extinguish the flame radiation, and (b) sufficient for suppression of an air/ethene
426 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

C

r

VP VP

IS DS V

o A

TS

4 5

SCOPE

DS - Driver Section, TS - Test Section, V - High Vacuum Gate Valve,

IS - Ignition Source, CP - Dual Circulating Pump, VP - Rotary Vacuiom

Pump, E - Exhaust, F - Fuel, O - Oxidizer, A - Agent, 1 - Triggering

Transducer, 2,3 - Piezo-electric Pressure Transducers, 4,5 - Fast Photodiodes, SCOPE - Collects Four Pressure and Visible Radiation Signals

Figure 49. Schematic diagram of detonation/deflagration tube facility.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 427

Figure 50. Photograph of detonation/deflagration tube facility.

I

t
428 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

1 - Low-noise Cable, 2 - Cable Connector, 3 - Piezo-electric

Pressiire Transducer, 4 - Stub, 5 - Tube, 6 - Retaining Nut,

7 - Holder, 8 - Seals, 9 - Seal Weld

Figure 51. Pressure transducer mounting.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 429

1 - Fast Photodiode, 2 - BNC Cable, 3 - Quartz Window,

4 - "O" Ring, 5 - Seal Weld, 6 - Plexiglas Cover, 7 - Barrel,

8 - Holder, 9 - Washer, 10 - Stub, 11 - Tube

Figure 52. Photodiode mounting.
430 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

555IC Pulse Input

+5V

DAC Test Circuit

ti t. h u t, k t,

[> Stage 1

1/4 7404IC

Input Enable

'O'-l, '1'- Stage X

Remote Module Located on Tube

+ 15V +5V

0

Output

Stage 2

Stage 3

Stage 4

Stage 5

Stage 6

Stage 7 +5V

1/2

Tri-State Buffer

1/2

Digital to Analog

Converter

15V

Oscilloscope

Data

Acquisition

Computer

Figure 53. Logic diagram of photodiode sequencing system.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 431

Figure 54. Rendering of high-speed turbulent flame entering test section.
432 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

a)

1% vol

At 273.6 [IS ^ 3.663 k Kz

t

t .

r<l 1

1 MM 1 1 1 1 1 1 1 1

b)

10% vol. C4F10

At 1.638 RS kt 963.4 Hz

Figure 55. Pressure traces comparing (a) non-suppressed and (b) suppressed flames using

FC-31-10.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 433

mixture at a total pressure of 100 kPa. Knowing the distance between the two pressure transducers,

the shock speed was determined from the time lag between the pressure rises. The pressure ratio was

tabulated from the initial pressure in the tube and either the first peak or the maximum pressure

increase. The sensitivity of the flame speed and pressure ratio to the voltage of the ignition system,

the mixing time of the components before ignition, the presence or absence of the gate valve, the

speed of opening the gate valve, and cleaning the tube between runs were all investigated.

Experiments were conducted with 100% nitrogen in the test section, a 5% ethene in air mixture

in the driver section, and the total pressure equal to 20, 50 and 100 kPa. The incident shock wave

velocity measured 2.2 m beyond the gate valve and 0.3 m from the end of the tube was 420 + 8 m/s

at all three total pressures. The pressure ratio based upon the initial pressure rise was 2.5 ± 0.5, and

about 3.0 + 1.0 based upon the peak increase.

As can be seen in Figure 56, no significant changes in shock speed occurred for a partial

pressure fraction (which is approximately equal to the mole fraction) of nitrogen in air greater than

40%. The region of no change was extended down to 30% for the 20 kPa case. When no suppressing nitrogen was added to the test section the shock velocities attained values of 780 and 1 170 m/s,

respectively, for the 20 kPa and 100 kPa experiments. The pressure ratio based upon the initial rise

is plotted in Figure 57. Pj/Pq increased dramatically at the same point as the velocity when the

partial pressure of nitrogen was decreased, reaching maxima of 26 and 18, respectively, for the 20

kPa and 100 kPa initial conditions when no nitrogen was added to the ethene/air mixture. The peak

pressure ratio, which normally did not correspond to the initial pressure pulse, exceeded 50: 1 for the

20 kPa and 50 kPa experiments. The 100 kPa experiments generated peak pressure ratios around

30:1.

The results of the preliminary parameter assessment led to an experimental protocol which

yielded flame speeds which were reproducible from run to run within about + 2%. Pressures downstream of the Shockwave had a higher variability (+ 20%) because of the complex shock

structures created by interactions with the spiral rod inserts.

4.5.5 Results and Analysis. Mach number and pressure ratio were the two dependent parameters

which were measured as a means to characterize the extent of flame suppression. The Mach number

was based upon the time it took for the pressure wave to travel the distance between the two pressure

transducers, normalized by the sonic velocity of the reactant gases in the test section. The pressure

ratio was evaluated from the average amplitude of the first pressure pulse to be recorded by each

transducer, normalized by the initial pressure. Consecutive pressure jumps occurred, as can be seen

in Figure 55, indicating that localized explosions in the mixing region between the spirals were

present. Individual runs were concluded before the shock wave reflected from the end plate arrived

back at the pressure transducers. The distance between the leading shock wave and the flame front was measured in some of the experiments from the time lag between the photodiode and pressure

transducer response at the same location. The primary independent variables were agent type and

concentration. The fuel/air equivalence ratio and the system absolute pressure were also varied for a number of agent experiments.

Table 13 summarizes the results with no suppressant in the test section, and compares the Mach

number, pressure ratio, and flame spacing to the situation in a "fully suppressed" flame, in which

pure nitrogen was located in the test section. Note that neither the Mach number nor pressure ratio go to zero for the fully suppressed case because total dissipation of the shock wave would require a

substantially longer tube. On the other hand, the flame radiation disappears entirely; hence, the

infinite separation distance.

The bromine atom in CF3Br is known to inhibit laminar flames by scavenging active radicals

from the chain-branching reactions. Experiments using halon 1301 were run to compare to an
434 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 56. Effect of initial pressure on the reduction of shock wave velocity by nitrogen.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 435

Figure 57. Effect of initial pressure on the reduction in pressure ratio by nitrogen.
436

Table 13.

SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Summary of flame parameters for ethene/air mixtures under suppression extremes

Amount of Initial Equivalence Mach Number Pressure Mame/Shock

N2 added to Pressure Ratio Ratio Separation

Test Section Distance, mm

0% 20 kPa 0.75 2.4 + 0.1 26 ± 1 a

50 kPa 0.75 2.6 ± 0.1 26 + 1 a

100 kPa 0.75 3.4 + 0.1 18 + 2 3 + 1

100 kPa 1.00 4.1 + 0.1 26 + 2 2 ± 1

100 kPa 1.25 4.5 ± 0.1 35 ± 3

a

100% 20 kPa 0.75 1.25 ± 0.05 2.5 ± 0.5 00

50 kPa 0.75 1.25 ± 0.05 2.0 + 0.5 00

100 kPa 0.75 1.25 + 0.05 2.5 + 0.5 00

100 kPa 1.00 1.3 ± 0.1 3.5 + 0.5 00

^ quantity not measured

inerting agent like N2, and to determine the suitability of the facility for assessing the effectiveness of

a wide range of agents for suppressing high speed turbulent flames and detonations. Figures 58 and

59 show the shock Mach numbers and the respective pressure ratios measured at three difi'erent total pressures as a function of the partial pressure fraction of CF3Br in the test section. The largest eff'ect of total pressure occurred between 50 and 100 kPa for halon, compared to between 20 and 50 kPa for N2. Halon 1301 suppressed the flame at a partial pressure fraction of 10% to the same extent as if the test section had been completely filled with nitrogen. An unusual behavior occurred in the

100 kPa experiments when the concentration was between 2% and 3%. Both the Mach number and

pressure ratio increased with the amount of CF3Br, followed by the expected decrease for large

concentrations. The reversal, while small, was greater than the uncertainty in the data. Halocarbons, unlike nitrogen, are known to promote the production of soot. To determine the

sensitivity of the shock velocity to soot contamination, a series of experiments with a 5 % ethene/air

mixture at a total pressure of 20 kPa was run for CF3Br partial pressure fractions between 0 and 10%

with and without cleaning the tube and spirals in the test section. The performance was similar, with

the maximum deviation between the cleaned and uncleaned results amounting to less than 35 m/s.

4.5.5.1 Lean Mixtures at 100 kPa. Attenuation of the shock speed and pressure increase by

the core agents and CF3I was measured with the ethene/air equivalence ratio fixed at 0.75 (5.0% by

volume C2H4), an initial absolute pressure of 100 kPa, and an initial temperature of 22 °C. It was

found that the amplitude and speed of the pressure wave, and the speed of the trailing flame, were all strongly dependent on the agent type and concentration. The flame always followed the shock wave

in such a way that both speeds were equal. However, when the amount of the agent in the mixture was increased, the distance between the shock and flame increased as well, up to around 100 mm as
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 437

Figure 58. Effect of initial pressure on the reduction in shock wave Mach number by CF3Br.
Figure 59. Effect of initial pressure on the reduction in pressure ratio by CF3Br.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 439

full suppression was approached. At the extinguishing concentration the radiation disappeared, which

indicated the absence of the flame. In that situation the pressure wave amplitude was attenuated by a factor of eight and the wave speed by a factor of three, similar to the results for nitrogen (see

Table 13).

The results for all the alternative agents are compared in Figures 60-63 . The amount of agent is expressed both as a mass fraction and as a partial pressure fraction. The Mach numbers and pressure

ratios at zero represent the pure combustible mixture with no flame suppressing agent present. One

can see that the concentrations necessary for total extinguishment for all the compounds are between

40 and 50% by mass. However, at low concentrations the Mach numbers and pressure ratios are higher for some agents than even the value for the pure combustible mixture. Because of the sheer

number of data, it is instructive to examine the results by the class of compound; i.e., FCs, HFCs,

HCFCs and the IFC (CF3I).

The FCs, as a class, were generally the best performers on both a partial pressure and mass

fraction basis. In fact, FC-116 was superior to CF3Br, and FC-318 was about equivalent. FC-218

and FC-31-10 slightly enhanced the pressure ratio in low concentration. This can be seen in Figures

64 and 65. FC-218 reduced the pressure ratio to less than 5:1 at a mass fraction of 0.29, which was

better than with CF3Br.

Adding hydrogen to the molecule had a significant effect on the performance of the HFCs,

shown in Figures 66 and 67. The HFC-32/125 mixture produced peak pressures more than double

the value for no suppressant. The Mach number was increased to its highest value of 4.1 when the mass fraction was 11 %. It wasn't until the mass fraction exceeded 38% that the HFC-32/125 mixture

became as eff'ective as nitrogen in reducing the speed and pressure build-up of the shock wave. The

two fluoropropanes, HFC-227ea and HFC-236fa produced the lowest pressure build-up of the HFCs

and did a good job of suppressing the shock speed.

The chlorine atom in the two HCFCs created an additional complexity because chlorine is a strong oxidizer. From Figure 68, one can see that HCFC-22 was the least effective on a mass basis

of all of the agents in fully suppressing the combustion wave Mach number, requiring a mass fraction

of over 50% in the test section. The maximum pressure ratio for HCFC-124 was 32:1 at a 23% mass

fraction (see Figure 69), which was exceeded only by HFC- 125 and the HFC-32/125 mixture (Figure

67).

The one IFC tested was CF3I. The Mach number and pressure build-up, shown in Figure 70, were cut about in half with partial pressure fractions in the test section of only 15 to 20%. None of

the other chemicals, including CF3Br, was able to accomplish that. Unfortunately, when the mass

fraction was increased to 30 % , the Mach number shot back up and the pressure ratio attained a value

of 21:1, which was greater than when no CF3I was present. This reversal, which was slight in the

bromine-containing halon 1301, changed what at first appeared to be the most eff'ective suppressing

agent into one of the least effective agents. It is known that iodine atoms can cause a catalytic effect

in some reactions by lowering the overall activation energy. At intermediate concentrations the

possibility also exists that the iodine (and bromine) recombined, reducing their impact on the

combustion process. It is not out of the question that in the cases under consideration both of these

effects became important.

4.5.5.2 Stoichiometric Mixtures. Changing the fuel/air mixture (/.^., the equivalence ratio)

changes the flame temperature and radical concentration significantly. Figure 71 demonstrates the

impact of equivalence ratio on the shock Mach number and pressure ratio with no agent in the test section. To see if the relative performance of the agents was dependent upon the fuel/air ratio a number of experiments were run under stoichiometric conditions. The equivalence ratio was based

upon the amount of ethene in the mixture and was not adjusted to account for the contribution of the
440 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

0 10 20 30 40 50

AGENT MASS FRACTION/ %

Figure 60. Reduction in shock Mach number as a function of mass fraction for all agents tested.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 441

Figure 61. Reduction in shock Mach number as function of partial pressure for all agents

tested.
442 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 62. Reduction in pressure ratio as a function of mass fraction for all agents tested.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 443

Figure 63. Reduction in pressure ratio as a function of partial pressure for all agents tested.
444 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

T

J 1 I I 1 1 I I \ 1 U 1 0 10 20 30 40 50

FC MASS FRACTION/ %

Figure 64. Reduction in shock Mach number vs mass fraction of FCs compared to CF3Br.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 445

Figure 65 . Reduction in pressure ratio as function of FC mass fraction compared to CF3Br.
446 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

A C

o c

y c H F

3 2 6 o c gHF^

V CF.Br

0 0 20 30

HFC MASS FRACTION/ %

40 50

Figure 66. Reduction in shock Mach number as function of HFC mass fraction, compared to CFgBr.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 447

Figure 67. Reduction in pressure ratio as a function of liydrofluorocarbon mass fraction,

compared to CF3Br.
448 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

0 10 20 30 40 50

HCFC MASS FRACTION/ %

Figure 68. Reduction in shock Mach number as a function of hydrochlorofluorocarbon mass

fraction, compared to CF3Br.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 449

HCFC MASS FRACTION/ %

Figure 69. Reduction in pressure ratio as a function of hydrochlorofluorocarbon mass fraction,

compared to CF3Br.
450 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 70. Reduction in shock wave Mach number and pressure ratio as a function of CF3I

mass fraction, compared to CF3Br.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 451

Figure 71 . Effect of fuel/air equivalence ratio on shock wave Mach number and pressure ratio.
452 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

agent to the fuel or oxidizer pool . The initial pressure and temperature remained the same for the

stoichiometric experiments: 100 kPa and 22 °C.

The results for the FCs are plotted in Figures 72-74. In all cases, the Mach number and

pressure ratio increased as the fuel/air mixture was changed from lean to stoichiometric. The shapes

of the curves, in the most part, were not significantly altered. The impact was greatest at low

concentrations, decreasing as the partial pressure fraction reached 8%. For FC-116, the biggest

difference occurred at a partial pressure fraction of 3 %

.

A distinctly different behavior occurred with the two HFCs shown in Figures 75 and 76. The

hydrogen atoms attached to the agent molecules had less of an enhancing effect under stoichiometric

conditions. The over-pressure was greatly reduced when the equivalence ratio was increased, leading

to a cross-over condition where both the Mach number and pressure ratio became less for the

stoichiometric condition. The implication of these data is that HFC-125 and the HFC-32/125 mixmre

increase in relative effectiveness under stoichiometric conditions.

Trifluoroiodomethane behaved no differently in suppressing the shock speed and pressure ratio under lean and stoichiometric conditions. From the data in Figure 77 it can be seen that the Mach

number and pressure ratio increased uniformly over the range of partial pressure fractions investigated

at an equivalence ratio of 1.0. The reversal of suppression effectiveness at partial pressure fractions

between 3% and 6% observed at an equivalence ratio of 0.75 also occurred for stoichiometric

combustion.

4.5.5.3 Ranking the Agents. The results which were gathered indicated the complexity of the

suppression process in the deflagration/detonation tube. Because one does not know a priori the

conditions in an actual dry bay fire zone, and because different initial conditions {i.e., pressure, temperature and fuel/air ratio) affect the amount of agent required for suppression to varying degrees,

a specific set of initial conditions was chosen at which all the agents were compared: 100 kPa,

22 °C, and 0.75 ethene/air equivalence ratio. The reduction in pressure ratio rather than the Mach

number was chosen as the measure of suppression because of its direct impact on the structural

integrity of the dry bay. The amount of halon 1301 required to reduce the pressure (a) to 10% and

(b) to one-half the maximum increase was used to normalize the results. Figure 78 displays three different performance parameters calculated under these conditions. The

flame suppression number (FSN) is defined as the mass fraction of an agent required for suppression

divided by the required mass fraction of halon 1301. The volume factor (VF) is defined as the

storage volume of the alternative agent necessary for suppression divided by the storage volume of the

CF3Br. The saturated liquid density of the agents at 20 °C was used to convert mass to volume.

Halon 1301 has a VF and FSN of unity by definition. For the other chemicals, the smaller these

values the better is the agent.

While there were some reversals depending upon the basis of evaluation (i.e., 90% reduction

FSN, 50% reduction FSN, or VF), Figure 78 shows that FC-218 was clearly the best performer

under the conditions tested, and that the HFC-32/125 mixture and HCFC-22 were the worst. The

FSN of FC-116 based upon a 50% reduction in pressure build-up was the lowest of all agents, even performing better than halon 1301. However, because hexafluoroethane does not condense at room

temperature its density is low, leading to a volume factor which is undesirably high. Nitrogen, which

is also a gas at room temperature, had a VF of 32. On a mass basis nitrogen performed almost as well as 1301. The CF3I had the third lowest volume factor of the alternative agents, although its FSN

values were close to the highest.

4.5.6 Conclusions and Recommendations from the Premixed Deflagration/Detonation Study. It

is necessary to emphasize that the experimental conditions in the deflagration/detonation tube differed
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 453

J ^ I , I < I , \ I [ 1 \ I Q

0 2 4 6 8 10 12

C2F6 PARTIAL PRESSURE FRACTION/ %

Figure 72. Effect of equivalence ratio on the reduction in Mach number by partial pressure of

FC-116.
454 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

C3F8 PARTIAL PRESSURE FRACTION/ %

Figure 73 . Effect of equivalence ratio on the reduction in shock Mach number and pressure

ratio by partial pressure of FC-218.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 455

Figure 74. Effect of equivalence ratio on the reduction in shock Mach number and pressure

ratio by partial pressure of FC-31-10.
456 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 75. EflFect of equivalence ratio on tlie reduction in shock wave Mach number and

pressure ratio by partial pressure of HFC- 125.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 457

Figure 76. Effect of equivalence ratio on the reduction of shock wave Mach number and

pressure ratio by partial pressure of HFC-32/125 mixture.
458 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Figure 77. Effect of equivalence ratio on the reduction of shock wave Mach number and

pressure ratio by partial pressure of CF3I.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 459

Figure 78. Flame suppression numbers and volume factors of the agents examined in the

detonation/deflagration tube.
460 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

significantly from those used in the cup burner, opposed-flow diffusion flame burner and turbulent

spray flame burner. The main qualitative difl'erence was the occurrence of a strong shock wave ahead

of the flame. That wave influenced the gasdynamic, thermodynamic and chemical state of the pure

combustible mixture in the driver section and the mixture containing an agent in the test section.

Another feature was a high level of turbulence within the flame due to its high speed and the

interactions with the spiral obstruction. The quantitative diff'erence was a supersonic regime (relative to the undisturbed mixture) of flame/shock propagation and strong pressure changes (due to confinement and shock) during the process. Thus the oxidizer, fuel and agent molecules underwent preliminary processes before entering the flame zone.

4.5.6.1 Conclusions. The following conclusions can be drawn based on the results obtained:

1 . Suppression of highly dynamic flames can be eff^ectively studied in the two-sectional tube,

permitting clear discrimination of performance among various alternative extinguishing agents.

2. The high-speed turbulent flame and the flame in the quasi-detonation wave under suppression

strictly follows the shock wave which is always ahead of it in such a way that its velocity is the same as the shock velocity. The distance between the flame and the shock increases with the

amount of an extinguishing agent. At extinguishment the flame disappears while the residual

shock still exists.

3. The suppression process of the lean ethene/air mixture is pressure dependent both for chemically

inert nitrogen and chemically active halon 1301. That may be related to the fact that the

oxidation mechanism of ethene is known to be pressure dependent.

4. The analysis of the suppression data for the lean ethene/air mixture leads to the division of the

alternative compounds into four general categories. These are, in order of decreasing efl'ective- ness: CF3Br and CF3I; perfluorocarbons; hydrofluorocarbons; and hydrochlorofluorocarbons.

Within each of the categories the agents can be ordered (on a mole fraction basis) approximately

inversely according to the molecular weight of the molecule.

5. The extinguishing concentrations of the more effective agents are around 10% by volume and

40% by weight, while less eff'ective ones are 15-30% by volume and 40-50% by weight. The

least eff'ective agent is the HFC-32/125 mixture, giving unusually high pressure ratios. HCFC-22

requires the highest extinguishing concentration of ail the alternatives.

6. The presence of a hydrogen-containing suppressant in the combustible mixture results in a

significant increase in pressure ratio relative to that for the pure combustible mixture. The

phenomenon occurs also for the compounds not containing hydrogen atoms at relatively lower

concentrations but the impact is not so dramatic. The impact is generally weaker for stoichiometric relative to lean mixtures. That may be related to the fact that the agent acts as an extra fuel,

causing the mixture to be richer. It is suspected that the effect may be associated with the

release of hydrogen during the process which enhances the branching steps in the ethene

oxidation mechanism. It is also not out of the question that a homogeneous autocatalysis occurs lowering the overall activation energy and enhancing the oxidation mechanism.
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 461

7. The behavior of CF3I is different from the other agents. At lower concentrations (2-3% by

volume) the performance is the best of all the alternatives. However, at intermediate concentrations (3-7% by volume) the performance is worsened significantly. Eventually at higher

concentrations, up to the extinguishing value of 10%, the performance is comparable with the

perfluorocarbons and CF3Br. The phenomenon is independent of the equivalence ratio of the

combustible mixture. The behavior at the intermediate concentrations may be attributed to the

catalytic effect caused by the iodine atoms.

4.5.6.2 Recommendations. The relative ranking of the agents in Figure 78 can be used as a guide to choose a replacement for halon 1301 for suppressing high-speed turbulent flames and quasidetonations. The influence of rich equivalence ratios up to 1.5, of other gaseous fuels, and of various

tube geometries on the performance of the selected agents should be explored to determine the

generality of the rankings. A more fully instrumented apparatus would allow the evolution of the

flame and shock all along the length of the tube to be followed, and the key intermediate species and

products of combustion to be ascertained.

The unusual behavior of CF3I in lean and stoichiometric ethene/air mixtures indicates a critical need for further research into the kinetics of the compound. It is recommended that a shock tube be

designed to investigate the dynamic and kinetic characteristics under high pressure and temperature

conditions as well as various mixture compositions. The set-up should include a special dump

chamber for collecting combustion products and optical access for spectroscopic measurements of

important species. In general, the application of this unique experimental apparatus for studying fire suppression

processes under highly dynamic conditions should provide the information necessary to establish a physical/chemical model of the phenomena, contributing to the fire safety of facilities exposed to fast flames, explosions and detonations.

4.6 Summary of Conclusions on Flame Suppression Effectiveness

1. The relative ranking of the agents depends upon whether one uses a mass basis or a volume

basis.

2. The relative ranking of the agents varies only slightly among the three non-premixed burners,

but a substantially different ranking results from the premixed deflagration/detonation

apparatus.

3. The relative ranking of the agents is not much affected by the fuel type or air temperature.

4. The quantity of agent required to suppress a flame varies with the type of burner generally in the following order:

cup burner = low strain OFDF > spray burner > high strain OFDF > deflagration tube

5. The quantity of agent required to suppress a non-premixed flame varies somewhat with the

type of fuel in the following order:

propane > heptane > JP-8 > JP-5 > HF-5606 > HF-83282
462 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

6. The quantity of agent required to suppress the turbulent spray flame increases with decreasing

rate of agent injection.

7. The quantity of agent required to attenuate the shock wave speed and pressure ratio in the

deflagration/detonation tube varies with equivalence ratio, sometimes increasing as one goes

from lean to stoichiometric and sometimes decreasing.

8. Sodium bicarbonate was the most effective suppressant of the three non-premixed flames,

requiring less mass than halon 1301. In general, the smaller size particles were more

eff'ective.

9. lodotrifluoromethane was more eff'ective than halon 1301 in suppressing the spray burner

flame, had a similar eff'ectiveness with the cup burner flames, and was less efl'ective in attenuating the shock speed and pressure in the detonation/deflagration apparatus.

10. None of the core agents performed as well as halon 1301 in any of the four flame suppression

studies, requiring from approximately one and a half to four times the storage volume as halon 1301 to suppress the diff'erent flames.

11. HCFC-124 was the generally the best performer of the core agents in suppressing the non- premixed flames, and FC-218 was judged the best for attenuating the shock speed and

pressure in the detonation/deflagration tube.

12. FC-116 was the poorest performer (on a liquid storage volume basis) of all the core agents

evaluated in the flame suppression tests; the HFC-32/125 mixture was the second poorest

performer, and led to the highest over-pressures in the lean detonation/deflagration uibe

study.

4.7 References

Baker, W.E., Cox, P. A., Westine, P.S., Kulesz, J.J. and Strehlow, R.A., Explosion Hazards

and Evaluation, Elsevier, 1983.

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Fire and Flanimability 5, 255 (1974a).

Bajpai, S.N., "Extinction of vapor fed diflijsion flames by Halons 1301 and 1211-Part I," ser No. 22430, Factory Mutual research Corporation, Norwood, Massachusetts, November (1974b).

Bui, M. and Seshadri, K. "Comparisons between experimental measurements and numerical

calculations of the structure of heptane-air difliision flames," Combust. Sci. Tech. 79, No 4-6, 293

(1991).

Chapman, W.R. and Wheeler, R.N.V., "The Propagations of Flame in Mixtures of Methane and

Air. Part IV. The Eff'ect of Restrictions in the Path of the Flame," Journal of the Chemical Society,

London, 1926.

Chelliah, H., Bui-Pham, M., Seshadri, K. and Law, C. K., "Numerical description of the

structure of counterflow heptane-air flames using detailed and reduced chemistry with comparisons to experiments," Twenty-Fourth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, Pennsylvania, p. 851 (1992).
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 463

Creitz, EC, "Inhibition of Diffusion Flames by Methyl Bromide and Trifluoromethyl Bromide

Applied to the Fuel and Oxygen sides of the reaction zone," Journal of Research of the National

Bureau of Standards (U.S.) 65A, 3S9, (1961).

Creitz, E.C., "Extinction of Fires by Halogenated Compounds - A Suggested Mechanism," Fire

Technology 5, 131 1972.

Das, A., Relationship Between Ignition Source Strength and Design Inerting Concentration of

Halon 1301, M.S. Thesis, Fire Protection Engineering Department, Worcester Polytechnic Institute,

1986.

Dodding, R.A., Simmons, R.F. and Stephens, A., "The Extinction of Methane - Air Diffusion

Flames by Sodium Bicarbonate Powders," Combustion and Flame 15, 313 (1977).

Dow-Corning, Midlane, MI, Communication of Test Results, 1993.

Ewing, C.T., Faith, F.R., Hughes, J.T. and Carhart, H.W., "Flame Extinguishment Properties

of Dry Chemicals: Extinction Concentrations for Small Diffusion Pan Fires," Fire Technology 25,

134 (1989).

Fendell, F. E., "Ignition and extinction in combustion of initially unmixed reactants," J. Fluid

Mechanics 27, 281 (1965)

Galbraith Laboratories, Knoxville, TN, Communication of Test Results, 1993.

Gann, R.G., Barnes, J.D., Davis, S., Harris, J.S., Harris, R.H., Herron, J.T., Levin, B.C.,

Mopsik, F.I., Notarianni, K.A., Nyden, M.P., and Ricker, R.E., "Preliminary Screening Procedures

and Criteria for Replacements for Halons 1211 and 1301," National Institute of Standards and

Technology, NIST Technical Note 1278, U.S. Government Printing Office, Washington, DC (1990).

Gmurczyk, G., Grosshandler, W., Peltz, M. and Lowe, D., "A Facility for Assessing Suppression Effectiveness in High Speed Turbulent Flames," Eastern States Section Conference/The Combustion Institute, Princeton, 1993.

Gorden, S. and McBride, B.J., "Computer Program for the Calculation of Complex Chemical

Equilibrium Compositions, Rocket Performance, Incident and Reflected Shocks, and Chapman-Jouget

Detonations," NASA, NASA SP-273,(1976).

Goscoyne Laboratories, Baltimore, MD, Communication of Test Results, 1993.

Grosshandler, W.L., Lowe, D., Rinkinen, W., and Presser, C, "A Turbulent Spray Burner for Assessing Halon Alternative Fire Suppressants," ASME Winter Annual Meeting, December, 1993.

Hamins, A., The Structure and Extinction of Diffusion Flames, Ph.D. Thesis, University of

California, San Diego, 1985.

Hamins, A., and Seshadri, K., "Prediction of overall chemical kinetic rate parameters for extinction of diffusion flames above multicomponent fuels," Combust. Sci. Tech. 38, 89 (1984a).

Hamins, A., and Seshadri, K., "Structure of counterflow diffusion flames burning multicomponent fuels," Twentieth Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, Pennsylvania, p. 1905 (1984b).

Hamins, A., and Seshadri, K., "The influence of alcohols on the combustion of hydrocarbon

fuels in diffusion flames," Combust. Flame 64, 43 (1986).

Hamins, A., and Seshadri, K., "The structure of diffusion flames burning pure binary, and

ternary solutions of methanol, heptane, and toluene," Combust. Flame 68, 295 (1987).

Hamins, A., Thridandam, H., and Seshadri, K., "Structure and extinction of a partially premixed

diffusion flame," Chem. Eng. Sci. 40, 2027 (1985).

Heinonen, E.W, Kirst, J. A. and Moussa, N.A., Fire/Explosion Protection Characterization and

Optimization: Alternative Dry Bay Fire Suppression Agent Screening, Medium-Scale Test Results,

Final Report, NUMERI OC 91/16, Report JTCG/AS-91-VR-005, 1991.
464 SECTION 4. FLAME SUPPRESSION EFFECTIVENESS

Hirst, R. and Booth, K., "Measurements of Flame-Extinguishing Concentrations," Fire

Technology 13, No. 4, (1977).

Kent, J.H., and Williams, F.A., "Extinction of Laminar Diffusion Flames for Liquid Fuels,"

Fifteenth Symposium (International) on Combustion, pp. 315-325, The Combustion Institute, Pitts- burgh, 1974.

Krishnamurthy, L., Williams, F.A., and Seshadri, K., "Asymptotic Theory of Diffusion Flame

Extinction in the Stagnation-Point Boundary Layer," Combust. Flame 26, 363 (1976).

Lee, J.H.S., "Fast Flames and Detonations," Chemistry of Combustion Processes, The

American Chemical Society, 1984.

Lee, J.H., Knystautas, R. and Chan, C.K., "Turbulent Flame Propagation in Obstacle-filled

Tubes," Twentieth Symposium (International) on Combustion, The Combustion Instimte, 1984.

Lefebvre, M.H., Nzeyimana, E., Van Tiggelen, P.J., Progress in Astronautics and Aeronautics

,

AIAA, 1992.

Linan, A., "The Asymptotic Structure of Counterflow Diffusion Flames for Large Activation

Energies," Acta Astronautica 1, 1007 (1974).

Lum, L., National Institute of Standards and Technology, Gaithersburg, MD, Personal Communication, 1993.

Miller, E. and McMillion, L.G., "The Suppression of Opposed-Jet Methane-Air Flames by

Methyl Bromide," Combustion and Flame 84, 37 (1992).

Milne, T.A., Green, C.L., and Benson, D.K.,"The Use of the Counterflow Diffusion Flame in Studies of Inhibition Effectiveness of Gaseous and Powder Agents," Combustion and Flame 15, 255

(1970).

Mitani, T., "Flame Retardant Effects of CF3Br and NaHC03," Combustion and Flame 50, 111

(1983).

Nettleton, M.A., Gaseous Detonations, Their Nature, Effects and Control, Chapman and Hall,

1987.

Peraldi, O., Knystautas, R. and Lee, J.H., "Criteria for Transition to Detonation in Tubes,"

Twenty-first Symposium (International) on Combustion, The Combustion Institute, p. 1629, 1986.

Peters, N., "Local premixing due to flame stretch and nonpremixed turbulent combustion,"

Combust. Sci. Tech. 30, 1 (1982).

Peters, N., "Laminar diffusion flamelets in nonpremixed turbulent combustion," Prog. Energy

Combust. Sci. 70, 319 (1984).

Pitts, W.M., Nyden, M.R., Gann, R.G., Mallard, W.G. and Tsang, W, "Construction of an

Exploratory List of Chemicals to Initiate the Search for Halon Alternatives," NIST Technical Note

1279, U.S. Government Printing Oflfice, Washington, DC, 1990.

Reid, R.C., Prausnitz, J.M., and Sherwood, T.K., The Properties of Gases and Liquids, Third

Edition, pp. 404-411, McGraw/Hill, New York, 1977.

Seshadri, K., Mauss, F, Peters, N. and Warnatz, J., " A flamelet calculation of benzene

formation in coflowing laminar diffusion flames," The Twenty-Third Symposium (International) on

Combustion, The Combustion Institute, Pittsburgh, Pennsylvania, 463 (1990).

Seshadri, K., and Williams, FA., "Effect of CF3Br on counterflow combustion of liquid fuel with diluted oxygen," ACS Symposium Series, No. 16, Halogenated Fire Suppressants, R.G. Gann

(Ed.), American Chemical Society, Washington, DC, p. 149 (1975).

Seshadri, K., and Williams, FA., "Structure and extinction of counterflow diffusion flames

above condensed fuels: comparison between poly(methylmethacrylate) and its liquid monomer, both

burning in nitrogen-air mixtures," / Polymer Sci. 16, 1755 (1978a).
SECTION 4. FLAME SUPPRESSION EFFECTIVENESS 465

Seshadri, K.. and Williams , F. A., "Laminar flows between parallel plates with injection of a

reactant at high Reynolds number," Int. J. Heat Mass Transfer 21, 251 (1978b)

Sheinson, R.S., Penner-Hahn, J.E., and Indritz, D., "The Physical and Chemical Action of Fire

Suppressants," Fire Safety Journal 75, 43 (1989).

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Smooke, M.D., Puri, I.K., and Seshadri, K., "A comparison between numerical calculations and

experimental measurements of the structure of a counterflow diffusion flame burning diluted methane

in diluted air," Twenty-First Symposium (International) on Combustion, The Combustion Institute,

Pittsburgh, Pennsylvania, p. 1783 (1986).

Strasser, A., Liebman, I., and Kuchta, J.M., "Methane Flame Extinguishment with Layered

Halon or Carbon Dioxide," Fire Technology 10, 25 (1974).

Tucker, D.M., Drysdale, D.D. and Rasbash, D.J., "The Extinction of Diffijsion Flames Burning

in Various Oxygen Concentrations by Inert Gases and Bromotrifluormethane, " Combustion and Flame

41, 293 (1981).

Van Wylen, G.J., and Sonntag, R.E., Fundamentals of Classical Thermodynamics, Second

Edition, John Wiley and Sons, p. 400, 1978.

Vazquez, I., Grosshandler, W., Rinkinen, W., Glover, M., and Presser, C, "Suppression of

Elevated Temperature Hydraulic Fluid and JP-8 Spray Flames," Fourth International Symposium on

Fire Safety Science, Ottawa, June, 1994.

Westbrook, C.K., Nineteenth Symposium (International) on Combustion, The Combustion

Institute, p. 127, 1982.

Westbrook, C.K., "Chemical Kinetics of Hydrocarbon Oxidation in Gaseous Detonations,"

Combustion and Flame 46, 191, 1982.

Williams, F.A., "A Unified View of Fire Suppression," Journal of Fire and Flammability 5, 54

(1974).

Williams, F.A., Combustion Theory, Second Edition, Addison Wesley, 2nd Edition, (1985).

Williams, FA. and Peters, J.W., "Studies of Dry Powder Extinguishment of Difflision Flames

for Condensed Fuels," Final Technical Report, National Bureau of Standards, NBS Grant No.

NB81NADA, June 15, 1982.

467

5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

FOR ADDITIONAL FIRE FIGHTING CHEMICALS

Marc R. Nyden, Gregory T. Linteris,

Building and Fire Research Laboratory

Donald R.F. Burgess, Jr., Phillip R. Westmoreland, Wing Tsang and Michael R. Zachariah

Chemical Science and Technology Laboratory

Replacements for the current commercial halons should posses a diverse set of properties which are

rarely found together in the same molecule. Thus, the ideal candidate for the replacement of halon

1301 would be a nontoxic gas which is reactive in flames and in the troposphere, yet at the same

time, inert in the stratosphere and in its storage environment. The present generation of replacements, as typified by the core candidates listed in Section 1, were selected on the basis of a compromise, whereby fire suppression efficiency was sacrificed to ensure acceptable environmental

properties. The research reported in this section was directed at developing the capability to predict

the fire suppression efi'ectiveness, propensity to generate corrosive combustion products, and environmental impact of a molecule on the basis of its structure. This is essential to the development of a

rational approach to the search for new and more efi'ective fire fighting chemicals.

The immediate objective of the research described in Section 5.1 was to provide a chemical basis

for understanding the relative fire suppression efficiencies of the core candidates. In this pursuit, a

kinetic mechanism, accounting for the interactions of substituted fluoromethanes and fluorethanes in hydrocarbon flames, was developed and incorporated into a computer model. The expectation is that

fire scientists will some day be able to use this model to predict the performance of replacement

candidates, thereby reducing the time and expense involved in laboratory screening.

The fire suppression efficiencies of the core candidates are poor compared to halon 1301. This

implies that much higher concentrations will be required to put out a fire. The levels of corrosive

by-products, particularly HCl and HP, generated in the extinction process are likely to be considerably higher, as well. The empirical study summarized in Section 5.2 was undertaken to identify the

factors which determine the potential of an agent to form acid gases in a combustion environment so

that the worst off"enders can be eliminated from further consideration

Additional research was conducted to expand the scope of agent screening beyond the core

candidates. This eff'ort is described in Section 5.3. Structure/activity relationships (SARs) for ozone

depletion and global warming were tested and developed. These were used, in conjunction with

experimental measurements of fire suppression efficiency and considerations of the commercial

availability and suitability for total flooding, to identify the most promising candidates from an

exhaustive list of possibilities.

5.1 Kinetics of Fluorine-Inhibited Hydrocarbon Flames

5.1.1 Project Objectives. The major objective of this subtask was to provide a chemical basis for rationalizing the relative degree of eff'ectiveness of each candidate agent. A fundamental understanding of the chemistry of these agents in hydrocarbon flames should facilitate identification of desired

characteristics of effective agents. That is, utilization of simple chemical concepts should enable
468 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

screening and selection of potential agents with minimal time and human resources. In order to accomplish the objective of this subtask, it was necessary to develop a chemical mechanism based on

elementary reaction steps for their destruction, their participation in and influence on hydrocarbon

flame chemistry, as well as for prediction of potential by-products of incomplete combustion.

Unfortunately, neither such a mechanism nor a review of the relevant chemistry existed prior to this study. Consequently, a significant effort was required in order to "simply" construct such a comprehensive mechanism prior to its utilization in any simulations.

The focus of this task was restricted to the chemistry involving only fluoromethanes and

fluoroethanes. This includes both those candidate agents specifically being considered as replacements

(CH2F2, CF3-CH2F, CF3-CHF2, CF3-CF3), as well as for all of the other possible fluoromethanes

(CH3F, CHF3, CF4) and fluoroethanes (CH3-CH2F, CH3-CHF2, CH3-CF3, CH2F-CH2F, CH2FCHF2, CHF2-CHF2). The two chlorine-substituted agents (CHF2CI, CF3-CHF2CI) and the larger

fluorinated hydrocarbon agents (C3Fg, C3F7H, C4FJ0, cyclo-C4Hg) were not specifically considered.

Including these additional candidate agents would significantly increase the complexity of the

chemistry that must be considered. However, the eff'ectiveness of each can be estimated to some

degree by analogy to the other agents that were studied. This can be done by using the qualitative

trends observed and the fundamental understanding of the chemistry as developed as part of this subtask.

The complete set of fluoromethanes and fluoroethanes were studied for two basic reasons. First,

when the four specific candidate agents being considered decompose in the flame, they generate a pool of fluorinated hydrocarbon stable species and radicals, which results in the formation of many of

the other fluoromethanes and fluoroethanes. Consequently, in order to describe the decomposition of

the four specific agents (and resultant chemistry) adequately, it is necessary to describe the chemistry

of all of the many intermediates and products that are created, including most of the other fluoromethanes and fluoroethanes.

For example, the lowest energy and primary decomposition pathway for one of the candidate

agents, CF3-CF3, involves dissociation of the C-C bond to form (two) •CF3 radicals. These •CF3

radicals will then react with methyl radicals, •CH3, which are present in significant concentrations in hydrocarbon flames. This radical-radical combination reaction has two decomposition channels whose

relative importance is dependent upon temperature and pressure. One channel results in the formation

of a fluoroethylene, CH2 = CF2 (and HF by-product). The other channel results in the formation of

another fluoroethane, CH3-CF3. In order to predict the products in the flame correctly, the magnitude and rate of heat release in the flame, and ultimately the eff'ectiveness of the added agent, it is also necessary to describe correctly the decomposition of these two additional stable fluorinated

hydrocarbon species (CH3-CF3, CH2 = CF2). When the decomposition channels for these two

molecules are considered and for all of the other reaction channels for •CF3, as well as for all of the

relevant chemistry for the other three specific candidate agents, very quickly most of the possible

fluoromethanes and fluoroethanes must be considered.

Second, given that as part of the overall project there was, unfortunately, no directly related experimental task to these simulations, it was imperative to provide a level of self-consistency to this work by considering a range of modeling parameters, including different reactor/flame geometries,

diff'erent fuels, different (potential) agents, and many other conditions. In doing so, it allowed us to develop confidence in the validity of the qualitative trends that we observed. Quantitative prediction

of the absolute or even relative effectiveness of the specific agents will require benchmarking of the

simulations with experimental measurements.

5.1.2 Background. There has been a significant amount of work over many years investigating the

eff^ectiveness of halogenated fire suppressants (Ellis, 1948; Simmons and Wolfhard, 1955; Rosser
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 469

et al., 1958; Friedman and Levy, 1963; Fenimore and Jones, 1963; Halpem, 1966; McHale et al,

1971; Day etal, 1971; Creitz, 1972; Biordi etal, 1973; Larsen, 1974; Brown, 1975; Gann, 1975;

Larsen, 1975; Skinner, 1975; Biordi et al, 1976; Biordi et al, 1978; Fristrom and Van Tiggelen,

1979; Dixon-Lewis, 1979; Westbrook, 1980; Safieh et al, 1982; Westbrook, 1983; Vandooren et al,

1988; Masri, 1992; Morris et al, 1992), as well as other types of fire suppressants (Jorrissen and

Meuwissen, 1924; Jorissen et al, 1932; Lask and Wagner, 1962; Ibiricu and Gaydon, 1964;

Fristrom, 1967; McHale, 1969; Creitz, 1970; Cotton and Jenkins, 1971; Hastie, 1973; Jensen and

Jones, 1978; Jensen and Jones, 1982). We will not review this body of work, but refer the reader to those relevant sources. However, in the following sections we will provide a brief overview of the

most relevant work relating to the agents being considered as part of this project. For the most part,

this means the chemistry of CF3Br flame suppression. We will also describe the general and specific

physical and chemical processes that are most important in influencing flame suppression and other

effects {e.g., ignition, promotion, by-products) in hydrocarbon flames. This will include those factors specifically related to the candidate agents.

The chemistr>' of CF3Br and its role in flame suppression has been reviewed in detail by others. A number of these works are referenced in the prior section. A large part of our work is based on

the pioneering work in this area by Biordi and coworkers (Biordi et al, 1978) and Westbrook (1983).

In earlier experiments on a range of candidates, CF3Br was identified as being very effective for extinguishing flames. However, its mechanism was not understood. Biordi and coworkers construct- ed the first flame-sampling molecular beam mass spectrometer in order to understand how CF3Br

inhibited combustion in hydrocarbon flames. In this work, they measured both stable and radical

species in methane flames that were doped with CF3Br. Many of the relevant elementary reactions

describing the decomposition of CF3Br, its chemistry, and its influence on hydrocarbon flames were

determined in this work. Westbrook developed the first comprehensive chemical mechanism to described in detail the chemistry of CF3Br. This mechanism was then used to model the inhibition of

CF3Br in hydrocarbon flames. As a result of this work and other work, it is generally agreed that flame suppression by bromine-containing compounds is a result of catalytic destruction of H atoms by

Br atoms in the flame. The ability of bromine to recycle in the chemical system in the flame is directly related to the weak molecular bonds formed by bromine. Thus, bromine when complexed in molecules such as CF3Br, CH3Br, HBr, and Br2, can through a number of reactions regenerate Br

atoms. It was also determined in these studies that •CF3, formed by decomposition of CF3Br, also removes H atoms by competition with radicals important to combustion {e.g., H, O, OH, HO2).

There are a variety of ways in which fire suppressants act in inhibiting hydrocarbon flames.

Most of these eff'ects are intimately related. For example, a heat loss means a temperamre decrease,

which causes the chemistry to slow, which means fewer radicals leading to product formation, which means less heat generated, which results in a temperature decrease and so on. One can separate

suppression efl'ects into two general categories: physical and chemical (although there is overlap). We

prefer a definition where chemical effects are directly related to the characteristics of the specific

molecule, such a H or F or CI substimtion, and physical effects are not. For example, heat capacity

to a first approximation is largely a function of the number of atoms in the molecule and their

connectivity, but not the identity of the molecule. Consequently, heat capacity can be considered an

effect of more of a physical nature than a chemical one. There are a number of physical effects that relate to fluid mechanics, mass transport, and heat

transpon processes that were not specifically studied as a part of our task (largely dealing with the

chemistry of these agents). Physical effects were addressed by other tasks in the overall project that consisted of measurements of PVT properties, discharge dynamics, dispersion mechanics, as well as measurements of extinction effectiveness for a number of more realistic geometries than can be
470 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

considered in this modeling task. We will not discuss these in any detail here and refer the reader to the sections dealing with other tasks. However, it is necessary to understand the impact of these

effects as they relate to our results in order to provide a framework for transfer of the results of our

simulations to more realistic conditions. Some of the fluid mechanics and mass transport effects include dilution, turbulent mixing, diffusive mixing, thermal diffiision, and buoyancy. Heat related

effects, which we could not directly address in our simulations, include the latent heat of vaporization

and non-adiabatic effects, such as radiative cooling or heat losses for flames attached to a cold

surface. There are a number of different types of chemical effects. Most of these involve different

competing factors. First, all of the fluorinated hydrocarbons will eventually decompose and then burn

(forming CO2, HjO, and HF). This liberates heat and increases flame temperatures (which of course speeds flame chemistry). On the other hand, the agents are large molecules with many atoms. Consequently, their high heat capacities may result in a decrease in temperature in the flame prior to complete combustion (which of course slows flame chemistry). The competition between these two

factors will be strongly dependent upon conditions; most important of which will probably be the

mechanics of mixing of the fuel and oxidizer (including difliision processes).

Another set of competing effects involves the fluorinated radicals produced by decomposition of

the agents. These radicals are slower to burn than their pure hydrocarbon analogues, because the C-F

bond is significantly stronger than the C-H bond. Consequently, reactions involving these radicals

may effectively compete with analogous pure hydrocarbon chemistry by creating less "flammable"

intermediates, thereby inhibiting combustion of the hydrocarbon fuel. For example, since the agents

are added to the air stream, their immediate decomposition products (radicals) are formed in oxygen

rich, relatively cold regions (preheat) of the flame. Consequently, these radicals may be involved in termination steps, such as •CHF2 + H02» -* CH2F2 + O2, slowing radical chain reactions and

inhibiting the flame. These radicals will also simply compete with hydrocarbon radicals for important

H, O, and OH radicals. On the other hand, these radicals can also react with stable molecules in the

colder air stream, such as O2, generating more radicals, such as O atoms, and thereby initiating

chemistry or promoting combustion of the fuel. These and other chemical effects (and the relative importance of each) are being considered as part of this work.

5.1.3 Reaction Engineering Approach

5.1.3.1 Species Thermochemistry. We have constructed a large comprehensive reaction set or "mechanism" for fluorinated hydrocarbon chemistry involving Cj and C2 stable and radical hydrocarbon species, including partially oxidized fluorinated hydrocarbons. Existing thermochemical data was

compiled and evaluated. Where little or no data existed for potential species of interest (most of the

radicals), we have estimated that thermochemistry using both empirical methods, such as group

additivity (Benson, 1976), and also through application of ab initio molecular orbital calculations

(Frisch et al., 1992; Curtiss et ai, 1991; Melius, 1990). Where heat capacity data existed, but only

over a limited temperature range, these data were fit to two splined polynomials which extrapolated to the theoretical high temperature limit and were well-behaved at intermediate temperatures {e.g..

Figure 1). In all cases (experimental, empirical, and ab initio), significant effort was made to utilize thermochemical data for each which was consistent (correct relative values) with data for all other

species. Figure 1 shows a typical fit to heat capacity data for one of the fluoroethanes. This figure is a (black & white) "screen dump" of the (color) display window of the interactive, graphics program

"NIST FITCP," which was developed as part of this subtask.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 47

Substance: CF2H-CF2H

50.

K 20

Break T; 1140.00

1000 1500

I '

'— '—

r

2000 2500 3000

TEttPERRTURE K 3500 -1000

Tbmp Incr: 20.00

4sao 5000

92MEL ;

Figure 1 . NIST FITCP. Typical fit to heat capacity data for a fluoroethane. This is a (black

& white) "screen dump" of the (color) display window of the interactive, graphics

program "NIST FITCP" developed as part of this subtask.
472 SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH

There are a number of general sources of relevant compiled and evaluated thermochemical data. These include The Chemical Thermodynamics of Organic Compounds (Stull et al., 1969), JANAF

Thermochemical Tables (Stull and Prophet, 1971), Thermochemical Data of Organic Compounds

(Pedley et al., 1986), TRC Thermodynamic Tables (Rodgers, 1989), Physical and Thermodynamic

Properties of Pure Chemicals (Daubert and Danner, 1985), NIST Structures and Properties Database

and Estimation Program (Stein et al., 1991). There are two compilations/evaluations of fluorinated

hydrocarbons in the Journal of Physical and Chemical Reference Data: "Ideal Gas Thermodynamic

Properties of Six Fluoroethanes" (Chen et al., 1975) and "Ideal Gas Thermodynamic Properties of

Halomethanes" (Kudchadker and Kudchadker, 1978). There are a few individual sources of more

recent data for thermochemistry that are relevant. These include "Thermochemistry of Fluorocarbon

Radicals" (Rodgers, 1978), "Hydrocarbon Bond Dissociation Energies" (McMillen and Golden,

1982), and "Halomethylenes: Effects of Halogen Substitution on Absolute Heats of Formation" (Lias

et al., 1985). There are a numerous other references with thermodynamic data for fluorinated

hydrocarbons that we have compiled as part of this work and will be discussed in more detail in a

later publication. It should be noted that for most of the radicals, we have relied upon recent ab

initio calculations of thermochemical data. This includes both ab initio calculations done as part of

this subtask of the project and those done previously by other workers. Tschuikow-Roux and

coworkers have calculated thermochemistry for the fluoroethyl radicals (Chen et al., 1990a, 1990b;

Chen et al., 1991a, 1991b). Melius has calculated thermochemistry for many stable and radical

fluorinated species (Melius, 1993). Nyden (1993) has used ab initio calculations to obtain

thermochemical data for a number of the fluoroethanes and fluoroethyl radicals. Details of our ab

initio calculations will be given in a subsequent section.

A listing of the thermochemistry that was used for fluorinated hydrocarbon species is given in Table 1 , consisting of enthalpies of formation and entropies (at standard state) and temperaturedependent heat capacities. Standard thermochemistry for hydrocarbon, hydrogen/oxygen, and

hydrogen/fluorine/oxygen species was used and is not included in this table.

5.1.3.1.1 Hydrocarbon and H/O/F Thermochemistry. We have used standard hydrogen/oxygen and hydrocarbon thermochemistry, most of which can be found in the JANAF tables

(Stull and Prophet, 1971). More recent thermochemistry for C2H3 and HCO has been utilized. Further refinement of this mechanism should include re-adjustment of any other thermochemistry or

rate constants that are based upon the older values for the heat of formation of these two species.

Thermochemistry for F, F2, and HE can be found in the JANAF tables as can data for F0«,

FOO», EOF, and HOE (The latter species were eventually excluded from the reaction set as the

simulations showed that their concentrations were negligible at high temperatures under hydrocarbon

flame conditions). The thermochemical data for F, F2, and HE are very well established. Although

there is a more recent value for the heat of formation of HE, the values contained in the JANAF table were used for consistency. This is because there are many thermochemical and rate data for fluorinated species that are based on the current JANAF recommendation for HE Further refinement

of this mechanism should include evaluation of this new value for HE and re-adjustment of any other

thermodynamic or kinetic data that are based on HE thermochemistry.

5.1.3.1.2 Cj Fluorinated Hydrocarbon Thermochemistry. Thermochemical data for the

fluoromethanes: CH3F, CH2F2, CHF3, and CF4 can be found in the JANAF tables (Stull and

Prophet, 1971). The thermodynamic properties of these fluoromethanes have also been re-examined

subsequently (Rodgers et al., 1974). As part of this subtask of the project, we have also calculated

this data using the BAC-MP4 ab initio method (Melius, 1990). Details of our ab initio calculations
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 473

will be given in a subsequent section. Melius (1993) has also calculated this thermochemical data for the fluoromethanes. It is necessary to have ab initio calculated values for these molecules with well

known thermochemical data in order to adequately benchmark ab initio calculations for those

molecules with unknown or uncertain thermochemistry. Entropy and heat capacity data for the

fluoromethanes are derived from spectroscopic data and are known to a very good level of certainty

(less than 0.5 J/mol/K for entropies). Heats of formation for CH2F2 and CF4 are the best known,

with levels of uncertainty of less than 1.5 kJ/mol. The values for CH2F2 and CF4 are derived from

their heats of combustion (which are dependent on the heat of formation of HF). These data also

correlate well with heats of formation for other fluorine-containing molecules. Heats of formation of

CH3F and CHF3 have somewhat higher uncertainties of less than 10 kJ/mol. The heat of formation

of CH3F is derived from appearance potential measurements (high level of uncertainty) and also from

bond additivity correlations with heats of formation for CH4 and the other fluoromethanes. The heat

of formation of CHF3 is obtained from heat of combustion data for CHF3, equilibrium data with other

compounds, and some kinetic data. The data derived from heat of combustion appear to be

complicated by competing reactions. Heats of formation for the fluoromethanes calculated using

BAC-MP4 are within 1-3 kJ/mol of the experimentally derived values (except for CF4, the reference

compound). A better BAC correction should be possible using correlations with all experimental data

for fluoromethanes and fluoroethanes rather than just one compound (CF4).

Thermochemical data for the perfluoromethyl radical (•CF3) can be found in the JANAF tables. More recent experimentally derived values for the heat of formation of •CF3 can be found elsewhere

(Rodgers, 1978; McMillen and Golden, 1982; Tsang, 1986). These suggest a value that is about 4

kJ/mol higher than the JANAF recommendation. All of these heat of formation data are derived from

kinetic experiments for the forward and reverse rates of several reaction. The JANAF recommendation is also based in part of correlations with heats of formation (and bond strengths) for other

fluorine-containing compounds. The entropy and heat capacity data for •CF3 in the JANAF table are derived from spectroscopic measurements with an uncertainty of about 5 J/mol/K in the entropy. The

other fluoromethyl radicals (•CH2F, •CHF2) are not in the JANAF tables. However, reliable,

experimentally derived values for these two compounds can be found in other evaluated sources (McMillen and Golden, 1982; Pickard and Rodgers, 1983).

As part of this subtask of the project, we have also calculated thermochemical data for the

fluoromethyl radicals using the BAC-MP4 ab initio method. These data have also been calculated by

Melius (1993). We have used the entropies and heat capacities for •CH2F and •CHF2 from the ab

initio calculations in the kinetic simulations part of this subtask, since there does not exist (to our knowledge) experimental values for these two radical species. Heats of formation for •CH2F, •CHF2,

and •CF3 calculated by this method are within 1-2 kJ/mol of the experimentally derived values. Although there are the experimentally derived values for •CH2F and •CHF2, we have used the ab

initio values in order to be consistent with other ab initio calculations. Undoubtedly, these data

should be critically evaluated in further refinements of the mechanism. This may consist of using ab

initio calculations benchmarked appropriately to experimental values. Thermochemical data for the fluoromethylenes, :CHF and :CF2, can be found in the JANAF

tables. Unfortunately (since :CHF and tCFj are very important species), there are significant

uncertainties in their heats of formation. The values for :CF2 are the best and are derived from

appearance potential, ionization potentials, kinetic data, and heats of reactions with heats of formation

for related molecules. The best value based on ionization potential data is -186(± 10) kJ/mol. The

kinetic data appear to provide the most consistent values. The JANAF recommendation is -182(_r 6)

kJ/mol (with a range of -172 to 190 kJ/mol). More recently Rodgers (1978) has recommended a
474 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

value based largely upon kinetic data of -187(±4) kJ/mol. Subsequently, Lias and coworkers (Lias

et al., 1985) have recommended a significantly lower value of -205(±13) kJ/mol. This recommendation was based solely on appearance potential and ionization data. Depending on what data were

used, values ranged from -172 to -213 kJ/mol could be obtained. Given the large uncertainty in values derived from appearance and ionization potentials and given that it is essential to use a heat of

formation for :CF2 that is consistent with kinetic data, we believe that the JANAF recommendation or possibly the more recent value of Rodgers should be used. To date we have employed the JANAF

recommendation until all of the experimental and ab initio data have been critically evaluated.

The uncertainty in the heat of formation for :CHF is even greater due to the lack of direct,

reliable data (heat of reaction, kinetic, or otherwise). The recommended JANAF value of 126(±29)

kJ/mol is based very roughly on bond energies from other molecules. More recently, Lin and

coworkers (Hsu et al., 1978) and Pritchard and coworkers (Pritchard et al., 1984) have independently recommended 163(±8) kJ/mol based upon heat of reaction and kinetic data. However, Lias and

coworkers again considered the kinetic and heat of reaction data to be unreliable and have used proton

affinity values to make a recommendation of 108(+13) kJ/mol, which is much lower than all other

recommendations. Again, since it is necessary to use a value that is consistent with kinetic data for neutral thermal species (not ionic species), we believe that either the JANAF value or the Lin and

Pritchard value should be used. To date we have employed the JANAF recommendation until all of

the experimental and ab initio data have been critically evaluated.

As part of this subtask of the project, we have also calculated thermochemical data for :CHF and

:CF2 using the BAC-MP4 ab initio method. These data have also been calculated by Melius (1993).

The calculated value for :CF2 is about 10-15 kJ/mol lower than the best experimentally derived

values. For :CHF, the BAC-MP4 ab initio value is about 25 kJ/mol higher than the JANAF value

and 10 kJ/mol lower than the value determined by Lin and coworkers (Hsu et al., 1978) and

Pritchard and coworkers (Pritchard et al., 1984). We have also calculated heats of formation for :CHF and :CF2 using G2 ab initio theory (Curtiss et al., 1991) and obtain better agreement with the

experimentally-derived values (within 5-10 kJ/mol of the JANAF recommendations).

Thermochemical data for most of the other possible Cj fluorinated hydrocarbons can be found in the JANAF tables. These include 'CF, CHF = 0, CF2 = 0, and •CF = 0. The biggest uncertainties

here are for CHF = 0 and •CF = 0 where no direct experimental data are available and, consequently,

their heats of formation were calculated using average bond dissociation energies from other related compounds. We have calculated thermochemistry for these carbonyl fluorides using the BAC-MP4

ab initio method, as well using the 02 ab initio method. For 'CF^O and CHF = 0, the ab initio values are about 10 kJ/mol and 20 kJ/mol lower than their respective JANAF recommendations. For

CF2 = 0, the ab initio values are about 20-40 kJ/mol higher than the JANAF recommendation. Given

the importance of these species and the uncertain thermochemistry, further refinement of this mechanism will require these values be re-examined. To date, we have used the JANAF values for consistency. There is an experimentally derived value for the heat of formation of CF30* (Batt and

Walsh, 1982). However, we have used our BAC-MP4 ab initio value for the sake of consistency

because we have relied solely on ab initio values for the other fluoromethoxy radicals and for a number of reactions involving CF30». Undoubtedly, all of these data should be re-evaluated in order

to provide the best values.

5.1.3.1.3 C2 Fluorinated Hydrocarbon Thermochemistry. The thermochemical data for C2

fluorinated hydrocarbons that were used in this work are given in Table 1

.
SECTION 5. FLAME INHIBrnON CHEMISTRY AND THE SEARCH 475

Thermodynamic properties for most of the fluoroethanes are derived from experimental data. This work is compiled and evaluated in the Journal of Physical and Chemical Reference Data (Chen

et al, 1975). These data and data for other fluoroethanes can be also found in the DIPPR compilation (Daubert and Danner, 1985).

There are experimentally derived thermochemical data (Rodgers, 1978) for a few of the

fluoroethyl radicals (CH3-CH2», CF3-CH2*, CF3-CF2*). Thermochemistry for all of the fluoroethyl

radicals have been calculated through the application of ab initio molecular orbital theory by

Tschuikow-Roux and coworkers (Chen et al., 1990a, 1990b; Chen et al, 1991a, 1991b). Thermochemistry for a number of the fluoroethyl radicals have also been calculated using the BAC-MP4 and

the G2 ab initio methods. We have used experimentally derived data, calculated some values using

group additivity, and also calculated some of the values using the BAC-MP4 ab initio method. To

dale, we have chosen to use those values for the complete set calculated by Tschuikow-Roux and

coworkers for consistency'. In their work, the experimental heat of formation data for the three

fluoroethyl radicals were employed. For the other fluoroethyl radicals, they determined heat of

formations by benchmarking their ab initio values to experimental data using a series of isodesmic

reactions. Future refinement of this mechanism will need to consider all of the ab initio calculations

with the relevant experimental data using bond additivity and group additivity methods, as well as using isodesmic reactions. This would yield the best, consistent thermochemistry for all of the

fluoroethyl radicals which are, whole or in part, experimentally derived.

Thermochemistry for CF2 = CF2 can be found in the JANAF tables and data for CH2 = CHF and

CH2 = CF2 in DIPPR (Daubert and Danner, 1985). For the three other fluoroethylenes

(CHF = CHF(F), CHF = CHF(Z), CHF = CF2), we have used thermochemistry from our ab initio calculations. Melius (1993) has also calculated this data. Again, as in the case of the fluoroethyl

radicals, future refinement of this mechanism would involve benchmarking the ab initio calculations

with the experimental data to provide a consistent set. There are not experimentally derived

thermochemical data (to our knowledge) for the fluorovinyl radicals. Consequently, we have used

values that we have calculated as part of this subtask the BAC-MP4 ab initio method.

The thermochemistry of the fluoroacetylenes (C2HF, C2F2) can be found in the JANAF tables. We have used these values. However, there are relatively large uncertainties in these data: +60

kJ/mol and ±20 kJ/mol in the heats of formation of C2HF and C2F2, respectively. Further

refinement of this mechanism should critically evaluated the sensitivity of the simulations to these

data. Fluoroketenes and the fluoroketyl radical can be formed through a number of channels

(analogous to simple hydrocarbon chemistry). To assess the importance of these species and relevant

reactions, we have included these species in the mechanism. There are not experimentally derived

data for these species. Consequently, we have had to rely upon values which we have calculated as part of this task using the BAC-MP4 ab initio method.

5.1.3.1.4 Ab Initio Thermochemistry. As part of this subtask, we have calculated thermochemistry for many fluorinated species through the application of ab initio molecular orbital theory. In this work, the GAUSSIAN series of programs was used (Frisch et al, 1992). In the BAC-MP4 ab

initio method (Melius, 1990), the geometries are optimized at the Hartree-Fock level, the relative energies are then calculated at the MP4 level, and finally the absolute energies are calculated by

applying a Bond Additivity Correction (BAC). The BAC is based on the type of bond {e.g. , C-H, CF), its bond length, the identity of all next nearest neighbors, and the bond length between the nearest neighbor and the next nearest neighbor. Both calculated and experimental derived (where available)

heats of formation for these compounds are summarized in Table 2 and Table 3 . A discussion of the
476 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

comparison between experimental heats of formation for these fluorinated hydrocarbons and the ab

initio values is presented elsewhere in sections dealing with each class of compounds. It is necessary

to have ab initio calculated values for molecules with well known thermochemical data in order to adequately benchmark ab initio calculations for those molecules with unknown or uncertain thermochemistry.

Thermochemical data for all of the species that we have calculated are given in Appendix A in a "JANAF style" tabular format. Values for moments of inertia, vibrational frequencies, bond

distances, bond angles, heat of formation at 0 K, heat of formation at standard state, and entropy at standard state are given for each species. Heat capacity, free energy of formation, and other

thermochemical quantities as a function of temperature are also given for each species. Many of the

species included in Appendix A were initially considered in the chemical mechanism, but subsequently eliminated to keep the reaction set to a manageable level, because the simulations indicated they did

not significantly contribute to the overall chemistry. Consequently, many of the species are not found

in Table 1 , which contains only those species employed in the final chemical kinetic mechanism used

in the premixed flame simulations. The importance of some of the species, such as perfluoromethanol

(CF3OH) or the fluoroacetaldehydes {e.g., CHF2-CHO), should be reexamined in future refinements

of this chemical mechanism. For example, they would undoubtedly play larger roles at lower

temperatures.

We have calculated thermochemical data using the BAC-MP4 ab initio method for all of the

possible Cj fluorinated hydrocarbons. This work includes the fluoromethanes (CH3F, CH2F2, CHF3,

CF4), fluoromethyls (•CH2F, 'CHFj, •CF3), fluoromethylenes (:CHF, :CF2), and fluoromethylidyne

(•CF). We have also calculated thermochemistry for most of the possible oxidized Cj fluorinated

hydrocarbons. This includes all of the carbonyl fluorides (CF2 = 0, CHFO, 'CFO), one of the

fluoromethanols (CF3OH), all of the fluoromethoxy radicals (CHjFO*, CHF20.), CF30»), one of the

hydroxyfluoromethyl radicals (•CF2OH), one of the fluoromethylperoxy radicals (CF300*), one of

the fluoromethylperoxides (CF3OOH), and fluoroformic acid (FCO2H). We note that there are energetically diff"erent, configurational isomers for the hydroxyfluoromethyl radicals and we calculated

values for each. Melius (1993) has also calculated thermochemistry for some of the species and for a number of related species including the other fluoromethanols (CH2FOH, CHF2OH) and the other

fluoromethylperoxy radicals (CH2F00*, CHF200*), including the energetically diff'erent, configurational isomers when possible.

Thermochemical data for some of the fluoroethanes (CH3-CH2F, CH3-CHF2, CH3-CF3, CHF2-

CHF2, CHF2-CF3), some of the fluoroethyl radicals (CHF2-CH2*, CH3-CF2*, CF3-CH2V CHF2-

CF2V CF3-CF2''), all of the fluoroethylenes (CH2 = CHF, CHF = CHF, CH2 = CF2, CHF = CF2,

CF2 = CF2), and all of the fluorovinyl radicals (CH2 = CF., CHF = CH., CHF = CF., CF2 = CH.,

CF2 = CF»), have been calculated by Melius (1993) using the BAC-MP4 ab initio method. We note

that CHF = CHF and two of the fluorovinyl radicals each have two energetically diff'erent, configurational isomers (Z and E). In addition. Melius (1993) has calculated thermochemistry for oxidized C2

fluorinated hydrocarbon species, including the fluoroacetaldehydes (CH2F-CHO, CHF2-CHO,

CF3HO) and the fluoroacetyl radicals (CH2F-C0v CHF2-C0v CF3-CO.); In addition to these

calculations, there have been ab initio calculations for the fluoroethanes and fluoroethyl radicals by

other workers using other methods. As indicated previously, Tschuikow-Roux and coworkers have

calculated thermochemistry for all of the fluoroethyl radicals in a series of papers (Chen et al., 1990a,

1990b; Chen et al., 1991a, 1991b). In addition, Nyden (1993) has used G2 ab initio calculations

(Curtiss et al., 1991) to obtain thermochemistry for a number of the fluoroethanes and fluoroethyl

radicals.
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 477

As part of this task, we have replicated the calculations of Melius using the BAC-MP4 method

for a number of these molecules. We have also calculated thermochemical data for a number of other

related species, which were necessary to have either to include directly in the chemical mechanism or

to benchmark (in combination with experimental data) other ab initio calculations. These include some of the other fluoroethanes (CH2F-CH2F, CH2F-CHF2, CH2F-CF3), the two fluoroacetylenes

(C2HF, C2F2, 'C^F), the two fluoroketenes/ethenones (CHF = C = 0, CF2 = C = 0), and the fluoroketyl radical (•CF = C = 0).

Using both the experimentally derived and the ab initio calculated thermochemistry values

presented in Tables 2 and 3, group additivity values (for heats of formation) for the fluoroethanes were determined which were consistent with the majority of the data. These values are relatively

useful, since employing the group additivity method of Benson (1976), one can obtain relatively good

thermochemistr}' in the absence of experimental or calculated data. For a complicated system such as being considered in this project (fluorinated-hydrocarbons), this is an extremely valuable tool. In determining group additivity values consistent with the majority of the data, the best correlations were

obtained when a correction for repulsive interactions between fluorines on adjacent carbons was

included. It was observed that the ab initio values often differed from the experimentally derived

values in a consistent fashion. In general, with increasing number of C-F bonds in the molecules, the

difference between the ab initio and the experimentally derived heats of formation also increased.

This ab initio values were, in general, lower than the experimentally derived values in the range of 5- 15 kJ/mol (depending on the ab initio method) per C-F bond. This trend was used to benchmark the

ab initio values when determining group additivity values consistent with all the data. Group values were also determined for chloroethanes for reference. All of these data are shown in Figure 2.

5.1.3.2 Reaction Kinetics. Utilizing the species identified as potentially important, a grid of

possible reactions was constructed. Existing chemical rate data involving these fluorinated species was then compiled and evaluated. Where rate data were available, but only over limited temperature

ranges or at different pressures (for unimolecular or chemically activated steps), RRKM (Robinson

and Holbrook, 1972) and QRRK (Dean and Westmoreland, 1987) methods were used to estimate the temperamre dependencies (at 1 atmosphere pressure) of the rates and to predict relative rates where

multiple product channels were possible. Where no rate data were available for potential reactions,

the rate constants were estimated by analogy to other hydrocarbon or substituted hydrocarbon

reactions. The rate constant prefactors were adjusted for reaction path degeneracy and the rate constant activation energies were adjusted empirically based on relative heats of reaction or relative bond energies {i.e., Evans-Polanyi relationships).

Initially, upper limits were used for estimated rate constants. If as a result of simulation under a

variety of conditions (using different agents, flame geometries, etc.). it was observed that a specific

reaction with an upper limit rate constant did not significantly contribute to the destruction or creation

of any of the species in the "mechanism," then that estimate continued to be used. If however, a

specific reaction contributed to the chemistry and its rate constant was an upper-limit estimate, then

its value was re-examined and possibly refined. For important contributing reactions where no good

analogy was available, where significant uncertainty existed in the barrier (generally reactions with

tight transition states and modest-to-large barriers), or where multiple energetically similar product

channels were possible, ab initio methods were used to calculate the geometries and energies of the

transition states. RRKM methods were then applied to obtain the temperature (and pressure)

dependence of the rate constant. The procedure used for iterative development of the mechanism is schematically shown in Figure 3.
478 SECTIONS. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Group Additivity for Halogenated Ethanes and Ethyl Radicals

00

d

oI

o

IL.o

0) B

h

O

H

-200 -

300 -

nuoro(Expl:)

O nuoro(MP4-BAC)

Q fluoro (HF-corr)

ffl chloro (Expt) e chloro (MP4-BAC)

® chloro (HF-corr)

-J. chloro (semi-empir1)

><; chloro (seml-empir2)

CH3-CH2*

CH3-CHF*

CH2F-CH2_^

CH2F-CHF*-

CH3-CF2* CH3-CH3

CHF2-CH2*J_

CH3-CH2F

CH2F-CF2

r^xjo* CH2F-CH2F CF3-CH2*^ CH3-CHF2

CHF2-CF2*

CF3-CHF*-

CF3-CF2*

CH2F-CHF2

CH3-CF3 "Best' Group Values

CHF2-CHF2

CH2F-CF3

CHF2-CF3

CF3-CF3

J.

CH3 - 12.4 CH2C1 - 16.2 CH2F - 52.1 CHC12 - 18.0 CHF2 -104.4 CC13 - 18.2 CF3 -159.7 CH2*

CHF*

CF2*

+ 39.5

- 3.6

- 53.0 CHCl* + CC12* + 28.8

24.4

X _L

300 -200 -100 0 Group Additivity Heat of Formation (kJ/mol/4. 18)

Figure 2. Empirically-Determined Halocarbon Thermochemistry. Group values were determined for predicting heats of formation of fluorinated and chlorinated ethanes

consistent with the majority of experimental and ab initio data.
SECTION 5. FLAME INHTBITION CHEMISTRY AND THE SEARCH . . .

Reaction Mechanism Development & Validation

Chemical Reaction Mechanism

Reaction Rates

Reactxjr Model

Post-Processor

Key Species

Important Reactions

Rate Measurements

Thermochemical Data

AHf Sf Cp

[^Transport Data

How Rate

Temperature

Concentration

Fuel

Agent

Rame Measurements

-Flame speeds

-Temperature profiles

-Concentration profiles

Figure 3. Reaction Mechanism Development and Validation.
480 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 1. Fluorinated Hydrocarbons: Enthalpy of Formation, Entropy, Heat Capacities

Fluorocarbon AH?(298) S°(298)

C°(T)

300 500 800 1000 1200 1500

ref

CH3F -234.3 222.8 37.50 51.26 68.90 77.30 83.34 90.21 a CH2F2 -450.6 246.6 42.96 58.91 76.33 83.55 88.84 94.35 a CHF3 -697.0 259.6 51.15 69.26 85.10 91.01 94.89 98.83 a CF4 -933.2 261 .3 61.30 80.67 94.49 98.73 101.41 103.67 a •CH2F -31.5 234.4 38.71 48.67 59.29 64.20 67.92 72.02 b •CHF2 -247.3 256.0 42.21 54.10 65.33 69.66 72.60 75,69 b •CF3 -470.3 265.0 49.94 63.30 73.24 76.37 78.27 79.92 a :CHF 125.5 223.2 34.63 39.25 45.21 47.93 50.05 52.43 a

:CF2 -182.0 240.7 39.01 46.45 52.35 54,18 55.28 56.31 a •CF 255.2 212.9 30.05 32.24 34.79 35.77 36,32 36.96 a CF30« -628.3 275.2 56.44 73.91 86.83 90,89 93.40 95.56 b CHF = 0 -376.6 246.7 40.50 51.65 63.12 67.92 71.20 74.70 a CF2 = 0 -638.9 258.8 47.41 60.49 70.82 74.39 76.71 78.78 a •CF = 0 -171.5 248.4 38.99 45.06 50.63 52.76 54.14 55.44 a CHF=C=0 -147.2 270.5 56.75 71.49 84.93 90.34 94.31 98.24 c CF2 = C = 0 -290.4 288.8 68.30 81.79 92.56 96.66 99.61 102.48 c •CF = C = 0(E) 69.0 276.2 56.8 65.85 73.77 76.59 94.38 98.22 c CH3-CHF2 -493.7 282.5 68.72 97.11 124.65 136.47 145.39 154.49 d CH2F-CHF2 -643.5 311.7 75.31 107.53 133.47 143.93 151.98 159.92 e CH3-CF3 -745.6 287.4 78.81 108.32 133.82 144.08 151.47 158.99 d CHF2-CHF2 -890.1 314.8 85.23 115.25 141.25 150.60 156.64 163,29 b CH2F-CF3 -895.8 316.2 86.58 118.04 143.15 152.09 158.17 164,43 d CHF2-CF3 -1104.6 333.8 96.09 127.53 151.13 158.85 164.06 168,60 d

CF3-CF3 -1343.9 332.1 106.79 139.13 160.34 166.68 169.89 172.97 a CH3-CHF« -72.4 274.0 58.79 79.75 103.33 114.06 122.34 130.91 e CH2F-CHF* -235.5 293.4 69.59 91.24 111.69 119.75 125.52 131.33 f CHF2-CH2* -277.2 297.8 67.06 90.20 111.94 121.10 128.00 135.04 e CH3-CF2* -302.5 290.4 67.34 89.22 111.30 120.80 127.83 135.52 e CH2F-CF2* -446.0 311.1 74.07 97.38 118.86 127.24 133.22 139.11 f CHF2-CHF* -451.9 310.7 76.14 99.54 119.95 127.91 133.70 139.35 f

CF3-CH2» -517.1 302.6 77.37 101.75 121.27 128.88 134.26 139.75 e CHF2-CF2* -660,2 328.4 84.32 107.31 127.04 134.10 138.83 143.26 f CFj-CHF* -688.3 326.3 85.84 109.68 128.34 135.01 139.45 143.76 f

CF3-CF2* -891.2 340.5 92.84 118.29 136.19 141.80 145.18 148.11 f CH2 = CHF -138.9 262.3 50.61 71.66 91.80 100.57 107.42 114.10 d CHF = CHF(Z) -301.2 268.6 58.06 79.62 99.16 107.13 112.47 118.31 b
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH ... 481

Table 1. Fluorinated Hydrocarbons: Enthalpy of Formation, Entropy, Heat Capacities

Fluorocarbon AH?(298) S°(298)

C°(T)

300 500 800 1000 11 700 ^uu Ljyjv

ref

CHF = CHF(£) -302.2 267.8 59.82 00 fii 1 IZ.OJ 110 nr\ Ho. /l) b CH2=CF2 -336.8 265.2 59.33 81 .55 inn 70 1lUn7/ . lAIH 111 1/1 1 lo. /U d CHF = CF2 -495.8 292.7 69.45 1iUn7/ .Ho 48 1 1 T 01 1 1 8 rfi"? 118.0/ 1 0/1 1 1 Z4 . 1

b CF2=CF2 -658,5 299.9 OU. I\J inn AO I

1 1IDS .DJ SI 1 Tfi 80 izj.sy 1 T/; 0*7 lio.o / a CHF = CH»(£) 124.3 258.0 so 61 67u / . in1 u 0 1 .JO 870 f^f^ / .00 09 9

yz.zi c CHF = CH»(Z) 123.0 257.0 50.74 67.33 81.76 87.92 92,44 c CH2=CF» 109.2 256,8 50,22 66.10 80.71 86.89 91.30 96.37 c CHF = CF»(£) -41.0 279.3 59.30 74.51 87,79 92,80 96.05 99.66 c CHF = CF»(Z) -42.7 279.3 59.30 74.51 87,79 92,80 96.05 99.66 c CF2 = CH» -67.8 277.0 59.20 76,36 89.73 94,73 97,78 100.81 c CF2 = CF» -216.3 300.5 68.26 83,42 95.36 99,41 101.77 104.39 c C2HF 125.5 231,5 52.48 62.17 69.69 73,22 75,83 78.82 a

C2F2 20.9 244,0 57.07 67.96 76.67 79,74 81,54 83.46 a NOTATION

AHf(298): enthalpy of formation (kJ/mol) at standard state (1 atmosphere; 298.15 K)

S°(298): entropy (kJ/mol) at standard state (1 atmosphere; 298.15 K)

Cp(T): heat capacity (J/mol/K) at constant pressure

at standard state (1 atmosphere; T = 300, 500, 800, 1000, 1200, and 1500 K)

REFERENCES

a. Stull and Prophet, 1971

b. Melius, 1993

c, this work

d. Daubert and Danner, 1985

e, Chen et ai, 1990a; Chen et al., 1990b

f, Chen et al., 1991a; Chen et al, 1991b
482 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 2. BAC-MP4 Thermochemical Computations (Cj): Enthalpy of Formation (kJ/mol)

FLUOROCARBON BAC-MP4 ref G2 ref expt ref CH3F -233.9 a,b -246.4 c -234.3 d CH2F2 -451.0 a,b -456.6 c -450.6 d CHF3 -699.6 a,b -716.6 c -697.1 d CF4 -934.3 a,b -962.7 Q -933.0 Hu •CH.F - 31.4 a,b - 32.6 e •CHFj -247.3 a,b -252.9 Q -247.7 Q •CF3 -471.5 a,b -483.0 c -470.3 d :CHF 152.7 b 131.8 a 125.5 d :CF2 -203.3 a,b -190.0 a -182.0 d •CF 236.4 b 255.2 d CHF = 0 -395.0 a,b -376.6 d

CF2 = 0 -598.3 a,b -620.1 a -638.9 d •CF = 0 -182.8 a,b -171.5 d

CH2F0* -194.6 b CHFjO* -405.8 b

CFjO* -628.4 a,b -655.6 f CF3OH -219.8 a,b

CH2F0H(£) -412.1 b

CHF20H(£) -684.5 b

-456.5 a,b -455.2

•CF20H(G) -463.2 a,b CF3OOH -807.5 a CHjFOO* -172.8 b CHF200* -401.2 b CFjOO* -627.6 a,b CF(0)OH -615.0 a

Enthalpy of Formation (kJ/mol) at sandaid sate (1 atmosphere; 298.15 K)

BAC-MP4: calculated (BAC-MP4 ab initio method: Melius, 1990) G2: calculated (G2 ab initio method: Curtiss et ai, 1991)

expt: experimentally derived

REFERENCES

a: this work

b: Melius, 1993

c: Nyden, 1993

d: Stull and Prophet, 1971

e: McMillen and Golden, 1982

f: Batt and Walsh, 1982
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . . 483

Table 3. BAC-MP4 Thermochemical Computations (C2): Enthalpy of Formation (kJ/mol)

FLUOROCARBON BAC-MP4 ref TSC ref G2 ref expt ref CH3-CH2F -272.4 a,b -284.1 c -263.2 f CH2F-CH2F -446.9 b -489.9 c -431.0 f

CH3-CHF2 -505.4 b -518.5 c -493.7 f CH2F-CHF2 -671.5 a -643.5 g CH3-CF3 -755.2 b -777.0 c -745.6 f CHF2-CHF2 -890.4 a,b

CH2F-CF3 -913.4 b -895.8 f

CHF2-CF3 -1124.2 b -1104.6 f

CF3-CF3 -1343.9 d

Z L

- 56.2 a -44.8 i c

CH3-CHF« -72.4 i - 79.1 c

CHjF-CHF* -235.6 i J -255.8 c

CHFj-CH,* -280.7 b -277.4 i -294.2 c

CHvCF,* -300.0 b -302.5 i -309.7 c -302.5 h

CH2F-CF2* -446.0 j CHF2-CHF* -451.9 j

CF3-CH2* -527.2 b -517.1 i -517.1 h

CHF2-CF2* -673.2 b -660.2 j

CF3-CHF* -688.3 j

CF3-CF2* -907.5 b -891.2 j -891.2 h

CH2 = CHF -139.3 a,b -138.9 f CHF = CHF(Z) -301.2 a,b CHF = CHF(£) -302.1 a,b

CH2 = CF2 -340.2 a,b -336.8 f CHF = CF2 -485.8 a,b

CF2 = CF2 -653.5 a,b -658.6 d

CHF = CH»(Z) 124.3 a,b CHF = CH»(-£) 123.0 b

CH2 = CF« 109.2 a,b CHF = CF»(Z) - 41.0 b
484 SECTION 5. FLAME INHmiTION CHEMISTRY AND THE SEARCH

Table 3. BAC-MP4 Thermochemical Computations (C2): Enthalpy of Formation (kJ/mol)

FLUOROCARBON BAC-MP4 ref TSC ref G2 ref expt ref CHF = CF»(£^ - 42.7 b

CF2 = CH» - 67.8 a,b

CF2 = CF» -216.3 a,b

C2HF 118.0 a 125.5 d

C2F2 31.8 a 20.9 d

•C.F 454.0 a CHF=C=0 -147.3 a

CF2 = C = 0 -290.4 a

•CF = C = 0(£^ 69.0 a

CH2F-CH0(Z) -322.6 b

CHjF-CHOiE) -328.9 b

CHF2-CH0(£) -525.1 b

CHF2-CH0(Z) -538.9 b

CF3-CHO -774.5 b

CHjF-CO'iZ) -169.9 b

CH2F-CO»(£0 -172.8 b

CHF2-C0»(Z) -377.4 b

CF3-C0»(Z) -610.0 b

CF3-C0»(£^ -611.3 b

Enthalpy of Formation (kJ/mol) at standanj state (1 atmosphere; 298.15 K)

BAC-MP4: calculated (BAC-MP4 ab initio method: Melius, 1990) TSC: calculated (ab initio method of references i,j) G2: calculated (G2 ab initio method: Curtiss et al.. 1991)

expt: experimentally derived REFERENCES

a: this work

b: Melius, 1993

c: Nyden, 1993

d: Stull and Prophet, 1971

e: McMillan and Golden, 1982

f: Daubert and Danner, 1985

g: Pedley, 1986

h: Rodgers, 1978

i: Chen et al., 1990a; Chen et al., 1990b

j: Chen et al. , 1991a; Chen et al. , 1991b
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 485

A listing of the rate constants in the reaction set or mechanism used in the simulations is given in

Appendix B. In addition, other reactions were also considered, but were observed not to contribute

under the conditions tested. The hydrocarbon and hydrogen/oxygen reaction subsets of the mechanism were derived from the Miller-Bowman mechanism (Miller and Bowman, 1989). Some

modifications to this accepted mechanism were made. Many of the relevant rate constants can be

found in the "NIST Chemical Kinetics Database" (Mallard et al, 1993).

5.1.3.1.1 Hydrocarbon/H/O/F Kinetics. The hydrogen/oxygen and hydrocarbon reaction

subsets of the mechanism are derived from the Miller-Bowman mechanism (Miller and Bowman,

1989) and consists of about 30 species and 140 reactions. Some modifications to this accepted

mechanism were made. All N-containing species (except N2) and reactions were removed. A

number of the rich species {e.g., C2H, C4H2) were eliminated from the mechanism in order to keep

the number of species in the mechanism to a manageable level. Eliminating these species is valid,

since the agent is (in general) added to the air (fuel lean). A number of species {e.g., CH3OH) were

also added to the mechanism. In addition to these addition and deletions, a number of rate constants

for a number of reactions {e.g., CH3-f-0H) were adjusted to provide correct fallolT and productchannel ratios. Most of the reactions involving F and HF with hydrogen- and oxygen-containing species have

been measured and the uncertainties in these values are relatively low. There are three reactions of

this type which were determined to participate in the chemistry under a variety of conditions. These

reactions are the combination of H and F to form HF (and the reverse decomposition) and the

hydrogen abstractions by F atoms from H2 and H2O. The rate expressions used for these contributing

reactions are given in Appendix B. The HF decomposition reaction has been measured only at temperatures above about 4000 K

(Jacobs et al., 1965; Blauer, 1968, Blauer et al., 1971). Although this reaction in the decomposition

direction is unimportant at typical flame temperatures, the reverse H -f F = HF combination must be

considered. Extrapolating the recommended value (Baulch et al., 1981) for decomposition to 1000 K

may result in an uncertainty of as much as a factor of ten, especially when considering non-simple

Arrhenius dependence to the rate and different third-body efficiencies. However, since many other

reactions (F -I- H2, H2O, RH) contribute to F atom destruction, the uncertainty in the absolute rate of

the forward or reverse reaction is most likely unimportant.

The hydrogen abstraction reactions of F atoms with Hj and H2O have been measured only near room temperature (Wurzberg and Houston, 1980; Stevens et al., 1989; Walther and Wagner, 1983).

These values were extended to higher temperatures by fitting the reported values to extended

Arrhenius expressions. For the H2 reaction, an expression with T^-^ dependence was chosen

consistent with the recommended value (Cohen and Westberg, 1983). For the H2O reaction, an

expression with T^^ dependence was chosen by analogy to other reactions.

There are a number of other reactions which were included in the mechanism, but were never observed to contribute to the chemistry. These reactions include the combination of F atoms to form

F2 and the hydrogen abstractions by F atoms from OH, HO2, and H2O2. The oxy-fluoro-species

F0«, HOF, F00», and F2O were also initially considered in the mechanism. However, given the

very low concentration of F atoms at high temperatures in the hydrocarbon/air flame, these species

are present in extremely low concentrations and do not contribute to the overall chemistry. The rate constants used for reactions involving these species will be detailed later in another publication.

5.1.3.2.2 Cj Fluorinated Hydrocarbon Kinetics. The Cj reaction subset of the mechanism

consists of about 15 species (fluorocarbons) and 200 reactions (including reactions with H, O, OH,

etc.). Both thermally and chemically activated decompositions are considered {e.g., CH2F2 ^ :CHF
486 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 4. RRKM Predictions of Product Channels for Fluoromethyl Combinations.

I
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 487

Figure 5. Comparison between Experimental and Calculated Rate Constants. Ab initio and

RRKM methods for calculating rate constants must be validated by comparison to known, experimentally-derived values.
488 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

+ HF and •CHF2 + H :CHF + HF). Fluoromethane decompositions via abstraction of H atoms

by H, O, and OH radicals are also considered.

There have been a number of measurements of the unimolecular decomposition of fluoromethanes (with HF elimination): at least two for CH3F (Politanskii and Shevchuk, 1967; Schug and

Wagner, 1973), at least one for CH2F2 (Politanskii and Shevchuk, 1968), and several for CHF3

(Tschuikow-Roux, 1965; Tschuikow-Roux and Marte, 1965; Modica and LaGraff, 1966; Politanskii

and Shevchuk, 1968; Biordi et ai, 1978; Schug et ai, 1979; Hidaka et al., 1991). In addition, there

have been quite a few measurements of the unimolecular decomposition of other halomethanes

(eliminating HF, HCl, or HBr) such as CHF2CI (Norton, 1957; Edwards and Small, 1964; Gozzo and

Patrick, 1964; Edwards and Small, 1965; Gozzo and Patrick, 1966; Barnes et ai, 1971; Kushina et

al., 1972; Schug et al., 1979; Zhitnev et al., 1990; Zhitnev et al., 1991), CHF2Br (Cox and

Simmons, 1971) , CHFCI2 (Kushina et al., 1972), and CHCI3 (Shilov and Sabirova, 1960; Schug et

al., 1979). All of these halomethane decomposition reactions have a small-to-moderate barrier in the

reverse direction (i.e., carbene insertion into HF, HCl, or HBr) of 10-40 kJ/mol. Consequently, all of the halomethane measurements are important from the point of evaluating the fluoromethane values

(both experimental and calculated) for consistency. Furthermore, the barriers-to-insertion of :CHF

and :CF2 in these reactions can be used to benchmark reactions of :CHF and :CF2 with many other

important molecules where there is no or little information available (i.e., the reactions of :CHF and

:CF2 with H2, H2O, CH4, C2H5, fluoromethanes, fluoromethyls, etc.). In this work, we have employed rate expressions for HF elimination from CH3F and CHF3 that are fits to extended Arrhenius form to the experimental data of Schug and Wagner (1973) for CH3F

and to the experimental data of Hidaka et al. (1991) for CHF3. These data were obtained, in general,

at different temperatures and pressures than are relevant to the atmospheric flame conditions

considered here. The experimental data were interpolated or extrapolated and fit using temperature

dependencies (T*') that were consistent with the experimental data and our RRKM calculations (using our BAC-MP4 barriers) for these systems. For HF elimination from CH2F2, we have employed a

rate expression from our RRKM/BAC-MP4 calculations, although there is reasonable experimental

data by Politanskii and Shevchuk (1968) that could be used. For the H2 elimination channels for CH3F and CH2F2 decompositions, we have used rate expressions from our RRKM calculations using our BAC-MP4 ab initio barriers. The H2 elimination pathways are minor channels and were included

for completeness. F atom eliminations from the fluoromethanes are negligible decomposition

channels, except for CF4, where it is the only possible pathway. For this reaction, we have used a

rate expression that is a fit to extended Arrhenius form to the room temperature value of Plumb and

Ryan (1986a, 1986b) using a reasonable temperature dependence (T"^ ^). Further refinement of this mechanism should include critical evaluation of all of the experimental data (for the HF elimination

channels) and the ab initio barriers. RRKM methods should be used to obtain rate expressions that are consistent with both rate and (especially) thermochemical data. Although there is some uncertainties in the rate expressions for pyrolysis of the fluoromethanes due to extrapolation of measurements

from other temperatures and pressures, these uncertainties are most likely acceptable since the

primary decomposition pathways in atmospheric flames are by H atom abstraction by H and OH

radicals and not by pyrolysis.

There have been no measurements to our knowledge for reactions involving chemically activated or "hot" fluoromethanes other than the room temperature measurements of the rate of reaction for CF3 + H ^ Products by Ryan and Plumb (1984) and Tsai and McFadden (1989). This class of

reactions is simply the radical combination of fluoromethyls and H atoms to form two product

channels, HF or H2 elimination, when the hot fluoromethane is not stabilized or returned to reactants.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 489

We have used RRKM methods with the BAC-MP4 ab initio barriers to insertion for :CHF and

:CF2 into HF and H2 and the energetics of the reactions pathways to estimate values for these various

reactions, as well as for the stabilized fluoromethane channels. The barriers to insertion for :CHF

and :CF2 into HF are derived from the experimental measurements for the unimolecular decomposition of the fluoromethanes. A example of a similar set of RRKM calculations for fluoroethane

decompositions (both thermal and chemically activated decompositions) is shown in Figure 4. These

RRKM calculations were done for all of the fluoromethanes and fluoroethanes (only four for the

fluoroethanes are shown in Figure 4).

We have also used the BAC-MP4 ab initio method (Melius, 1990) to provide estimates of the

energy and structure of the transition state for some of these reactions, which agree well with the

experimental values for the HF reactions. For insertion into H2, where there are no experimental

data, ab initio calculations were used to estimate the energy and structure of each transition state. Comparisons between experimental and ab initio calculated values for a few of these reactions are shown in Figure 5. Although there are no experimental measurements at flame temperatures for chemically activated fluoromethane decompositions and these are primary pathways for destruction of

fluoromethyl radicals, the corresponding uncertainties in the rate expressions are small since these

combinations are barrierless and occur nearly on every collision. There have been a quite a number of measurements for H atom abstractions from fluoromethanes

by H radicals, but only a few for H atom abstraction by O or OH radicals. For H atom abstraction

by H radicals, there have been at least 3 measurements for CH3F (Hart et al., 1974; Westenberg and

deHaas, 1975; Aders et al., 1975), at least one measurement for CH2F2 (Ridley et al., 1972), and

quite a few for CHF3 (Ayscough and Polanyi, 1956; Pritchard et al., 1956; Skinner and Ringrose,

1965; Amphlett and Whittle, 1967; Arthur and Bell, 1968; Fagarash and Moin, 1968, Kibby and

Weston, 1968; Berces et al., 1972; Kondratiev, 1972; Arthur et al., 1975; Arthur and Bell, 1978).

We should note that many of the measurements for the CHF3 reactions are actually measurements of

the reverse rate or CF3 + H2 -» CHF3 + H. Two of the citations (Kondratiev, 1972; Arthur and

Bell, 1978) are evaluations of the experimental data. We have also calculated the structure and

energy of each transition state for these H atom abstraction reactions from the fluoromethanes using

the BAC-MP4 ab initio method. The energy barriers are included in Table 4 and compare well with

the experimental values. For the CH3F + H reaction, all of the workers cited above incorrectly identified the reaction as abstraction of F atoms instead of H atoms. These workers only measured the disappearance of the

reactants and simply assigned the product channel by analogy to the CH3Br + H reaction, where it is known that the halogen atom (Br) is abstracted. However, the C-F bond is much stronger than the CBr bond or even the C-H bond. Consequently in CH3F, the H atom instead of the F atom, is abstracted. Our ab initio calculations (see Table 4) also support this argument where abstractions of

H atoms from the fluoromethanes by H radicals were calculated to have energy barriers of 49.4,

40.6, 53.6 kJ/mol for the CH3F, CH2F2, CHF3 series, respectively. These same calculations yield

barriers-to-abstraction of F atoms of 131.4, 142.7, 168.6, 171.1 kJ/mol for the CH3F, CH2F2, CHF3,

CF4 series, respectively. This is a significant difference and clearly supports assignment of H atom

abstraction has the dominant channel.

In this work, we have used our fits to extended Arrhenius form to the experimental data of

Westenberg and deHaas (1975), Ridley et al. (1972), and Arthur and Bell (1978) for H atom

abstraction from CH3F, CH2F2, and CHF3, respectively. A temperature dependence of T-^-^ was used

in analogy to the recommended value (Tsang and Hampson, 1986) for H atom abstraction from

methane (CH4). The experimental data were all obtained at modest temperatures (600-900 K). The

uncertainties in extrapolation of this data to flame decomposition temperatures is most likely
490 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 6. Empirically-Determined Activation Energies. In the absence of experimental data,

the rate constant for a reaction can be estimated relatively well using its heat of

reaction and known rate constants for other similar reactions.
SECTION 5. FLAME INfflBmON CHEMISTRY AND THE SEARCH 491

Figure 7. Typical Reaction Pathways for CH2F2 and CHF3.
492 SECTION 5. FLAME INHmiTION CHEMISTRY AND THE SEARCH

acceptable. We note that based on our limited evaluation, the trends in the rate expressions for this homologous series appear to be consistent with changes in the number of H atoms and the C-H bond

strengths or heats of reaction {e.g.. Figure 6).

For the F atom abstraction reactions from CH3F, CH2F2, and CHF3 by H radicals, we have

employed rate expressions derived from our BAC-MP4 ab initio calculations. The F atom abstraction

pathways are negligible channels and these reactions were included simply for completeness in development of this reaction set. However, for reaction of H atoms with CF4, the only possible

pathway is F atom abstraction. For this reaction, we have used the experimentally derived rate expression of Kochubei and Moin (1969, 1971).

There have only been a few measurements of H atom abstractions from fluoromethanes by O

radicals. Parsamyan and coworkers have measured the rate of reaction for CH3F (Parsamyan et al.

,

1967) and for CH2F2 (Parsamyan and Nalbandyan, 1968). Jourdain et al. (1978) and Miyoshi et al. (1993) have measured the rate of reaction for CHF3. In our work, we have used our fits to extended

Arrhenius form to the experimental data using a temperature dependence of T* ^ in analogy to the

recommended value (Tsang and Hampson, 1986) for H atom abstraction from methane (CH4). We

note that we have not used the more recent value for CHF3 + O by Miyoshi et al. (1993), because it appears that this data may be complicated by the CHF3 :CF2 + HF decomposition reaction at the

highest temperatures. This should be examined in more detail, since our conclusion was based on a

limited evaluation of the data. There have been a few measurements of H atom abstractions from fluoromethanes by OH

radicals. Cohen and Benson (1987a, 1987b) have used transition-state-theory calculations to analyze

and predict rate coefficients for a series of halogen-substituted methanes and ethanes. Much of their

analysis is based on the experimental data of Jeong and Kaufman (1982), but is also consistent with

other measurements for fluoromethanes (Howard and Evenson, 1976; Clyne and Holt, 1979; Nip et

al., 1979; Talukdar et al., 1991). In our work, we have used the recommendations of Cohen and

Benson (1987a, 1987b). Since these recommendations are based on experimental measurements are

relatively low temperatures (about 300-500 K) and these reactions are primary decomposition

pathways for the fluoromethanes and significantly higher temperatures, it would be valuable to have

experimental measurements of these rates near flame temperatures. We note that based on our limited

evaluation, the trends in the rate expressions for this homologous series appear to be consistent with

changes in the number of H atoms and the C-H bond strengths or heats of reaction (e.g.. Figure 6).

There have been quite a few measurements of H atom abstraction from methane (CH4) by F

atoms at or near room temperature (Wagner et al., 1971; Pollock and Jones, 1973; Williams and

Rowland, 1973; Manning et al., 1975; Smith et al., 1977; Clyne and Hodgson, 1983; Pagsberg et

al., 1988). In our work, we have used a fit to extended Arrhenius form to the recommended rate constant of Atkinson et al. (1992) using a reasonable temperature dependence (T*^^) in order to extend the rate expression to flame temperatures. Although there is some uncertainty in extrapolating

the rate constant to flame temperatures, this uncertainty is most likely unimportant since this reaction occurs on almost every collision and there are many other reactions (e.g., F -I- H2, F + H2O, and F

+ other hydrocarbons) that contribute to F atom destruction. For completeness in the reaction set (although it is unlikely that they will contribute), we have also included the reactions for H atom

abstractions from the fluoromethanes by F atoms. There have also been a number of m.easurements

for these reactions at or near room temperature for CH3F (Pollock and Jones, 1973; Smith et al.,

1977; Manocha et al., 1983), for CH2F2 (Pollock and Jones, 1973; Smith et al., 1977; Manocha et

al., 1983, Clyne and Hodgson, 1985; Nielsen et al., 1992), and for CHF3 (Pollock and Jones, 1973;

Goldberg and Schneider, 1976; Smith et al., 1977; Clyne and Hodgson, 1983; Maricq and Szente,

1992). For these reactions, rate expressions were used where the rate constant prefactor relative to
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 493

that recommended for CH4 + F (Atkinson et al, 1992) was adjusted to account for reaction path

degeneracy {i.e., fewer number of H atoms) and the activation energy was adjusted such that the rate was consistent with the measurements at room temperature. Use of extended Arrhenius form in these cases is not justified, because of the lack of temperature-dependent experimental measurements.

There have been quite a few measurements of metathetical reactions of methyl/fluoromethyl

radicals with methane/fluoromethanes. These will not be reviewed here. In our work, we have used

the recommendations of Kerr and Parsonage (1976) which are consistent with the majority of the

experimental data and empirical relationships for barriers to reactions (such as shown in Figure 6).

The recommended values are largely based on the pioneering work in this area by Pritchard and

coworkers {e.g., Pritchard et al., 1965), Whittle and coworkers {e.g.. Chamberlain and Whittle,

1972), and Arthur and coworkers {e.g., Arthur and Bell, 1978). We note that based on our limited

evaluation, the trends in the rate expressions for this homologous series appear to be consistent with

changes in the number of H atoms and the C-H bond strengths or heats of reaction {e.g.. Figure 6). Although there have been no experimental measurements of metathetical reactions of vinyl

radicals (C2H3) with the fluoromethanes, one can estimate their rates by analogy to the methyl radical

(CH3) reactions. We have used rate expressions for these reactions where the activation energy was

reduced by 10% (due to the roughly 6 kJ/mol decrease in the heat of reaction) consistent with the

empirical relationship demonstrated in Figure 6. Given the abundance of H and OH radicals in hydrocarbon flames and the somewhat higher

barrier for H atom abstraction by methyl/fluoromethyl and vinyl radicals, these are secondary reaction

pathways. However, under pyrolytic conditions they may contribute (especially the C2H3 reactions

which have the lowest barriers) and, consequently, for completeness should be retained in the reaction

set. The experimental measurements for H atom abstractions by the methyl/fluoromethyl radicals were all made at relatively low temperatures (about 300-600 K). Extrapolation of these measurements

to flame decomposition temperatures may introduce significant uncertainty in the rates especially since

these reactions should have considerable non-simple Arrhenius temperature dependencies. In further

refinement of this mechanism, these data should be critically evaluated. Experimental measurements

at significantly higher temperatures would also be extremely valuable.

Fluoromethyl radicals are destroyed by three general pathways whose relative importance are

sensitive to conditions. 1) Fluoromethyl radicals can combine with H atoms forming chemically

activated fluoromethanes which may either be stabilized or eliminate HF or H2 (creating methylene/fluoromethylenes). This class of reactions was discussed with the fluoromethane chemistry.

2) Fluoromethyl radicals can react with oxygen-containing species {i.e., O2, O, OH) resulting in the

formation of fluoromethoxy radicals and carbonyl fluoride species (as well as elimination byproducts). 3) Fluoromethyl radicals can also combine with methyl (CH3) or fluoromethyl radicals

forming chemically activated fluoroethanes which may either be stabilized or eliminate HF (creating

ethylene/fluoroethylenes). This class of reactions will be discussed with the fluoroethane chemistry.

There have been a number of measurements of the reaction of CF3 with O2 at or near room

temperature (Vedeneev et al. 1978; Ryan and Plumb, 1982; Caralp et al, 1986; Cooper et al., 1988;

Orlando and Smith, 1988), but none (to our knowledge) for reaction of the other fluoromethyl

radicals with O2. At low temperatures the only possible product pathway is formation of the

fluoromethylperoxy radical. These type of radical species are known to play a role in atmospheric

chemistry. At high temperatures in a flame, these species will be present in significantly smaller

concentrations and there are other possible product pathways for the fluoromethyl -I- O2 reactions.

By analogy to the CH3 -I- O2 reaction, fluoromethoxy radicals (and O atom by-product) can be

formed upon dissociation of the 0-0 bond. The fluoromethylperoxy radical complex may also

rearrange and undergo a internal abstraction reaction (of H atoms) creating fluoroformaldehyde

species (CHF = 0 or CF2 = 0) and OH. In contrast to that assumed for the analogous chloromethyl
494 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

radical + Oj-* (chloro)formaldehyde + CIO reactions (Ho et al., 1992), abstraction of F atoms is not possible because of the much stronger C-F bond. In our work we have assumed that at flame

temperatures, the primary product channel is dissociation of the complex to fluoromethoxy radicals

and O atoms. This assumption should be examined in more detail in further refinements of this mechanism. There are not any measurements (to our knowledge) for these reactions. For the flame

chemistry, this is a minor channel. However, at lower temperatures and/or under ignition conditions,

this class of reactions should be pathways which contribute to the chemistry.

The other various existing data and our many estimations, calculations, and evaluations for other

reactions involving Cj fluorocarbon species will not be discussed in further detail here. In many

homologous series of reactions, missing rate constants {i.e., no experimental data) were estimated by

analogy to other reactions. For example, an empirical relationship between the activation energy and

the heat of reaction for H atom abstraction reactions was determined. This is shown in Figure 6. As

a result, in the absence of experimental data, the rate constant for a reaction can be estimated

relatively well using its heat of reaction and known rate constants for other similar reactions.

Typical reaction pathways for the decomposition of two potential agents, CH2F2 and CHF3, are shown in Figure 7. Briefly, fluoromethane decompositions via abstraction of H atoms by O and OH

radicals are also considered with abstractions by OH and H the major decomposition pathways. The

fluoromethyls produced via these reactions are destroyed by several pathways whose relative importance are sensitive to conditions. These pathways include reactions with H radicals, CH3

radicals, and oxygen-containing species (O2, O, OH). The products of the latter reactions consist of

carbonyl fluorides {i.e., CF2 = 0, CHF = 0, •CF = 0) and HF or other elimination products {e.g.,

•CHF2 + OH CHF = 0 + HF). In the simulations, it was observed almost exclusively that any

reaction channel with an HF product was the dominant channel. The fluoromethylenes (:CHF, rCFj) were largely created by combination of fluoromethyls and H radicals via chemically activated

fluoromethanes (and HF elimination). The fluoromethylenes were predominantly destroyed similarly

by combination with H radicals via chemically activated fluoromethylenes (and HF elimination

creating ^CH and ^CF). 'CF radicals created here were largely consumed by reactions with H2O and

O2 resulting in CHF = 0 and 'CF^O formation.

5.1.3.2.3 C2 Fluorinated Hydrocarbon Kinetics. The reaction subset consisted of about 40

species and 400 reactions. This reaction set will not be described here in detail. Briefly, the

fluoroethane destruction pathways (like fluoromethanes) consist of thermally and chemically activated

decompositions and H atom abstraction reactions. Fluoroethyl radicals can react with H radicals (like

fluoromethyls) creating fluoroethylenes (via chemically activated fluoroethanes and HF elimination).

Fluoroethyl radicals can also react with oxygen-containing species (O2, O, OH) resulting in the

formation of oxidized fragments {e.g., CF3-CF2* + O -* •CF3 + CF2 = 0). Fluoroethylenes

(produced from thermally and chemically activated fluoroethane decompositions) are predominantly

destroyed via reaction with O radicals resulting in the formation of oxidized fragments {e.g.,

CH2=CF2 + O ^ •CH = 0 -I- •CHF2). Fluoroethylenes are also destroyed to a lesser degree through

H atom abstraction by radicals such as OH, resulting in formation of fluorovinyl radicals {e.g.,

CH2 = CF2 + OH CF2==CH« -t- H2O). Fluorovinyl radicals (like fluoromethyl and fluoroethyl

radicals) are destroyed via reactions with H radicals, as well as with oxygen-containing species.

However, it was observed that the fluorovinyl radicals established a dynamic equilibrium with the

parent fluoroethylenes, irrespective of the specific creation and destruction pathways.

The kinetics of decomposition of most of the fluoroethanes (HF elimination) has been measured

in a comprehensive series of work by Tschuikow-Roux and coworkers (Tschuikow-Roux et al., 1970;

Tschuikow-Roux and Quiring, 1971; Tschuikow-Roux et al., 1971; Millward et al., 1971; Millward
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 495

f3P(:a!:fi

-1 .0

-2.

3.0

4. D

5.D

7. BDE-2 8.DDE-2 8.20E-2

TIME (sec)

CF2 :

40E-2

NET CHEMICnL

7.8DE-2 8.4DE-2

Figure 8. Typical concentration profiles of an agent (1.9% CHF3) and its decomposition

intermediates and products in a lean (0=0.65), premixed methane/air flame.

i
496 SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH

and Tschuikow-Roux, 1972; Sekhar and Tschuikow-Roux, 1974). Data for HF elimination from the

other fluoroethanes have been obtained by Kerr and Timlin (1971) and Trotman-Dickenson and

coworkers (Cadman et al.. 1970). Fluoroethane decomposition kinetics have also been measured by a few other workers (Kochubei et al., 1980; Mitin et al., 1988). In our work, we have used the rate expressions for these reactions as reported. However, the measurements by Tschuikow-Roux and

coworkers were at significantly higher pressures (3-5 atm.). Consequently, our rates should be

considered as upper limits to the rates in flames at atmospheric conditions. Future refinement of this mechanism should use RRK methods benchmarked to the experimental data to provide better values

for the temperature (and pressure) dependencies of these unimolecular decomposition reactions.

There have been quite a few measurements of H atom abstractions from fluoroethanes by OH

radicals. Cohen and Benson (1987a, 1987b) have used transition-state-theory calculations to analyze

and predict rate coefficients for a series of halogen-substituted methanes and ethanes. Much of their

analysis is based on the experimental data of Clyne and Holt (1979) and Jeong et al. (1984). Other

experimental data included in their analysis was from the measurements by Howard and Evenson

(1976), Handwerk and Zellner (1978), Nip et al. (1979), and Martin and Paraskevopoulos (1983). In our work to date, we have used the recommendations of Cohen and Benson (1987a, 1987b). For the

three asymmetric fluoroethanes (CH3-CH2F, CH3-CHF2, CH2F-CHF2) where there are diff'erent functional H substitutions, we have estimated the relative branching ratios. Recently there have been

a number of good measurements for these reactions for a number of the fluoroethanes by Huie and

coworkers (Liu et al., 1990; Zhang et al., 1992), by Ravishankara and coworkers (Talukdar et al.,

1991; Gierczak et al., 1991), and by Nielsen (1991). Based on some of the more recent measure- ments there are newer recommendations by Cohen and Westberg (1991) for some of these reactions.

The biggest changes are for reactions involving CH3-CHF2 and CHF2-CF3. However, the changes in the rate expressions are only significant at temperatures well below flame temperatures (because T*' dependence). Since these recommendations are based on experimental measurements are relatively

low temperatures (about 300-500 K) and these reactions are primary decomposition pathways for the

fluoroethanes at significantly higher temperatures, it would be valuable to have experimental

measurements of these rates near flame temperatures. We note that based on our limited evaluation,

the trends in the rate expressions for this homologous series appear to be consistent with changes in the number of H atoms and the C-H bond strengths or heats of reaction {e.g.. Figure 6). There have been no measurements (to our knowledge) for abstraction of H atoms from the

fluoroethanes by H and O radicals. We have estimated the rates for these reactions by analogy to those for H atom abstraction from the fluoroethanes by OH radicals. Each prefactor and activation

energy was scaled relative its OH reaction by the same amount as in series of H atom abstraction

from the unsubstituted-ethane (C2H6) by H, O, and OH radicals. There have been quite a few measurements of the reactions of O atoms with fluoroethylenes.

These will not be reviewed here. In our work, we have used our fits to extended Arrhenius form to the recommendations of Cvetanovic (1987) in order to extrapolate the low temperature values (300-

500 K) to flame temperatures. A temperature dependence of T^ -^ was used in analogy to other

reactions. The recommended values by Cvetanovic are largely based on work in this area by Herron

and Huie (1973), Jones and Moss (1974), Atkinson and Pitts (1977), and Gutman and coworkers

(Park et al., 1984). Typical creation and destruction paths for CH2 = CHF are shown in Figure 8. For this class of reactions, it is generally understood that the dominant pathway is where the

products are the fluoromethyl and (fluoro)formyl radical {e.g. CH2 = CHF -I- O HCO -I- •CH2F)

following dissociation of the chemically activated fluoroethylene oxide formed by O atom attack on

the double bond; that is, the O atom "adds" to the side with the least number of electronegative

substitutions {i.e., F) and the H atom "migrates" to the other side. The numerous other possible
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 497

channels are generally considered to be minor pathways: fluorine-substitute analogs of 1) stabilized

ethylene oxide, 2) stabilized acetaldehyde, 3) formaldehyde + methylene, 4) acetyl radical + H, 5)

ketene + H2, and 6) vinyl radical + OH. The latter, abstraction of H atom, is a separate reaction

from the first four (addition/elimination). For completeness in the reaction set, we have included the

latter abstraction reaction using rate expressions based on analogy to the value estimated by Fontijn

and coworkers (Mahmud et al., 1987) for the unsubstituted ethylene reaction. It should be noted that

for pefluoroethylene, the only possible channel is CF2 = CF2 + O -* CF2 = 0 + :CF2 (i.e., no H

migration possible). It should also be noted that for CH2 = CHF there are two possible channels

("addition" of the O atom to one side or the other). We have used an estimated additional 4 kJ/mol

for "addition" of the O atom to the fluorinated carbon. This is consistent with a upper limit

measurement at room temperature for this reaction by Gutman and coworkers (Slagle et al., 1974).

Given that this is a primary decomposition pathway for the fluoroethylenes and that the rate expression we are using are based on experimental measurements at low temperatures (300-500 K), it would be very valuable to have measurements of these reactions and product channels at near flame

temperatures.

Most of the rate expressions used in our work for reactions involving fluoroethyl radicals,

fluoroethylenes, fluorovinyl radicals, and fluoroacetylenes were estimated by analogy to the reactions

for the unsubstituted hydrocarbon species. The rate constant prefactors were adjusted to account for reaction path degeneracy {i.e., number of H atoms). To date, we have not adjusted the activation

energies. Clearly, for H atom abstraction reactions, the barriers should change with C-H bond

strengths and for unimolecular and chemically activated reactions, the rates should be very sensitive to changes in stabilization with difi'erent number of F atom substitutions. Future modifications of this mechanism should address this issue. In our work, we have used rate expressions for the fluorovinyl radicals by analogy to that for C2H5 -I- O2 ^ C2H4 -I- HO2 measured by Bozzelli and Dean (1990). We have used the recommendation of Baulch et al. (1992) for C2H5 -I- O ^ CH2O + CH3 to estimate the analogous reactions for fluorovir^l radicals. In our work, we have used rate expressions for the reactions of fluoroethylenes

with H atoms by analogy to the recommendations of Tsang and Hampson (1986) for the two pathways

C2H4 4- H C2H5 and C2H4 -I- H ^ C2H3 -I- H2. For abstraction of H atoms from the fluoro- ethylenes by OH radicals, we have used our fit (with an estimated T^-^ dependence) to the recommendation of Baulch et al. (1992) for C2H4 -I- OH -* C2H3 + H2O, which is based on a measurement by

Tully (1988). Westmoreland (1992) has calculated the temperature (and pressure) dependencies of

the rate for the chemically activated reaction C2H3 -f O2 -* CH2O + HCO. We have used these

values for the analogous fluorovinyl radical reactions. In our work, we have used the recommendations of Warnatz (1984) and Tsang and Hampson (1986) for the C2H3 -H O ^ Products and C2H3 -I- OH Products reactions, respectively, for the analogous fluorovinyl radical reactions. The

recommendation for the first reaction by Warnatz is based on measurements by Heinemann et al. (1988). For reaction of H atoms with the two fluoroacetylenes (C2HF, C2F2), have used rate expressions derived from the recommendation of Warnatz (1984) for C2H2 + H -* C2H3, which is based on measurements by Payne and Stief (1976).

5.1.3.3 Reactor Models. This large set of rate constants or "mechanism" was used to model

the fluorinated hydrocarbon chemistry under a variety of conditions including diff'erent fuels,

equivalence ratios, diff'erent agents, agent concentrations, and for a variety of reactor geometries.

Since it is not yet possible to model a turbulent, chemically reacting flow, we relied upon employing

a number of geometries in order to provide a good picture of realistic conditions. Simple plug flow

conditions (1-D/time, no diffiision) were used to test a wide variety of conditions and to "debug" the

mechanism since these types of calculations can be performed relatively quickly. The plug flow
498 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 4. BAC-MP4 Transition State Calculations: Activation Energy for Reaction (kJ/mol)

TRANSITION STATE AH?t E CH3F -* :CH2 + HF 120.8 368.6

:CHF + H2 *182.4 418.8

CH2F2 -» :CHF + HF -121.8 327.2

-:CF2 + H2 -51.9 406.7

-* :CH2 + F2 *346.4 792.0

CHF3 -:CF2 + HF -388.3 313.8

-* :CHF + F2 *52.3 754.4

CF3OH CF2 = 0 + HF -741.8 180.3

CH2F0* •CH = 0 + HF -109.6 87.4

CH2F0* -* CF2 = 0 + F -19.2 177.8

CHF20* •CF = 0 + HF -94.6 311.3

CHF20» -* CF2 = 0 + H -328.9 77.0

•CF20H(E) -» •CF = 0 + HF -300.0 156.5 ^CF2= 0 + H -315.2 141.3

-* CHF20* -301.9 154.6

CIA PM P -1- HP -4.6 270.3

LHF2-LHr2 ^ /~*T I C? -* LH2r2 + iCFj

CH4 + :CF2 ^ CH3-CHF2 39.7 317.9

CH3F + H -* •CH2F + H2 28.9 50.2 ^ •CH3 + HF 110.9 131.4

CH2F2 + H -» •CHF2 + H2 -189.1 40.6 ^ •CH2F + HF -87.0 142.7

CHF3 + H 'CFj + H2 -431.0 53.6

-* 'CHFj + HF -317.6 168.6

CF4 + H •CF3 + HF -545.2 171.1

CF2 = 0 + H ^ •CF = 0 + HF -229.7 150.6

-* 'CFjOH -315.1 65.3

(G2) -368.2 33.9

-* CHFjO* -302.1 78.2
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . . 499

Table 4. BAC-MP4 Transition State Calculations: Activation Energy for Reaction (kJ/mol)

TRANSITION STATE AH? Ea CF2 = 0 + H2O -* FC(0)OH + HF

(G2)

-718.8

-710.4

121.3

151.5

CHj-CHF* CH2 = CH» + HF 216.7 289.1

CH2F-CH2» -» CH2 = CHF + H 98.3 154.4

BAC-MP4 ab initio calculations (Melius, 1990) G2 ab initio calculations (Curtis et al., 1991)

AH?: enthalpy of formation (kJ/mol) at standard state (1 atmosphere; 298.15 K)

E^: activation energy (icJ/mol) at 298.15 K

*: spin contaminated

calculations provide essentially the limit of maximum concentration gradients since diffusion or the

interaction between reactant-like and product-like molecules is neglected. A limited set of freely

propagating, premixed flame calculations (1-D/time and distance, axial diffusion) were done to provide more realistic concentration and temperature gradients in the model. This was the primary

method of calculating the effects of agents on flame speed and extinction. In addition, a limited set of continually stirred tank reactor (CSTR), burner-stabilized premixed

flame, and opposed-flow diffusion flame (OFDF) calculations was also done. The CSTR calculations can be considered as providing an upper limit to the importance of diffusion since reactants,

intermediates, and products chemically interact or are completely mixed. This is a case of essentially

minimal concentration gradients and may provide some analogies to turbulent reacting flows. The

burner-stabilized, premixed flame more closely mimics a number of realistic burning geometries such

as a flame near a fuel line leak or possibly a pool fire where the surfaces in these cases can provide a

large heat sink to the flame. The opposed-flow diffusion flame geometry provides a somewhat

realistic model of flame conditions which exist where a burning fuel jet or rapidly heated expanding

fuel vapor comes in contact with the surrounding air. Unfortunately, these calculations which provide

the most realistic picture of diffusion-dominated conditions, are also the most time-consuming

computationally. Consequently, only a limited number of these have been done. It is expected that more calculations will be performed when a greater degree of certainty has been developed in the

reaction mechanism or when experimental measurements in the OFDF geometry become available. Schematic representations of these reactor geometries are shown in Figure 9.

5.1.4 Fluorinated Hydrocarbon Chemistry and Flame Suppression

5.1.4.1 Reaction Path Analysis. The fluoromethanes decomposed primarily through H atom

abstraction by OH radicals in the premixed flame with H atom abstraction by H radicals the major

secondary pathway (on the order of 10-20%). H atom abstraction by O radicals and unimolecular

decomposition eliminating HF are the next important minor decomposition channels (generally

contributing less than 10%). These minor pathways only begin to contribute as secondary channels at the highest temperatures in the premixed flame and the unimolecular decomposition channel at high
500 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

REACTOR MODELS

No Transport

Reactants

Time

Temperature

Residence

Time,

Temperature

Phig Flow Reactor (PFE) Perfectly Stirred Reactor (PSR)

1- D Transport

BURNED MIXTURE

REAC TION ZONE

I

PREMIJ^ED MDOURE

FUEUOKIDIZER

(UNBURNED)

Premixed Flame

COOLED POROUS

PLUG BURNER

Oxidizer

Flame

stagnation

Plane

Stagnation Point

Streamline

Fuel

Counter Flow Diffusion Flame

Figure 9. Reactor Models.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 501

temperatures/short residence times under plug flow conditions. It is likely that the unimolecular

decomposition channel also plays a role in a diffiision flame due to the lack of fuel (and consequently

H and OH radicals) in the preheat, oxygen side of the flame. Abstraction by methyl, fluoromethyl,

and vinyl radicals are negligible in the premixed flame although they may be minor channels in plug

flow or difi\ision flames. Abstraction of F atoms by H radicals is negligible for CH3F, CH2F2, and

CHF3, since the H atom abstraction channel has a significant lower barrier (30-50 kJ/mol) than the F

atom abstraction channel (120-160 kJ/mol). However, in the case of CF4, F atom abstraction by H

radicals is the only possible abstraction channel. Similarly, since there is no HF elimination channel,

the only possible unimolecular decomposition pathway for CF4 is elimination of F atoms. There are two possible channels for production of fluoromethanes. The first involves the

insertion of methylene (:CH2) into HF (reverse reaction of the unimolecular decomposition) and the

second is disproportionation between fluoromethyl radicals to form fluoromethanes and :CF2.

Although latter channel has significant experimental basis (Pritchard et al., ), there is some uncertainty with regard to this reaction, because there is also evidence to suggest that the reverse reaction or

insertion of :CF2 (a closed shell species) should have a modest barrier. Channels of the first type can

contribute small amounts to "recycling" of F mass in the system. For example, CH2F2 can decompose via H atom abstraction reactions to form •CHF2, which can react with H radicals to create :CHF

and HF (by elimination from chemically activated CH2F2). The :CHF formed via this pathway can

then insert into HF (which is present in large concentrations in the system) and consequently reform

CH2F2. However, in the cases we have tested, this channel contributes generally less than 10% to the net rate of reaction for CH2F2. On the other hand, this channel may be significant in predicting

by-products of incomplete combustion. For example, CH3F and the intermediates and products of its decomposition will always be formed in hydrocarbon flames irrespective of fluorocarbon starting

material simply because the HF (always present in the fluorinated hydrocarbon-doped flame) will react with singlet methylene (':CH2) (always present in hydrocarbon flames). The latter channel

involving the disproportionation between fluoromethyl radicals to form fluoromethanes and :CF2

should be investigated further. There are three general channels for the destruction of the fluoromethyl radicals: (1) reaction

with H radicals creating (fluoro)methylenes and HF - via chemically activated fluoromethanes; (2)

reaction with O and OH radicals creating elimination products (i.e., H, HF) and (fluoro)formaldehydes (CH2O, CHF = 0, CF2 = 0) - via both thermal and chemically activated fluoromethoxy radicals;

and (3) reaction with CH3 radicals creating (fluoro)ethylenes and HF - via chemically activated

fluoroethanes.

The relative importance of these three general channels depends upon fuel, agent, flame

geometry, and other conditions. For example, under more realistic conditions in a diffiision flame

where the agent is added to the air stream, the importance of the CH3 combination route should be

substantially diminished (as should the H combination to some degree), because of the lack of fuel (mixed) in the air. On the other hand the OH combination route should significantly increase in importance due to its high dispersion in the flame due to the abundance of H2O in the flame.

Similarly, the O atom reaction route should decrease in importance. Furthermore, other otherwise

minor oxidative routes such as •CH2F-I-02 and •CH2F-I-HO2 may begin to contribute in the preheat,

oxygen rich side of the flame. For ignition delays under plug flow conditions these oxidative routes should be very important as contributors (like analogous reactions for pure hydrocarbons systems).

At high agent concentrations, autocombination of fluoromethyl radicals are contributing

reactions. For CF3 + CF3, there is no HF elimination channel possible and, consequently, CF3-CF3 is the sole product (at the highest temperatures this rate falls-off' slightly). For •CHF2 -I- •CHF2, the

elimination charmel is diminished somewhat in importance, since the CHF2-CHF2 with four heavy F
502 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

atoms can effectively stabilized the heat of reaction and compete as a product channel. This trend can

be seen in Figure 4 (RRKM calculations).

The fluoromethyl radicals are primarily formed by H atom abstractions from the fluoromethanes.

However, there are several other channels which can contribute to their formation and should be

discussed in detail with reactions classified as C2 chemistry. For example, the reactions CH2 = CHF

+ O ^ •CH2F + HCO or CHF2-CF2* + H ^ •CHF2 + •CHF2 contribute to the formation of

fluoromethyl radicals. Similarly, there are a number of other decomposition channels which can be

classified as C2 or C3 chemistry such as •CH2F + C2H4 •CH2-CH2-CH2F. These should be

investigated further.

5.1.4.2 Freely Propagating, Premixed Flame Simulations. Adiabatic, freely propagating,

premixed flame calculations (Kee et ai, 1985) were performed utilizing the reaction mechanism

outlined above. Typically, fuel lean CH4/air mixtures were simulated in order to be most sensitive to flame speed changes and (in a practical sense) since agents are added to the air supply. It would be more realistic to simulate agent effects in a diffiision flame. We have performed a few opposed flow

diffusion flame calculations; however, those results are preliminary.

A summary of the effects on adiabatic flame temperature and speed are shown in Figure 10 and

Figure 11 for addition of a variety of "agents" to a CH4/air flame (equivalence ratio of 0.65). In

addition to potential fluorinated hydrocarbon agents (CF4, CHF3, CH2F2, CF3-CF3, CF3-CF2H, CF3-

CFH2), other species were added to investigate the effect of heat capacity and heat release on changes

in flame speed and temperature. These reference "agents" include N2, H2O, CO2, HF, and CH4. In order to correct for differences in heat capacities for the different "agents", the amount of added

"agent" was normalized to an equivalent amount of N2, adjusting for relative heat capacities at 1500

K. For example, addition of 1 % CF4 (with Cp= 105 J/mol/K) would be roughly equivalent to addition of 3% N2 (with Cp = 35 J/mol/K). Using heat capacities at other temperatures (1000-2000 K)

had little impact on the relative normalized mole fractions. In Figure 10, it can be seen that the effect of the various agents on flame temperature can be

bracketed by addition of inert molecules (N2, H2O, CO2, HF), where there is a decrease in flame

temperature, and by addition of more fuel (CH4), where there is a large increase in flame temperature. The decrease in flame temperature upon addition of the inerts is due to dilution and increased

heat capacity of the mixture. The increase in flame temperature upon addition of more fuel is due to increased heat release as the fuel lean mixture becomes more rich. Addition of agents which are more fuel-like results in larger increases in flame temperature. All of the fluorinated hydrocarbons

are fuels since they all eventually decompose, burn, and form CO2, H2O, and HF (liberating heat).

At one extreme is CF4, very little of which decomposes in the flame, and consequently, there is only

a small increase in flame temperature relative to addition of inerts. On the other extreme is CH2F2,

which completely burns forming highly exothermic products CO2 and HF. In Figure 11, a range of

effects on flame speed for the various added agents can be observed. For the inert molecules and the

nearly inert fluorinated hydrocarbons (CF4, CF3-CF3), a decrease in flame speed is observed

consistent with dilution of the mixture. On the other extreme, for CH4 (the fuel) and CH2F2 (a

slightly poorer fuel), an increase in flame speed is observed. Of the various agents considered, only

CHF3 was seen to have any chemical effect in flame suppression. Figure 10 and Figure 11 show that although there is an increase in flame temperature upon CHF3 addition, there is also a decrease in flame speed relative to inert molecule addition. Inspection of reaction pathways for CHF3 and other

agents (see Figure 7) reveals that a significant amount of decomposed CHF3 results in the formation

of a relatively unreactive perfluorocompound, carbonyl fluoride (CF2 = 0). Decomposition of

CF2 = 0 occurs only very slowly via reaction with either H radicals (H addition 4- HF elimination) or
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 503

Figure 10. Effect of Agents on Flame Temperature.
504 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

0.00 0.02 0.04 0.06

Agent Mole Fraction (normalized to Cp for N2)

Figure 1 1 . Effect of Agents on Flame Speed.
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 505

with H2O (through a hot fluoroformic intennediate). For all other agents, this CF2 = 0 bottleneck is avoided.

5.1.5 Discussion. These simulations are consistent with qualitative trends observed in experimental

measurements of agent effectiveness in the work by others in other parts of the overall project at NIST. However, in order to quantitatively predict agent effectiveness, experimental validation of the

mechanism will be necessary. As indicated in the last section, the premixed flame simulations demonstrate that differences in

the relative influence of candidate agents on hydrocarbon flames can be explained largely by simple

differences in their relative heat capacities and relative amounts of heat release due to agent combustion. That is, flame extinction upon addition of each of the potential agents is only due to simple

dilution of the flame by increased heat capacity of the mixture (consequently, reduced flame

temperamres). This result is not surprising and was the expected behavior in the absence of any

simulations or experimental measurements of agent effectiveness.

The majority of the candidate agents are perfluorocarbons, which are essentially inert relative to the fuel, and therefore by definition, cannot (chemically) inhibit flames. These inert "agents" can

extinguish flames by dilution; however, so can equivalent amounts of completely inert molecules,

such as N2 or COj, as well as sufficient excess oxygen (O2) in premixed flames. In contrast to completely inert molecules, the nearly inert perfluorocarbons will eventually

decompose (at least partially) under flame conditions (at the highest temperatures and long times),

resulting in exothermic products COj and HP (liberating heat) and increasing flame temperatures. - However, since the perfluorocarbons decompose much slower than the hydrocarbon fuels, their

decomposition intermediates and products cannot participate in the hydrocarbon combustion chemistry. Consequently, these "agents" cannot directly influence flame chemistry, but can only increase

flame temperatures; that is, relative to inert molecules these agents can only indirectly promote

combustion (through heat release) and cannot suppress flame processes. This expected behavior was observed in the simulations (i.e., Figures 10 and 11). For example,

addition of CF3-CF3 increases both flame temperature and speed relative to addition of inert molecules. Inspection of concentration profiles and reaction pathways in the premixed flame

simulations for the perfluorocarbons CF4 and CF3-CF3 reveal that the small fraction of these

compounds that decompose, do so in the post-flame zone (subsequent to the important combustion

chemistry). The larger the perfluorocarbon, the more readily it decomposes; that is, a larger fraction

of it decomposes for a given set of conditions. Consequently, more heat is released and more

enhancement of the flame is expected. This trend can be seen in Figures 10 and 11, where both

flame temperature and speed are increased relative to inert molecules, going from CF4 to the larger

and less stable CF3-CF3. It should be noted that in accordance with this behavior, one of the

perfluorocarbons, cyclo-C4Fg, is expected to have a larger enhancement than the corresponding

alkane, because it is likely to decompose more rapidly (to two molecules of CF2 = CF2) and have more combustible intermediates (thereby liberating more heat, increasing flame temperature and

speed). In our work, we did not simulate the effect of agents with three or more carbons atoms,

because of the additional complexity of the chemistry that must be considered. Consequently, we

have no simulations to quantitatively support the above predicted qualitative behavior of the large

molecules.

The influence of addition of the hydrofluorocarbon agents to hydrocarbon flames can best be

described in reference to the related perfluorocarbons. The perfluorocarbons are relatively inert molecules, because they are relatively slow to decompose even under flame conditions. Hydrogen

atom substitution of these molecules will significantly increase their reactivity and enable them to decompose prior to or during combustion (as opposed to in post-flame zones). However, for the
506 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

same reason, hydrogen atom substitution should significantly increase their fuel characteristics. That

is, the hydrofluorocarbons will contribute excess heat when they decompose and their products react further to form exothermic products CO2 and HF. Of course, these agents will also contribute

hydrogen atoms (fuel) to the chemistry. These hydrofluorocarbon agents may also decompose

sufficiently "early" that the radicals they produce may enhance ignition chemistry (promotion).

However, the fluorinated radicals, which are less combustible than the corresponding hydrocarbon

radicals, may also compete for important radicals in the flame, thereby slowing flame propagation

(inhibition). In order for the hydrofluorocarbons to inhibit flames, these competitive reactions must

be suflJicient to counteract both the enhancement of flame chemistry due the hydrogen atomsubstitution and the heat release due to combustion of the added agent. The simulations indicate that

this radical scavenging eff'ect is never dominant.

The influence of addition of the hydrochlorofluorocarbon agents to hydrocarbon flames can best

be described in reference to their hydrofluorocarbon analogues (e.g., CF3-CHFCI versus CF3-CHF2).

The primary decomposition pathway for the hydrofluorocarbons is through H atom abstraction by H

or OH radicals in the flame. In general, HF elimination is a minor contributing pathway except

under pyrolytic conditions where it may be dominant. However, for the hydrochlorofluorocarbons,

decomposition by HCl elimination is an energetically feasible pathway. Consequently, these

molecules may have a dilTerent chemical effect on the flame, if HCl (or other Cl-substituted

molecules) influence the chemistry. One of the primary reasons for the effectiveness of CF3Br is that HBr that is formed is recycled to Br atoms (which scavenge H atoms in the flame) through the

energetically favorable reaction HBr + OH ^ Br + H2O (AH^xn-'l^O kJ/mol). On the other hand F

atoms do not catalytically destroy H atoms because HF, once formed, is never consumed. This is because the reaction HF + OH ^ F + H2O (AHrjj„== +75 kJ/mol) is essentially energetically

prohibited. However, the analogous reaction HCl + OH CI + H2O (AH^j^j^ = -65 kJ/mol) is energetically favorable. Therefore, CI atoms may catalytically destroy H atoms in the flame (like Br

atoms) and consequently, the hydrochlorofluorocarbon agents may be more efl'ective than their

hydrofluorocarbon analogues. The one exception to this may be CHF2CI. In this case, HCl

elimination yields :CF2 (relatively reactive), while for CF3-CHFCI, HCl elimination should yield

CF2 = CF2 (relatively unreactive). The :CF2 species may be formed sufficiently "early" in the flame

that it may enhance ignition chemistry (potential source of radicals). If this is the case, CHF2CI

could promote, rather than inhibit the flame. In our work, we did not simulate the eff'ect of agents

with chlorine atoms (simultaneous with fluorine atoms), because of the additional complexity of the

chemistry that must be considered. Consequently, we have no simulations to quantitatively support

this predicted qualitative behavior for the hydrochlorofluorocarbon agents.

We did not perform any substantial number of simulations with other fuels. However, as discussed above, the diff'erences in the relative influence of candidate agents on hydrocarbon flames can be explained largely by simple diff'erences in their relative heat capacities and relative amounts of

heat release due to agent combustion. That is, flame extinction upon addition of each of the potential

agents is only due to simple dilution of the flame by increased heat capacity of the mixture and

consequently, reduced flame temperatures. Therefore, a fuel with a higher adiabatic flame temperature should require more added agent in proportion with its higher energy content to slow the flame

speed by a given amount (or completely extinguish the flame). We did not perform a sufficient number of simulations to quantitatively support this predicted qualitative behavior between fuels.

5.1.6 Recommendations. Of all the potential agents considered, only trifluoromethane (CHF3) was

observed in the simulations to inhibit hydrocarbon flames to any degree. All of the other potential

agents either had no influence (inert molecules), were promoters (H atom substituted), or were fiiels themselves (such as CH2F2). Consequently, we would recommend CHF3 (although it was not one of
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 507

the 12 candidate agents). However, there may be other properties that would preclude its use such as

toxicity or corrosiveness.

Of all of the specific candidate agents, we would recommend perfluoroethane (CF3-CF3), since our simulations indicate that it will extinguish flames through dilution on a par with based on their

relative heat capacities. CF3-CF3 may be preferable to N2 because it is denser than air or because of

other properties. Based on our simulations, the single H atom-substituted hydrofluorocarbons should

always be less eff'ective than their perfluorocarbon analogues (e.g., CF3-CF',H versus CF3-CF3)

because their fuel-like characteristics (heat release) will always dominate any radical scavenging.

Based on the eff'ectiveness of CHF3 in the simulations and knowledge of its mechanism for inhibition, there are other potential agents which should be considered. Any molecule that decomposes readily to yield perfluoromethyl radicals (•CF3) could provide some degree of chemical inhibition

of flame processes. That is, any molecule CF3-X which has a weak bond (C-X) between the CF3

group and any other group (denoted by X). CF3-Br is the best example (although undesirable because

of its ODP). CF3-I is another simple example. We would also recommend perfluoroacetone (CF3-

C(0)-CF3) and hexafluoroazomethane (CF3-N = N-CF3). These potential agents are known to be good

sources of perfluoromethyl radicals (•CF3) and may decompose sufficiently "early" in the flame that the •CF3 radicals generated could participate in ignition chemistry. The •CF3 radicals can be a sink

for terminating radical production through the reaction •CF3 -I- HO2 -* CHF3 + O2, which is a contributing reaction (but not the dominant one) in the effectiveness of CF3Br flame inhibition. The

•CF3 radicals will also simply compete for all radicals in the flame. These molecules are amenable to simulations, since their decomposition products (•CF3, CO, and N2) are already contained in the

mechanism which has been developed. These agents should also be readily available for use in experimental determinations of their potential effectiveness. However, there may be other properties

that would preclude their use such as toxicity, corrosiveness, or instability.

5.2 Acid Gas Formation in Halogenated Hydrocarbon-Inhibited Flames

5.2.1 Objective and Background. The objective of this part of the research project was to obtain an understanding of the chemical and physical process of acid gas formation in inhibited flames in order to predict the amount of acid gas formed in suppressed dry bay and engine nacelle fires. This

is a continuing project, and this section of the report describes results of preliminary work to date. The halogen acid or hydrogen halide HX (where X represents the halogen) is a thermodynamically stable product in mixtures containing hydrogen and halogen atoms. Since the halogenated

hydrocarbon molecules of the present inhibitors can decompose in flames through thermal decomposition or radical attack, formation of the stable halogen acid is likely. The proposed alternatives to halon 1301 (CF3Br), primarily fluorinated and chlorinated hydrocarbons, are required in much higher

concentrations than 1301 (Section 4). Consequently, they have the potential to have correspondingly

higher concentrations of decomposition by-products. The acid gases hydrogen fluoride (HF) and

hydrogen chloride (Hcl) are believed to be the most corrosive of these products.

In order to understand the formation rates of acid gases in dry bay and engine nacelle fires it is necessary to examine the thermodynamics and chemical kinetics relevant to the formation of the acid

gases as well as the eff'ects of the flow field and mixing on the chemistry. An engine nacelle fire may

be similar to a steady turbulent spray diffusion flame, whereas a dry bay fire may resemble a rapidly

advancing turbulent premixed flame. Because suppression of the dry bay fires is rapid, occurring in a time of under 100 ms, it is also be necessary to consider transient effects on the acid gas formation

process.
508 SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH

Formation of acid gases in inhibited hydrocarbon flames has been studied for many years. The

research can be categorized as either global measurements of HF produced in suppressed fires, or

detailed flame structure measurements. Burdon et ai, ( 1955) ignited mixtures of fuel, air, and

CH3Br in flasks, analyzed the products and found copious amounts of HBr. Numerous premixed low

pressure flame studies (e.g. Wilson, 1965; Biordi et al., 1973; Safieh et ai, 1982;, and Vandooren et

al,. 1988) used mass spectroscopy to measure the profiles of hydrogen halides and other products in hydrogen, carbon monoxide, and hydrocarbon flames inhibited by CH3Br, CF3Br, and CF3H. These

studies indicated conversion efficiencies of the halogens in the inhibitor into halogen acids on the

order of unity. Acid gas formation in hydrocarbon-air pool fires suppressed by CF3Br has been

studied by Sheinson et al. (1981, 1982). These studies, in test volumes of 1.7 and 650 m?, stressed

the difficulties in probe sampling for acid gases. The latter study described an in situ IR absorption

method for measuring HBr and HF. To overcome these limitations and also provide time-resolved

acid gas concentration data. Smith et al. (1993) developed a new HX sampling technique and obtained

HX and inhibitor concentrations as functions of time for discharge of CF3Br into a 56 m-^ space. In a

series of experiments with a variety of fuels and halogenated inhibitors, Yamashika (1973) showed

that the extinction time for a compartment fire sprayed with inhibitor is dependent upon the discharge

rate and room volume. He then showed (Yamashika, 1974) that the amounts of hydrogen halides and

carbonyl halides are also dependent upon the discharge rate. Using a simple model of acid gas

formation based on the steady-state rates, he developed a model of transient acid gas formation to explain his results. In more recent studies, Ferreira et al. (1992a,b) injected CF3Br, C3HF7, and C4FJ0 into an

enclosure fire and measured the HF produced using ion-selective electrodes. Di Nenno et al. ( 1993)

introduced halon alternatives into compartment fires and measured the HF, HCl, and COF2 produced

using Fourier transform infrared spectroscopy. These studies again confirmed the importance of

injection rate and fuel consumption rate on the amount of acid gas produced. Filipczak (1993)

introduced CF2ClBr and CF3Br into a methane flame and measured the O2, CO2, H2O, HF, HCl,

HBr, and unreacted inhibitor using a mass spectrometer. Hoke and Herud (1993) are currently

developing a fast-response ion-selective electrode for measuring HF and HCl produced in extinguished fires in crew compartments of combat vehicles.

Previous research related to understanding acid gas formation in inhibited flames can be seen to include both detailed flame structure measurements and global measurements of HF produced in suppressed fires. The inhibited low-pressure premixed flame studies provide detailed data on the

concentration profiles of major species as a function of position in the flame for inhibition by CH3Br,

CF3Br, and CF3H. These results provide the basis for obtaining a good understanding of the

underlying chemical kinetics of the formation of acid gases. The second category of experiments is global measurements of the amount of acid gas formed in suppressed diffusion flames. These studies

provide important information on the magnitude of the acid gases produced and allow a comparison of

the relative amount of acid gases formed by new halon alternatives. In addition, some progress has

been made (Yamashika, 1974; Smith et al., 1993) in developing engineering models of acid gas

formation in suppressed flames. There remains a need to develop a fundamental basis for interpreting

the data on acid gas formation in flames suppressed by halon alternatives, and to understand the

chemical kinetic rates of acid gas formation in flames inhibited by these alternative agents. In

particular, there exists a need to understand the relationship between fuel and inhibitor type, flame

characteristics, and the concentrations of by-products formed.

5.2.2 Experimental Approach. The formation of toxic and corrosive by-products in flames

suppressed by halogenated hydrocarbons may be controlled by transport rates of the inhibitor into the
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 509

flame, chemical kinetic rates, and equilibrium thermodynamics. These phenomena in turn will be

affected by the fuel type, local stoichiometry, inhibitor type and concentration, and the characteristics

of the flow field (mixing rates, strain, and stabilization mechanisms). The approach in this research

is to study the influence of key parameters (including the fuel type, inhibitor type and concentration,

rate of inhibitor injection, and flame type) through systematic experiments on laboratory-scale flames.

Inhibitor will be added to the fuel or air stream of co- and counter-flow diffiision flames and

premixed flames under steady-state and transient conditions. Inhibitor type, concentration, and

injection rates will be varied as will fuel type. The diflPusion flames will be operated under both

laminar and turbulent conditions to vary the mixing rates, and premixed flames will be studied over a range of fuel-air equivalence ratios. This is an ongoing project. At this point, the measurements of

steady-state acid gas formation rates in inhibited co-flow propane-air diffiision flames have been

completed. The apparatus used to obtain these data, the results, and their interpretation are presented

below.

The experiments were performed with a propane-air co-flow diffusion flame. Two burner

types were used. The first was modelled after the cup burner described by Booth et al. (1973) and

Bajpai (1974). The experimental arrangement is shown in Figure 12. The burner consists of a 28

mm diameter pyrex cup positioned concentrically in a 120 mm diameter 450 mm tall chimney at about 150 mm from the base. In these experiments with propane, the cup burner was modified for use with a gaseous fuel (liquid fuels will be tested in future experiments). The cup was filled with 1 mm diameter glass beads and covered with a stainless steel screen. The second burner consisted of a 25 cm long pyrex tube with a 0.50 mm diameter opening positioned concentrically and at the same

height as the cup burner, with the same chimney. The cold flow Reynolds number based on the exit velocity in the tube was 10,500. This second burner, referred to here as the jet burner, was designed

to provide turbulent mixing of the inhibitor in the air stream with the fuel. The air used was shop compressed air (filtered and dried) which was passed through an 0.01

micron filter to remove aerosols and particulates, a carbon filter to remove organic vapors, and a desiccant bed to remove water vapor. The fuel gas was propane (Matheson, CP grade) at flow rate of

0.114 1/min at 21 °C. Gas flows were measured with rotameters (Matheson 1050 series) which were

calibrated with bubble and dry (American Meter Co. DTM-200A and DTM-325) flow meters.

Inhibitor gases were of difl'erent purities from various suppliers.

Before measuring HF in the product gases, the concentration of inhibitor in the air stream

necessary to extinguish the flame was determined. The inhibitor was then added to the co-flowing air stream at a concentration of either 50 or 90% of the extinguishing concentration, and the product

gases were sampled for acid gas. In one series of experiments with the cup burner, the inhibitor was

added to the gaseous propane stream at 70% of the concentration which was found to extinguish the

flame.

A wet chemistry technique was used to measure the HF and HCl concentrations in the exhaust

gases from the co-flow diffiision flames. A glass furmel was placed over the chimney and the exhaust

gases passed through the 4.0 cm diameter neck. A quartz probe, centered in the neck, extracted a measured fraction of the product gases (approximately 0.5%), and directed the gases through

polyethylene sample lines to polyethylene impingers filled with water which trapped the acid gases. The sample flow was continued for a total collection time of sixty seconds. The quartz probe and

sample lines were washed with water which was returned to the impinger. The sample was tested for F' and CI' using ion-selective electrodes (Orion models 96-09 and 96-17B). It should be noted that since COF2 is know to hydrolyze rapidly in the presence of water, this technique for acid gas measurement includes F' from both HF and COF2. To reduce the effects of sampling losses reported

by other investigators, a quartz probe and polyethylene sample lines were used, the distance from the
510 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

chimney top to the bubbler was kept small ( ~ 10 cm) and the sample lines were washed with the

bubbler fluid immediately after the sample was collected.

5.2.3 Model for Acid Gas Formation. A model for the amount of acid gas formed in an inhibited

diffusion flame can be developed in a manner analogous to the determination of the local equivalence

ratio and structure for difliision flames. In the classic Burke-Schumann analysis (Burke and

Schumann, 1928), the chemical reaction is assumed to occur at a sheet. This location serves as a sink

for the fuel and oxidizer which are assumed to diffiise there in stoichiometric proportions based on

complete reaction to the most stable products (i.e., HF formation is assumed to be controlled by

equilibrium thermodynamics rather than chemical kinetics). This model will be referred to as the

stoichiometric model.

In the case of a flame inhibited by halogenated hydrocarbons, a thermodynamic analysis shows

that in equilibrium, the inhibitor readily breaks down to CO2, HX, and H2O. Consequently, the

inhibitor is assumed to be consumed like a fuel and form the most stable products. The assumptions

used in the stoichiometric model are:

1. there is complete reaction of the inhibitor molecule with fuel and air to the most stable products;

2. the inhibitor in the air stream which by-passes the flame sheet does not decompose through

interaction with the post-combustion product gases;

3. there is no loss of acid gases to the chimney walls; and

4. the product gases are perfectly mixed.

An equation for the reaction of an arbitrary hydrocarbon with air and arbitrary halogenated hydrocarbon inhibitor is:

6CO2 + XHp + r\HF + (\>HCl + eCOF^ + nCOCl^ + ^O^ + 3.76y^

In this equation a is specified as is (3 when inhibitor is present only in the fuel stream. For inhibitor

present in the air stream, (3 is determined by the concentration of inhibitor in the air stream and the

ratio of the diffusion rates for oxygen and inhibitor.

l/(p/-)-[c + l/4(J-e-y)]

Where p is the ratio of the binary diffiision coefficient of the inhibitor in nitrogen to that of oxygen

with nitrogen and r is the ratio of the concentration of inhibitor to oxygen in the air stream. An atom

balance for all species provides all of the unknown coefficients, and an estimate of amount of acid gas

formed per mole of fuel (tj + (l))/a is then readily found.

5.2.4 Results and Discussion. The acid gases produced were measured at inhibitor concentrations

of 50 and 90% of the concentration of inhibitor found to extinguish the flame when the inhibitor was

added to the co-flowing air stream in the cup burner and jet burner, and at 70% of the extinction

concentration for inhibitor added to the propane fuel stream for the cup burner. Table 5 lists the

extinction concentrations for each agent for inhibitor added to the air stream of both burners, and for
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 511

inhibitor added to the fuel stream of the cup burner. As the table indicates, the jet burner flame

typically requires about 50% less inhibitor in the air stream to extinguish the flame than the cup

burner, even for identical fuel and air flows, although there are notable exceptions: CFgBr, which

required about one fifth as much inhibitor in the jet burner than in the cup burner, and C2HF5 and the

CH2F2/C2H2F4 mixture which had nearly the same extinction concentrations. In addition to providing

the necessary extinction conditions for specification of inhibitor flows at 50 and 90% of extinction,

these results also demonstrate the sensitivity of the extinction conditions to the burner geometry.

The measured HF for these difiiision flames is shown in Figures 13 and 14. These figures

present the HF produced (moles/min) for each inhibitor for the five burner/inhibitor combinations.

The total flow was about 1 mole/min. For each inhibitor, the measured HF is plotted for the cup

burner and for the turbulent jet burner with inhibitor present in the air stream at 50 and 90% the

extinction value, and for inhibitor added to the fuel stream in the cup burner at 70 % of the value

necessary to cause extinction. The figure indicates that the amount of HF varies with both the flame

type and the inhibitor type by about a factor of five for each. Note that the fuel and air flows are held constant in these data. Hydrogen chloride was also measured in these experiments and the

results are qualitatively the same as for HF. For clarity of presentation, however, only the HF results are presented.

In order to provide insight into the controlling parameters in inhibited diffusion flames, the data

of Figures 13 and 14 are presented in an alternative form in Figures 15 to 26. In these figures, the

amount of HF produced is plotted as a function of the inhibitor concentration in air. The symbols

represent the experimental data, while the lines marked F and H represent estimates of the fluxes of

fluorine and hydrogen into the reaction zone based on the stoichiometric model described above.

Figure 15 shows the measured and estimated HF production rates in a propane-air diffusion

flame for CjF^ in the cup and jet burners. The squares are the experimental data for the cup (c) and

jet (j) burners at 50 and 90% of extinction, and with inhibition addition to the fuel stream (f) of the

cup burner of 70% of the extinction concentration. The estimated uncertainty on the HF are +10%

of the value for each data point. The curve labeled F in Figure 15 is the maximum fluoride atom

mass flux into the reaction sheet of the diflFusion flame calculated using the stoichiometric model

described above. The curve labeled F' in Figure 15 is the fluoride mass flux when the diffiision rate of the inhibitor relative to oxygen is modified to account for preferential difliision of oxygen relative to the inhibitor using the molecular weight correction factor

In this equation, , , and W- are the molecular weights of nitrogen, oxygen, and inhibitor.

The predicted fluorine and hydrogen fluxes are based on actual experimental flows which vary

slightly from run to run. The slight variations in flows cause the slight discontinuities in the F and H

curves as in Figure 16. Qualitatively, the curves F and F' are seen to increase with increasing inhibitor concentration in

air, and the mass flux of inhibitor into the reaction zone is lower when a lower rate of diffusion is used for the inhibitor. The curves labeled H and H' (coincident for C2F6) show the estimated

hydrogen atom flux into the reaction zone as a function of inhibitor concentration in the air stream. Since this inhibitor does not contain hydrogen, all of the hydrogen is from the propane, and
512 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Table 5 . Extinction concentrations in percent for halon alternatives added to a co-flow propaneair cup burner and jet burner flame

Inhibitor Extinction Concentration

in air (%)

cup jet

CF3Br 4.3 (± 0.1) 0.8 (+ 0.02)

C3H2F6 7.2 (+ 0.2) 4.0 (±0.1)

CH2F2/C2H2F4 15.2 (± 0.5) 15.5 (± 0.5)

5.0 (+ 0.2) 3.2 (± 0.1)

C2HCIF4 8.6 (+ 0.3) 4.2 (± 0.1)

C2H2F4 11.1 (± 0.3) 9.5 (+ 0.3)

C2HF5 10.2 (+ 0.3) 6.2 (± 0.2) CHCIF2 13.8 (± 0.4) 6.7 (± 0.2)

C3HF7 7.6 (± 0.2) 4.2 (± 0.1)

C2F6 9.4 (+ 0.3) 3.8 (± 0.1)

C3F8 7.5 (+ 0.2) 3.8 (± 0.1)

C4F8 7.6 (± 0.2) 5.1 (± 0.2)

increasing inhibitor in the air stream does not increase the hydrogen flux into the flame. One would

expect that the HF production rate would not be greater than the estimated flux of F or H into the

reaction zone. For this inhibitor, the flame appears to be hydrogen limited above about 5% C2F5 in

the air stream; however, when there is not enough hydrogen, the most stable product is COF2, which

is know to rapidly hydrolyze in the presence of water, and would also appear as F in the bubbler.

Consequently, the hydrogen limit may or may not exist (depending upon whether the kinetics are fast enough to form COF2 in the hydrogen-limited case).

Also shown in the figure are the experimentally measured HF production rates for the jet and

cup burners (labeled c and j respectively) at 50 and 90% of the extinction concentration of C2Fg. As

indicated, the measured quantities of HF are lower than both the fluorine and hydrogen limits, and the

measured values are closer to the estimated limits when the effects of preferential diffrision are included as described above. The results for the inhibitor addition to the fuel stream in the cup

burner (labeled f) are plotted at an inhibitor concentration of zero so that they can be included in the

figure.

When viewed in this manner, the behavior of the alternative inhibitors falls into three categories. In the first category are those inhibitors (C2F5, C3Fg, C4F10, C3F8, C2HF5, and C3HF7; Figures 15

to 20) where, at the highest inhibitor concentration tested (cup burner at 90% of extinction ), the

estimated hydrogen flux into the reaction zone is lower than the fluorine flux, and is not a strong

function of the inhibitor concentration. For these inhibitors, the HF produced does not increase

significantly when the inhibitor concentration in the air stream increases above that necessary for a hydrogen/fluorine ratio in the reaction zone of about unity (the region of where the lines marked F

and H or F' and H' cross in Figures 15 to 26). The second category includes those inhibitors
SECTION 5 . FLAME INHIBITION CHEMISTRY AND THE SEARCH 513

(C2H2F4. C2HCIF4. C3H2F6. and CH2F2/C2H2F4; Figures 21 to 24) for which the estimated H and F

fluxes are about equal. For these inhibitors, the amount of HF produced increases with increasing

inhibitor concentration in the air, but the highest concentration tested corresponds roughly to a unity

F/H ratio in the reaction zone. The last category consists of CF3Br and CHCIF2 (Figures 25 and 26)

for which the estimated hydrogen flux is much higher than fluorine flux, and there is estimated

always to be more hydrogen than fluorine in the reaction zone. For these inhibitors, the HF produced

is always increasing with higher agent concentration in the air stream. Although the stoichiometric model is very simple and is only expected to provide an upper limit on the amount of HF formed, it is instructive to investigate the possible reasons that the measured HF

production rates might be lower than the estimates. Lower HF may be measured in the experiments

due to experimental difficulties, for example: loss of HF to the chimney walls, loss in the sampling

system, HF undetected by the ion-selective electrodes, or imperfect mixing in the product gases. Based on exploratory parametric tests, these loss mechanisms are considered to be of secondary

importance. The predicted values of the HF production do not include chemical kinetic limitations

and the estimates of transport rates into the reaction zone are only approximate. Additional experiments will be performed to allow examination of these important parameters.

5.2.5 Future Work. In order to eliminate the error associated with the estimates of the inhibitor

transport rate into the reaction zone, experiments will be performed with inhibitor added to the reactant stream in a premixed burner. Adding known concentrations of inhibitor to a premixed flame

and measuring the HF production rates will eliminate the need to estimate transport rates of inhibitor

into the flame and allow a better assessment of kinetic limitations to HF formation. The premixed

flame will also be used to perform measurements of the burning rate reduction with addition of

inhibitor. These measurements will be used to validate chemical kinetic mechanisms of the reaction

of hydrocarbons, halogens, and air. Measurements of the final product species in inhibited flames

will provide insight into the kinetic mechanisms and can be used for comparison with kinetic

modeling results. Finally, detailed flame structure measurements in inhibited counterflow difiiision flames will allow comparison with the results of detailed numerical calculations of the flame structure

to provide a sound fundamental basis for estimates of HF production rates in suppressed dry bay and

engine nacelle fires. The approach to be used in developing a model for HF production in these fires

is similar to that of Yamashika (Yamashika, 1974). The experimentally determined steady-state

production rates for HF will be used to predict the time-dependent production rate of HF. The

important parameters then become the value of the extinction concentration for a given fire and the

time it takes the concentration to reach this concentration. Since chemical kinetic limitations appear

to aff'ect HF formation rates, detailed flame structure measurements together with numerical

modelling of the kinetics are necessary to provide the background for development of a global model

of HF formation.

Although results on acid gas production in JP8 cup burner flames have not yet been obtained, it

is possible to use the propane results and the model described above to obtain estimates of the amount

of acid gas (HF and HCl) that would be formed in a turbulent JP8 flame suppressed by the proposed

alternative agents, and use this estimate as a basis for ranking the agents with respect to acid gas

formation. The quantity of HX (sum of HF and HCl) formed is estimated from the steady state production rates predicted by the stoichiometric model for JP8 fuel with suppressant present at 50%

of the extinction value. Kinetic limitations are estimated from those indicated by the propane results. Table 6 provides the rank ordering of the agents and gives the fraction of acid gas the agent forms in
514 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 6. Relative ranking of alternative inhibitors with respect to steady state hydrogen halide

production rates

Inhibitor Relative Steady State Acid Gas Production

CFjBr 0.05 (± 0.01)

C2HCIF4 0.16 (± 0.03)

C2F6 0.21 (± 0.03)

C4F10 0.27 (± 0.04)

CHCIF2 0.31 ( + 0.05)

0.32 (± 0.05)

C3F8 0.34 (± 0.05)

C3HF7 0.35 (± 0.05)

C2HF5 0.41 (± 0.06)

C3H2F6 0.43 (± 0.06)

C2H2F4 0.74 (± 0.10)

CH2F2/C2H2F4 1.0 ( + 0.10)

Steady state compared to the agent which has the highest acid gas formation rate (CH2F2/C2HF5).

The agent CF3Br is listed for comparison. In practice, the amount of acid gas formed in a suppressed

fire will be dependent upon the time it takes the agent to reach the extinction concentration. For

example, the rank ordering in the list below (which is based on steady-state results) can be greatly

affected by the injection system used and the mixing characteristics of the particular agent once it is released.
SECTION 5. FLAME ESTHIBITION CHEMISTRY AND THE SEARCH 515

To Vacuum

LLi Filter

Sample Bottle

Cup Burner

Rotameter

Cold Trap

t

Bubble Flow Meter

Figure 12. Experimental apparatus for co-flow diffusion flame studies of acid gas formation in inhibited propane-air flames.
516 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 13. Measured HF production rates in co-flow propane-air diffusion flames. Data are shown for cup and jet burners at 50 and 90% of the extinction concentration for agent added to the air stream, and at 70% in the fuel stream in the cup burner.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 517

70%' fuel 50% iet 90% jet 50% cup 90% cup

Burner/Inhibitor Configuration

Figure 14. Measured HF production rates in cup and jet burners. Note that the curves for CF2H2/C2HF5 and C2H2F4 are reduced by a factor of 4.
518 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 15. HF produced vs. C2F5 mole fraction for cup (c) and jet (j) burners at 50 and 90% of

extinction, and with inhibitor addition to the fuel stream (f) of cup burner at 70% of

extinction.
SECTION 5. FLAME INmBITION CHEMISTRY AND THE SEARCH 519

Figure 16. HF produced as a function of C3Fg mole fraction in air stream. Label definitions as

in Figure 15. Discontinuities in curves are due to slight variations in flows from

run to run.
520 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 17. Moles of HF produced as a function of the C4FJ0 mole fraction in the air stream. Label definitions as in Figure 15.
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 521

0.00 0.05 0.10 0.15

C4F8 Mole Fraction in Air Stream

Figure 18. Moles of HF produced as a function of the C4F8 mole fraction in the air stream. Label definitions as in Figure 15.
522 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 19. Moles of HF produced as a function of the C2HF5 mole fraction in the air stream. Label definitions as in Figure 15.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 523

Figure 20. Moles of HF produced as a function of the C3HF7 mole fraction in the air stream. Label definitions as in Figure 15. -
524 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 21. Moles of HF produced as a function of the C2H2F4 mole fraction in the air stream. Label definitions as in Figure 15.
SECTION 5. FLAME INHmiTION CHEMISTRY AND THE SEARCH 525

Figure 22. Moles of HF produced as a function of the C2HCIF4 mole fraction in the air stream. Label definitions as in Figure 15.
526 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Figure 23. Moles of HF produced as a function of the C3H2F6 mole fraction in the air stream. Label definitions as in Figure 15.
SECTION 5. FL\ME INHIBITION CHEMISTRY AND THE SEARCH 527

Figure 24. Moles of HF produced as a function of the CF2H2/C2HF5 mole fraction in the air stream. Label definitions as in Figure 15.
528 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

0.10 -

c

E

0.00 0.05 0.10 0.15

CHF2CI Mole Fraction in Air Stream

Figure 25 . Moles of HF produced as a function of the CHF2CI mole fraction in the air stream. Label definitions as in Figure 15.
SECTION 5. FLAME ESfHIBrnON CHEMISTRY AND THE SEARCH 529

Figure 26. Moles of HF produced as a function of the CF3Br mole fraction in the air stream. Label definitions as in Figure 15.
530 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

5.3 Alternative Low ODP Chemicals with High Fire Suppression Effectiveness

The major objectives of this research were to broaden the scope of alternative agent screening beyond

the core candidates listed in Section 1 and to develop and test structure/activity relationships for environmental impact. The motivation for this pursuit was supplied by the following considerations

which could be expected to affect the outcome of the alternative agents program:

1. the original list of compounds might not include a viable candidate;

2. guidelines for environmental impact might change without warning; and

3. a new class of environmentally safe fire extinguishants might be identified.

5.3.1 Background. The mechanism by which halons extinguish fires is thought to involve both

chemical and physical processes (Sheinson et al., 1989 and Pitts et al., 1990). Simulations using

detailed kinetic models (Westbrook, 1983) indicate that hydrogen halides (HX), which form when

halons are introduced into flames, promote radical recombination via HX + H (OH) H2 (H2O) +

X. This inhibits the combustion process by reducing the concentration of radicals in the reaction zone. At some point, the rate of recombination exceeds the rate of propagation and the flame is extinguished. Physical suppression results from the removal of thermal energy from the flame by the

vibrational excitation of the agent. This mechanism is universal in the sense that all molecules absorb

heat, however, it is much less efficient than chemical suppression for the low molecular weight

compounds which are most suitable for total flooding applications. Those halocarbons which contain

Br or I are chemically active and are considerably more effective (per unit mass) than the corresponding F and Cl-containing compounds which act predominantly as physical suppressants (see Table 7 in Section 4).

The current commercial halons pose a serious threat to the ozone layer. The determining factor

is whether they are sufliciently stable to make it into the stratosphere where over 90% of the column

height of atmospheric ozone is located (Finlayson-Pitts and Pitts, 1986). Once they enter the

stratosphere, halon molecules are photolyzed by short wavelength UV radiation. The halogen atoms

released in this process are capable of catalyzing the destruction of ozone through the following

mechanism:

xo +0 - + X,

where X is CI, Br or I. The precise definition of ozone depletion potential (ODP) (Wuebbles, 1983) is

AOJRX) ... ODP(RX) =

, (5)

AO^iCFCl^)

where AO3 is the amount of ozone destroyed by emission of compound RX over its entire atmospheric lifetime. This quantity is computed by modeling molecular transport, typically as a function of

both latitude and altitude, through the troposphere and into the stratosphere. Chemical and photolytic

processes, which can remove the compound from the atmosphere, are accounted for by kinetic
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 531

mechanisms which are parameterized on the basis of measured rate constants. The computed ODP

depends on the following critical factors:

1 . the fraction of molecules which reach the stratosphere, which in turn, depends on how long the

molecule survives before it is oxidized or otherwise destroyed in the troposphere;

2. the number of CI, Br and I atoms in the molecule; and

3. their relative efficiencies in catalyzing the destruction of O3 according to the mechanism

summarized in Equation (1).

The catalytic efficiency of Br is approximately 40 times greater than that of CI (Pyle et al.,

1991). The conversion of halogen oxides into free halogen atoms via

BrO + CIO - CI + Br +0 (6)

is responsible for a large part of this enhancement. The kinetic constants for the analogous reactions

of 10 with CIO and BrO have not been measured. There is an indication, however, that low levels of

I may greatly enhance the ozone destruction efficiencies of both CI and Br (Huie, 1993), which are

far more prevalent than I in the stratosphere.

The ozone depletion potential of fluorocarbons is essentially zero because, unlike the other

halogens, atomic F is quickly converted to the hydrogen halide, HF, which is chemically and

phololytically inert. On the other hand, any organic molecule which contains F will absorb infrared

radiation in the atmospheric window region between 800 and 1400 cm"^ and will, therefore, have a nonzero global warming potential (GWP). DiflFerences in the GWPs of these compounds are determined primarily by how long they survive in the troposphere.

On the basis of this analysis it is apparent that the most effective suppressants will contain the ozone depleting halogens Br or I. These compounds will almost certainly contain F, as well, since

the substitution of F for H atoms is usually necessary to reduce flammability and increase volatility.

On this basis, it is clear that low values of both ODP and GWP will be realized only in compounds

which have short tropospheric lifetimes (t]). Thus, in the process of identifying replacement agents,

their propensity to engage in one or more of the following processes which can remove these

chemicals from the troposphere was a major concern:

1. reaction with OH radicals;

2. photolysis by long wavelength (> 290 nm) UV-VIS radiation; and

3. hydrolysis and/or solvation in H2O.

5.3.2 Methods. An exploratory list of 103 potential halon alternatives representing 10 distinct

chemical families was cited in NIST TN 1279 (Pitts et al., 1990). This document served as the point

of departure and was augmented as new possibilities arose during the course of this investigation.

A short list was constructed from the most promising exploratory compounds. These candidates were selected on the basis of fire suppression effectiveness , suitability for total flooding applications

in aircraft dry bays and engine nacelles, commercial availability and ODP. Other properties,

including GWP and acute toxicity were also taken into consideration.

The results of cup burner and other fire suppression measurement techniques were available for many of the exploratory compounds (Purdue Research Foundation, 1950; Sheinson, et al. 1989;

Zallen, 1992). Additional cup burner measurements (see Table 7 of Section 4) were conducted on

representative compounds from chemical families which were rated highly with respect to the other
532 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 7. ODP, GWP and t, of Selected HCFCs

Formula HCFC ''1 ODP GWP

CHCI2CF3 123 1.6 0.016 0.19

CH3CCI3 NA 6.3 0.11 0.024 CHCIFCF3 124 6.6 0.019 0.10 CH3CCI2F 141b 7.8 0.080 0.092 CHCIF2 22 15.3 0.051 0.36

CH3CCIF2 142b 19,1 0.056 0.37

Values taken from Fischer et al., 1989.

selection criteria but for which there were no data relating to their eflFectiveness as flame suppressants.

Compounds requiring suppression concentrations significantly higher than the core candidates were

eliminated from further consideration.

Volatility is a critical factor in agent dispersal. Compounds having boiling points above ambient

temperature (25 °C), with the possible exception of powders, are not suitable for total flooding

applications and were not considered any further. This investigation was primarily directed at finding near-term solutions to the urgent problem

created by the production phaseout of halon 1301. The time and expense involved in developing the

capability to produce a new agent was considered prohibitive. Consequently, any compound

appearing on the exploratory list which could not be purchased from a commercial vendor was

eliminated from further consideration. It will be beneficial to remove this limitation in future

investigations directed at identifying second-generation halon replacement candidates.

The hypothesis that T| is a reasonable predictor of both ODP (Herron, 1990) and GWP (Herron;

1990) is confirmed by the data listed in Table 7 and plotted in Figures 27 and 28. Qualitative

assessments of T|, and consequently of environmental impact, can be made on the basis of SARs

pertaining to the specific chemical family under consideration. Thus, for example, saturated

halocarbons (SH) can react with OH if they have a labile C-H bond. The presence of certain

functional groups, such as an ether (HE) or carbonyl (HK) oxygen, increase the absorption cross- section for UV-VIS radiation and can activate C-H bonds for abstraction of H by OH (Zhang et al. , 1992a). Similarly, OH radicals are known to attack ir bonds in unsaturated halocarbons (UH). Many

inorganic compounds, including the silicon, germanium (SG), sulfur (S) and phosphorous (P) halides

have a tendency to hydrolyze, while metallic compounds (M) are frequently water soluble. The

presence of a heavy halogen, particularly 1 (HI), shifts the absorption spectrum of the compound to the red and thereby increases the likelihood that these compounds will photolyze in the troposphere.

This is especially true if there are multiple halogens bonded to the same carbon.

More reliable assessments of t, can be made on the basis of laboratory measurements. At NIST,

rate constants for reaction with OH (k^^ ) have been measured by flash photolysis resonance

fluorescence (Kurylo and Braun, 1976). In this technique, a small fraction of the OH radicals

generated by photolysis of H2O are excited by exposure to a high intensity lamp. The time dependent

concentrations are monitored by detecting the emissions of resonant radiation from OH at 308 nm

with a photomultiplier. A discharge flow technique coupled with electron paramagnetic resonance

detection of OH has also been used to measure the OH rate constants for many halogen-containing

compounds (Orkin and Khamaganov, 1993a). In either case, the rate constants for the pseudo first- order reaction {i.e. [RX] > > [OH]) are given by
SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH 533

CM

dOO

Figure 27. ODP, normalized by dividing the computed value by the number of CI atoms in the

molecule, plotted as a function of tropospheric lifetime.
534 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

M CM T- O

d o o d

dMO

Figure 28. GWP plotted as a function of tropospheric lifetime.
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 535

[RX\ dt

where RX denotes the compound under investigation and square brackets are used to indicate

concentrations. According to Ahrrenius theory, the temperature dependence of the rate constant is given in Equation (8).

k = A exp(^), (8)

RT

where R is the gas constant and T is the temperature in Kelvins. The kinetic constants, A and E^, are the bimolecular collision frequency and activation energy (energy barrier for the reaction), respectively.In lieu of performing simulations with an atmospheric model, tropospheric lifetimes were

estimated from

= ro'j^, (9)

' ^,(277)

where ky^Q{211) and k5(277) are the rate constants for the reactions of methyl chloroform (CH3CCI3)

and the alternative agent with OH at 277 K and 7.0 is the lifetime (in years) of CH3CCI3 (Zhang et

al., 1992b). Methyl chloroform is used as a standard because it has an intermediate T|, which is useful for modeling the atmospheric transport of a wide range of compounds, and it is known to have

only anthropogenic sources. Kinetic constants for some representative compounds are listed in

Table 8. Experimental measurements of OH reactivity are very expensive. Painstaking efforts have to be

taken to ensure sample purity because trace amounts of a reactive contaminant will invalidate the

results. In the present investigation, measured rate constants were supplemented with values obtained

from correlations between t, and C-H bond dissociation energies (BDEs). The basis of these

correlations is the hypothesis that the rate determining step in the reaction between OH and saturated

halocarbons is the abstraction of a H to form H2O. This process requires breaking a C-H bond in the

halocarbon. The natural logarithm of T] (Fischer et al., 1989) is plotted as a function of the smallest

C-H BDE (McMillan and Golden, 1982) in the molecule for a series of HFCs in Figure 29. The

linearity of this plot confirms the validity of this hypothesis.

Experimental determinations of BDEs are difficult to make and are subject to large errors, however, recent advances in the theory and application of molecular quantum mechanics have made

the calculation of chemically accurate BDEs in small halocarbons almost routine (Curtiss et al.,

1991). The C-H and C-X BDEs were computed for some small molecules using a variety of ab initio methods including: local (LDA) and nonlocal density functional approximations (NLDA); Hartree- Fock (HF) and HF corrected to fourth order in the Moller-Plesset perturbation expansion of the

energy (MP4); and the Gaussian-Gl (Gl) and -G2 (G2) procedures. The results of these calculations

are listed in Tables 9 and 10. The rms error from the experimental BDEs (McMillan and Golden,

1982) for the fluoro- and chloromethanes corresponding to the most accurate calculations (G2) are

less than 5 kJ/mole and 10 kJ/mole, respectively.
Figure 29. Tropospheric lifetime plotted as a function of C-H BDE.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . . 537

Table 8. OH Kinetic Constants and T]'s of Selected Compounds

Compound A (X 10'^)^ E,/R (K) ri(years)''

Substitued Methanes CH4 7.0 2061 12.8 CHF3 0.49 2115 192

CHF2CI 0.81 1516 14.0 CHFjBr 0.93 1326 6.8

Substituted Ethanes'^ CH3CH3 7.8 1020 0.27

CH2FCF3 1.03 1588 14.3

CH2BrCF3 0.85 1113 3.4

CHClBrCFj 1.28 995 1.6

Substituted Ethenes CH2=CH2 2.03 -411 0.0047 CH2=CHBr 1.8 -405 0.0068

CF2 = CCln 1.5 -480 0.0062

a. In units of cm^ molecule s .

b. From Equation 9.

c. Orkin and Khamaganov, 1993a and 1993b.

Table 9. C-H and C-F BDEs in the Fluoromethanes, kJ/mol

Bond LDA' NLDA' HF'' MP4'' Gl'^ G2'' Exp.*^

CH3-H 494 464 362 458 434 435 440

CH2F-H 464 447 353 440 422 424 424

CHF2-H 458 496 354 441 422 424 432

CF3-H 482 461 462 462 443 445 447

CH3-F 570 478 273 447 462 463 460

CH2F-F 599 502 322 488 503 505 500

CHF2-F 621 565 359 520 535 537 534

CF3-F 629 510 374 533 548 551 546

aBiosym Technologies, Inc., 1992.

bFrisch, M.J., et al , 1992. cMcMillanand Golden., 1982.
538 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 10. C-H and C-Cl BDEs in the Chloromethanes, kJ/mol

Bond LDA^ NLDA^ HF'' MP4'' Gl'' 02^

CH3-H 494 464 362 458 434 435 440

CH2CI-H

CH2CI-H

4654 4444 3513 4184 4164 4164 4224

CHCI2-H 431 Aid 419 338 401 A r\ t 401 401 414

/^/^l T T CCI3-H 415 401 322 385 390 390 401

CH3-CI 376 351 226 340 349 347 354

CH2CI-CI 334 320 200 318 337 334 335

CHCI2-CI 304 286 165 291 321 315 325

CCI3-CI 272 247 128 265 301 297 306

^Biosym Technologies, Inc., 1992.

^Frisch, M.J. et al., 992.

'^McMillan and Golden, 1982.

Table 11. Photolytic Rate Constants and t{s of Selected Compounds^

Compound K T] (years)''

CF2ClBr 2.2 X

10-^ 14

CF2Br2 2.3 X

10"^ 1.3

CF3I 6.4 X

10"^ 0.0047

^Herron, 1990 and references therein.

''From Equation 1 1

.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 539

Rate constants for photolysis were obtained from

= ffodX, (10)

where the integral is over the wavelength range of atmospherically accessible solar radiation, f is the

solar flux of photons, and a is the measured UV-VIS absorption cross-section. Values for representative compounds are listed in Table 11. In this approximation, the quantum yield for photodissociation

is assumed to be unity. Tropospheric lifetimes were estimated from

where the factor in the numerator converts from, seconds to years.

5.3.2.2 Exploratory List. The plan of work for this task was based on a strategic organization

of the candidates into four working categories. The first of these, hereafter referred to as Category I, comprise the 41 halocarbons identified in NIST TN 1279 which contain at least one CI, Br or I atom.

These compounds are listed in Table 12. Unfortunately, all Category I compounds have some

potential for ozone depletion and global warming. The expectation was, however, that the environmental impact of some of these compounds would be acceptable due to their reactivity with OH

radicals, or their tendencies to undergo photolysis in the troposphere.

The Categor>' II candidates, which are listed in Table 13, are the complementary set of

halocarbons which contain F as the only halogen. As a group, fluorocarbons are much less effective

as fire suppressants than are the analogous compounds containing Br or I (see Table 7 of Section 4).

All of the compounds in this category are considered to have zero ODP; however, GWP was an

important consideration.

Category III consists of the halogen-containing inorganic compounds identified in NIST TN 1279

(Table 14). Many of these compounds, for example Cr02Cl2 and SnC^, are highly toxic and are, therefore, not viable halon replacement candidates. The justification for their further consideration was that some are extremely effective flame inhibitors. The expectation was that new and important

principles would be learned in the process of evaluating and comparing their effectiveness as fire suppressants and that these insights could be immediately applied to further the primary objectives of

this project.

The non-halogenated inorganic compounds identified in NIST TN 1279 were placed in Category

rv. They are listed in Table 15. All have zero ODP and minimal GWP.

5.3.3 Results. The selection process was carried out in two steps. First, compounds were

eliminated from the exploratory list on the basis of boiling point (BP), commercial availability (CA),

and other readily available data. Core candidates (indicated by the presence of an asterisk) were

evaluated elsewhere in this report (see Section 10.2) and were not explicitly included in this process.
540 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . .

Table 12. Category I Compound;

Compound Formula Chemical Family Basis for Elimination or Further ConsideiBtion

CBr2F2 SH ODP"

SH bp''

CHC1F2* SH see Section 10.2 SH BP

^2 2 3 SH BP

C2HCIF4* SH see Section 10.2

^2"3^'2' SH BP

C2H3CIF2 SH FSN"^ SH GDP

CHBrClF SH BP

C2HBrClF3 SH BP

C2HBrClF3 SH BP

C2HBrF4 SH O A d CA"

C2H2BrF3 SH BP

SH BP

\-2ii2"'2' 2 SH BP

SH BP

CjHBrFg SH BP

^3^2^^2^4 SH BP

^3^2^^2^4 SH BP

CHiBrFc SH BP

HK BP

BrCF^rOOrFo HK BP

CF3COOCH2Br HK BP

BrrF^CF = CF^ UH BP

UH BP

UH BP

PHF^orF^rHFri OV-ZV^i TV_-111 1 HE BP -

PPn PHPIOPHF ^ n\- v_ nr 2 HE BP

CF2BrOCF2Br HE CA

CF2BrOCF3 HE CA

CBrF20CF2CHFCl HE BP
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH ... 541

Table 12. Category I Compounds

Compound Formula Chemical Family Basis for Elimination or Further Consideration

BrC4F70 HE BP

CF3I HI short list

CF2CII HI CA

CF2BrCF2l HI BP

CF2ICF2I HI BP

CH3I HI BP

CH3CH2I HI BP

CH3CH2CH2I HI BP

CF3CF2CF2I HI BP

a. ozone depletion potential

b. boiling point

c. fire suppression number (see the definition given in Section 4.5.5.3)

d. commercial availability

* core candidate
542 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Table 13. Category II Compounds

("'nmnni V_. \,) 1 IIL/\JlillU i nH Family

CF4 SH

C2F6* SH

C3F8* SH

C4F1O SH

C4F8* SH

CHF3 SH

C2HF5' SH

C2H2F4 SH

CF3COCF3 HK

CF3COOCOCF3 HK

(iC3F7)2CO HK

CF3COOCH3 HK

C3F6 UH

C4F8 UH

CFjC^Hs UH

CF3CH = CF2 UH

CF2CH = CHCF3 UH

CF3CH=CH2 UH

(C4Fg)CH = CH(C4F9) UH

(CF3)2C = C(CF3)2 UH

CHF2OCHF2 HE

CHF2OCF3 HE

C4F9C4F7O HE

C4F8O HE

CF3SCF3 HE

(CF3)3N HE

Basis for Elimination or Further Consideration

FSN

see Section 10.2

see Section 10.2

see Section 10.2

see Section 10.2 FSN

see Section 10.2

see Section 10.2 FSN

BP

BP

BP

short list

FSN

BP

short list FSN

FSN

BP

BP

FSN

FSN

BP

BP

CA

CA
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . . 543

Table 14. Category III Compounds

Compound Chemical Family Basis for Elimination or Further Consideration

S FSN

SF5CI S FSN

SFgBr S CA

PF3 P hazardous

PCI3 P BP

PF2Br P CA

POF3 P hazardous

POCI3 P BP

P0F2Br P BP

SiCl4 SG BP

SiCl3F SG CA

SiF4 SG short list

SiBrF3 SG CA

SiBr3p SG BP

Si(CH3)3Cl SG BP

SiCH3Cl3 SG BP

Si(CH3)3CH2Cl SG BP

GeCl4 SG BP

Cr02Cl2 M BP

SnCl4 M BP

TiCl4 M BP

The major reason for the elimination of each exploratory compound is noted next to its chemical

formula in Tables 12-15. Although values for ODP and relative fire suppression efficiency, as indicated by fire suppression number (FSN) (Section 4.5.5.3), could not be found for every com- pound on the exploratory list, a reasonable assessment could usually be made on the basis of

measured values obtained for other members from the same chemical family. Those candidates which

were considered to merit further study, as well as, a few additional possibilities which arose during

the course of this investigation, were included in a short list of compounds. These compounds were

evaluated in the second step of the search process.
544 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Table 15. Category IV Compounds

<^ompouna Chemical Family Basis for Elimination or Further Consideration

Si(CH3)4 SG BP

Ge(CH3)4 SG CA

NaHC03 M see Section 10.2

NaC2H3C02 M BP

KHCO3 M BP

K2C204«H20 M BP

KC2H3CO2 M BP

K2C5H702«'/2H20 M BP

Cr(C5H702)3 M BP

Pb(C2H5)4 M BP

Fe(C0)5 M BP

N2 I FSN

CO2 I FSN

Ar I FSN

5.3.3.1 Development of the Short List. The fire suppression characteristics of the chlorocarbons are comparable to their fluorocarbon analogs (see Table 7 of Section 4). Two of these

compounds, CF2 = CC1F and CHFCI2, appeared to offer promising atmospheric properties, as well. The ozone destruction efficiency of CI is significantly less than that of Br. Thus, for example, the

ODP of CHF2CI is only 0.055, whereas the ODP of CHF2Br is 1.4 (Pyle et a/. ,991). The presence

of a X bond ensures that the r, of CF2 = CC1F will be short (see Table 8) which suggests that both the

ODP and GWP will be low. The computed BDE for the C-H bond in CHFClj is 398 kJ/mole which

is 17 kJ/mole less than the C-H BDE in CHF2CI. On the basis of the correlation in Figure 29, this effect should correspond to an order of magnitude reduction in T]. Consequently, the ODP of

CHFCI2 should be significantly lower than 0.055 even though it contains twice as many ozone

depleting halogens as CHF2CI.

The brominated methanes, particularly CF2Br2 and CFBr3, were of further interest. The

presence of multiple Br atoms ensures a low FSN and imparts a significant cross-section for the

absorption of long wavelength UV-VIS radiation which, in turn, increases the likelihood that these

compounds will be photolyzed in the troposphere (see Table 13). The possibility that one of the

brominated ethanes might be viable was raised by recent results obtained for CF3CHBrCl (Orkin and

Khamaganov, 1992b). The best estimate ODP for this compound is 0.14 (Pyle et a/., 1991) which is almost at the upper limit of acceptability. The brominated ethenes, CF2 = CFBr and CF2 = CHBr, are

reactive with OH and were also expected to have acceptable ODPs.
SECTION 5. FLAME DJHIBITION CHEMISTRY AND THE SEARCH . . . 545

Table 16. The Short List of Candidates

Family Chemical Formula Basis for Elimination or

Category I

Chlorinated Methanes roJN (2.9)

Chlorinated Ethenes

CiUIIlIllalcU IViclllollCo pup Rr Rr PFRr \JUr \l.'t, L.ZJ)

BP (107 °C)

Brominated Ethanes CH2BrCF3 BP (26 °C)

Brominated Ethenes CF2 = CFBr, CF2 = CHBr FSN (2.3, 1.9)

lodocarbons CF3I recommended

Category II

Fluorinated Ethenes CF2 = CFCF3 FSN (2.5)

Category III

Silicon Compounds SiF4 FSN (3.3)

Nitrogen Compounds NF3 FSN (oxidizer)

The iodocarbons are effective suppressants (Nimitz and Lankford, 1993) and they are known to be light sensitive, which suggests that they will be photolyzed in the troposphere. The iodine analog

of halon 1301, CF3I, was considered particularly attractive because it also has a low boiling point

(-22.5 °C) which makes it suitable for total flooding applications.

All of the compounds in Category II have zero ODP; however, the high GWP of the saturated

fluorocarbons on the exploratory list might play an important role in determining whether they can be

used as fire extinguishants. Many of these compounds are core candidates and were already

undergoing thorough testing. Thus, in an efl'ort to avoid a duplication of effort, the decision was made to focus on the unsaturated fluorocarbons. Particular attention was placed on compounds, such

as CF2 = CHCF3 and CF2 = CFCF3, where the C-F bond might be weakened by resonance stabiliza- tion in the radical. The rationale was the hypothesis that the exceptionally strong C-F bonds, which

characterize most of the fluorocarbons, might have an adverse effect on their effectiveness as flame

suppressants.

Similar reasoning was used to justify the further consideration of two inorganic halides, SiF4 and

NF3. The former compound readily hydrolyzes to HE and silanol in the atmosphere, whereas the NF bonds in the latter molecule are exceptionally weak.

5.3.3.2 Recommendations. The short list is presented in Table 18 along with the primary

reason for the elimination of each candidate. The FSNs of the chlorinated methanes and ethenes, the

fluorinated ethenes, and SiF4 were noticeably worse than the corresponding values obtained for the

leading candidates from the core list. The brominated ethenes were a little better, but not enough to warrant the recommendation of a Br-containing compound. NF3 was disqualified because it behaved
546 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

as an oxidizer resulting in a visible increase in the intensity of the cup burner flame. Two of the

compounds on the short list, CHFjBr and CF2Br2, were eliminated because their ODPs were

unacceptably high, as were the boiling points of CFBrj and CH2BrCF3. The boiling point of the

former compound was mistakenly reported as 10.7 °C in some of the relevant literature (Zallen,

1992). The actual boiling point of CFBr3 is 107 °C. On the basis of the reasoning presented in this section, it was determined that CF3I is the only candidate which merits further consideration with

respect to the near term applications of the sponsor. The concentration of CF3I required to extinguish the cup burner flame was 3.2 + 0.3% by

volume. This is almost identical to the value obtained for halon 1301 (3.1%). The UV-VIS cross- section for CF3I is displayed in Figure 30. The short of this molecule is attributable to the tail of

this curve which extends to wavelengths beyond 290 nm where there is an appreciable solar flux at ground-level

.

Detonation and deflagration suppression tests (see Section 4.5) have indicated that CF3I is the

most eff"ective agent tested (including halon 1301) at mole fractions less than about 3%. Its eff'ective- ness in suppressing detonation, however, diminished at higher concentrations, eventually becoming

one of the least efi'ective agents at concentrations exceeding 6%. The C-I bond in CF3I is very weak.

The results obtained from photofragmentation translational spectroscopy indicate that the corresponding BDE is about 223 kJ/mole (Felder, 1992). This suggests that reactivity, particularly when in the

presence of metals which can function as heterogeneous catalysts, may present a significant problem.

Indeed, the results of coupon corrosion tests have indicated the possibility of aggressive interactions

between CF3I and both AM355 and CD 172 (see Section 7.2).

Atmospheric modeling of CF3I was performed by the Atmospheric and Geophysical Sciences

Division at Lawrence Livermore National Laboratory. This efl'ort was intended to be a first attempt

at a solution to a diflRcult problem. Computed values for the ODP of CF3I , particularly for releases

at altitude, should be considered as preliminary estimates due to the presence of a number of serious

data gaps in the stratospheric chemistry of iodine. These include (Miziolek, 1993):

1 . the absence of rate constants and branching ratios for the critical reactions of 10 with BrO and

CIO;

2. no measurements of the rate of photolysis of 10; and

3. no studies of the temperature dependencies of the reactions of 10 with HO2 and O.

The calculations performed at LLNL were based on the extended lOx chemistry which uses bromine

analog rate constants for the reactions of 10 with CIO, OH and NO3, as well as, for the reaction of

HI with O. The values obtained for the ODP and tropospheric lifetime corresponding to a groundlevel release were 0.011 and 1.15 days, respectively. These findings are consistent with the previous

estimates made by Susan Solomon. The LLNL group also performed a simulated high altitude release

of CF3I at 19.5 km (over latitudes extending from 23.5 to 58.2 degrees N). They calculated an ODP

of 13.7 and a lifetime of 9.05 hours. Please note that the ODP value for the release of CF3I at altitude is very close to the ground level ODP of CF3Br ( ~ 16). This is an artifact of the extended

lOx chemistry used in the LLNL model which assumes that the stratospheric chemistry of I is essentially identical to that of Br.

5.3.4 Conclusions. CF3I is a viable candidate for the replacement of halon 130 L This is the first iodine-containing compound to be seriously considered. Other members of this family may have

applications as fire extinguishants, and as refrigerants and foam blowing agents, as well. Every elTort should be made to resolve the uncertainties in the stratospheric chemistry of iodine so that there is a

clear path to examine all of the potential uses of these compounds.
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 547

CD n r- N ro

r- V- CM P4 CM CN

I I I I I I o o o o o o

UIO 'UO]']09£ SSOJ3

Figure 30. Absorption Cross-Section for CF3I. The smooth curve was obtained by fitting the

discrete absorption data to 3 Gaussians (dotted lines).
548 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

The research reported in this section is the first step in a rationale approach to the search for new

and more effective fire fighting chemicals. Additional research, directed at developing the capability

to predict key properties on the basis of molecular structure, is ongoing and will determine the

prospects for identifying a new generation of zero ODP, low GWP fire suppressants.

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and Nugaav,T .B.K., "Pulse laser pyrolysis - mathematical model and investigation of difluorochloromethane conversion kinetics, " Vestn. Mosk. Univ. Ser. Khim. 31, 329 (1990).

Zhitnev, Y.N., Zakharchenko, A.V., Mordkovich, N.Y., Nugaev, T.B.H., Tveritinova, E.A.,

and Timofeev, V.V., "The pulsed homogenous laser pyrolysis: determination of arrhenius parameters

of CHCIF7 destruction basing on the model of physical and chemical processes," Laser Chem. 11,

71-81 (1991).

SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 561

Appendix A. Tabulation of Ab Initio Thermochemical Data for Fluorinated

Hydrocarbons.

Thermochemical data calculated using the BAC-MP4 ab initio method for fluorinated hydrocarbons

(and a few other reference species) are presented in a "JANAF style" tabular format. Values for principal moments of inertia (I^, Ij,, L in kg m"^), fundamental vibrational frequencies in cm"'),

equilibrium bond distances (Rg in 10 m), equilibrium bond angles (6^ in degrees), heat of formation

at 0 K (AHf! 0 K, 1 atmosphere in kJ mol'^), heat of formation at standard state (AH^: 298.15 K, 1 atmosphere in kJ mol"'), and entropy at standard state (S°: 298.15 K, 1 atmosphere in J K'^ mol"')

are given for each species. Heat capacity (0° in J K"' moK^), free energy of formation (AG" in kJ

mol"'), equilibrium constant for formation (Log Kf, dimensionless), and other ideal gas phase thermal

functions (as a function of temperature) are also given for each species.
562 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

HF

AH2(0 K) = -273.22 kJ mol"^

AH^(298.15 K) = -273.29 kJ mol"^

S°(298.15 K) = 173.37 J K'^ mol"^

Moments of Inertia (10"^^ kg m^) 1.321.32

Vibrational Frequencies (cm"^)

3890.2

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Fl-Hl 0.911

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

|- J K"^ mol"^ -| I kJ mol"^

Ta. //ifi\. S° H°-H AU o

f AG^ Log

J J w io J> . J o - 4 O X - 9 7 7 9 9 -274. 30 ^ 9 ^9 nn\j \j 0 Q 1 0 X O -L . _ o^ . o o - 9 7 7 9 9 -274 . 68 7 1 / X 74

^ \j ^ -3 1 n X o o . _ 1X . 4 n - 9 7 7 9 P, -275 . 06 c; 7 A 7

7J nU n\J 9 Q . X u n n - 9 7 7 7 n -275 . 39 4^ 7/ 7J DR nU 9 Q . 1X nu X / o . 04. nX . C 1 O X - 9 7 7 7 A -275 . 77 4^ 1X -L D 400 29 . 10 181

.

92 2 . 96 -273 . 42 -276 . 10 36 05

450 29 . 10 185 . 35 4 . 42 -273 . 51 -276 . 44 32 09

500 29 . 12 188 . 42 5 . 87 -273 . 63 -276 . 73 28 91

600 29 . 17 193 . 73 8 . 79 -273 . 88 -277 . 36 24 14

700 29 .28 198 . 23 11

.

71 -274 . 18 -277 . 90 20 74

800 29 .47 202 . 16 14 . 65 -274 .47 -278 . 40 18 18

900 29 . 74 205 . 64 17 . 61 -274 .81 -278 . 86 16 18

1000 30 . 07 208 . 79 20 . 60 -275 . 10 -279 . 32 14 59

1100 30 .45 211

.

68 23 . 62 -275 .39 -279 . 70 13 28

1200 30 . 84 214 . 34 26 . 69 -275 .68 -280 . 08 12 19

1300 31 . 24 216 . 83 29 . 79 -275 . 93 -280 . 45 11 27

1400 31 . 63 219 . 16 32 . 94 -276 . 19 -280 . 79 10 48

1500 32 . 01 221

.

35 36 . 12 -276 . 44 -281

.

12 9 79

1600 32 .38 223 . 43 39. 34 -276 . 65 -281

.

42 9 19

1700 32 . 71 225 . 40 42 . 59 -276 . 86 -281

.

71 8 66

1800 33 . 03 227 . 28 45 . 88 -277 . 02 -281

.

96 8 18

1900 33 .33 229 . 07 49 . 20 -277 . 23 -282 . 25 7 76

2000 33 .60 230 . 79 52 . 54 -277 .40 -282 . 50 7 38

2100 33 . 85 232 . 44 55 . 91 -277 . 52 -282 . 75 7 03

2200 34 . 08 234 . 02 59 . 31 -277 .69 -283 . 01 6 72

2300 34 .29 235 . 54 62 . 73 -277 . 82 -283 . 26 6 43

2400 34 .48 237 . 00 66 . 17 -277 . 94 -283 . 47 6 17

2500 34 . 66 238 . 41 69. 63 -278 . 07 -283 . 72 5 93

2600 34 . 83 239 . 78 73 . 10 -278 . 15 -283 . 93 5 70

2700 34 . 98 241

.

09 76 . 59 -278 .28 -284 . 14 5 50

2800 35 . 12 242 . 37 80 . 10 -278 . 36 -284 . 39 5 30

2900 35 .25 243 . 60 83 . 62 -278 .45 -284 . 60 5 13

3000 35 .37 244 . 80 87 . 15 -278 . 53 -284 . 80 4 96
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 563

F2

AH|(0 K) = 1.21 kJ mol"^

AHf(298.15 K) -' = 1.13 kJ mol"^

S°{298.15 K) = 201.42 J K'^ mol'^

Moments of Inertia (10'^^ kg m^)

28.53 28.53

Vibrational Frequencies (cm'-"-) 1111.7 _

Bond Distances (10'^° m) Bond Angles (°) Dihedral Angles (°) F1-F2 1.345

Enthalpy Reference Temperature

Standard State Pressure = P

1

1 J mol 1 _ T/K C° s

0 150 29 I2 181 . 18

200 29 . 28 189 . 58

250 29 . 67 196 . 15

300 30 . 25 201 .61

350 30 . 94 206 . 32

400 31 . 63 210 . 50

450 32 . 2 8 214 .26

500 32 . 87 217 .69

600 33 . 85 223 . 78

700 34 . 58 229 . 05

800 35 . 12 233 . 71

900 35 . 53 237 . 87

1000 35 . 85 241 .63

1100 36 . 09 245 . 06

1200 36 .29 248 .21

1300 36 .44 251 . 12

1400 36 . 57 253 . 82

1500 36 .67 256 .35

1600 36 .76 258 . 72

1700 36 . 83 260 . 95

1800 36 .89 263 . 06

1900 36 . 94 265 . 05

2000 36 . 99 266 . 95

2100 37 . 02 268 . 76

2200 37 . 06 270 .48

2300 37 . 09 272 . 13

2400 37 . 12 273 . 70

2500 37 . 14 275 . 22

2600 37 . 16 276 . 68

2700 37 . 17 278 . 08

2800 37 . 19 279 .43

2900 37 . 21 280 . 74

3000 37 . 22 282 . 00

= Tj. = 298.15 K

1 atmosphere

kJ mol'^

H°-H°(T^) AHf AG° Log

-4 37 1 .21 1 . 34 -0 47

-2 92 1 . 17 1 . 38 -0 36

-1 44 1 . 17 1 . 42 -0 30

0 05 1 . 13 1

51 -0 26

1 59 1 . 09 1 . 55 -0 23

3 15 1 . 05 1 . 63 -0 21

4 75 0 . 96 1 . 72 -0 20

6 38 0 . 92 1 . 80 -0 19

9 72 0 . 84 1 . 97 -0 17

13 14 0 . 75 2 . 13 -0 16

16 62 0 . 71 2 . 34 -0 15

20 16 0 . 63 2 . 55 -0 15

23 73 0 . 59 2 . 80 -0 15

27 33 0 . 54 3 . 01 -0 14

30 94 0 . 50 3 . 22 -0 14

34 58 0 .46 3 . 47 -0 14

38 23 0 .46 3 . 68 -0 14

41 89 0 .42 3 . 93 -0 14

45 57 0 .42 4 . 14 -0 14

49 25 0 .38 4 . 39 -0 14

52 93 0 .38 4 . 64 -0 13

56 62 0 .38 4 . 85 -0 13

60 32 0 . 33 5 . 10 -0 13

64 02 0 .33 5 . 36 -0 13

67 72 0 .33 5 . 56 -0 13

71 43 0 . 29 5 . 82 -0 13

75 14 0 .29 6 . 07 -0 13

78 86 0 .29 6 . 32 -0 13

82 57 0 .29 6 . 53 -0 13

86 29 0 .29 6 . 78 -0 13

90 01 0 .25 7 . 03 -0 13

93 73 0 .25 7 . 28 -0 13

97 45 0 .25 7 . 49 -0 13
564 SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH

CH3F

AH°(0 K) = -225.48 kJ mol"^

AH^(298.15 K) = -233.83 kJ mol"^

S°(298.15 K) = 231.46 J K'^ mol-^

Moments of Inertia (10'^^ kg m^)

5.25 31.84 31.84

Vibrational Frequencies (cm"^)

1059.4 1171.3 1171.4 1474.7

1476.1 2886.7 2958.1 2958.6

1476 .

1

Bond Distances

Cl-Hl

Cl-Fl

C1-H2

C1-H3

(10 10 m)

1 . 082

1 . 365

1 . 082

1 . 082

Bond Angles (°) Hl-Cl-Fl 109.13

H1-C1-H2 109.84

H1-C1-H3 109.84

Dihedral Angles (°)

F1-C1-H2

F1-C1-H3

119 . 54

•119 . 54

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

- J mol"^ -

I I kJ mol"^

T/K (c° S° H°-H°(T^) AH0

f AG° Log Kf

150 33 .S2 207 . 76 -5 . 15 -229 .33 -221

.

79 77 ., 24

200 33 . 82 217 . 40 -3 . 47 -230 .79 -219 . 07 57 .. 22

250 35 . 22 225 . 07 -1

.

75 -232 . 34 -216 . 02 45 .. 13

300 37 . 59 231

.

69 0 . 07 -233 .89 -212 . 63 37 .. 02

350 40 .67 237 . 71 2 . 03 -235 .39 -208 . 99 31

.

. 19

400 44 . 14 243 . 37 4 . 14 -236 . 81 -205 . 14 26 ., 79

450 47 . 73 248 . 77 6 . 44 -238 .20 -201 . 12 23 .. 34

500 51 .27 253 . 99 8 . 92 -239 .41 -196 . 94 20 .. 57

600 57 . 91 263 . 93 14 . 38 -241 . 58 -188 . 28 16 .,39

700 63 . 81 273 . 31 20 . 47 -243 . 34 -179 . 24 13 ..38

800 68 . 99 282 . 18 27 . 12 -244 . 76 -170 . 00 11

.

. 10

900 73 .49 290 . 57 34 . 25 -245 . 85 -160 . 58 9 ..32

1000 77 .41 298 . 52 41 . 80 -246 .65 -151

.

00 7 ,. 89

1100 80 . 81 306 . 06 49. 71 -247 .27 -141

.

38 6 .. 71

1200 83 . 75 313 . 22 57 . 94 -247 .69 -131

.

67 5 .. 73

1300 86 .30 320 . 03 66 . 45 -247 . 94 -121

.

92 4 .. 90

1400 88 . 52 326 . 51 75 . 19 -248 . 07 -112 . 17 4 .. 18

1500 90 . 44 332 . 68 84 . 14 -248 . 07 -102 . 38 3 ,. 57

1600 92 . 12 338 . 57 93 . 27 -248 . 03 -92 . 63 3 .. 02

1700 93 . 59 344 . 20 102 . 56 -247 . 86 -82 . 89 2 .. 55

1800 94 . 88 349 . 59 Ill

.

98 -247 .69 -73 . 14 2 .. 12

1900 96 . 01 354 . 75 121

.

53 -247 .40 -63 . 47 1 .. 74

2000 97 . 02 359 . 70 131 . 19 -247 . 11 -53 . 81 1 ..41

2100 97 . 91 364 . 46 140 . 93 -246 . 77 -44 . 22 1 .. 10

2200 98 . 70 369 . 03 150 . 76 -246 .40 -34 . 64 0 ,. 82

2300 99 .40 373 . 43 160 . 67 -245 . 98 -25 . 15 0 .. 57

2400 100 . 04 377 . 68 170 . 64 -245 . 56 -15 . 65 0 .. 34

2500 100 . 60 381

.

77 180 . 67 -245 . 10 -6 . 23 0 .. 13

2600 101 . 11 385 . 73 190 . 76 -244 . 60 3 . 18 -0 .. 06

2700 101 . 58 389 . 55 200 . 89 -244 . 09 12 . 51 -0 ,. 24

2800 102 . 00 393 . 25 211 . 07 -243 . 59 21 . 84 -0 ..41

2900 102 . 38 396 . 84 221 . 29 -243 . 05 31 . 17 -0 .. 56

3000 102 . 73 400 . 32 231 . 55 -242 . 50 40 . 46 -0 .. 70
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 565

CH2F2

ah2(0 k)

AHf(298.15 K)

S°{298,15 K)

Moments of Inertia

16.51 77.17

= -443 .25 kJ mol"^

= -451 . 06 kJ mol"^

= 252 . 16 J K"^ mol"^

(10"'*'^ kg m^)

88 .30

Vibrational Frequencies (cm"-'-) 509.7 1105.4 1122.6 1164.4

1463.6 1529.2 2941.9 3006.3

1258 .

Bond Distances (10"-'-° m)

Cl-Hl 1.078

Cl-Fl 1.338

C1-F2 1.338

C1-H2 1.078

Bond Angles (°) Hl-Cl-Fl 108.93

H1-C1-F2 108.91

H1-C1-H2 112.48

Dihedral Angles (°)

F1-C1-F2

F1-C1-H2

118 .26

120. 89

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^^ = 1 atmosphere

1- J K"^ mol'^ -| I kJ mol"^

i / K. n ~ n AH0

f AG° Log '^f X D u . 0 ^ z9 9z 0 . D QJ7 , U D -436 . 85 X _) z . 1 0 . -L Z

0 n n ^ u u 0 bQ Q 0 -3 z J b . -3 Q n A A Q . 1 u -433 . 25 X X J . 1 c . Lz> 250 39 . 63 244 . 91 -1 . 99 -449 . 78 -429 . 36 89 ,. 70

300 43 . 09 252 . 43 0 . 08 -451 . 12 -425 . 18 74 .. 02

350 47 . 01 259 . 36 2 . 33 -452 .37 -420 . 74 62 .. 79

400 51 . 07 265 . 90 4 . 78 -453 . 55 -416 . 18 54 .. 35

450 55 . 07 272 . 15 7 . 43 -454 .63 -411

45 47 .. 76

500 58 . 86 278 . 15 10 . 28 -455 . 64 -406 . 64 42 ..48

600 65 .63 289 . 50 16 . 52 -457 .27 -396 . 73 34 ,. 54

700 71 . 32 300 . 06 23 . 38 -458 . 57 -386 . 52 28 .. 84

800 76 . 08 309. 90 30 . 75 -459 . 57 -376 . 18 24 .. 56

900 80 . 09 319 . 10 38 . 57 -460 .28 -365 . 68 21 .. 22

1000 83 .48 327 . 72 46 . 75 -460 . 78 -355 . 10 18 .. 55

1100 86 .35 335 . 81 55 . 25 -461 . 16 -344 . 47 16 .. 36

1200 88 . 81 343 . 43 64 . 01 -461 .33 -333 . 80 14 .. 53

1300 90 . 90 350 . 63 72 . 99 -461 .41 -323 . 09 12 ,, 98

1400 92 . 70 357 . 43 82 . 18 -461 .37 -312 . 42 11 ., 66

1500 94 .26 363 . 88 9153 -461 .29 -301

71 10 .. 51

1600 95 . 60 370 . 01 10102 -461 . 08 -29100 9 .. 50

1700 96 . 77 375 . 84 110 . 64 -460 .87 -280 . 37 8 ..61

1800 97 .80 38140 120 . 37 -460 . 57 -269 . 70 7 .. 83

1900 98 . 69 386 . 71 130 . 20 -460 . 24 -259 . 12 7 .. 12

2000 99 .48 391

79 140 . 11 -459 . 86 -248 . 57 6 ,.49

2100 100 . 18 396 . 66 150 . 09 -459 .49 -238 . 07 5 .. 92

2200 100 .80 40134 160 . 14 -459 . 07 -227 . 61 5 ,.40

2300 101 . 36 405 . 84 170 . 25 -458 .61 -217 . 19 4 .. 93

2400 101 . 85 410 . 16 180 . 41 -458 . 15 -206 . 82 4 ,. 50

2500 102 .29 414 . 32 190 . 61 -457 .65 -196 . 52 4 ,. 11

2600 102 .69 418 . 35 200 . 87 -457 . 14 -186 . 23 3 ,. 74

2700 103 . 05 422 . 23 211

15 -456 . 64 -175 . 98 3 ,.40

2800 103 . 38 425 . 98 221

48 -456 . 10 -165 . 77 3 ,. 09

2900 103 .68 429 . 61 231

83 -455 . 55 -155 . 56 2 . 80

3000 103 . 95 433 . 14 242 . 21 -455 . 01 -145 . 35 2 . 53
566 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHF3

AH°(0 K) = -692.49 kJ mol"^

AHf{298.15 K) = -699.45 kJ mol'^

S°(298.15 K) = 268.72 J K"^ mol"^

Moments of Inertia (10"*^ kg m^)

78.94 79.02 144.03

Vibrational Frequencies (cm"-'-) 491.7 492.1 680.7 1126.9 1185.7

1186.0 1413.4 1413.6 3036.6

Bond Distances (10"^° m) Bond Angles {°) Dihedral Angles (' Cl-Hl 1.074

Cl-Fl 1.316 Hl-Cl-Fl 110.44

C1-F2 1.316 H1-C1-F2 110.44 F1-C1-F2 120.03

C1-F3 1.316 H1-C1-F3 110.44 F1-C1-F3 -120.03

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

j - J K"^ mol"^ -

I I kJ mol"^

T /K C° S° H°-H AH0

f AG?

J

J \j * J-1 Z'Q 7 7 q 03VJ .J - 6 52 D J* 0 . 2 2 -682 . 54 P.D 1/ 200 41 . 66 250 . 32 -4 . 55 -697 . 39 -677 . 85 111 02

250 46 .48 260 . 13 -2 . 35 -698 .48 -672 . 83 140 57

300 51 .40 269 . 04 0 . 10 -699 .48 -667 . 64 116 24

350 56 .25 277 . 33 2 . 79 -700 . 36 -662 . 29 98 83

400 60 . 86 285 . 14 5 . 72 -701 . 15 -656 . 80 85 76

450 65 . 11 292 . 56 8 . 87 -701 .87 -651

.

24 75 59

500 68 . 96 299 . 62 12 . 22 -702 .45 -645 . 59 67 44

600 75 . 48 312 . 80 19 . 46 -703 . 33 -634 . 17 55 21

700 80 .63 324 . 84 27 . 27 -703 . 96 -622 . 62 46 46

800 84 . 73 335 . 88 35 . 55 -704 . 38 -610 . 95 39 89

900 88 . 01 346 . 05 44 . 19 -704 . 59 -599 . 23 34 78

1000 90 . 70 355 . 47 53 . 13 -704 .67 -587 . 52 30 69

1100 92 . 90 364 . 23 62 . 31 -704 .63 -575 . 72 27 34

1200 94 . 74 372 . 39 71 . 70 -704 . 50 -563 . 96 24 55

1300 96 .28 380 . 03 81 . 25 -704 . 33 -552 . 20 22 19

1400 97 . 58 387 . 22 90 . 94 -704 . 08 -540 . 45 20 16

1500 98 .69 393 . 99 100 . 76 -703 . 75 -528 . 69 18 41

1600 99 . 64 400 . 39 110 . 68 -703 .41 -516 . 98 16 88

1700 100 .46 406 . 45 120 . 68 -703 . 04 -505 . 30 15 53

1800 101 . 17 412 . 22 130 . 77 -702 . 62 -493 . 67 14 32

1900 101 . 79 417 . 71 140 . 91 -702 .20 -482 . 04 13 25

2000 102 . 33 422 . 94 151

.

12 -701 . 74 -470 . 49 12 29

2100 102 . 81 427 . 94 161

.

38 -701 . 24 -458 . 98 11 42

2200 103 .23 432 . 74 171 . 68 -700 . 74 -447 . 56 10 63

2300 103 .61 437 . 33 182 . 02 -700 .23 -436 . 14 9 90

2400 103 . 94 441 . 75 192 . 40 -699 .69 -424 . 80 9 25

2500 104 .24 446 . 00 202 . 81 -699 . 15 -413 . 50 8 64

2600 104 . 51 450 . 09 213 . 25 -698 .60 -402 . 25 8 08

2700 104 . 75 454 . 04 223 . 71 -698 . 06 -391 . 04 7 56

2800 104 . 97 457 . 86 234 . 20 -697 .47 -379 . 82 7 09

2900 105 . 17 461 . 55 244 . 71 -696 . 89 -368 . 65 6 64

3000 105 .35 465 . 11 255 . 23 -696 .30 -357 . 48 6 22
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 567

CF4

AH?(0 K)

AHf(298.15 K)

S°{298.15 K)

Moments of Inertia

142.60 142.61

= -928 . 60 kJ mol"^

= -934 .37 kJ mol"^

= 273 .27 J K"^ mol'^

(lO"'*'^ kg m^)

142 .61

Vibrational Frequencies (cm"-'^) 422.0 422.0 609.8 609.8 609.8

895.5 1314.2 1314.2 1314.2

Bond Distances (10" m)

Cl-Fl 1.302

C1-F2 1.302

C1-F3 1.302

C1-F4 1.302

Bond Angles (°) F1-C1-F2 109.47

F1-C1-F3 109.47

F1-C1-F4 109.47

Dihedral Angles (°)

F2-C1-F3 120.00

F2-C1-F4 -120.00

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

|- J K"^ mol"^ -| I kJ mol"^

t/k ( S H -H0 / m \ (Tj.) AH0

f AG? Log Kf

150 40 .79 238 . 66 - 7 . 64 - 932 . 20 -913 . 07 317 ,. 94

200 48 . 23 25141 -5 . 42 -933 . 12 -906 . 55 236 ,. 76

250 55 .45 262 . 96 -2 . 82 -933 . 83 -899 . 85 188 ,. 00

300 61 . 97 273 . 66 0 . 11 -934 .37 -893 . 03 155 ,.48

350 67 . 73 283 . 65 3 . 36 -934 . 79 -886 . 13 132 .. 24

400 72 . 75 293 . 03 6 . 87 -935 . 08 -879 . 18 114 ,.80

450 77 . 09 301

85 10 . 62 -935 .25 -872 . 20 101

. 24

500 80 . 83 310 . 18 14 . 57 -935 . 33 -865 . 21 90 ..38

600 86 . 78 325 . 47 22 . 97 -935 .38 -85119 74 .. 10

700 91 . 16 339 . 19 3188 -935 .21 -837 . 22 62 ..47

800 94 .40 35159 41

17 -934 .87 -823 . 24 53 .. 75

900 96 . 85 362 . 85 50 . 73 -934 . 50 -809 . 27 46 .. 97

1000 98 . 72 373 . 16 60 . 51 -934 . 08 -795 . 34 41. 54

1100 100 . 17 382 . 64 70 . 46 -933 . 58 -78145 37 ,. 11

1200 101 . 32 39140 80 . 54 -933 . 07 -767 . 60 33 .,41

1300 102 .25 399 . 56 90 . 72 -932 . 57 -753 . 75 30 ..28

1400 103 . 00 407 . 16 100 . 98 -932 . 03 -739 . 94 27 ..61

1500 103 .62 414 . 29 Ill

32 -931 .48 -726 . 18 25 ..29

1600 104 . 13 420 . 99 12170 -930 . 90 -712 . 45 23 .. 26

1700 104 . 56 427 . 32 132 . 14 -930 .35 -698 . 77 21 ,.47

1800 104 . 93 433 . 30 142 . 61 -929 . 77 -685 . 13 19 ,.88

1900 105 .24 438 . 99 153 . 12 -929 . 18 -671

57 18 ,. 46

2000 105 . 51 444 . 39 163 . 66 -928 .60 -658 . 06 17 ,. 19

2100 105 . 74 449 . 55 174 . 22 -927 . 97 -644 . 59 16 ,. 03

2200 105 . 94 454 . 47 184 . 80 -927 .38 -63120 14 .. 99

2300 106 . 12 459 . 18 195 . 41 -926 . 76 -617 . 85 14 ,. 03

2400 106 . 28 463 . 70 206 . 03 -926 . 17 -604 . 55 13 ,. 16

2500 106 .42 468 . 04 216 . 66 -925 . 54 -591

28 12 ,. 35

2600 106 . 54 472 . 22 227 . 31 -924 . 92 -578 . 10 11 ,.61

2700 106 . 65 476 . 24 237 . 97 -924 .29 -564 . 92 10 .. 93

2800 106 . 75 480 . 12 248 . 64 -923 . 62 -551 . 79 10 ,.29

2900 106 . 84 483 . 87 259 . 32 -922 . 99 -538 . 69 9 ,. 70

3000 106 . 92 487 . 49 270 . 01 -922 . 36 -525 . 59 9 ,. 15
568 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

•CHjF

AH2(0 K) = -21.51 kJ mol'^

AH°(298.15 K) = -31.44 kJ mol"^

S°(298.15 K) = 234.32 J K"^ mol"^

Moments of Inertia (10 kg m^)

3.15 26.82 29.53

Vibrational Frequencies (cm'^)

771.2 1134.2 1143.5 1443.5 2962.3

3087 . 6 Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Cl-Fl 1.331

Cl-Hl 1.073 Fl-Cl-Hl 113.95

C1-H2 1.073 F1-C1-H2 113.95 H1-C1-H2 144.84

Enthalpy Reference Temperature = = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

]- J K"^ mol"^ -| I kJ mol"^

T/K C

lO S° H°-H°(T^) AHo

f AG? Log L. 150 33 . |8 210 . 04 -5 . 28 -29 .25 -30. 2 9 10 .. 55

200 34 . 60 219 . 82 -3 . 58 -29 . 96 -30 . 54 7 .. 98

250 36 . 40 227 . 72 -1

.

80 -30 . 71 -30 . 63 6 ..40

300 38 . 74 234 . 56 0 . 07 -31 .46 -30 . 59 5 .. 32

350 41 . 31 240 . 72 2 . 07 -32 . 22 -30 . 42 4 .. 54

400 43 . 90 246 . 41 4 . 20 -32 .89 -30 . 12 3 .. 93

450 46 . 37 251 . 72 6 . 46 -33 . 56 -29 . 75 3 ,.45

500 48 . 68 256 . 73 8 . 84 -34 . 18 -29 . 33 3 ,. 06

600 52 . 76 265 . 97 13 . 92 -35 .27 -28 . 28 2 ..46

700 56 . 25 274 . 38 19 . 37 -36 . 19 -27 . 07 2 .. 02

800 59 . 25 282 . 09 25 . 15 -37 . 03 -25 . 69 1

.

.68

900 61 . 88 289 . 22 31

.

21 -37 . 70 -24 . 23 1

.

.41

1000 64 . 17 295 . 86 37 . 51 -38 .28 -22 . 64 1

.

. 18

1100 66 . 18 302 . 08 44 . 03 -38 . 79 -21 . 00 1

,

. 00

1200 67 . 94 307 . 91 50 . 74 -39 .20 -19 . 33 0 .. 84

1300 69 . 47 313 . 41 57 . 61 -39 . 54 -17 . 57 0 .. 71

1400 70 . 81 318 . 61 64 . 63 -39 . 83 -15 . 82 0 .. 59

1500 71 . 99 323 . 54 71 . 77 -40 . 12 -14 . 02 0 ..49

1600 73 . 02 328 . 21 79 . 02 -40 .33 -12 . 22 0 ..40

1700 73 . 93 332 . 67 86 . 37 -40 . 50 -10 . 42 0 .. 32

1800 74 . 73 336 . 92 93 . 81 -40 .67 -8 . 62 0 ..25

1900 75 . 44 340 . 98 101

.

31 -40 . 79 -6 . 86 0 ,. 19

2000 76 . 07 344 . 87 108 . 89 -40 . 92 -5 . 10 0 ,. 13

2100 76 . 63 348 . 59 116 . 52 -41 . 05 -3 . 35 0 ,. 08

2200 77 . 12 352 . 17 124 . 21 -41 . 09 -1 . 63 0 ,. 04

2300 77 . 57 355 . 61 131

.

95 -41 . 17 0 . 08 0 ,. 00

2400 77 . 97 358 . 92 139 . 72 -41 .21 1 . 76 -0 .. 04

2500 78 . 33 362 . 10 147 . 54 -41 .25 3 . 39 -0 ,. 07

2600 78 . 66 365 . 18 155 . 39 -41 .30 5 . 02 -0 ,. 10

2700 78 . 95 368 . 16 163 . 27 -41 . 34 6 . 65 -0 ,. 13

2800 79 . 22 371

.

03 171

.

18 -41 . 34 8 . 28 -0 ,. 15

2900 79 . 46 373 . 82 179 . 11 -41 . 34 9 . 87 -0 ,. 18

3000 79 . 68 376 . 52 187 . 07 -41 . 34 11

.

51 -0 ,.20
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 569

• CHF.

AH2(0 K)

AH°(298.15 K)

S°(298.15 K)

= -243 . 76 kJ mol"^

= -247 . 27 kJ mol"^

= 255 . 94 J K"^ mol"

Moments of Inertia (10

12.20 74.78 85.62

kg

Vibrational Frequencies (cm"-'^) 521.4 1041.3 1150.0 1180.8 1343.1

3005 .

3

Bond Distances (10"-'-° m)

Cl-Hl 1.076

Cl-Fl 1.314

C1-F2 1.314

Bond Angles (°) Hl-Cl-Fl

H1-C1-F2

113 . 77

113 . 77

Dihedral Angles C

F1-C1-F2 128 . 66

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p = 1 atmosphere

J K"^ mol"^ - 1 kJ mol -1 i/K r° fl ~ n

0

f AGf Log Kf

X D U

"3 A . / D

o "3 n u o

c - D . 0 /I cr

. 6 D -244 . 01 Q A 0 4

Q "7 y / o A n J o . / Z z4 U . -3 - J . Q CZ 0 D -

'~y A c Z 4 b n 0 . Uz -243 . 51 6 3 D y

2 5 0 3 9 . 2 7 24 8. 7 9 - 1 . 96 - 24 6 . 6 5 -242 . 80 50 73

300 42 .25 256 . 21 0 . 08 -247 .27 -242 . 00 42 13

350 45 .43 262 . 96 2 . 27 -247 . 90 -241 . 08 35 98

400 48 . 55 269 . 23 4 . 62 -248 .45 -240 . 12 31 35

450 51 .48 275 . 12 7 . 12 -248 . 95 -239 . 07 27 75

500 54 . 15 280 . 68 9 . 11 -249 .41 -237 . 94 24 86

600 58 . 70 290 . 98 15 . 41 -250 .20 -235 . 64 20 51

700 62 .33 300 . 31 21

.

47 -250 . 83 -233 . 17 17 40

800 65 . 25 308 . 83 27 . 86 -251 . 33 -230 . 62 15 06

900 67 .63 316 . 65 34 . 51 -251 . 75 -227 . 99 13 23

1000 69 .60 323 . 89 41 . 37 -252 . 13 -225 . 27 11 77

1100 71 . 24 330 . 60 48 . 41 -252 .42 -222 . 51 10 57

1200 72 .63 336 . 86 55 . 61 -252 . 63 -219 . 74 9 56

1300 73 . 80 342 . 72 62 . 93 -252 . 84 -216 . 94 8 72

1400 74 . 80 348 . 23 70 . 36 -253 . 01 -214 . 10 7 99

1500 75 . 65 353 . 41 77 . 89 -253 . 17 -211

.

25 7 36

1600 76 .39 358 . 32 85 . 49 -253 .30 -208 . 36 6 80

1700 77 . 03 362 . 97 93 . 16 -253 .38 -205 . 52 6 31

1800 77 . 59 367 . 39 100 . 89 -253 .47 -202 . 67 5 88

1900 78 . 08 371 . 60 108 . 68 -253 . 55 -199 . 87 5 49

2000 78 . 51 375 . 62 116 . 51 -253 . 59 -197 . 07 5 15

2100 78 .89 379 . 46 124 . 38 -253 .63 -194 . 30 4 83

2200 79 .22 383 . 14 132 . 29 -253 . 68 -191

.

54 4 55

2300 79 . 53 386 . 66 140 . 22 -253 . 72 -188 . 82 4 29

2400 79 . 79 390 . 05 148 . 19 -253 . 76 -186 . 10 4 05

2500 80 . 03 393 . 32 156 . 18 -253 . 76 -183 . 43 3 83

2600 80 . 24 396 . 46 164 . 19 -253 . 76 -180 . 79 3 63

2700 80 . 44 399 . 49 172 . 23 -253 . 76 -178 . 11 3 45

2800 80 . 62 402 . 42 180 . 28 -253 . 76 -175 . 48 3 27

2900 80 . 78 405 . 25 188 . 35 -253 . 76 -172 . 84 3 11

3000 80 . 92 407 . 99 196 . 43 -253 . 72 -170 . 21 2 96
570 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

• CF3

AH|(0 K) = -469.24 kJ mol"^

AHf(298.15 K) = -471.90 kJ mol"^

S°(298.15 K) = 264.64 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

75.31 75.31 145.53

Vibrational Frequencies (cm

490.6 490.6 676.5 1086.2

1285 . 9 1285 . 9

Bond Distances (10 10 m)

Cl-Fl

C1-F2

C1-F3

1 . 301

1 . 301

1 .301

Bond Angles (°) Xl-Cl-Fl 107.60

X1-C1-F2 107.60

X1-C1-F3 107.60

Dihedral Angles

F1-C1-F2

F1-C1-F3

120 . 00

-120 . 00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

]- J K"^ mol"^ - I i kJ mol"^

T/K S° H°-H°(T^) AH0

f AG? Log Kf

150 37 .^2 235 . 13 -6 . 47 -470 . 74 -464 . 76 161

,

. 84

200 41 . 58 246 . 43 -4 . 51 -471 .20 -462 . 71 120 .. 84

250 46 . 02 256 . 19 -2 . 31 -471 . 62 -460 . 57 96 ,. 23

300 50 . 22 264 . 95 0 . 09 -471 . 91 -458 . 36 79 ,. 80

350 54 . 09 272 . 99 2 . 70 -472 . 16 -456 . 10 68 ,. 07

400 57 . 57 280 . 45 5 . 49 -472 . 33 -453 . 80 59 ..26

450 60 . 64 287 . 41 8 . 45 -472 . 50 -451 . 50 52 ..41

500 63 . 32 293 . 94 11 . 55 -472 . 62 -449 . 19 46 ,. 92

600 67 . 62 305 . 88 18 . 11 -472 . 75 -444 . 51 38 .. 70

700 70 .80 316 . 56 25 . 04 -472 . 79 -439 . 82 32 ,. 82

800 73 . 16 326 . 17 32 . 24 -472 . 83 -435 . 14 28 ..41

900 74 . 94 334 . 90 39 . 65 -472 . 79 -430 . 41 24 .. 98

1000 76 .31 342 . 87 47 . 22 -472 . 75 -425 . 68 22 .. 23

1100 77 . 37 350 . 19 54 . 91 -472 . 71 -420 . 91 19 .. 99

1200 78 . 22 356 . 96 62 . 68 -472 . 71 -416 . 14 18 .. 11

1300 78 . 89 363 . 25 70 . 54 -472 .67 -411 . 37 16 .. 53

1400 79 . 44 369 . 12 78 . 46 -472 .62 -406 . 56 15 ., 17

1500 79 . 89 374 . 61 86 . 42 -472 . 58 -401 . 79 13 ., 99

1600 80 . 26 379 . 78 94 . 43 -472 . 58 -397 . 02 12 .. 96

1700 80 . 58 384 . 66 102 . 47 -472 . 54 -392 . 25 12 ,. 05

1800 80 . 84 389 . 27 110 . 55 -472 . 50 -387 . 52 11 ,. 24

1900 81 . 07 393 . 65 118 . 64 -472 . 50 -382 . 79 10 ., 52

2000 81 .27 397 . 81 126 . 76 -472 .46 -378 . 11 9 .. 87

2100 81 .43 401

.

78 134 . 90 -472 . 42 -373 . 42 9 ..29

2200 81 . 58 405 . 57 143 . 05 -472 .42 -368 . 78 8 .. 76

2300 81 . 71 409 . 20 151

.

21 -472 . 37 -364 . 18 8 ..27

2400 81 . 83 412 . 68 159 . 39 -472 . 33 -359 . 61 7 .. 83

2500 81 . 93 416 . 02 167 . 57 -472 .29 -355 . 05 7 .. 42

2600 82 . 01 419 . 24 175 . 77 -472 .25 -350 . 54 7 .. 04

2700 82 . 10 422 . 34 183 . 98 -472 .21 -346 . 02 6 ., 69

2800 82 . 17 425 . 32 192 . 19 -472 . 16 -341

.

50 6 ..37

2900 82 . 24 428 . 21 200 . 41 -472 . 12 -336 . 98 6 .. 07

3000 82 .30 431

.

00 208 . 64 -472 . 08 -332 . 46 5 .. 79
SECTION 5. FLAME INHIBrnON CHEMISTRY AND THE SEARCH 571

:CHF (singlet)

AH|(0 K) = 152.59 kJ mol"^

AHf(298.15 K) = 152.77 kJ mol"^

S°{298.15 K) = 222.83 J K"^ mol'^

Moments of Inertia (10"*^ kg m^)

1.72 22.40 24.12

Vibrational Frequencies (cm'-'-) 1189.2 1404.2 2732.5

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Cl-Fl 1.295

Cl-Hl 1.104 Fl-Cl-Hl 102.80

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I j kJ mol"^

t/k C S

ttO tt H -HO / m \ (Tj.) AHo

f AG° Log

150 3 3 .27 199 . 73 -4 . 9 9 153 . 09 145 . 35 - 50 .. 61

200 33 . 41 209 . 32 - 3 . 33 153 . 09 142 . 72 -37 .. 28

250 33 . 84 216 . 81 - 1 . 64 152 . 97 140 . 16 -29 ., 28

300 34 . 62 223 . 04 0 . 06 152 . 76 137 . 57 -23 ,. 95

350 35 . 66 228 . 46 1

.

82 152 . 51 135 . 02 -20 .. 15

400 36 . 85 233 . 30 3 . 63 152 .21 132 . 55 -17 .. 31

450 38 . 09 237 . 71 5 . 51 151 . 92 130 . 08 -15 .. 10

500 39 . 33 241

.

78 7 . 44 151 . 59 127 . 65 -13 ..33

600 41 . 66 249 . 17 11. 49 150 . 92 122 . 88 -10 .. 70

700 43 . 72 255 . 74 15 . 77 150 .25 118 . 24 -8 .. 82

800 45 .53 261

.

70 20 . 23 149 . 58 113 . 72 -7 .. 42

900 47 . 08 267 . 16 24 . 86 148 . 95 109 . 29 -6 ,. 34

1000 48 .41 272 . 19 29 . 64 148 . 32 104 . 93 -5 ,.48

1100 49 . 56 276 . 86 34 . 53 147 . 70 100 . 71 -4 .. 78

1200 50 . 54 281

.

21 39 . 54 147 . 11 96 . 48 -4 ..20

1300 51 .38 285 . 29 44 . 64 146 . 52 92 . 38 -3 .. 71

1400 52 . 10 289 . 13 49 . 81 145 . 94 88 . 28 -3 ..29

1500 52 . 72 292 . 74 55 . 06 145 .35 84 . 27 -2 .. 93

1600 53 .26 296 . 16 60 . 35 144 . 77 80 . 25 -2 .. 62

1700 53 . 73 299 . 41 65 . 71 144 . 18 76 . 27 -2 .. 34

1800 54 . 14 302 . 49 71 . 10 143 . 59 72 . 34 -2 .. 10

1900 54 . 50 305 . 43 76 . 53 143 . 05 68 . 41 -1 ,. 88

2000 54 . 81 308 . 23 82 . 00 142 .47 64 . 48 -1 .. 68

2100 55 . 09 310 . 91 87 . 49 141 . 92 60 . 50 -1 .. 51

2200 55 .33 313 . 48 93 . 01 141 . 34 56 . 57 -1

,

. 34

2300 55 . 56 315 . 94 98 . 56 140 . 79 52 . 63 -1 ..20

2400 55 . 75 318 . 31 104 . 12 140 .21 48 . 70 -1 .. 06

2500 55 . 93 320 . 59 109 . 71 139 . 66 44 . 77 -0 .. 94

2600 56 . 08 322 . 79 115 . 31 139 . 08 40 . 84 -0 .. 82

2700 56 .22 324 . 91 120 . 93 138 . 53 36 . 90 -0 .. 71

2800 56 .35 326 . 95 126 . 55 137 . 99 32 . 97 -0 .. 62

2900 56 .47 328 . 93 132 . 19 137 .40 29 . 08 -0 ,. 52

3000 56 . 58 330 . 85 137 . 85 136 . 86 25 . 23 -0 .. 44
572 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

rCFj (singlet)

AH?(0 K) = -203.84 kJ mol"^

AHf(298.15 K) = -203.33 kJ mol"^

S°{298.15 K) = 240.43 J K"^ mol"^

Moments of Inertia (10"*^ kg m^)

9.35 64.95 74.30

Vibrational Frequencies (cm'^)

651.0 1155.5 1240.8

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Cl-Fl 1.283

C1-F2 1.283 F1-C1-F2 104.47

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J mol"^ -| i kJ mol"^

T/K (C° S° H°-H°(T^) AHo

f AG? Log Kf

150 33 .Bl 215 . 80 -5 . 35 -203 .30 -208 . 07 72 ..45

200 35 .20 225 . 72 -3 . 63 -203 .26 -209 . 66 54 .. 76

250 36 . 96 233 . 76 -1 . 82 -203 .30 -211 . 29 44 .. 15

300 38 . 94 240 . 67 0 . 07 -203 . 34 -212 . 92 37 ,. 07

350 40 . 97 246 . 82 2 . 07 -203 .43 -214 . 51 32 .. 02

400 42 . 90 252 . 42 4 . 17 -203 . 51 -216 . 15 28 .. 22

450 44 . 66 257 . 58 6 . 36 -203 . 59 -217 . 74 25 ,.27

500 46 .23 262 . 37 8 . 63 -203 .68 -219 . 33 22 .. 91

600 48 . 79 271 . 04 13 . 39 -203 . 93 -222 . 46 19 ..37

700 50 . 71 278 . 70 18 . 37 -204 . 18 -225 . 56 16 .. 83

800 52 . 14 285 . 57 23 . 51 -204 .43 -228 . 61 14 .. 93

900 53 . 22 291 . 78 28 . 78 -204 . 76 -231

.

58 13 ..44

1000 54 . 04 297 . 43 34 . 15 -205 . 14 -234 . 51 12 ..25

1100 54 .69 302 . 62 39. 58 -205 . 52 -237 . 40 11

.

.27

1200 55 .20 307 . 39 45 . 08 -205 . 94 -240 . 20 10 ..46

1300 55 .61 311

.

83 50 . 62 -206 .40 -242 . 96 9 .. 76

1400 55 . 94 315 . 97 56 . 20 -206 . 86 -245 . 68 9 .. 17

1500 56 . 22 319 . 83 61

.

81 -207 . 36 -248 . 40 8 ..65

1600 56 .45 323 . 47 67 . 44 -207 . 86 -251 . 04 8 ..20

1700 56 . 64 326 . 90 73 . 10 -208 .36 -253 . 72 7 ..79

1800 56 .80 330 . 14 78 . 77 -208 .87 -256 . 35 7 ..44

1900 56 . 94 333 . 21 84 . 46 -209 .41 -258 . 95 7 .. 12

2000 57 . 06 336 . 14 90 . 16 -209 . 95 -261 . 58 6 .. 83

2100 57 . 16 338 . 92 95 . 87 -210 .50 -264 . 18 6 .. 57

2200 57 .25 341 . 59 101

.

59 -211 . 04 -266 . 81 6 ,. 33

2300 57 . 33 344 . 13 107 . 32 -211 . 58 -269 . 45 6 .. 12

2400 57 .40 346 . 57 113 . 06 -212 . 13 -272 . 04 5 .. 92

2500 57 .46 348 . 92 118 . 80 -212 .67 -274 . 68 5 ,. 74

2600 57 . 52 351

.

18 124 . 55 -213 .22 -277 . 32 5 .. 57

2700 57 . 56 353 . 35 130 . 30 -213 . 76 -279 . 95 5 ..42

2800 57 .61 355 . 44 136 . 06 -214 .30 -282 . 55 5 ..27

2900 57 .65 357 . 46 141

.

82 -214 . 85 -285 . 10 5 .. 14

3000 57 .68 359 . 42 147 . 59 -215 .39 -287 . 65 5 .. 01
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 573

CHF=0

AH2(0 K) = -378.57 kJ mol"^

AHf(298.15 K) ' = -382.28 kJ mol'^

S°{298.15 K) = 246.15 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

8.84 68.99 77.83

Vibrational Frequencies (cm"-'-) 659.1 1050.8 1115.1 1374.1 1876.5

2996 .

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Cl-01 1.164

Cl-Fl 1.314 01-Cl-Fl 123.03

Cl-Hl 1.081 01-Cl-Hl 126.96

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^^ = 1 atmosphere

I - J K'^ mol"^ -I

j kJ mol"^

T/K ( S° H°-H°(T^) AHo

f AG? Log Kf

150 33 .12 22126 -5 . 42 -380 .20 -377 . 40 131

.41

200 35 .38 231

20 -3 . 69 -380 . 91 -376 . 35 98 ..29

250 37 . 57 239 . 32 -1

87 -381 .62 -375 . 14 78 ,. 38

300 40 . 21 246 . 40 0 . 08 -382 .29 -373 . 80 65 .. 08

350 43 . 03 252 . 81 2 . 15 -382 . 96 -372 . 38 55 .. 57

400 45 . 86 258 . 74 4 . 38 -383 . 59 -370 . 83 48 ..42

450 48 . 57 264 . 29 6 . 74 -384 . 13 -369 . 24 42 .. 86

500 51 . 12 269 . 55 9. 23 -384 .68 -367 . 56 38 ,.40

600 55 . 64 279 . 28 14 . 58 -385 .60 -364 . 09 31.69

700 59 . 43 288 . 15 20 . 33 -386 .35 -360 . 45 26 .. 90

800 62 . 61 296 . 30 26 . 44 -386 . 98 -356 . 69 23 ..29

900 65 . 26 303 . 83 32 . 84 -387 . 52 -352 . 88 20 ..48

1000 67 .49 310 . 83 39 . 48 -387 . 98 -348 . 95 18 ..23

1100 69 .37 317 . 35 46 . 33 -388 . 36 -344 . 97 16 .. 38

1200 70 . 96 323 . 45 53 . 34 -388 .69 -340 . 95 14 .. 84

1300 72 .32 329 . 19 60 . 51 -388 . 94 -336 . 94 13 ,. 54

1400 73 .47 334 . 59 67 . 80 -389 .20 -332 . 84 12 .,42

1500 74 .47 339 . 69 75 . 20 -389 .40 -328 . 74 11

,45

1600 75 . 32 344 . 53 82 . 69 -389 . 57 -324 . 64 10 .,60

1700 76 . 07 349 . 12 90 . 26 -389 . 74 -320 . 54 9 .,85

1800 76 . 72 353 . 49 97 . 90 -389 . 87 -316 . 44 9 ., 18

1900 77 .28 357 . 65 105 . 60 -389 . 95 -312 . 38 8 ., 59

2000 77 . 78 361 . 63 113 . 35 -390 . 07 -308 . 32 8 ., 05

2100 78 . 22 365 . 43 12116 -390 . 16 -304 . 26 7 ., 57

2200 78 .61 369 . 08 129 . 00 -390 .20 -300 . 24 7 ., 13

2300 78 . 96 372 . 58 136 . 88 -390 .28 -296 . 27 6 ., 73

2400 79 .27 375 . 95 144 . 79 -390 .33 -292 . 29 6 ,, 36

2500 79 . 55 379 . 19 152 . 73 -390 .37 -288 . 32 6 ,, 02

2600 79 . 80 382 . 32 160 . 69 -390 .41 -284 . 43 5 ,, 71

2700 80 . 02 385 . 33 168 . 69 -390 .41 -280 . 50 5 ,. 43

2800 80 . 22 388 . 25 176 . 70 -390 .45 -276 . 60 5 ,. 16

2900 80 .41 391 . 07 184 . 73 -390 .45 -272 . 67 4 ,. 91

3000 80 . 58 393 . 79 192 . 78 -390 .45 -268 . 74 4 ,.68
574 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF2=0

AH2(0 K) = -595.34 kJ mol'^

AH^(298.15 K) = -598.38 kJ mol'^

S°(298.15 K) = 258.58 J K'^ mol"^

Moments of Inertia (10"^^ kg m^)

68.97 69.00 137.97

Vibrational Frequencies (cm"^)

563.3 610.1 779.0 976.6 1305.5

1953 . 0 Bond Distances (10"^° m) Bond Angles {°) Dihedral Angles

Cl-01 1.157

Cl-Fl 1.290 01-Cl-Fl 125.86

C1-F2 1.290 01-C1-F2 125.86

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K'^ mol"^ -I I kJ mol"^

t/k X / IN. C° S° H°-H AH o

f AG?

15 0 3 5 ll 2 3 0. 72 - 6 . ]_ -]_ - 5 96 . 93 - 590 . 1

1

^ \J Zj A Q

9^ nn\j \j .J ? 2 41. 4

1

- 4 - R Q 7 -587 . 77 -L O J D u 250 43 .37 250 . 60 -2 . 18 -597 . 98 -585 . 30 122 29

300 47 . 51 258 . 88 0 . 09 -598 .40 -582 . 75 101 46

350 51 . 30 266 . 49 2 . 56 -598 . 73 -580 . 15 86 58

400 54 .69 273 . 57 5 . 21 -598 . 98 -577 . 48 75 41

450 57 . 68 280 . 18 8 . 02 -599 .23 -574 . 80 66 72

500 60 .31 286 . 40 10 . 97 -599 .40 -572 . 12 59 76

600 64 . 64 297 . 80 17 . 23 -599 .65 -566 . 68 49 33

700 67 . 98 308 . 02 23 . 87 -599 . 82 -561 . 16 41 87

800 70 . 58 317 . 27 30 . 80 -599 . 94 -555 . 64 36 28

900 72 .61 325 . 71 37 . 97 -600 . 07 -550 . 07 31 92

1000 74 . 22 333 . 45 45 . 31 -600 . 11 -544 . 46 28 44

1100 75 . 51 340 . 59 52 . 80 -600 . 15 -538 . 86 25 59

1200 76 . 55 347 . 20 60 . 40 -600 . 24 -533 . 25 23 21

1300 11 .40 353 . 36 68 . 10 -600 . 28 -527 . 60 21 20

1400 78 . 11 359 . 13 75 . 88 -600 . 28 -521 . 91 19 47

1500 78 .69 364 . 54 83 . 72 -600 . 32 -516 . 26 17 98

1600 79 . 18 369 . 63 91

.

61 -600 . 36 -510 . 62 16 67

1700 79 . 60 374 . 44 99 . 55 -600 . 36 -504 . 97 15 51

1800 79 . 95 379 . 00 107 . 53 -600 .40 -499 . 32 14 49

1900 80 . 26 383 . 33 115 . 54 -600 .40 -493 . 71 13 57

2000 80 . 53 387 . 46 123 . 58 -600 .40 -488 . 11 12 75

2100 80 . 76 391 . 39 131

.

65 -600 .40 -482 . 54 12 00

2200 80 . 96 395 . 15 139. 73 -600 .45 -477 . 02 11 33

2300 81 . 14 398 . 76 147 . 84 -600 .45 -471 . 49 10 71

2400 81 .29 402 . 21 155 . 96 -600 .40 -466 . 01 10 14

2500 81 .43 405 . 53 164 . 10 -600 .40 -460 . 57 9 62

2600 81 . 55 408 . 73 172 . 24 -600 .40 -455 . 14 9 14

2700 81 .67 411 . 81 180 . 41 -600 .36 -449 . 70 8 70

2800 81 . 77 414 . 78 188 . 58 -600 . 36 -444 . 26 8 29

2900 81 . 86 417 . 66 196 . 76 -600 . 32 -438 . 86 7 90

3000 81 . 94 420 . 43 204 . 95 -600 . 28 -433 . 42 7 55
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 575

• CF=0

AH2(0 K) = -183.43 kJ mol'^

AHf(298.15 K) = -182.87 kJ mol"^

S°(298.15 K) = 248.32 J K"^ mol'^

Moments of Inertia (10"'*^ kg m^)

4.19 70.99 75.18

Vibrational Frequencies (cm"-'-) 632.0 1081.1 1915.2

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles {°)

Cl-01 1.152

Cl-Fl 1.304 01-Cl-Fl 128.26

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K 1 S° H°-H AHo

f AG° Log Kf

150 34 .DO 223 . 66 -5 . 36 -182 . 88 -188 . 66 65 ,. 70

200 35 . 33 233 . 61 -3 . 63 -182 . 84 -190 . 62 49 ,. 78

250 36 . 97 241

.

66 -1

.

82 -182 . 84 -192 . 59 40 ,. 24

300 38 .67 248 . 55 0 . 07 -182 . 88 -194 . 56 33 ,. 87

350 40 .30 254 . 64 2 . 05 -182 . 92 -196 . 52 29 ,.33

400 41 . 84 260 . 12 4 . 10 -182 . 97 -198 . 49 25 ,. 92

450 43 .27 265 . 14 6 . 23 -183 . 05 -200 . 46 23 ,. 26

500 44 . 57 269 . 76 8 . 43 -183 . 13 -202 . 38 21

.

. 14

600 46 . 85 278 . 10 13 . 00 -183 . 38 -206 . 23 17 .. 95

700 48 . 71 285 . 46 17 . 78 -183 . 64 -210 . 08 15 ..67

800 50 . 22 292 . 07 22 . 73 -183 . 97 -213 . 80 13 .. 96

900 51 .44 298 . 06 27 . 82 -184 .35 -217 . 48 12 ..62

1000 52 .42 303 . 53 33 . 01 -184 . 77 -221

.

12 11 ,. 55

1100 53 . 22 308 . 57 38 . 29 -185 .23 -224 . 68 10 ..67

1200 53 . 88 313 . 23 43 . 65 -185 . 69 -228 . 20 9 .. 93

1300 54 .43 317 . 56 49. 07 -186 . 15 -231

.

67 9 ..31

1400 54 .88 321

.

61 54 . 53 -186 .69 -235 . 10 8 .. 77

1500 55 .25 325 . 41 60 . 04 -187 . 19 -238 . 45 8 ,. 30

1600 55 . 57 328 . 99 65 . 58 -187 . 74 -241 . 79 7 .. 89

1700 55 . 84 332 . 36 71 . 15 -188 .28 -245 . 14 7 .. 53

1800 56 . 08 335 . 56 76 . 75 -188 . 82 -248 . 45 7 .. 21

1900 56 . 28 338 . 60 82 . 37 -189 .37 -251 . 75 6 .. 92

2000 56 .45 341

.

49 88 . 00 -189 . 91 -255 . 01 6 .. 66

2100 56 .61 344 . 25 93 . 65 -190 . 50 -258 . 32 6 ..42

2200 56 . 74 346 . 89 99 . 32 -191 . 04 -261

.

58 6 .,21

2300 56 . 86 349 . 41 105 . 00 -191 .63 -264 . 89 6 .. 02

2400 56 . 96 351

.

83 110 . 70 -192 . 17 -268 . 19 5 .. 84

2500 57 . 05 354 . 16 116 . 39 -192 . 76 -271 . 50 5 ,.67

2600 57 . 14 356 . 40 122 . 11 -193 .30 -274 . 76 5 .. 52

2700 57 .21 358 . 56 127 . 82 -193 . 89 -278 . 07 5 .. 38

2800 57 .28 360 . 64 133 . 54 -194 .43 -281 . 33 5 ,.25

2900 57 . 34 362 . 65 139 . 28 -195 . 02 -284 . 60 5 ,. 13

3000 57 .39 364 . 59 145 . 01 -195 . 56 -287 . 82 5 ,. 01
576 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CH2F0*

AH|(0 K)

AH^(298.15 K)

S°(298.15 K)

186 . 73 kJ mol"^

194 .40 kJ mol'^

= 257 . 00 J K" mol"

Moments of Inertia (10 kg m ) 15.36 75.54 85.58

Vibrational Frequencies (cm"^)

527.3 911.9 1047.1 1108.4

1412.9 1439.9 2885.3 2937.7

1198 .

1

Bond Distances (10

Cl-01 1.345

Cl-Fl 1.354

Cl-Hl 1.084

C1-H2 1.084

10 m) Bond Angles (°) 01-Cl-Fl 112.61

01-Cl-Hl 108.34

01-C1-H2 108.34

Dihedral Angles (°)

Fl-Cl-Hl

F1-C1-H2

119 .89

119 . 89

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

t/k C

lO nOS

ttO tt H -H (Tj.) AHo

f AG° Log Kf

150 34 . 80 23 0. 58 - 5 . 78 -190 . 54 -180 . 83 62 ,, 98

200 37 . 20 240 . 89 -3 . 98 -191 . 88 -177 . 44 46 .. 34

250 40 . 56 249 . 53 -2 . 05 -193 . 22 -173 . 68 36 ,.29

300 44 . 61 257 . 28 0 . 08 -194 .43 -169 . 70 29 .. 54

350 48 . 97 264 . 48 2 . 42 -195 . 56 -165 . 48 24 ,. 70

400 53 . 32 271 . 31 4 . 98 -196 . 56 -161 . 17 21

.

. 04

450 57 . 45 277 . 83 7 . 75 -197 .48 -156 . 69 18 ,. 19

500 61 . 27 284 . 08 10 . 72 -198 . 24 -152 . 13 15 .. 89

600 67 . 93 295 . 86 17 . 19 -199 .49 -142 . 84 12 ..43

700 73 . 42 306 . 76 24 . 27 -200 .41 -133 . 30 9.. 95

800 77 . 98 316 . 87 31 . 84 -201 . 04 -123 . 68 8 ,. 08

900 81 . 79 326 . 28 39 . 84 -201 .46 -113 . 97 6 .. 62

1000 85 . 00 335 . 07 48 . 18 -201 . 67 -104 . 22 5 .. 44

1100 87 . 71 343 . 30 56 . 82 -201 . 75 -94 . 43 4 ,.48

1200 90 . 03 351

.

03 65 . 71 -201 . 71 -84 . 60 3 .. 68

1300 92 . 00 358 . 32 74 . 81 -201 . 59 -74 . 77 3 ., 00

1400 93 . 70 365 . 20 84 . 10 -201 . 42 -64 . 98 2 ..42

1500 95 . 16 371

.

71 93 . 55 -201 . 12 -55 . 19 1

.

. 92

1600 96 . 42 377 . 90 103 . 13 -200 . 79 -45 . 40 1

.

.48

1700 97 . 52 383 . 78 112 . 83 -200 . 46 -35 . 65 1

,

. 10

1800 98 . 48 389 . 38 122 . 62 -200 . 04 -25 . 94 0 ..75

1900 99. 32 394 . 73 132 . 52 -199 . 62 -16 . 28 0 ..45

2000 100 . 06 399 . 84 142 . 49 -199 . 16 -6 . 69 0 .. 17

2100 100 . 71 404 . 74 152 . 52 -198 .66 2 . 89 -0 .. 07

2200 101 . 29 409 . 44 162 . 63 -198 . 15 12 . 38 -0 ..29

2300 101

.

81 413 . 95 172 . 78 -197 .61 21 . 84 -0 .. 50

2400 102 . 27 418 . 30 182 . 99 -197 . 07 31

.

25 -0 ..68

2500 102 . 69 422 . 48 193 . 23 -196 . 52 40 . 63 -0 ,.85

2600 103 . 06 426 . 51 203 . 52 -195 . 98 49 . 96 -1

,

. 00

2700 103 . 40 430 . 41 213 . 85 -195 . 39 59 . 25 -1

.

. 15

2800 103 . 70 434 . 18 224 . 20 -194 . 81 68 . 49 -1

.

.28

2900 103 . 98 437 . 82 234 . 58 -194 . 18 77 . 74 -1

.

.40

3000 104 . 23 441. 35 245 . 00 -193 . 59 86 . 94 -1

.

. 51
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 577

CHFjO*

AH|(0 K)

Ah5(298.15 K)

S°(298.15 K)

Moments of Inertia

76.21 78.48

= -398 . 90 kJ mol"^

= -405 . 62 kJ mol'^

= 274 . 61 J K"^ mol'^

(lO"^"^ kg m^)

141 .68

Vibrational Frequencies (cm

443.3 499.4 644.7 1014.8 1147.3

1154.8 1355.1 1390.9 2961.3

Bond Distances (lO"-"-" m)

Cl-01 1.346

Cl-Hl 1.080

Cl-Fl 1.327

C1-F2 1.327

Bond Angles (°) 01-Cl-Hl 106.89

01-Cl-Fl 111.12

01-C1-F2 111.12

Dihedral Angles {°)

Hl-Cl-Fl 120.39

H1-C1-F2 -120.39

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K C° S° H°-H°(T^) AHo

f AG? Log Kf

150 37 .81 244 . 21 -6 . 68 -402 . 58 -389 . 53 135 65

200 42 .61 255 . 73 -4 . 67 -403 . 76 -385 . 01 100 56

250 47 . 71 265 . 79 -2 . 42 -404 . 76 -380 . 24 79 45

300 52 . 85 274 . 94 0 . 10 -405 . 64 -375 . 30 65 34

350 57 . 84 283 . 47 2 . 87 -406 .39 -370 . 20 55 25

400 62 . 51 291

.

50 5 . 88 -407 . 06 -365 . 01 47 66

450 66 . 77 299 . 11 9 . 11 -407 . 56 -359 . 74 41 76

500 70 .58 306 . 35 12 . 55 -407 . 98 -354 . 43 37 02

600 76 . 96 319 . 80 19 . 94 -408 . 57 -343 . 67 29 92

700 81 . 95 332 . 06 27 . 89 -408 . 90 -332 . 84 24 84

800 85 . 89 343 . 27 36 . 29 -409 . 03 -322 . 00 21 02

900 89 . 04 353 . 57 45 . 04 -408 . 99 -311

.

08 18 05

1000 91 .61 363 . 09 54 . 08 -408 . 86 -300 . 20 15 68

1100 93 . 72 371

.

92 63 . 35 -408 .61 -289 . 28 13 74

1200 95 .47 380 . 16 72 . 81 -408 . 32 -278 . 40 12 12

1300 96 . 93 387 . 86 82 . 43 -407 . 98 -267 . 52 10 75

1400 98 . 17 395 . 09 92 . 19 -407 .61 -256 . 65 9 58

1500 99 .22 401

.

90 102 . 06 -407 . 19 -245 . 81 8 56

1600 100 .13 408 . 33 112 . 03 -406 . 73 -235 . 02 7 67

1700 100 . 90 414 . 43 122 . 08 -406 . 22 -224 . 26 6 89

1800 101 . 57 420 . 21 132 . 21 -405 . 72 -213 . 55 6 20

1900 102 . 16 425 . 72 142 . 39 -405 . 22 -202 . 88 5 58

2000 102 .68 430 . 97 152 . 64 -404 .68 -192 . 30 5 02

2100 103 . 12 435 . 99 162 . 93 -404 . 13 -181 . 71 4 52

2200 103 . 52 440 . 80 173 . 26 -403 . 55 -171

.

21 4 07

2300 103 . 88 445 . 41 183 . 63 -403 . 00 -160 . 75 3 65

2400 104 . 19 449 . 84 194 . 04 -402 .42 -150 . 37 3 27

2500 104 .47 454 . 10 204 . 47 -401 . 83 -140 . 04 2 93

2600 104 . 73 458 . 20 214 . 93 -401 .20 -129 . 70 2 61

2700 104 . 96 462 . 16 225 . 41 -400 . 62 -119 . 45 2 31

2800 105 . 16 465 . 98 235 . 92 -399 . 99 -109 . 20 2 04

2900 105 . 35 469 . 67 246 . 45 -399 . 36 -98 . 99 1 78

3000 105 . 52 473 . 25 256 . 99 -398 . 74 -88 . 78 1 55
578 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CFjO*

AH?(0 K) = -623.37 kJ mol"^

AHf(298.15 K) = -628.33 kJ mol"^

S°(298.15 K) = 292.57 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

136.13 139.58 144.70

Vibrational Frequencies (cm"-'^) 224.6 410.9 572.2 583.2 607.1

883.4 1274.8 1277.7 1309.3

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles

Cl-01 1.353

Cl-Fl 1.308 01-Cl-Fl 111.34

C1-F2 1.308 01-C1-F2 111.33 F1-C1-F2 120.51

C1-F3 1.307 01-C1-F3 106.73 F1-C1-F3 -119.74

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^^ = 1 atmosphere

1 - J K'^ mol"^ -

| I kJ mol"^

T/K S° H°-H°(T_) AHo AG° Log Kf

150 44 .19 255 . 71 -8 . 12 -626 . 68 -609 . 94 212 ,.39

200 51 . 68 269 . 51 -5 . 71 -627 . 43 -604 . 25 157 ,. 81

250 58 .40 281 . 78 -2 . 96 -627 . 98 -598 . 44 125 ,. 03

300 64 .49 292 . 97 0 . 12 -628 . 35 -592 . 50 103 ,. 16

350 69 . 89 303 . 33 3 . 48 -628 . 56 -586 . 55 87 ,. 53

400 74 . 63 312 . 98 7 . 10 -628 .69 -580 . 53 75 .. 81

450 78 . 76 322 . 01 10 . 93 -628 .69 -574 . 55 66 ..69

500 82 . 30 330 . 49 14 . 96 -628 .65 -568 . 56 59 .,39

600 87 . 95 346 . 03 23 . 49 -628 . 35 -556 . 60 48 ..45

700 92 . 10 359 . 91 32 . 51 -627 .89 -544 . 67 40 .. 64

800 95 . 17 372 . 42 41 . 87 -627 . 35 -532 . 83 34 .. 79

900 97 .49 383 . 77 51 . 51 -626 . 76 -521

.

03 30 .. 24

1000 99 .25 394 . 14 61 . 35 -626 . 18 -509 . 32 26 ,. 60

1100 100 . 63 403 . 66 71 . 35 -625 . 51 -497 . 60 23 .. 63

1200 101 . 72 412 . 47 81

.

47 -624 . 88 -485 . 93 21

.

. 15

1300 102 . 59 420 . 65 91

.

69 -624 . 25 -474 . 30 19 .. 06

1400 103 .29 428 . 28 101

.

98 -623 . 58 -462 . 75 17 ., 26

1500 103 . 88 435 . 42 112 . 34 -622 . 96 -451

.

20 15 .. 71

1600 104 . 36 442 . 14 122 . 75 -622 .29 -439 . 74 14 ., 35

1700 104 . 77 448 . 48 133 . 21 -621 . 66 -428 . 32 13 ., 16

1800 105 . 11 454 . 48 143 . 71 -620 . 99 -416 . 94 12 ., 10

1900 105 .41 460 . 17 154 . 23 -620 . 36 -405 . 60 11 ., 15

2000 105 . 66 465 . 59 164 . 79 -619 . 73 -394 . 34 10 ..30

2100 105 . 88 470 . 75 175 . 36 -619 . 06 -383 . 13 9 ., 53

2200 106 . 07 475 . 67 185 . 96 -618 .40 -372 . 00 8 .,83

2300 106 . 24 480 . 39 196 . 58 -617 . 77 -360 . 91 8 .,20

2400 106 . 38 484 . 92 207 . 21 -617 . 10 -349 . 87 7 ., 61

2500 106 . 52 489 . 26 217 . 85 -616 .43 -338 . 90 7 ,, 08

2600 106 .63 493 . 44 228 . 51 -615 .80 -327 . 94 6 ., 59

2700 106 . 73 497 . 47 239 . 18 -615 . 13 -317 . 06 6 ., 13

2800 106 .83 501 . 36 249 . 86 -614 .46 -306 . 19 5 ., 71

2900 106 . 91 505 . 11 260 . 54 -613 . 79 -295 . 35 5 . 32

3000 106 . 99 508 . 73 271

.

24 -613 . 12 -284 . 51 4 . 95
SECTION 5. FLAME INHIBrnON CHEMISTRY AND THE SEARCH 579

•CFjOH

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

Moments of Inertia

76.29 78.01

= -451 .87 kJ mol"^

= -456 .47 kJ mol"^

= 298 . 07 J K"^ mol"^

(10"^^ kg m^)

146 . 52

Vibrational Frequencies (cm"-"-) 30.9 500.4 514.0 623.0 987.3

1215.0 1267.2 1301.6 3570.6

Bond Distances (lO"-"-" m)

Cl-01 1.324

Cl-Fl 1.318

C1-F2 1.317

01-Hl 0.954

Bond Angles (°) 01-Cl-Fl 114.03

01-C1-F2 114.35

Cl-Ol-Hl 111.32

Dihedral Angles (°)

F1-C1-F2 126.17

F2-01-H1 68.56

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

|- J K"^ mol"^ -| I kJ mol"^

T/K S H -H AHo

f AG° Log Kf

150 4 5 . 43 262 . 55 - 7 . 78 - 4 54 . 55 -444 . 26 154 ,. 6 9 200 50 . 13 276 . 2 5 - 5 . 40 -455 . 30 -440 . 70 115 ., 10

250 55 . 09 287 . 97 -2 . 77 -455 . 97 -437 . 02 91

.

.30

300 59 . 94 298 . 44 0 . 11 -456 .47 -433 . 17 75 ,.42

350 64 .49 308 . 03 3 . 22 -456 .89 -429 . 28 64 .. 06

400 68 . 64 316 . 92 6 . 55 -457 .23 -425 . 35 55 .. 54

450 72 . 34 325 . 22 10 . 08 -457 . 44 -421 . 37 48 .. 91

500 75 . 59 333 . 01 13 . 78 -457 .60 -417 . 35 43 ..60

600 80 . 90 347 . 28 21

.

61 -457 . 73 -409 . 32 35 .. 63

700 84 . 95 360 . 08 29 . 92 -457 . 73 -401 . 29 29 .. 94

800 88 . 11 371

.

63 38 . 58 -457 . 56 -393 . 25 25 ., 67

900 90 .63 382 . 16 47 . 52 -457 .35 -385 . 18 22 .. 36

1000 92 . 70 391. 82 56 . 69 -457 . 10 -377 . 15 19 .. 70

1100 94 .43 400 . 74 66 . 05 -456 . 77 -369 . 11 17 ., 53

1200 95 . 90 409 . 02 75 . 57 -456 .43 -361

.

12 15 .. 72

1300 97 . 15 416 . 75 85 . 22 -456 . 06 -353 . 13 14 .. 19

1400 98 . 23 423 . 99 94 . 99 -455 . 64 -345 . 14 12 ,. 88

1500 99 . 17 430 . 80 104 . 86 -455 .22 -337 . 23 11 ., 74

1600 99 . 99 437 . 22 114 . 82 -454 . 76 -329 . 32 10 .. 75

1700 100 . 71 443 . 31 124 . 85 -454 .30 -321

.

46 9 .. 88

1800 101 . 34 449 . 08 134 . 96 -453 . 84 -313 . 63 9 ., 10

1900 101 . 90 454 . 58 145 . 12 -453 . 34 -305 . 85 8 .,41

2000 102 .40 459 . 82 155 . 34 -452 . 83 -298 . 11 7 .. 79

2100 102 . 84 464 . 83 165 . 60 -452 .29 -290 . 45 7 ..22

2200 103 . 24 469 . 62 175 . 90 -451 . 79 -282 . 84 6 ., 71

2300 103 . 59 474 . 21 186 . 25 -451 . 24 -275 . 27 6 ., 25

2400 103 . 91 478 . 63 196 . 62 -450 . 66 -267 . 73 5 .. 83

2500 104 .20 482 . 88 207 . 03 -450 . 11 -260 . 29 5 .. 44

2600 104 . 46 486 . 97 217 . 46 -449 . 53 -252 . 84 5 .. 08

2700 104 . 69 490 . 92 227 . 92 -448 . 94 -245 . 43 4 .. 75

2800 104 . 91 494 . 73 238 . 40 -448 . 36 -238 . 07 4 .. 44

2900 105 . 10 498 . 41 248 . 90 -447 . 77 -230 . 75 4 .. 16

3000 105 . 28 501 . 98 259 . 42 -447 . 14 -223 . 43 3 .. 89
580 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHjFOH hocf trans

AH?(0 K) = -400.07 kJ mol"^

AHf(298.15 K) = -412.09 kJ mol"^

S°{298.15 K) = 252.25 J K"^ mol"^

Moments of Inertia (10'^^ kg m^)

17.26 77.89 89.86

Vibrational Frequencies (cm"-'-) -115.0 507.5 1024.4

1237.7 1250.2 1482.3

2917.0 3671.3

Bond Distances (10"-'^° m)

Cl-01 1.377

Cl-Fl 1.344

Cl-Hl 1.084

C1-H2 1.084

01-H3 0.947

1125.5 1150.2

1536.4 2875.1

Bond Angles (°) Dihedral Angles (°) 01-Cl-Fl 106.83

01-Cl-Hl 111.76 Fl-Cl-Hl 117.94

01-C1-H2 111.76 F1-C1-H2 -117.94

C1-01-H3 110.14 F1-01-H3 180.00

Enthalpy Reference Temperature = T^. = 2 98.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K C° S° H°-H°{T^) AH0

f AG? Log Kf

150 34 .I9 225 . 93 -5 . 75 -406 . 06 -387 . 40 134 90

200 37 . 10 236 . 25 -3 . 96 -408 . 11 -380 . 91 99 48

250 40 .23 244 . 84 -2 . 03 -410 . 16 -373 . 88 78 11

300 44 .23 252 . 52 0 . 08 -412 . 17 -366 . 43 63 80

350 48 . 78 259 . 68 2 . 41 -414 . 05 -358 . 69 53 53

400 53 . 50 266 . 50 4 . 96 -415 . 76 -350 . 70 45 80

450 58 . 14 273 . 07 7 . 75 -417 .35 -342 . 50 39 75

500 62 . 53 279 . 43 10 . 77 -418 . 82 -334 . 13 34 90

600 70 . 37 291 . 54 17 . 43 -421 . 33 -316 . 98 27 59

700 77 . 01 302 . 90 24 . 81 -423 .38 -299 . 45 22 34

800 82 .63 313 . 56 32 . 80 -425 . 09 -281

.

58 18 39

900 87 .43 323 . 58 41

.

30 -426 . 43 -263 . 55 15 30

1000 91 . 55 333 . 01 50 . 26 -427 . 56 -245 . 39 12 82

1100 95 . 11 341

.

90 59 . 60 -428 .44 -227 . 07 10 78

1200 98 .20 350 . 32 69 . 27 -429 . 15 -208 . 66 9 08

1300 100 . 88 358 . 28 79 . 22 -429 . 70 -190 . 20 7 64

1400 103 . 22 365 . 85 89 . 43 -430 . 16 -171

.

67 6 41

1500 105 . 26 373 . 04 99 . 86 -430 .45 -153 . 13 5 33

1600 107 . 05 379 . 89 110 . 47 -430 . 70 -134 . 60 4 39

1700 108 . 62 386 . 43 121 . 26 -430 . 87 -116 . 06 3 57

1800 110 . 01 392 . 68 132 . 19 -430 . 95 -97 . 53 2 83

1900 111 . 24 398 . 66 143 . 26 -430 . 99 -78 . 99 2 17

2000 112 .33 404 . 40 154 . 44 -430 . 95 -60 . 50 1 58

2100 113 . 30 409 . 90 165 . 72 -430 . 91 -42 . 01 1 04

2200 114 . 16 415 . 19 177 . 09 -430 . 78 -23 . 56 0 56

2300 114 . 94 420 . 28 188 . 55 -430 . 66 -5 . 15 0 12

2400 115 . 64 425 . 19 200 . 08 -430 . 53 13 . 22 -0 29

2500 116 .26 429 . 92 211 . 67 -430 . 32 31

.

55 -0 66

2600 116 . 83 434 . 50 223 . 33 -430 . 12 49 . 87 -1 00

2700 117 . 34 438 . 91 235 . 04 -42 9 . 86 68 . 16 -1 32

2800 117 . 81 443 . 19 246 . 80 -429 .61 86 . 44 -1 61

2900 118 . 24 447 . 33 258 . 60 -429 .36 104 . 73 -1 89

3000 118 . 62 451

.

34 270 . 44 -429 . 07 123 . 01 -2 14
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 581

CHjFOH hocf cis

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

= -408 . 78 kJ mol"^

= -420 . 73 kJ mol"^

= 252 .71 J K" mol -1

Moments of Inertia (10 ^"^ kg m^]

18.72 76.24 89.66

Vibrational Frequencies (cm" ) -375.3 490.1 997.9 1120.0

1239.3 1257.0 1460.0 1542.4

2963.3 3672.2

1150 .

2913 .

Bond Distances (10

Cl-01 1.376

Cl-Fl 1.360

Cl-Hl 1.081

C1-H2 1.081

01-H3 0.948

•10 Bond Angles (°) 01-Cl-Fl

01-Cl-Hl

01-C1-H2

C1-01-H3

109 . 55

110 .31

110 .31

108 . 83

Dihedral Angles

Fl-Cl-Hl 118.65

F1-C1-H2 -118.65

F1-01-H3 0.00

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

]- J K'^ mol"^ -| I kJ mol"^

T/K (nO S° H°-H°(Tj.) AHf AG? Log Kf

150 35 226 . 22 -5 . 79 -414 . 72 -396 . 14 137 .. 94

200 37 . 34 236 . 61 -3 . 98 -416 . 81 -389 . 66 101

. 76

250 40 . 51 245 . 26 -2 . 04 -418 . 82 -382 . 63 79 .. 94

300 44 . 53 252 . 99 0 . 08 -420 . 78 -375 . 22 65 ..33

350 49 . 08 260 . 19 2 . 42 -422 .67 -367 . 52 54 ..85

400 53 . 80 267 . 05 4 . 99 -424 . 38 -359 . 53 46 .. 95

450 58 .41 273 . 66 7 . 80 -425 . 97 -351

37 40 .. 78

500 62 . 77 280 . 04 10 . 83 -427 .40 -343 . 00 35 .. 83

600 70 . 55 292 . 19 17 . 51 -429 . 91 -325 . 93 28 ., 37

700 77 . 12 303 . 57 24 . 90 -431 . 96 -308 . 44 23 .. 02

800 82 .68 314 . 25 32 . 90 -433 . 63 -290 . 70 18 ., 98

900 87 .43 324 . 27 41 . 41 -435 . 01 -272 . 71 15 ., 83

1000 91 . 52 333 . 69 50 . 36 -436 . 10 -254 . 60 13 ., 30

1100 95 . 06 342 . 59 59 . 69 -436 . 98 -236 . 35 11 .. 22

1200 98 . 12 350 . 99 69 . 36 -437 . 73 -218 . 03 9 ..49

1300 100 . 80 358 . 95 79 . 31 -438 .27 -199. 62 8 .. 02

1400 103 . 13 366 . 51 89 . 50 -438 . 73 -18121 6 .. 76

1500 105 . 17 373 . 70 99 . 92 -439 . 07 -162 . 72 5 .,67

1600 106 . 96 380 . 54 110 . 53 -439 .28 -144 . 22 4 .. 71

1700 108 . 53 387 . 08 121 . 31 -439 .45 -125 . 77 3 ,. 86

1800 109 . 92 393 . 32 132 . 23 -439 . 57 -107 . 28 3 ., 11

1900 111 . 15 399 . 30 143 . 29 -439 . 61 -88 . 83 2 .. 44

2000 112 . 24 405 . 03 154 . 46 -439 . 57 -70 . 37 1 .. 84

2100 113 .22 410 . 53 165 . 73 -439 . 53 -51

97 1 ..29

2200 114 . 09 415 . 81 177 . 10 -439 .45 -33 . 60 0 ,. 80

2300 114 . 87 420 . 91 188 . 54 -439 . 32 -15 . 23 0 .,35

2400 115 . 57 425 . 81 200 . 07 -439 . 15 3 . 10 -0 .. 07

2500 116 .20 430 . 54 211 . 66 -438 . 99 21 . 34 -0 ..45

2600 116 . 77 435 . 11 223 . 30 -438 . 78 39 . 62 -0 ,. 80

2700 117 .29 439 . 53 235 . 01 -438 . 57 57 . 86 -1

. 12

2800 117 . 76 443 . 80 246 . 76 -438 . 32 76 . 07 -1

. 42

2900 118 . 19 447 . 94 258 . 56 -438 . 06 94 . 27 -1

. 70

3000 118 . 58 451

95 270 . 40 -437 . 77 112 . 47 -1

, 96
582 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHjFOH hocf gauche

AH?(0 K)

AHf(298.15 K)

S°(298.15 K)

-419.24 kJ mol"^

-430.32 kJ mol"^

257 .60 J K"^ mol"^

Moments of Inertia (10

18.02 78.78 89.64

kg m-^

Vibrational Frequencies (cm"-"^) 351.4

1244 . 4 2990 .

6

528 . 7 1362 . 7 3657 .

1

1019 .

3

1446 . 2 1068 .

8

1522 .

1

1132 . 5 2903 . 0

Bond Distances (10"^°

Cl-01 1.364

Cl-Fl 1.361

Cl-Hl 1.077

C1-H2 1.083

01-H3 0.948

m) Bond Angles

01-Cl-Fl

01-Cl-Hl

01-C1-H2

C1-01-H3

110 . 75

107 .47

112 .40

109 . 90

Dihedral Angles (°)

Fl-Cl-Hl 118.27

F1-C1-H2 -119.38

F1-01-H3 63.74

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K"^ mol"^ -I I kJ mol"^

T/K 1C° S° H°-H°(T^) AHo

f AG? Log Kf

150 38 .24 227 . 38 -6 . 56 -425 . 09 -406 ,.68 141

,

.61

200 42 . 16 239 . 00 -4 . 55 -426 . 94 -400 ,.28 104 ,. 54

250 46 . 56 248 . 97 -2 . 34 -428 . 73 -393 ,.46 82 ,.20

300 51 . 52 257 . 92 0 . 10 -430 . 37 -386 ..27 67 ,.26

350 56 . 76 266 . 25 2 . 79 -431 . 87 -378 ,. 86 56 ,. 54

400 61 . 97 274 . 16 5 . 74 -433 .21 -371

.

.20 48 ,.47

450 66 . 89 281 . 72 8 . 95 -434 . 38 -363 .,42 42 ,. 18

500 71 .41 288 . 98 12 . 39 -435 .43 -355 ,. 51 37 ,. 14

600 79 . 18 302 . 68 19. 92 -437 . 06 -339 ..41 29 .. 55

700 85 .49 315 . 34 28 . 17 -438 .27 -323 ,. 00 24 ,. 10

800 90 .69 327 . 08 36 . 99 -439 . 11 -306 ..48 20 ,. 01

900 95 . 05 338 . 01 46 . 31 -439 . 70 -289 .. 78 16 ,. 82

1000 98 . 76 348 . 21 56 . 03 -440 . 03 -273 .. 05 14 ..26

1100 101 . 96 357 . 77 66 . 10 -440 . 16 -256 .. 23 12 .. 17

1200 104 . 73 366 . 75 76 . 47 -440 .20 -239 ..41 10 .,42

1300 107 . 13 375 . 23 87 . 10 -440 . 07 -222 ,. 59 8 ,, 94

1400 109 . 24 383 . 24 97 . 96 -439 . 86 -205 .. 73 7 .. 68

1500 111 . 08 390 . 84 109 . 01 -439 . 57 -188 .. 95 6 .. 58

1600 112 .69 398 . 06 120 . 23 -439 . 19 -172 ,. 13 5 .. 62

1700 114 . 12 404 . 93 131

.

60 -438 . 73 -155 ,.39 4 .. 77

1800 115 .38 411 . 49 143 . 11 -438 .27 -138 .. 70 4 .. 02

1900 116 .49 417 . 83 154 . 73 -437 . 73 -122 ., 17 3 .. 36

2000 117 .48 423 . 99 166 . 46 -437 . 19 -105 .. 86 2 .. 76

2100 118 .37 429 . 89 178 . 28 -436 . 56 -89 ..66 2 ..23

2200 119 . 16 435 . 56 190 . 18 -435 . 93 -73 .. 51 1 .. 75

2300 119 . 87 441

.

01 202 . 15 -435 .30 -57 ..45 1 ., 31

2400 120 . 51 446 . 27 214 . 20 -434 .63 -41

.

. 51 0 .. 90

2500 121 . 08 451

.

34 226 . 30 -433 . 92 -25 .. 56 0 ., 53

2600 121 .60 456 . 23 238 . 45 -433 .21 -9 .. 75 0 ., 20

2700 122 . 07 460 . 96 250 . 66 -432 . 50 6 ., 02 -0 ., 12

2800 122 . 50 465 . 53 262 . 90 -431 . 75 21

.

. 76 -0 ..41

2900 122 . 89 469 . 96 275 . 19 -430 . 99 37 .,45 -0 ..67

3000 123 .25 474 . 26 287 . 51 -430 . 24 53 .. 09 -0 ., 92
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 583

CHF2OH hcoh trans

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

674 .25 kJ mol"^

684 . 31 kJ mol"^

275.29 J K

-1 mol"

Moments of Inertia (10 kg m^)

80.43 81.32 144.72

Vibrational Frequencies (cm"-'-) 300.8 498.6 532.3

1118.6 1192.0 1309.9

3040.2 3630.2

640 . 5 1384 . 2 1025 .

3

1437 .

7

Bond Distances (10

Cl-01 1.340

01-Hl 0.951

C1-H2 1.073

Cl-Fl 1.337

C1-F2 1.337

-10 m) Bond Angles (°) Cl-Ol-Hl

01-C1-H2

01-Cl-Fl

01-C1-F2

109 . 83

109 . 50

110 . 98

110 . 98

Dihedral Angles

H1-C1-H2 180.00

H2-C1-F1 121.21

H2-C1-F2 -121.21

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K'^ mol'^ -I

I kJ mol"^

T/K S° H°-H°{Tj.) AHf AG° Log Kf

150 41 .48 240 . 87 -7 . 52 -679 . 98 -658 . 19 229 18

200 47 .87 253 . 81 -5 . 29 -681 . 62 -650 . 70 169 94

250 54 .35 265 . 22 -2 . 75 -683 . 08 -642 . 83 134 31

300 60 . 68 275 . 67 0 . 11 -684 . 34 -634 . 71 110 50

350 66 .65 285 . 45 3 . 28 -685 . 42 -626 . 34 93 47

400 72 . 12 294 . 67 6 . 74 -686 . 34 -617 . 89 80 68

450 77 . 00 303 . 43 10 . 46 -687 . 14 -609 . 27 70 72

500 81 .30 311

.

74 14 . 42 -687 . 72 -600 . 61 62 74

600 88 .39 327 . 18 22 . 93 -688 .69 -583 . 08 50 76

700 93 . 90 341 . 21 32 . 08 -689 .27 -565 . 43 42 19

800 98 .30 354 . 03 41 . 73 -689 .61 -547 . 69 35 76

900 101 . 92 365 . 81 51 . 78 -689 . 73 -529 . 86 30 75

1000 104 . 95 376 . 70 62 . 17 -689 . 73 -512 . 00 26 74

1100 107 . 53 386 . 82 72 . 83 -689 . 61 -494 . 13 23 46

1200 109 . 75 396 . 27 83 . 73 -689 .40 -476 . 26 20 73

1300 111 .67 405 . 13 94 . 84 -689 . 10 -458 . 44 18 42

1400 113 . 34 413 . 47 106 . 12 -688 . 73 -440 . 58 16 44

1500 114 . 80 421

.

53 117 . 56 -688 .31 -423 . 09 14 73

1600 116 . 08 429 . 17 129 . 14 -687 .85 -405 . 64 13 24

1700 117 .20 436 . 43 140 . 83 -687 .35 -388 . 28 11 93

1800 118 .20 443 . 35 152 . 62 -686 . 80 -370 . 95 10 76

1900 119 . 08 449 . 94 164 . 51 -686 .22 -353 . 76 9 72

2000 119 . 86 456 . 24 176 . 49 -685 . 59 -336 . 64 8 79

2100 120 . 56 462 . 28 188 . 53 -684 . 96 -319 . 57 7 95

2200 121 . 19 468 . 07 200 . 64 -684 . 34 -302 . 63 7 18

2300 121 . 75 473 . 63 212 . 80 -683 . 62 -285 . 73 6 49

2400 122 .25 478 . 98 225 . 02 -682 . 95 -268 . 91 5 85

2500 122 . 70 484 . 12 237 . 28 -682 . 24 -252 . 17 5 27

2600 123 . 11 489 . 09 249 . 59 -681 . 53 -235 . 52 4 73

2700 123 .49 493 . 88 261 . 93 -680 . 78 -218 . 91 4 23

2800 123 .83 498 . 52 274 . 31 -680 . 03 -202 . 34 3 77

2900 124 . 14 503 . 00 286 . 73 -679 .27 -185 . 81 3 35

3000 124 .42 507 . 35 299 . 16 -678 .48 -169 . 33 2 95
584 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHFjOH hcoh gauche

AH2(0 K)

AHf(298.15 K)

S°(298.15 K)

662 . 24 kJ mol"^

671 . 82 kJ mol"^

279.18 J K"^ mol"^

Moments of Inertia (10

79.20 81.24

47 kg m'^ 146 .47

Vibrational Frequencies (cm"

199.0 485.7 489.9

1096.2 1199.5 1264.6

2971.2 3666.2

679 . 7 1403 . 9 1066 . 7 1457 . 5

Bond Distances (10

Cl-01 1.348

01-Hl 0.949

C1-H2 1.079

Cl-Fl 1.321

C1-F2 1.339

10 Bond Angles (°) Cl-Ol-Hl

01-C1-H2

01-Cl-Fl

01-C1-F2

110 . 08

113 .85

107 . 51

110 . 36

Dihedral Angles

H1-C1-H2

H2-C1-F1

H2-C1-F2

68 . 50

121 . 02

121 . 31

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K C

o S° H°-H°(T^) AHo

f AG° Log Kf

150 44 . l3 244 . 12 -7 . 75 -667 . 72 -646 . 39 225 08

200 49 . 88 257 . 57 -5 . 40 -669 .23 -639 . 06 166 90

250 54 . 93 269 . 15 -2 . 77 -670 .61 -631

.

37 131 91

300 59 . 97 279 . 55 0 . 11 -671 . 87 -623 . 37 108 53

350 64 . 99 289 . 12 3 . 26 -672 . 95 -615 . 17 91 81

400 69 . 82 298 . 09 6 . 65 -673 . 96 -606 . 85 79 24

450 74 . 35 306 . 57 10 . 28 -674 .80 -598 . 40 69 46

500 78 . 49 314 . 60 14 . 13 -675 . 55 -589 . 86 61 62

600 85 . 61 329 . 55 22 . 40 -676 . 72 -572 . 58 49 84

700 91

.

37 343 . 62 31

.

30 -677 . 56 -555 . 38 41 44

800 96 . 08 356 . 56 40 . 72 -678 . 10 -538 . 23 35 14

900 99 . 99 368 . 51 50 . 56 -678 .48 -521

.

03 30 24

1000 103 . 27 379 . 59 60 . 76 -678 . 64 -503 . 84 26 32

1100 106 . 08 389 . 92 71 . 25 -678 . 73 -486 . 64 23 11

1200 108 . 48 399 . 58 82 . 00 -678 . 64 -469 . 49 20 44

1300 110 . 56 408 . 65 92 . 98 -678 .48 -452 . 37 18 18

1400 112 . 36 417 . 20 104 . 14 -678 .23 -435 . 30 16 24

1500 113 . 93 425 . 27 115 . 47 -677 . 93 -418 . 27 14 57

1600 115 . 30 432 . 92 126 . 95 -677 . 56 -401 . 33 13 10

1700 116 . 51 440 . 19 138 . 55 -677 . 14 -384 . 43 11 81

1800 117 . 57 447 . 10 150 . 27 -676 .68 -367 . 61 10 67

1900 118 . 51 453 . 70 162 . 08 -676 . 18 -350 . 83 9 64

2000 119 . 35 460 . 01 173 . 98 -675 .63 -334 . 18 8 73

2100 120 . 09 466 . 05 185 . 97 -675 . 05 -317 . 57 7 90

2200 120 . 75 471 . 83 198 . 02 -674 .46 -301 . 04 7 15

2300 121 . 35 477 . 39 210 . 13 -673 . 83 -284 . 55 6 46

2400 121 . 88 482 . 75 222 . 30 -673 . 16 -268 . 19 5 84

2500 122 . 37 487 . 90 234 . 52 -672 .49 -251

.

88 5 26

2600 122 . 80 492 . 86 246 . 78 -671 . 82 -235 . 64 4 73

2700 123 . 20 497 . 66 259 . 09 -671 . 11 -219 . 45 4 25

2800 123 . 55 502 . 29 271 . 43 -670 .40 -203 . 30 3 79

2900 123 . 88 506 . 78 283 . 80 -669 .69 -187 . 19 3 37

3000 124 . 18 511 . 12 296 . 21 -668 . 94 -171 . 13 2 . 98
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 585

CF3OH

AH2(0 K) = -910.77 kJ mol"^

AHf(298.15 K) = -919.44 kJ mol"^

S°(298.15 K) = 290.10 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

143.10 144.11 145.53

Vibrational Frequencies (cm"-'-) 233.3 427.8 440.7 583.1 607.5

617.1 888.9 1123.5 1234.9 1323.7

1414.0 3654.0

Bond Distances (10'-'^° m) Bond Angles (°) Dihedral Angles

Cl-01 1.331

Cl-Fl 1.320 01-Cl-Fl 111.91

C1-F2 1.320 01-C1-F2 111.91 F1-C1-F2 119.76

C1-F3 1.304 01-C1-F3 108.83 F1-C1-F3 -120.12

01-Hl 0.950 Cl-Ol-Hl 110.26 Fl-Ol-Hl -59.80

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K"^ mol"^ - II kJ mol"^

I / IS. nO 0 11 ~ n

0

f AG° Log

150 46 :io 250 . 71 -8 . 70 -916 .21 -890 . 48 310 08

200 54 . 91 265 . 20 -6 . 17 -917 . 55 -881

.

69 230 27

250 63 . 14 278 . 35 -3 . 22 -918 . 64 -872 . 62 182 32

300 70 . 58 290 . 53 0 . 13 -919 .48 -863 . 37 150 32

350 77 .21 301

.

92 3 . 83 -920 . 10 -854 . 00 127 45

400 83 . 05 312 . 62 7 . 84 -920 . 52 -844 . 54 110 28

450 88 . 13 322 . 70 12 . 12 -920 . 81 -835 . 04 96 93

500 92 . 53 332 . 22 16 . 64 -921 . 02 -825 . 55 86 24

600 99 . 61 349 . 75 26 . 26 -921 . 07 -806 . 47 70 20

700 104 . 93 365 . 52 36 . 51 -920 . 86 -787 . 39 58 75

800 109 . 01 379 . 81 47 . 21 -920 . 44 -768 . 35 50 17

900 112 . 22 392 . 84 58 . 27 -919 .89 -749 . 35 43 49

1000 114 . 81 404 . 81 69 . 63 -919 .27 -730 . 40 38 15

1100 116 . 95 415 . 85 81

.

22 -918 . 56 -711

.

49 33 78

1200 118 . 74 426 . 11 93 . 01 -917 . 76 -692 . 66 30 15

1300 120 .26 435 . 68 104 . 96 -916 . 88 -673 . 88 27 07

1400 121 .55 444 . 63 117 . 06 -916 . 00 -655 . 13 24 44

1500 122 . 67 453 . 06 129 . 27 -915 . 08 -636 . 47 22 16

1600 123 . 65 461 . 01 141 . 59 -914 . 12 -617 . 85 20 17

1700 124 .49 468 . 53 153 . 99 -913 . 16 -599 . 36 18 41

1800 125 .24 475 . 67 166 . 48 -912 . 11 -580 . 91 16 86

1900 125 .89 482 . 46 179 . 04 -911 . 11 -562 . 50 15 46

2000 126 .47 488 . 93 191 . 66 -910 . 02 -544 . 21 14 21

2100 126 . 98 495 . 11 204 . 33 -908 . 97 -526 . 01 13 08

2200 127 . 44 501

.

03 217 . 05 -907 . 84 -507 . 85 12 06

2300 127 . 85 506 . 70 229 . 81 -906 . 76 -489 . 82 11 12

2400 128 . 22 512 . 16 242 . 62 -905 . 63 -471 . 83 10 27

2500 128 . 56 517 . 40 255 . 46 -904 . 50 -453 . 92 9 48

2600 128 . 86 522 . 44 268 . 33 -903 .33 -436 . 06 8 76

2700 129 . 13 527 . 31 281 . 23 -902 .20 -418 . 27 8 09

2800 129 . 38 532 . 01 294 . 16 -901 . 02 -400 . 53 7 47

2900 129 . 60 536 . 56 307 . 11 -899 .81 -382 . 84 6 90

3000 129 . 81 540 . 95 320 . 08 -898 . 64 -365 . 18 6 36
586 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CFjOO*

AH2(0 K) = -620.53 kJ mol"^

AH°(298.15 K) = -627.46 kJ mol"^

S°(298.15 K) = 314.37 J K"^ mol"^

Moments of Inertia (10"^"^ kg m^)

144.76 249.01 251.97

Vibrational Frequencies (cm'^)

122.1 278.9 420.0 442.3 571.0

593.0 686.4 878.3 1124.2 1251.0

1287.7 1340.3

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) Cl-01 1.372

01-02 1.314 Cl-01-02 110.50

Cl-Fl 1.300 01-Cl-Fl 106.06 02-Cl-Fl 180.00

C1-F2 1.304 01-C1-F2 111.13 02-C1-F2 60.61

C1-F3 1.304 01-C1-F3 111.13 02-C1-F3 -60.61

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -j I kJ mol"^

T/K 1%S° H°-H°{T^) AHo

f AG? Log Kf

150 53 269 . 64 -9 . 87 -625 .42 -596 . 93 207 . 86

200 62 . 72 286 . 11 -6 . 98 -626 .39 -587 . 27 153 . 37

250 71 . 63 301 . 09 -3 . 63 -627 . 06 -577 . 43 120 . 64

300 79 . 61 314 . 87 0 . 15 -627 . 47 -567 . 52 98 . 81

350 86 . 58 327 . 67 4 . 30 -627 . 64 -557 . 52 83 . 20

400 92 . 55 339 . 62 8 . 77 -627 . 64 -547 . 56 71 . 50

450 97 . 59 350 . 82 13 . 52 -627 .47 -537 . 56 62 ,.40

500 101 . 82 361

.

32 18 . 51 -627 .22 -527 . 60 55 ,. 12

600 108 . 32 380 . 49 29 . 04 -626 .47 -507 . 77 44 .20

700 112 . 89 397 . 54 40 . 12 -625 . 55 -488 . 06 36 ,.42

800 116 . 17 412 . 84 51 . 60 -624 . 50 -468 . 48 30 ,. 59

900 118 . 57 426 . 67 63 . 35 -623 .46 -449 . 03 26 ,. 06

1000 120 . 37 439 . 26 75 . 32 -622 .41 -429 . 65 22 ,. 44

1100 121 . 75 450 . 80 87 . 45 -621 . 37 -410 . 37 19 ,.49

1200 122 . 83 461 . 44 99 . 70 -620 . 32 -391 . 12 17 ,. 03

1300 123 .68 471 . 30 112 . 03 -619 . 32 -372 . 00 14 ., 95

1400 124 .37 480 . 49 124 . 45 -618 . 31 -352 . 92 13 .. 17

1500 124 . 93 489 . 16 136 . 93 -617 . 31 -334 . 05 11 ,. 63

1600 125 .39 497 . 43 149 . 46 -616 .35 -315 . 43 10 ..30

1700 125 . 78 505 . 23 162 . 03 -615 .38 -296 . 90 9 .. 12

1800 126 . 10 512 . 61 174 . 64 -614 .42 -278 . 49 8 .. 08

1900 126 . 38 519 . 62 187 . 27 -613 . 50 -260 . 20 7 .. 15

2000 126 . 62 526 . 28 199 . 93 -612 . 54 -241

.

96 6 .. 32

2100 126 . 83 532 . 63 212 . 61 -611 . 62 -223 . 84 5 .. 57

2200 127 . 01 538 . 69 225 . 31 -610 . 70 -205 . 85 4 ..89

2300 127 . 16 544 . 50 238 . 03 -609 . 78 -187 . 90 4 ..27

2400 127 .30 550 . 07 250 . 76 -608 . 86 -170 . 08 3 .. 70

2500 127 .42 555 . 42 263 . 50 -607 .89 -152 . 34 3 .. 18

2600 127 . 53 560 . 56 276 . 26 -606 . 97 -134 . 64 2 .. 70

2700 127 . 62 565 . 51 289 . 02 -606 . 09 -117 . 03 2 ..26

2800 127 . 71 570 . 30 301 . 79 -605 . 17 -99 . 45 1 .. 86

2900 127 . 79 574 . 91 314 . 57 -604 .25 -81 . 96 1 ..48

3000 127 . 86 579 . 37 327 . 36 -603 .33 -64 . 48 1

.

. 12
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 587

CF3OOH

AH|{0 K)

AHf(298.15 K)

S°(298.15 K)

= -797.30 kJ mol"^

= -807.48 kJ mol-i

= 314 . 50 J K'^ mol

Moments of Inertia (10 kg m^]

146 . 17 260.61 262 .43

Vibrational Frequencies (cm"-"-) 138.6 256.6 289.0 429.1 437.1

571.9 604.2 673.8 879.1 1067.5

1262.5 1293.6 1318.6 1438.0 3629.8

Bond Distances (10 10 m Cl-01

01- 02

Cl-Fl

C1-F2

C1-F3

02 - HI

351

388

305

306

314

0 . 952

Bond Angles

Cl-01-02

01-Cl-Fl

01-C1-F2

01-C1-F3

01-02-Hl

108 . 42

106 . 36

112 . 36

111.48

102 .23

Dihedral Angles (°)

02-Cl-Fl

02-C1-F2

02-C1-F3

C1-02-H1

179 . 75

60 . 07

-61.40

104 . 50

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

i - J K"^ mol"^ -I I kJ mol"^

T/K r,0 S° H°-H°(T^) AH0

f AGf Log Kf

150 57 lo 265 . 92 -10 . 77 -804 . 16 -766 . 89 267 04

200 68 .35 283 . 91 -7 . 62 -805 . 63 -754 . 25 196 98

250 78 . 04 300 . 13 -3 . 96 -806 . 72 -741 . 24 154 87

300 86 . 75 315 . 03 0 . 16 -807 . 51 -728 . 06 126 76

350 94 . 56 328 . 95 4 . 69 -808 . 01 -714 . 79 106 67

400 101 . 50 342 . 05 9 . 59 -808 . 31 -701 . 49 91 60

450 107 . 57 354 . 36 14 . 81 -808 .43 -688 . 18 79 88

500 112 . 81 365 . 97 20 . 32 -808 .39 -674 . 84 70 50

600 121 . 14 387 . 31 32 . 02 -807 .89 -648 . 19 56 43

700 127 .21 406 . 46 44 . 45 -807 . 09 -621

66 46 39

800 131 . 70 423 . 75 57 . 40 -806 . 05 -595 . 26 38 86

900 135 . 08 439 . 46 70 . 75 -804 . 88 -568 . 94 33 02

1000 137 . 70 453 . 83 84 . 40 -803 . 62 -542 . 79 28 35

1100 139 . 79 467 . 06 98 . 29 -802 . 32 -516 . 68 24 53

1200 141 .48 479 . 29 112 . 36 -800 . 98 -490 . 70 21 36

1300 142 . 88 490 . 67 126 . 59 -799 . 60 -464 . 84 18 68

1400 144 . 05 501

31 140 . 95 -798 .22 -439 . 07 16 38

1500 145 . 05 511

28 155 . 42 -796 . 80 -413 . 38 14 39

1600 145 . 90 520 . 67 169 . 98 -795 . 42 -387 . 77 12 66

1700 146 . 64 529 . 54 184 . 62 -794 . 00 -362 . 29 11 13

1800 147 . 28 537 . 93 199 . 33 -792 . 53 -336 . 90 9 78

1900 147 . 85 545 . 91 214 . 10 -791 . 11 -311 . 62 8 57

2000 148 . 34 553 . 51 228 . 92 -789 .69 -286 . 44 7 48

2100 148 . 78 560 . 76 243 . 78 -788 . 22 -26133 6 50

2200 149 . 17 567 . 69 258 . 69 -786 . 76 -236 . 35 5 61

2300 149 . 51 574 . 33 273 . 63 -785 .29 -21146 4 80

2400 149 . 82 580 . 81 288 . 61 -783 . 83 -186 . 94 4 07

2500 150 . 10 587 . 06 303 . 62 -782 .37 -162 . 55 3 40

2600 150 .36 593 . 07 318 . 65 -780 . 90 -138 . 24 2 78

2700 150 . 58 598 . 86 333 . 70 -779 .40 -114 . 01 2 21

2800 150 . 79 604 . 46 348 . 78 -777 . 93 -89 . 87 1 68

2900 150 . 98 609 . 87 363 . 88 -776 .42 -65 . 81 1 19

3000 151 . 15 615 . 10 378 . 99 -774 . 96 -41 . 80 0 73
588 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CH3-CH2F

AH?(0 K) = -256.98 kJ mol'^

AH|(298.15 K) = -272.35 kJ mol"^

S°(298.15 K) = 265.04 J K"^ mol"^

Moments of Inertia (10"*^ kg m^)

22.84 88.06 100.54

Vibrational Frequencies (cm"^)

243.8 392.5 783.4 866.8 1043.2

1106.9 1168.9 1269.4 1381.2 1417.4

1452.5 1468.9 1502.8 2868.8 2895.9

2924.5 2929.2 2947.8

Bond Distances (lO"-'-" m)

C1-C2 1.512

C2-H1 1.084

C2-H2 1.084

C2-H3 1.086

C1-H4 1.083

C1-H5 1.083

Cl-Fl 1.373

Bond Angles (°) C1-C2-H1 110.47

C1-C2-H2 110.48

C1-C2-H3 110.34

C2-C1-H4 111.55

C2-C1-H5 111.56

C2-C1-F1 109.54

Dihedral Angles {°)

HI -C2 -H2 119 95

HI -C2 -H3 239 98

HI -CI -H4 59 17

HI -CI -H5 180 93

HI -CI -Fl 300 06

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p = 1 atmosphere

I- J K"^ mol"^ -1 ] kJ mol'^

T/K C

0 S° H°-H°(T^) AH0

f AG? Log Kf

150 42 . I2 231

.

14 -7 . 39 -264 .85 -244 . 05 84 98

200 47 . 14 244 . 05 -5 . 17 -267 . 44 -236 . 77 61 83

250 52 . 92 255 . 19 -2 . 69 -269 . 99 -228 . 86 47 81

300 59 . 58 265 . 41 0 . 11 -272 .42 -220 . 45 38 38

350 66 . 68 275 . 10 3 . 25 -274 . 72 -211

.

67 31 59

400 73 . 79 284 . 45 6 . 75 -276 . 77 -202 . 55 26 45

450 80 . 60 293 . 52 10 . 61 -278 .65 -193 . 22 22 43

500 86 . 99 302 . 33 14 . 79 -280 .33 -183 . 68 19 19

600 98 . 41 319 . 20 24 . 09 -283 . 13 -164 . 10 14 29

700 108 . 16 335 . 10 34 . 44 -285 .27 -144 . 10 10 75

800 116 . 55 350 . 09 45 . 71 -286 . 86 -123 . 76 8 08

900 123 . 78 364 . 24 57 . 77 -287 . 98 -103 . 26 5 99

1000 130 . 03 377 . 61 70 . 50 -288 . 70 -82 . 59 4 31

1100 135 . 44 390 . 25 83 . 81 -289 . 11 -61

.

80 2 93

1200 140 . 12 402 . 24 97 . 62 -289 . 24 -40 . 96 1 78

1300 144 . 18 413 . 62 Ill

.

87 -289 . 16 -20 . 13 0 81

1400 147 . 70 424 . 43 126 . 49 -288 . 86 0 . 71 -0 03

1500 150 . 75 434 . 83 141 . 44 -288 . 44 21

.

38 -0 74

1600 153 . 42 444 . 84 156 . 68 -287 . 90 41

.

84 -1 37

1700 155 . 75 454 . 40 172 . 16 -287 .23 62 . 22 -1 91

1800 157 . 80 463 . 55 187 . 87 -286 . 52 82 . 51 -2 39

1900 159 . 59 472 . 30 203 . 76 -285 .68 102 . 68 -2 82

2000 161 . 19 480 . 70 219 . 81 -284 .80 122 . 72 -3 20

2100 162 . 60 488 . 77 236 . 02 -283 . 88 142 . 63 -3 55

2200 163 . 85 496 . 53 252 . 36 -282 . 92 162 . 42 -3 86

2300 164 . 98 504 . 00 268 . 82 -281 . 92 182 . 09 -4 14

2400 165 . 98 511 . 19 285 . 38 -280 . 83 201 . 67 -4 39

2500 166 . 88 518 . 13 302 . 04 -279 . 74 221 . 08 -4 62

2600 167 . 69 524 . 84 318 . 78 -278 .65 240 . 41 -4 83

2700 168 . 43 531

.

32 335 . 60 -277 . 52 259 . 66 -5 02

2800 169 . 09 537 . 60 352 . 49 -276 .35 278 . 86 -5 20

2900 169 . 70 543 . 67 369 . 44 -275 . 18 297 . 98 -5 37

3000 170 . 25 549 . 56 386 . 45 -273 . 97 317 . 06 -5 52
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 589

CH3-CH2CI

AH2(0 K) = -98.58 kJ mol"^

AHf(2 98.15 K) = -113.66 kJ mol"^

S°(298.15 K) = 276.16 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

26.35 153.85 169.86

Vibrational Frequencies (cm"-'^) 248.4 319.4 641.6 761.9 944.9

1055.8 1062.6 1247.5 1305.9 1398.2

1455.6 1463.7 1468.7 2869.1 2925.6

2938.8 2945.4 2990.0

Bond Distances (lO'-'-" m)

C1-C2 1.517

C2-H1 1.083

C2-H2 1.083

C2-H3 1.086

C1-H4 1.079

C1-H5 1.079

Cl-Cll 1.799

Bond Angles (°) C1-C2-H1 111.05

C1-C2-H2 111.05

C1-C2-H3 109.22

C2-C1-H4 111.79

C2-C1-H5 111.80

C2-C1-C11 111.49

Dihedral Angles (°)

HI -C2 -H2 120 85

HI -C2 -H3 240 43

HI -CI -H4 58 34

HI -CI -H5 180 81

HI -CI -Cll 299 58

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p = 1 atmosphere

!- J K'^ mol"^ - I I kJ mol"^

T/K r<0 S° H°-H°(T^) AH0

f AGf Log Kf

150 43 :i7 240 . 51 -7 . 80 -106 .40 -85 . 65 29 82

200 49 . 58 253 . 94 -5 . 48 -108 . 95 -78 . 37 20 47

250 56 . 04 265 . 72 -2 . 85 -111 . 42 -70 . 50 14 73

300 63 . 15 276 . 55 0 . 12 -113 . 76 -62 . 13 10 82

350 70 . 51 286 . 83 3 . 44 -115 . 90 -53 . 43 7 97

400 77 . 71 296 . 70 7 . 14 -117 . 86 -44 . 43 5 80

450 84 .49 306 . 23 11

.

18 -119 . 58 -35 . 19 4 09

500 90 . 76 315 . 45 15 . 56 -121 . 13 -25 . 77 2 69

600 101 . 79 332 . 98 25 . 20 -123 .68 -6 . 53 0 57

700 111 . 11 349 . 37 35 . 87 -125 . 56 13 . 14 -0 98

800 119 . 06 364 . 72 47 . 40 -126 . 94 33 . 10 -2 16

900 125 . 92 379 . 14 59 . 68 -127 . 86 53 . 22 -3 09

1000 131 .85 392 . 72 72 . 60 -128 .45 73 . 47 -3 84

1100 137 . 00 405 . 53 86 . 08 -128 . 70 93 . 81 -4 45

1200 141 .46 417 . 64 100 . 04 -128 . 74 114 . 18 -4 97

1300 145 . 33 429 . 12 114 . 41 -128 . 53 134 . 60 -5 41

1400 148 . 70 440 . 01 129 . 14 -128 . 16 154 . 98 -5 78

1500 151 .63 450 . 37 144 . 19 -127 .65 175 . 35 -6 11

1600 154 . 19 460 . 24 159 . 51 -127 . 07 195 . 64 -6 39

1700 156 .44 469 . 79 175 . 06 -126 . 36 215 . 69 -6 63

1800 158 .41 478 . 96 190 . 83 -125 . 56 235 . 52 -6 83

1900 160 . 15 487 . 74 206 . 78 -124 .68 255 . 27 -7 02

2000 161 .69 496 . 16 222 . 89 -123 . 76 274 . 89 -7 18

2100 163 . 05 504 . 24 239 . 15 -122 . 80 294 . 39 -7 32

2200 164 .26 512 . 01 255 . 53 -121 . 80 313 . 76 -7 45

2300 165 .35 519 . 49 272 . 03 -120 . 75 333 . 00 -7 56

2400 166 . 32 526 . 69 288 . 63 -119 . 66 352 . 13 -7 66

2500 167 .20 533 . 65 305 . 32 -118 . 53 371

.

12 -7 75

2600 167 . 99 540 . 36 322 . 09 -117 .40 390 . 03 -7 84

2700 168 . 70 546 . 85 338 . 94 -116 .23 408 . 82 -7 91

2800 169 . 35 553 . 13 355 . 86 -115 . 06 427 . 56 -7 98

2900 169 . 93 559 . 21 372 . 83 -113 . 85 446 . 27 -8 04

3000 170 .47 565 . 11 389 . 87 -112 .63 464 . 93 -8 09
590 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHjF-CHj.

AH°(0 K)

AH°(298.15 K)

S°{298.15 K)

-46 . 02 kJ mol"^

-56 . 22 kJ mol"^

275.18 J K"^ mol"

Moments of Inertia (10

20.06 82.42 96.41

kg

Vibrational Frequencies (cm" ) 158.7 382.9 439.8 831.2 956.3

1060.3 1100.4 1227.7 1388.3 1426.0

1486.1 2855.7 2906.9 2973.9 3070.7

Bond Distances (lO'-"-" C1-C2 1.489

Cl-Fl 1.371

Cl-Hl 1.084

C2-H2 1.074

C2-H3 1.074

C1-H4 1.087

m) Bond Angles (°) C2-C1-F1

C2-C1-H1

C1-C2-H2

C1-C2-H3

C2-C1-H4

110 . 04

111 . 36

119 .68

119 . 22

111 . 93

Dihedral Angles (°)

Fl-Cl-Hl

F1-C2-H2

F1-C2-H3

F1-C1-H4

240 . 16

163 . 95

-32.85

119 . 10

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

j - J K"^ mol"^ -

| I kJ mol"^

T/K %S° H°-H°(T^) AHo

f AG? Log Kf

150 45 238 . 96 -7 . 95 -51 . 42 -40 . 04 13 . 94

200 50 . 91 252 . 83 -5 . 53 -53 . 09 -36 . 02 9 .41

250 56 .40 264 . 77 -2 . 85 -54 . 73 -31

.

59 6 . 60

300 62 .27 275 . 57 0 . 11 -56 .27 -26 . 86 4 ,.68

350 68 . 26 285 . 62 3 . 38 -57 . 70 -21

.

88 3 .27

400 74 . 12 295 . 11 6 . 94 -58 . 99 -16 . 74 2 ,. 19

450 79 .64 304 . 17 10 . 79 -60 . 17 -11

.

42 1

,

. 33

500 84 . 77 312 . 83 14 . 90 -61 . 17 -6 . 02 0 ,. 63

600 93 . 85 329 . 11 23 . 84 -62 . 84 5 . 10 -0 ..45

700 101 . 57 344 . 17 33 . 62 -64 . 14 16 . 53 -1

.

.23

800 108 .19 358 . 18 44 . 12 -65 . 02 28 . 12 -1

,

. 83

900 113 . 90 371

.

26 55 . 23 -65 . 61 39. 83 -2 .. 31

1000 118 .85 383 . 52 66 . 87 -65 . 94 51

.

63 -2 .. 70

1100 123 . 14 395 . 05 78 . 98 -66 . 02 63 . 47 -3 .. 01

1200 126 . 86 405 . 93 91

.

48 -65 . 94 75 . 35 -3 .,28

1300 130 . 09 416 . 22 104 . 33 -65 . 73 87 . 28 -3 ., 51

1400 132 . 90 425 . 96 117 . 49 -65 .35 99 . 16 -3 .. 70

1500 135 .35 435 . 22 130 . 90 -64 .89 Ill

.

00 -3 ..87

1600 137 .49 444 . 02 144 . 54 -64 . 35 122 . 84 -4 .. 01

1700 139 . 36 452 . 42 158 . 39 -63 . 72 134 . 60 -4 .. 14

1800 141 . 00 460 . 43 172 . 41 -63 . 05 146 . 27 -4 ,. 24

1900 142 .45 468 . 09 186 . 59 -62 .30 157 . 90 -4 .. 34

2000 143 . 73 475 . 43 200 . 89 -61 . 55 169 . 41 -4 .. 42

2100 144 .87 482 . 47 215 . 33 -60 . 71 180 . 83 -4 .. 50

2200 145 . 88 489 . 24 229 . 86 -59 . 83 192 . 17 -4 .. 56

2300 146 . 79 495 . 74 244 . 50 -58 . 95 203 . 43 -4 ,.62

2400 147 .60 502 . 00 259 . 22 -58 . 03 214 . 56 -4 ..67

2500 148 . 33 508 . 05 274 . 02 -57 . 11 225 . 60 -4 ,. 71

2600 148 . 99 513 . 88 288 . 88 -56 . 15 236 . 61 -4 ., 75

2700 149 . 58 519 . 51 303 . 81 -55 . 15 247 . 53 -4 .. 79

2800 150 . 12 524 . 96 318 . 80 -54 . 18 258 . 36 -4 ., 82

2900 150 . 61 530 . 24 333 . 84 -53 . 14 269 . 20 -4 ., 85

3000 151 . 06 535 . 35 348 . 92 -52 . 13 280 . 04 -4 ., 88
SECTIONS. FLAME INHIBmON CHEMISTRY AND THE SEARCH 591

CH2=CHF

AH|(0 K) = -131.00 kJ mol"^

AHf(298.15 K) = -139.21 kJ mol"^

S°(298.15 K) = 252.27 J K"^ mol"^

Moments of Inertia (10'^^ kg m^)

12.50 77.57 90.08

Vibrational Frecpiencies (cm"-'-) 467.3 714.3 915.5

1148.6 1300.9 1393.4

3060.3 3086.1

Bond Distances (lO"-"-" m)

C1-C2 1.309

Cl-Fl 1.329

Cl-Hl 1.072

C2-H2 1.073

C2-H3 1.074

922.0 976.7

1690.0 3003.1

Bond Angles (°) Dihedral Angles (°) C2-C1-F1 122.38

C2-C1-H1 125.70

C1-C2-H2 119.75

C1-C2-H3 121.50

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ - I I kJ mol'

T/K Co S° H°-H AHo

f AG? Log Kf

150 35 . Is 223 . 78 -6 . 26 -134 . 89 -129 . 45 45 ., 08

200 39 . 66 234 . 58 -4 . 38 -136 . 36 -127 . 49 33 ..29

250 44 . 65 243 . 95 -2 . 28 -137 . 86 -125 . 10 26 .. 14

300 50 . 26 252 . 58 0 . 09 -139 . 24 -122 . 47 21 ,, 32

350 55 . 98 260 . 76 2 . 75 -140 . 54 -119 . 62 17 ,.85

400 61. 47 268 . 60 5 . 69 -141 . 75 -116 . 61 15 ..22

450 66. 57 276 . 14 8 . 89 -142 .80 -113 . 43 13 .. 17

500 7123 283 . 39 12 . 34 -143 . 76 -110 . 16 11 .. 51

600 79 . 31 297 . 12 19 . 88 -145 . 44 -103 . 34 9 .. 00

700 86 . 01 309 . 87 28 . 15 -146 . 73 -96 . 27 7 .. 18

800 9165 321

73 37 . 05 -147 . 78 -88 . 99 5 ,,81

900 96 . 47 332 . 81 46 . 45 -148 . 57 -8155 4 ,, 73

1000 100 . 60 343 . 19 56 . 32 -149 .20 -74 . 01 3 .. 87

1100 104 . 17 352 . 95 66 . 56 -149 . 66 -66 . 36 3 .. 15

1200 107 . 26 362 . 15 77 . 13 -150 . 00 -58 . 66 2 .. 55

1300 109 . 93 370 . 84 88 . 00 -150 .21 -50 . 88 2 .. 05

1400 112 . 25 379 . 08 99 . 11 -150 .33 -43 . 14 1 .. 61

1500 114 . 27 386 . 89 110 . 44 -150 .37 -35 . 31 1 .. 23

1600 116 . 04 394 . 33 12196 -150 .37 -27 . 57 0 .. 90

1700 117 . 58 401 . 41 133 . 64 -150 .33 -19 . 79 0 ,. 61

1800 118 . 94 408 . 17 145 . 47 -150 . 21 -12 . 09 0 ,.35

1900 120 . 14 414 . 63 157 . 42 -150 . 04 -4 . 39 0 ,. 12

2000 121 . 20 420 . 82 169 . 49 -149 . 83 3 . 18 -0 ,. 08

2100 122 . 14 426 . 76 18166 -149 . 62 10 . 75 -0 ,.27

2200 122 . 98 432 . 46 193 . 91 -149 .37 18 . 24 -0 ,.43

2300 123 . 73 437 . 95 206 . 25 -149 . 08 25 . 65 -0 ,. 58

2400 124 . 41 443 . 23 218 . 66 -148 . 74 32 . 97 -0 ,. 72

2500 125 . 01 448 . 32 23113 -148 .45 40 . 25 -0 ,. 84

2600 125 . 56 453 . 23 243 . 66 -148 . 07 47 . 49 -0 ,. 95

2700 126 . 05 457 . 98 256 . 24 -147 . 74 54 . 68 -1 . 06

2800 126 . 50 462 . 57 268 . 87 -147 .36 61

84 -1

. 15

2900 126 . 90 467 . 02 28154 -146 . 94 68 . 99 -1. 24

3000 127 . 28 471 . 33 294 . 25 -146 . 52 76 . 19 -1 . 33
592 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CH2=CF2

AH2(0 K) = -332.84 kJ mol"^

AHf(298.15 K) = -340.11 kJ mol"^

S°(298.15 K) = 270.73 J K"^ mol'^

Moments of Inertia (10"*^ kg m^)

74.36 78.66 153.02

Vibrational Frequencies (cm"-'-) 423.0 528.3 629.6 712.8 864.8

918.6 951.6 1332.4 1386.2 1747.8

3026.7 3113.8

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.305

Cl-Fl 1.303 C2-C1-F1 125.19

C1-F2 1.303 C2-C1-F2 125.27

C2-H1 1.070 C1-C2-H1 119.93

C2-H2 1.071 C1-C2-H2 119.82

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I i kJ mol"^

T/K C

o S° H°-H°(T^) AHo

f AG? Log Kf

150 38 . l3 238 . 25 -7 . 18 -336 . 64 -327 . 98 114 21

200 44 . 95 250 . 18 -5 . 10 -337 . 94 -324 . 93 84 86

250 52 . 09 260 . 98 -2 . 67 -339 . 11 -321 . 58 67 19

300 59 . 09 271

.

10 0 . 11 -340 . 16 -318 . 03 55 37

350 65 . 58 280 . 70 3 . 23 -341 . 04 -314 . 30 46 91

400 71

.

40 289 . 85 6 . 66 -341 . 83 -310 . 49 40 54

450 76 . 58 298 . 57 10 . 36 -342 . 50 -306 . 56 35 58

500 81

.

17 306 . 88 14 . 31 -343 . 09 -302 . 59 31 61

600 88 . 83 322 . 38 22 . 82 -344 . 01 -294 . 47 25 64

700 94 . 94 336 . 55 32 . 02 -344 .68 -286 . 23 21 36

800 99 . 93 349 . 56 41 . 77 -345 . 18 -277 . 82 18 14

900 104 . 06 361

.

58 51 . 97 -345 . 51 -269 . 37 15 63

1000 107 . 54 372 . 73 62 . 56 -345 . 72 -260 . 83 13 62

1100 110 . 49 383 . 12 73 . 47 -345 .85 -252 . 21 11 98

1200 113 . 00 392 . 84 84 . 64 -345 . 89 -243 . 59 10 60

1300 115 . 15 401

.

97 96 . 05 -345 . 85 -234 . 93 9 44

1400 117 . 00 410 . 58 107 . 66 -345 . 77 -226 . 27 8 44

1500 118 . 60 418 . 71 119 . 44 -345 . 64 -217 . 61 7 58

1600 119 . 99 426 . 40 131 . 38 -345 .47 -208 . 99 6 82

1700 121

.

20 433 . 72 143 . 44 -345 . 26 -200 . 37 6 16

1800 122 . 26 440 . 68 155 . 61 -345 . 01 -191

.

84 5 57

1900 123 . 19 447 . 31 167 . 88 -344 . 76 -183 . 30 5 04

2000 124 . 01 453 . 65 180 . 25 -344 .47 -174 . 89 4 57

2100 124 . 74 459 . 72 192 . 69 -344 . 13 -166 . 48 4 14

2200 125 . 38 465 . 54 205 . 19 -343 . 80 -158 . 20 3 76

2300 125 . 96 471 . 12 217 . 76 -343 .46 -149 . 95 3 41

2400 126 . 48 476 . 49 230 . 38 -343 . 09 -141 . 80 3 09

2500 126 . 94 481

.

67 243 . 05 -342 . 71 -133 . 72 2 79

2600 127 . 36 486 . 65 255 . 77 -342 .33 -125 . 65 2 52

2700 127 . 73 491 . 47 268 . 52 -341 . 92 -117 . 65 2 28

2800 128 . 08 496 . 12 281 . 31 -341 . 50 -109 . 70 2 05

2900 128 . 39 500 . 62 294 . 14 -341 . 08 -101 . 75 1 83

3000 128 . 67 504 . 98 306 . 99 -340 . 62 -93 . 76 1 63
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 593

CHF=CHF fccf cis

AH?(0 K)

AH^{2 98.15 K)

S°(298.15 K)

Moments of Inertia

38.51 139.95

= -294 . 43 kJ mol'^

= -301.29 kJ mol"^

= 274 .29 J K"^ mol

(10"^"^ kg m^)

178 .47

Vibrational Frequencies (cm

225.8 491.8 747.9

1004.3 1110.8 1254.7

3070.6 3094.3

-1^ 803 . 9 1380 .

3

905 . 7 1759 .

7

Bond Distances (10'^°

C1-C2 1.307

Cl-Fl 1.324

Cl-Hl 1.070

C2-H2 1.070

C2-F2 1.324

m) Bond Angles

C2-C1-F1

C2-C1-H1

C1-C2-H2

C1-C2-F2

122 . 58

123 . 07

123 . 07

122 . 58

Dihedral Angles (°)

F1-C2-F2 0 . 00

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

!- J K'^ mol"^ - I I kJ mol'^

T/K S° H°-H°(T^) AHo

f AG? Log Kf

150 41 .39 241

.

12 -7 . 29 -297 . 94 -289 . 70 100 88

200 46 .30 253 . 68 -5 . 10 -299 . 11 -286 . 81 74 91

250 52 . 07 264 . 62 -2 . 65 -300 .29 -283 . 63 59 26

300 58 . 19 274 . 65 0 . 11 -301 . 33 -280 . 29 48 80

350 64 . 22 284 . 08 3 . 17 -302 .29 -276 . 73 41 30

400 69 . 88 293 . 03 6 . 52 -303 . 13 -273 . 09 35 66

450 75 . 04 301

.

56 10 . 15 -303 . 88 -269 . 32 31 26

500 79 . 67 309 . 71 14 . 02 -304 . 55 -265. 47 27 73

600 87 . 53 324 . 96 22 . 39 -305 .60 -257 . 65 22 43

700 93 . 84 338 . 94 31

.

47 -306 .39 -249 . 62 18 63

800 99 . 00 351

.

82 41

.

12 -306 . 98 -241 . 46 15 77

900 103 . 28 363 . 74 51

.

24 -307 .40 -233 . 22 13 53

1000 106 . 87 374 . 81 61 . 76 -307 .69 -224 . 89 11 75

1100 109 . 91 385 . 14 72 . 60 -307 . 90 -216 . 48 10 28

1200 112 .49 394 . 81 83 . 72 -307 . 98 -208 . 07 9 06

1300 114 . 70 403 . 91 95 . 09 -307 . 98 -199. 62 8 02

1400 116 .60 412 . 48 106 . 65 -307 . 94 -191

.

13 7 13

1500 118 .25 420 . 58 118 . 39 -307 . 86 -182 . 67 6 36

1600 119 . 67 428 . 26 130 . 29 -307 . 73 -174 . 22 5 69

1700 120 . 91 435 . 55 142 . 33 -307 . 52 -165 . 81 5 09

1800 122 . 00 442 . 50 154 . 47 -307 . 31 -157 . 40 4 57

1900 122 . 96 449 . 12 166 . 72 -307 . 11 -149 . 12 4 10

2000 123 .80 455 . 45 179 . 06 -306 . 81 -140 . 83 3 68

2100 124 . 54 461

.

51 191

.

48 -306 . 52 -132 . 63 3 30

2200 125 .20 467 . 32 203 . 97 -306 .23 -124 . 52 2 96

2300 125 . 79 472 . 90 216 . 51 -305 .89 -116 . 48 2 64

2400 126 . 32 478 . 26 229 . 12 -305 . 56 -108 . 49 2 36

2500 126 . 80 483 . 43 241

.

78 -305 . 18 -100 . 54 2 10

2600 127 .22 488 . 41 254 . 48 -304 . 80 -92 . 68 1 86

2700 127 . 61 493 . 22 267 . 22 -304 .39 -84 . 85 1 64

2800 127 . 96 497 . 87 280 . 00 -303 . 97 -77 . 07 1 44

2900 128 .28 502 . 36 292 . 81 -303 . 55 -69 . 29 1 25

3000 128 . 57 506 . 72 305 . 65 -303 . 13 -61

.

50 1 07
594 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHF=CHF fccf trans

AH?(0 K)

AHf(298.15 K)

S°(298.15 K)

= -295 . 64 kJ mol"^

= -302 . 24 kJ mol"^

= 273 . 56 J K'^ mol -1

Moments of Inertia

14.20 204.16

(10-^"^ kg m^)

218 .35

Vibrational Frequencies (cm'

308.1 337.6 538.5

1128.5 1144.3 1269.3

3062.2 3071.4

846 . 3 1269 . 3 928 . 0 1747 .

6

Bond Distances (10

C1-C2 1.306

Cl-Fl 1.329

Cl-Hl 1.070

C2-H2 1.070

C2-F2 1.329

-10 m) Bond Angles (°) C2-C1-F1

C2-C1-H1

C1-C2-H2

C1-C2-F2

120 . 24

125 . 28

125 . 28

120 . 25

Dihedral Angles (°)

F1-C2-F2 180 . 00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

1

1 mol" 1 _ 1

1

1

1 kJ mol"^ 1

1 T/K S'

0 H°-H°(T^) AHo

f AG° Log Kf

150 42 .^6 239 . 06 -7 . 58 -299 . 16 -290 . 66 101 ,. 21

200 48 .41 252 . 14 -5 . 30 -300 .29 -287 . 65 75 ,. 13

250 54 . 12 263 . 55 -2 . 74 -301 . 33 -284 . 43 59 ,. 42

300 59 . 93 273 . 93 0 . 11 -302 .29 -281

.

00 48 .. 92

350 65 .63 283 .60 3 . 25 -303 . 17 -277 . 44 41 ,.40

400 71 . 01 292 . 72 6 . 67 -303 . 93 -273 . 76 35 .. 75

450 75 . 96 301 . 37 10 . 34 -304 . 64 -269 . 99 31

,

, 34

500 80 .45 309 .61 14 . 25 -305 . 26 -266 . 14 27 .. 80

600 88 . 10 324 . 98 22 . 70 -306 . 23 -258 . 28 22 ..49

700 94 .30 339 . 04 31

.

83 -306 . 98 -250 . 29 18 ..68

800 99 .39 351 . 98 41

.

52 -307 . 52 -242 . 13 15 ., 81

900 103 .61 363 . 93 51

.

68 -307 . 94 -233 . 93 13 ,. 58

1000 107 . 16 375 . 04 62 . 22 -308 . 19 -225 . 60 11

.

. 78

1100 110 . 16 385 .40 73 . 09 -308 . 36 -217 . 23 10 .. 32

1200 112 . 72 395 . 10 84 . 24 -308 .40 -208 . 82 9 ., 09

1300 114 . 91 404 .21 95 . 62 -308 .40 -200 . 41 8 ., 05

1400 116 . 79 412 . 79 107 . 21 -308 . 36 -191

.

96 7 .. 16

1500 118 .42 420 . 91 118 . 97 -308 . 24 -183 . 51 6 ..39

1600 119 . 83 428 .60 130 . 88 -308 . 07 -175 . 10 5 ., 72

1700 121 . 06 435 . 90 142 . 93 -307 . 90 -166 . 73 5 ., 12

1800 122 . 13 442 . 85 155 . 09 -307 . 65 -158 . 36 4 .,60

1900 123 . 08 449 .48 167 . 35 -307 .40 -150 . 08 4 .. 13

2000 123 . 91 455 . 81 179 . 70 -307 . 15 -141

.

88 3 ,. 71

2100 124 .65 461 . 88 192 . 13 -306 . 81 -133 . 72 3 .. 33

2200 125 .30 467 . 69 204 . 63 -306 . 52 -125 . 65 2 ., 98

2300 125 . 88 473 .27 217 . 19 -306 . 19 -117 . 61 2 ,.67

2400 126 .41 478 . 64 229 . 80 -305 . 81 -109 . 66 2 ,.39

2500 126 . 88 483 . 81 242 . 47 -305 .43 -101 . 80 2 .. 13

2600 127 . 30 488 . 80 255 . 18 -305 . 06 -93 . 97 1 ..89

2700 127 .68 493 .61 267 . 93 -304 . 64 -86 . 15 1

,

.67

2800 128 . 02 498 .26 280 . 71 -304 . 22 -78 . 41 1 ,.46

2900 128 . 34 502 . 76 293 . 53 -303 . 80 -70 . 67 1 ..27

3000 128 . 62 507 . 11 306 . 38 -303 .38 -62 . 93 1.. 10
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 595

CHF=CF2

AH2(0 K) = -480.24 kJ mol"^

AH^{298.15 K) -• = -485.56 kJ mol"^

S°(298.15 K) = 292.39 J K"^ mol"^

Moments of Inertia (10"'*^ kg m^)

76.88 212.39 289.27

Vibrational Frequencies (cm"-"-) 223.0 306.8 468.9 585.8 605.0

815.9 921.9 1144.8 1265.0 1368.2

1824.4 3105.2

Bond Distances (10"^° m) Bond Angles {°) Dihedral Angles (°) C1-C2 1.304

Cl-Fl 1.298 C2-C1-F1 125.42

C1-F2 1.303 C2-C1-F2 122.87

C2-H1 1.068 C1-C2-H1 123.39

C2-F3 1.327 C1-C2-F3 120.55

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

1

1 J K mol 1 11

1

1 kJ mol"^ 1

1 T/K s

0 ttO tt H -HO / m \ (Tj.) AHo

f AG? Log Kf

150 47 .30 253 . 10 -8 . 66 -483 . 50 -471 . 66 164 24

200 55 . 00 267 . 77 -6 . 10 -484 . 34 -467 . 60 122 12

250 62 . 34 280 . 84 -3 . 16 -485 . 01 -463 . 42 96 82

300 69 . 13 292 . 82 0 . 13 -485 .60 -459 . 07 79 93

350 75 .30 303 . 94 3 . 74 -486 . 06 -454 . 68 67 85

400 80 . 82 314 . 36 7 . 65 -486 .43 -450 . 20 58 79

450 85 . 72 324 . 17 11 . 81 -486 . 72 -445 . 72 51 73

500 90 . 05 333 .44 16 . 21 -486 . 98 -441 . 16 46 09

600 97 .23 350 . 51 25 . 59 -487 .27 -432 . 04 37 61

700 102 . 85 365 . 94 35 . 60 -487 . 44 -422 . 88 31 55

800 107 . 31 379 . 98 46 . 12 -487 .48 -413 . 67 27 01

900 110 . 90 392 . 83 57 . 04 -487 .44 -404 . 38 23 47

1000 113 . 83 404 . 67 68 . 27 -487 .31 -395 . 10 20 64

1100 116 . 24 415 . 63 79 . 78 -487 . 14 -385 . 81 18 32

1200 118 . 26 425 . 84 91

51 -486 . 93 -376 . 48 16 39

1300 119 . 95 435 .37 103 . 42 -486 . 72 -367 . 15 14 75

1400 121 .39 444 . 32 115 . 50 -486 .43 -357 . 82 13 35

1500 122 . 61 452 . 73 127 . 70 -486 . 18 -348 . 53 12 14

1600 123 . 66 460 . 68 140 . 01 -485 . 85 -339 . 24 11 07

1700 124 . 56 468 .21 152 . 42 -485 . 55 -330 . 03 10 14

1800 125 .35 475 . 35 164 . 92 -485 . 22 -320 . 83 9 31

1900 126 . 03 482 . 15 177 . 49 -484 . 84 -311

71 8 57

2000 126 .63 488 . 62 190 . 12 -484 . 51 -302 . 67 7 90

2100 127 . 16 494 . 82 202 . 81 -484 . 13 -293 . 67 7 30

2200 127 . 63 500 . 75 215 . 55 -483 . 71 -284 . 76 6 76

2300 128 . 05 506 .43 228 . 34 -483 . 34 -275 . 93 6 27

2400 128 .42 511 . 88 241 . 16 -482 . 92 -267 . 15 5 81

2500 128 . 75 517 . 13 254 . 02 -482 . 50 -258 . 45 5 40

2600 129 . 05 522 .19 266 . 91 -482 . 08 -249 . 83 5 02

2700 129 . 32 527 . 07 279 . 83 -481 . 62 -241

21 4 67

2800 129 . 56 531 . 77 292 . 77 -481 . 16 -232 . 67 4 34

2900 129 . 78 536 . 32 305 . 74 -480 . 70 -224 . 10 4 04

3000 129 . 98 540 . 73 318 . 73 -480 . 24 -215 . 52 3 75
596 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF2=CF2

AH?(0 K) = -650.28 kJ mol"^

AHf(298.15 K) = -653.67 kJ mol"^

S°(298.15 K) = 311.22 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

148.83 252.80 401.64

Vibrational Frequencies (cm"^)

198.9 199.9 382.0 431.3 530.1

540.3 563.3 777.2 1163.6 1341.2

1355.2 1915.3

Bond Distances (lO"-*^" m) Bond Angles i°) Dihedral Angles (°) C1-C2 1.300

Cl-Fl 1.301 C2-C1-F1 123.44

C1-F2 1.302 C2-C1-F2 123.45

C2-F3 1.301 C1-C2-F3 123.45

C2-F4 1.301 C1-C2-F4 123.43

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K C

lO S° H°-H°(T^) AHo

f AG° Log Kf

150 54 . ^9 265 . 02 -10 . 18 -653 . 08 -637 . 64 222 .. 03

200 64 . 99 282 . 25 -7 . 17 -653 .46 -632 . 41 165., 17

250 73 . 43 297 . 68 -3 . 70 -653 . 62 -627 . 18 131.. 04

300 80 . 51 311

.

72 0 . 15 -653 .67 -621

.

95 108 ,, 28

350 86 . 53 324 . 59 4 . 33 -653 . 62 -616 . 72 92 .. 03

400 91

.

74 336 . 4 9 8 . 79 -653 . 54 -611

.

49 79 ,. 85

450 96 . 26 347 . 56 13 . 49 -653 .42 -606 . 30 70 .,37

500 100 . 21 357 . 92 18 . 41 -653 .25 -601

.

12 62 .. 79

600 106 . 65 376 . 78 28 . 76 -652 . 83 -590 . 78 51 .,43

700 Ill

.

57 393 . 61 39 . 69 -652 . 37 -580 . 53 43 ., 32

800 115 . 36 408 . 77 51

.

04 -651 . 87 -570 . 28 37 ,,23

900 118 . 29 422 . 53 62 . 73 -651 .41 -560 . 07 32 ., 50

1000 120 . 60 435 . 12 74 . 68 -650 . 90 -549 . 90 28 ., 72

1100 122 . 43 446 . 70 86 . 83 -650 .40 -539 . 74 25 .,63

1200 123 . 90 457 . 42 99 . 15 -649 . 90 -529 . 57 23 ., 05

1300 125 . 10 467 . 39 Ill

.

60 -649 . 44 -519 . 44 20 .,87

1400 126 . 08 476 . 70 124 . 16 -648 . 98 -509 . 32 19 ., 00

1500 126 . 90 485 . 42 136 . 81 -648 . 52 -499 . 23 17 ., 38

1600 127 . 58 493 . 63 149 . 54 -648 . 06 -489 . 19 15 ., 97

1700 128 . 16 501 . 39 162 . 33 -647 .60 -479 . 19 14 ., 72

1800 128 . 65 508 . 73 175 . 17 -647 . 14 -469 . 24 13 ,, 62

1900 129 . 08 515 . 69 188 . 05 -646 .68 -459 . 36 12 ..63

2000 129 . 44 522 . 32 200 . 98 -646 . 22 -449 . 57 11

,

, 74

2100 129 . 76 528 . 65 213 . 94 -645 . 76 -439 . 86 10 ,, 94

2200 130 . 04 534 . 69 226 . 93 -645 .30 -430 . 20 10 .,21

2300 130 . 28 540 . 48 239 . 95 -644 . 84 -420 . 66 9 ., 55

2400 130 . 50 546 . 02 252 . 99 -644 .38 -411 . 16 8 ,, 95

2500 130 . 69 551 . 36 266 . 05 -643 . 88 -401 . 71 8 .,39

2600 130 . 86 556 . 48 279 . 13 -643 . 42 -392 . 38 7 ,,88

2700 131

.

01 561

.

43 292 . 22 -642 . 91 -383 . 05 7 .,41

2800 131

.

15 566 . 20 305 . 33 -642 .41 -373 . 76 6 ., 97

2900 131

.

28 570 . 80 318 . 45 -641 . 91 -364 . 47 6 ., 56

3000 131

.

39 575 . 25 331

.

58 -641 .41 -355 . 18 6 .. 18
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 597

CHF=CH« fcch trans

AH?{0 K)

AH|(298.15 K)

S°(298.15 K)

= 126 . 86 kJ mol"^

= 123 . 11 kJ mol"^

= 257 . 01 J K"^ mol'

Moments of Inertia (10 kg m^)

9 . 65 75 . 95 85 .60

Vibrational Frequencies (cm

481.1 643.7 756.2 809.8

1246.5 1474.5 3017.9 3103.5

1080 . 2

Bond Distances (lO"-"-" m)

C1-C2 1.322

Cl-Hl 1.074

Cl-Fl 1.330

C2-H2 1.068

Bond Angles (°) C2-C1-H1 125.44

C2-C1-F1 121.98

C1-C2-H2 132.85

Dihedral Angles (°)

Hl-Cl-Fl

F1-C2-H2

180 . 00

180 . 00

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K'^ mol"^ - I I kJ mol"^

T/K C° S° H°-H°(T^) AHo

f AG? Log Kf

150 36 .24 227 . 94 -6 . 38 125 . 19 121 . 71 -42 38

200 40 . 54 238 . 94 -4 . 47 124 .43 120 . 62 -31 51

250 45 . 61 248 . 53 -2 . 31 123 . 72 119 . 70 -25 01

300 50 . 76 257 . 32 0 . 09 123 . 09 118 . 95 -20 71

350 55 . 60 265 . 52 2 . 75 122 . 51 118 . 24 -17 64

400 59 . 99 273 . 23 5 . 64 122 . 01 117 . 61 -15 36

450 63 .89 280 . 52 8 . 74 121 . 55 117 . 03 -13 58

500 67 .35 287 . 42 12 . 02 121 . 17 116 . 52 -12 17

600 73 . 15 300 . 20 19 . 05 120 .46 115 . 60 -10 06

700 77 . 83 311 . 78 26 . 59 119 .87 114 . 81 -8 57

800 81 . 73 322 . 41 34 . 56 119 .37 114 . 14 -7 45

900 85 . 05 332 . 24 42 . 89 118 . 95 113 . 51 -6 59

1000 87 . 91 341 . 36 51

.

51 118 . 62 112 . 97 -5 90

1100 90 .38 349 . 86 60 . 41 118 .28 112 . 47 -5 34

1200 92 . 51 357 . 82 69 . 53 118 . 03 112 . 05 -4 88

1300 94 .36 365 . 29 78 . 86 117 . 78 Ill . 67 -4 49

1400 95 . 94 372 . 34 88 . 35 117 .61 Ill . 34 -4 15

1500 97 .30 379 . 00 97 . 99 117 .44 Ill . 00 -3 87

1600 98 .46 385 . 31 107 . 75 117 .28 110 . 67 -3 61

1700 99 .45 391

.

31 117 . 63 117 . 15 110 . 33 -3 39

1800 100 .29 397 . 01 127 . 61 117 . 07 109 . 96 -3 19

1900 101 . 00 402 . 45 137 . 66 116 . 94 109 . 58 -3 01

2000 101 . 60 407 . 64 147 . 78 116 . 86 109 . 12 -2 85

2100 102 . 11 412 . 61 157 . 96 116 . 78 108 . 66 -2 70

2200 102 . 53 417 . 36 168 . 19 116 .69 108 . 11 -2 57

2300 102 .89 421 . 92 178 . 46 116 .61 107 . 57 -2 44

2400 103 .20 426 . 30 188 . 77 116 .48 106 . 94 -2 33

2500 103 .45 430 . 52 199 . 10 116 .40 106 . 27 -2 22

2600 103 . 66 434 . 58 209 . 46 116 .32 105 . 56 -2 12

2700 103 . 84 438 . 49 219 . 85 116 . 23 104 . 85 -2 03

2800 103 . 99 442 . 27 230 . 25 116 . 11 104 . 10 -1 94

2900 104 . 11 445 . 92 240 . 67 115 . 98 103 . 39 -1 86

3000 104 .22 449 . 45 251

.

10 115 . 90 102 . 72 -1 . 79
598 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHF=CH» fcch cis

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

127 . 78 kJ mol"-^

= 124 . 11 kJ mol"^

257 . 95 J K" mol"

Moments of Inertia

12.29 71.42

10"*' kg m-^) 83 . 71

437 . 4 1252 .

8

629.4

1452 .

6

825 . 3 3062 .

1

853 .

1

3100 .

3

1053 .

2

Bond Distances (10 -10 C1-C2

Cl-Hl

Cl-Fl

C2-H2

1 . 321

1 . 071

1.336

1 . 068

Bond Angles (°) C2-C1-H1 125.39

C2-C1-F1 121.96

C1-C2-H2 135.69

Dihedral Angles (°)

Hl-Cl-Fl

F1-C2-H2

180 . 00

0 . 00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K"^ mol"^ -I I kJ mol"^

t/k 1C° S° H°-H AHo

f AG?

15 0 3 6 1

3

228 . 77 - 6 . 4 0 12 6 . 15 122 . 55 - 42 . 200 40 . 76 239 . 86 -4 . 47 125 . 44 121 . 42 -31 .. 71

250 45 .62 249 . 49 -2 . 31 124 . 73 120 . 50 -25 ,. 17

300 50 .62 258 . 26 0 . 09 124 . 10 119 . 62 -20 ,.83

350 55 .39 266 . 43 2 . 74 123 . 51 118 . 91 -17 .. 74

400 59 . 76 274 . 12 5 . 62 122 . 97 118 . 24 -15 ..44

450 63 . 66 281 . 39 8 . 71 122 . 51 117 . 61 -13 ,.65

500 67 . 13 288 . 27 11 . 97 122 . 13 117 . 03 -12 .. 23

600 72 . 94 301

.

02 18 . 98 121 . 38 116 . 02 -10 .. 10

700 77 . 62 312 . 58 26 . 51 120 . 79 115 . 19 -8 .. 59

800 81 . 51 323 . 16 34 . 46 120 .29 114 . 43 -7 ..47

900 84 . 82 332 . 95 42 . 76 119 . 83 113 . 72 -6 ,. 60

1000 87 . 66 342 . 05 51 . 38 119 .45 113 . 14 -5 ,. 91

1100 90 . 14 350 . 53 60 . 25 119 . 12 112 . 55 -5 ..35

1200 92 .28 358 . 46 69 . 35 118 .83 112 . 09 -4 ,. 88

1300 94 . 14 365 . 92 78 . 65 118 . 57 Ill

.

63 -4 ..49

1400 95 . 76 372 . 96 88 . 12 118 .37 Ill

.

21 -4 .. 15

1500 97 . 15 379 . 61 97 . 74 118 .20 110 . 83 -3 .. 86

1600 98 .35 385 . 92 107 . 50 118 . 03 110 . 46 -3 .. 61

1700 99 .39 391

.

90 117 . 36 117 . 91 110 . 04 -3 .. 38

1800 100 .27 397 . 61 127 . 32 117 . 78 109 . 62 -3 ., 18

1900 101 . 02 403 . 04 137 . 37 117 .65 109 . 16 -3 ,. 00

2000 101 .67 408 . 24 147 . 49 117 . 57 108 . 66 -2 ., 84

2100 102 . 22 413 . 20 157 . 68 117 .49 108 . 11 -2 ,,69

2200 102 . 68 417 . 97 167 . 91 117 .40 107 . 53 -2 ., 55

2300 103 . 07 422 . 54 178 . 19 117 .32 106 . 86 -2 .,43

2400 103 .40 426 . 93 188 . 51 117 . 24 106 . 19 -2 ., 31

2500 103 .68 431 . 15 198 . 87 117 . 15 105 . 44 -2 ., 20

2600 103 . 91 435 . 22 209 . 25 117 . 07 104 . 68 -2 ., 10

2700 104 . 11 439 . 14 219 . 66 117 . 03 103 . 89 -2 ., 01

2800 104 .27 442 . 93 230 . 08 116 . 94 103 . 09 -1

.

, 92

2900 104 .41 446 . 59 240 . 52 116 . 86 102 . 30 -1

.

, 84

3000 104 . 52 450 . 12 250 . 98 116 . 73 101

.

55 -1.. 77
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 599

CH2=CF«

ah2(0 k)

AHf{298.15 K)

S°(298.15 K)

Moments of Inertia

7.56 81.01

= 112 .80 kJ mol"^

= 109 . 10 kJ mol'^

= 256 . 78 J K"^ mol"

:iO-'*'^ kg m^; 88 .57

Vibrational Frequencies (cm"-'^) 427.0 613.9 796.9 905.2

1366.6 1536.6 2975.1 3076.3

1092 . 7

Bond Distances (10''-'-° m)

C1-C2 1.324

Cl-Hl 1.076

C1-H2 1.072

C2-F1 1.306

Bond Angles (°) C2-C1-H1 120.95

C2-C1-H2 119.84

C1-C2-F1 128.27

Dihedral Angles

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I

j kJ mol"^

T/K S° H°-H°(Tj.) AHf AG° Log Kf

150 36 .11 227 . 59 -6 . 39 111 . 13 107 . 74 -37 51

200 40 .85 238 . 73 -4 . 46 110 .42 106 . 65 -27 86

250 45 . 51 248 . 35 -2 . 30 109 . 75 105 . 77 -22 10

300 50 . 25 257 . 09 0 . 09 109 . 08 104 . 98 -18 28

350 54 . 78 265 . 18 2 . 72 108 .45 104 . 31 -15 56

400 58 . 97 272 . 78 5 . 56 107 . 91 103 . 68 -13 54

450 62 . 75 279 . 94 8 . 61 107 .40 103 . 14 -11 97

500 66 . 15 286 . 73 1183 106 . 94 102 . 68 -10 72

600 71 . 94 299 . 30 18 . 74 106 . 15 101 . 80 -8 86

700 76 .67 310 . 73 26 . 17 105 .44 101 . 13 -7 55

800 80 .62 32120 34 . 03 104 . 85 100 . 54 -6 57

900 83 . 96 330 . 86 42 . 26 104 . 31 100 . 12 -5 81

1000 86 . 84 339 . 82 50 . 80 103 . 85 99 . 75 -5 21

1100 89 . 32 348 . 18 59 . 60 103 .47 99. 50 -4 72

1200 91 .47 356 . 01 68 . 63 103 . 09 99 . 33 -4 32

1300 93 .35 363 . 36 77 . 87 102 . 80 99 . 20 -3 99

1400 94 . 99 370 . 29 87 . 27 102 . 51 99 . 12 -3 70

1500 96 .42 376 . 84 96 . 83 102 . 26 99 . 08 -3 45

1600 97 .70 383 . 08 106 . 53 102 . 05 98 . 99 -3 23

1700 98 . 82 389 . 04 116 . 34 101 . 88 98 . 87 -3 04

1800 99 . 81 394 . 72 126 . 26 101 . 71 98 . 74 -2 87

1900 100 .69 400 . 14 136 . 28 101 . 55 98 . 58 -2 71

2000 101 .47 405 . 33 146 . 37 101 .42 98 . 32 -2 57

2100 102 . 17 410 . 29 156 . 54 101 . 34 98 . 07 -2 44

2200 102 . 79 415 . 06 166 . 77 101 .25 97 . 78 -2 32

2300 103 . 34 419 . 64 177 . 07 101 . 17 97 . 40 -2 21

2400 103 . 82 424 . 05 187 . 41 101 . 13 96 . 99 -2 11

2500 104 . 25 428 . 30 197 . 81 101 . 09 96 . 52 -2 02

2600 104 .63 432 . 39 208 . 24 101 . 09 96 . 02 -1 93

2700 104 . 96 436 . 35 218 . 71 101 . 04 95 . 48 -1 85

2800 105 .25 440 . 17 229 . 20 101 . 04 94 . 93 -1 77

2900 105 . 50 443 . 86 239 . 73 101 . 04 94 . 39 -1 70

3000 105 . 73 447 . 44 250 . 29 101 . 04 93 . 89 -1 63
600 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHF=CF» fccf cis

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

= -38 . 79 kJ mol'^

= -41 . 10 kJ mol"^

= 281.12 J K'^ mol"^

Moments of Inertia (10

32 .42 146 .49

-47 kg m^)

178 . 91

Vibrational Frequencies (cm"-"^) 204.1 401.4 718.9 727.9

1120.4 1325.5 1579.2 3097.9

945 . 6

Bond Distances (10

C1-C2 1.324

Cl-Hl 1.068

Cl-Fl 1.330

C2-F2 1.300

10 m) Bond Angles (°) C2-C1-H1 123.49

C2-C1-F1 121.44

C1-C2-F2 128.59

Dihedral Angles ('

Hl-Cl-Fl

F1-C2-F2

180 . 00

0 . 00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K C

lO S° H°-H°(T^) AHo

£ AGf Log r 150 42 . la 247 . 25 -7 . 41 -40 . 00 -40 . 96 14 ., 27

200 47 . 64 260 . 26 -5 . 15 -40 .42 -41

.

25 10 .. 77

250 52 . 62 271 . 44 -2 . 65 -40 . 79 -41 . 46 8 .,66

300 57 . 55 281 . 47 0 . 10 -41 . 13 -41 . 63 7 ..25

350 62 . 21 290 . 70 3 . 10 -41 .42 -41 . 71 6 ..23

400 66 . 48 299 . 29 6 . 32 -41 .67 -41

.

80 5 ..46

450 70 . 31 307 . 34 9 . 74 -41 .88 -41

.

88 4 ,. 86

500 73 . 71 314 . 92 13 . 34 -42 . 09 -41

.

88 4 ..38

600 79 . 37 328 . 86 21 . 00 -42 .38 -41 . 88 3 ..65

700 83 . 81 341

.

42 29 . 17 -42 .63 -41

.

80 3 .. 12

800 87 . 35 352 . 83 37 . 73 -42 .89 -41

.

67 2 .. 72

900 90 . 22 363 . 27 46 . 61 -43 . 05 -41

.

42 2 ..40

1000 92 . 58 372 . 88 55 . 75 -43 .22 -41 . 17 2 .. 15

1100 94 . 55 381

.

77 65 . 11 -43 .39 -40 . 84 1 .. 94

1200 96 . 21 390 . 05 74 . 64 -43 . 56 -40 . 42 1 .. 76

1300 97 . 62 397 . 77 84 . 33 -43 . 72 -40 . 00 1

.

.61

1400 98 . 83 405 . 02 94 . 15 -43 .89 -39. 54 1

.

.48

1500 99 . 87 411

.

85 104 . 09 -44 . 02 -39 . 08 1

.

.36

1600 100 . 78 418 . 30 114 . 11 -44 . 14 -38 . 58 1

.

. 26

1700 101

.

57 424 . 40 124 . 23 -44 .27 -38 . 07 1

.

. 17

1800 102 . 27 430 . 19 134 . 42 -44 .39 -37 . 61 1

.

. 09

1900 102 . 90 435 . 70 144 . 67 -44 . 52 -37 . 20 1

.

. 02

2000 103 . 46 440 . 96 154 . 98 -44 . 64 -36 . 78 0 .. 96

2100 103 . 96 445 . 98 165 . 35 -44 . 73 -36 . 40 0 .. 91

2200 104 . 42 450 . 79 175 . 76 -44 . 81 -36 . 07 0 .. 86

2300 104 . 83 455 . 42 186 . 21 -44 . 85 -35 . 86 0 .. 81

2400 105 . 20 459 . 89 196 . 71 -44 . 94 -35 . 69 0 ,. 78

2500 105 . 54 464 . 19 207 . 24 -44 . 98 -35 . 61 0 ,. 74

2600 105 . 85 468 . 34 217 . 80 -45 . 02 -35 . 56 0 ,. 71

2700 106 . 14 472 . 34 228 . 40 -45 . 02 -35 . 52 0 ,.69

2800 106 . 39 476 . 20 239 . 02 -45 . 02 -35 . 52 0 .. 66

2900 106 . 63 479 . 94 249 . 66 -45 . 02 -35 . 48 0 ,. 64

3000 106 . 84 483 . 56 260 . 32 -45 . 02 -35 . 44 0 ,. 62
SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH 601

CHF=CF» fccf trans

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

Moments of Inertia

10.31 207.95

= -40.46 kJ mol'^

= -42 . 52 kJ mol"^

= 279.28 J K"^ mol"^

:iO"'^^ kg m^]

218 .26

Vibrational Frequencies (cm ) 295.9 297.9 518.3 684.6

1163.1 1259.9 1568.4 3051.6

1049 .

1

Bond Distances (10

C1-C2 1.327

Cl-Hl 1.072

Cl-Fl 1.327

C2-F2 1.304

10 m) Bond Angles (°) C2-C1-H1 124.86

C2-C1-F1 120.07

C1-C2-F2 125.81

Dihedral Angles ('

Hl-Cl-Fl

F1-C2-F2

180 . 00

180 . 00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K'^ mol"^ -1 j kJ mol"^

t/k %S

T T O TT H -H O / rri \ (Tj.) AB o

f AG? Log Kf

150 4 4 . 03 244 . 16 - 7 . 6 9 - 4

1

. 71 -42 . 22 14 6 9 200 49 . 58 257 . 62 - 5 . 3 5 -42 . 01 -42 . 34 11 06

250 54 . 62 269 . 25 -2 . 74 -42 .30 -42 . 43 8 87

300 59 . 32 279 . 65 0 . 11 -42 . 51 -42 . 51 7 40

350 63 .69 289. 14 3 . 18 -42 . 76 -42 . 51 6 35

400 67 . 70 297 . 93 6 . 47 -42 . 93 -42 . 55 5 55

450 71 . 32 306 . 12 9 . 95 -43 . 10 -42 . 51 4 94

500 74 . 55 313 . 81 13 . 60 -43 .26 -42 . 51 4 44

600 79 . 97 327 . 91 21

.

33 -43 .47 -42 . 38 3 69

700 84 .26 340 . 57 29. 55 -43 . 68 -42 . 26 3 15

800 87 . 68 352 . 05 38 . 15 -43 .85 -42 . 01 2 74

900 90 .46 362 . 54 47 . 07 -44 . 02 -41 . 76 2 42

1000 92 . 75 372 . 19 56 . 23 -44 . 18 -41

.

42 2 16

1100 94 . 65 381

.

12 65 . 60 -44 .35 -41

.

05 1 95

1200 96 .25 389 . 42 75 . 14 -44 .48 -40 . 63 1 77

1300 97 .60 397 . 17 84 . 84 -44 . 64 -40 . 12 1 61

1400 98 . 75 404 . 44 94 . 65 -44 . 77 -39 . 66 1 48

1500 99 . 73 411

.

28 104 . 58 -44 . 94 -39 . 12 1 36

1600 100 .58 417 . 74 114 . 60 -45 . 10 -38 . 62 1 26

1700 101 . 32 423 . 86 124 . 69 -45 .23 -38 . 12 1 17

1800 101 . 96 429 . 66 134 . 85 -45 .40 -37 . 66 1 09

1900 102 . 53 435 . 18 145 . 08 -45 . 52 -37 . 20 1 02

2000 103 . 02 440 . 45 155 . 35 -45 .69 -36 . 82 0 96

2100 103 .46 445 . 47 165 . 67 -45 .81 -36 . 44 0 91

2200 103 .85 450 . 29 176 . 04 -45 . 94 -36 . 11 0 86

2300 104 .20 454 . 91 186 . 44 -46 . 07 -35 . 86 0 81

2400 104 . 51 459 . 34 196 . 87 -46 . 19 -35 . 65 0 78

2500 104 . 79 463 . 60 207 . 34 -46 .28 -35 . 44 0 74

2600 105 . 04 467 . 71 217 . 83 -46 .40 -35 . 31 0 71

2700 105 .27 471 . 67 228 . 34 -46 .48 -35 . 19 0 68

2800 105 .48 475 . 49 238 . 88 -46 .61 -35 . 06 0 65

2900 105 .67 479 . 19 249 . 43 -46 .69 -34 . 94 0 63

3000 105 .85 482 . 77 260 . 01 -46 .78 -34 . 81 0 61
602 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF2=CH»

AH2(0 K)

Ah5(298.15 K)

S°(298.15 K)

64 . 94 kJ mol"^

67.66 kJ mol"^

276 . 92 J K" mol -1

Moments of Inertia (10"'*^ kg m^)

70 . 89 76 . 08 146 . 97

416 . 4 903 . 8 519 . 1 1278 . 3 528 . 2 1585 . 6 624 . 3 3128 .

7

818 . 7

Bond Distances (10

C1-C2 1.311

Cl-Fl 1.309

C1-F2 1.304

C2-H1 1.066

10 m) Bond Angles (°) C2-C1-F1 125.05

C2-C1-F2 124.78

C1-C2-H1 133.92

Dihedral Angles {°)

F1-C1-F2

F1-C2-H1

180 . 00

0 . 00

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p = 1 atmosphere

1 _

1 mol

1 _ 1 1 _ _ kJ mol"^ - 1

1 T/K 1c° S'

D H°-H°(T^) * r

AHo

f AG? Loq f 150 39 .ll 243 .63 -7 . 33 .TO4. R -66 . 90 23 . 30

200 46 . 3 9 255 . 96 - 5 . 19 -66 . 99 -66 . 99 17 . 49

250 53 . 12 267 . 05 -2 . 70 -67 .40 -66 . 99 13 ,. 99

300 59 . 19 277 .29 0 . 11 -67 . 66 -66 . 90 11 .65

350 64 .48 286 . 81 3 . 20 -67 . 86 -66 . 82 9 ,. 97

400 69 . 04 295 . 70 6 . 54 -67 . 99 -66 . 73 8 .. 71

450 72 . 97 304 . 04 10 . 08 -68 . 12 -66 . 57 7 ,. 73

500 76 .37 311 . 91 13 . 82 -68 . 16 -66 . 48 6 ,. 94

600 81 . 94 326 .39 21 . 72 -68 . 24 -66 . 23 5 ,. 77

700 86 . 26 339 .37 30 . 12 -68 . 24 -65 . 98 4 ,. 92

800 89 . 70 351 . 13 38 . 90 -68 . 24 -65 . 69 4 ..29

900 92 .46 361 . 84 47 . 99 -68 . 24 -65 . 35 3 ,. 79

1000 94 . 70 371 . 70 57 . 32 -68 . 24 -64 . 98 3 ,.39

1100 96 . 50 380 .80 66 . 86 -68 .20 -64 . 56 3 ,. 07

1200 97 . 96 389 .25 76 . 56 -68 . 20 -64 . 10 2 .. 79

1300 99 . 14 397 . 13 86 . 41 -68 . 20 -63 . 64 2 .. 56

1400 100 . 08 404 . 50 96 . 36 -68 . 24 -63 . 18 2 .. 36

1500 100 . 84 411 . 43 106 . 40 -68 . 24 -62 . 68 2 .. 18

1600 101 . 46 417 . 94 116 . 52 -68 . 32 -62 . 17 2 .. 03

1700 101 . 95 424 . 11 126 . 70 -68 .37 -61 . 67 1

.

. 90

1800 102 . 35 429 . 94 136 . 92 -68 . 45 -61 . 21 1

.

. 78

1900 102 .68 435 .47 147 . 18 -68 . 58 -60 . 79 1

.

.67

2000 102 . 94 440 . 75 157 . 48 -68 . 70 -60 . 42 1

.

. 58

2100 103 . 15 445 . 77 167 . 80 -68 . 83 -60 . 08 1

.

.49

2200 103 . 32 450 . 57 178 . 15 -68 . 95 -59 . 75 1

.

. 42

2300 103 .46 455 . 16 188 . 51 -69 . 12 -59 . 50 1

,

.35

2400 103 . 58 459 . 57 198 . 88 -69 . 33 -59 . 29 1

.

.29

2500 103 .67 463 . 79 209 . 27 -69 . 50 -59 . 12 1

.

. 24

2600 103 . 75 467 . 86 219 . 66 -69 . 71 -58 . 99 1

.

. 19

2700 103 . 81 471 . 77 230 . 07 -69 . 91 -58 . 87 1

.

, 14

2800 103 . 86 475 . 55 240 . 48 -70 . 12 -58 . 79 1

.

, 10

2900 103 . 91 479 . 19 250 . 89 -70 .37 -58 . 62 1

.

, 06

3000 103 . 94 482 . 71 261

.

31 -70 . 58 -58 . 45 1

.

. 02
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH . . . 603

CF2=CF»

AH2{0 K) = -215.73 kJ mol"^

AHf(298.15 K) = -216.36 kJ mol'^

S°(298.15 K) = 300.48 J K"^ mol"^

Moments of Inertia (10"'*^ kg m^)

72.80 217.89 290.68

Vibrational Frequencies (cm'-"-) 204.0 279.2 414.0 459.3 588.3

864.4 1203.3 1318.0 1663.5

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.319

Cl-Fl 1.304 C2-C1-F1 124.76

C1-F2 1.303 C2-C1-F2 122.90 F1-C1-F2 180.00

C2-F3 1.302 C1-C2-F3 126.52 F1-C2-F3 0.00

Enthalpy Reference Temperature = T^. = 2 9 8.15 K

Standard State Pressure = p^ = 1 atmosphere

1 _ J K

^ mol -L 1 1

1 kJ mol"^ 1 T/K S

_^

1 H°-H°(T^) AHo

f AG? Log Kf

150 49 .70 260 .20 -8 . 84 -216 .65 -214 . 14 74 56

200 57 . 00 275 . 54 -6 . 17 -216 .61 -213 . 30 55 71

250 63 . 09 288 . 94 -3 . 16 -216 . 52 -212 . 55 44 40

300 68 .28 300 . 90 0 . 13 -216 .35 -211 . 79 36 88

350 72 . 81 311 . 77 3 . 66 -216 . 19 -211

.

12 31 50

400 76 . 81 321 . 75 7 . 40 -215 . 98 -210 . 46 27 48

450 80 . 32 330 . 99 11 . 33 -215 . 81 -209 . 79 24 35

500 83 .41 339 .60 15 . 42 -215 .60 -209 . 16 21 85

600 88 .47 355 . 24 24 . 02 -215 .27 -207 . 99 18 10

700 92 . 33 369 . 15 33 . 07 -214 . 93 -206 . 82 15 43

800 95 .29 381 . 64 42 . 45 -214 . 64 -205 . 64 13 43

900 97 . 58 392 . 97 52 . 09 -214 .39 -204 . 47 11 87

1000 99 .37 403 .30 61 . 94 -214 . 18 -203 . 30 10 62

1100 100 . 81 412 .80 71 . 95 -214 . 01 -202 . 05 9 59

1200 101 . 98 421 . 58 82 . 09 -213 .89 -200 . 83 8 74

1300 102 . 95 429 . 73 92 . 33 -213 . 76 -199 . 53 8 02

1400 103 . 76 437 . 35 102 . 66 -213 .68 -198 . 24 7 40

1500 104 .45 444 .47 113 . 06 -213 . 64 -196 . 94 6 86

1600 105 . 04 451 . 18 123 . 53 -213 . 59 -195 . 64 6 39

1700 105 . 55 457 . 57 134 . 06 -213 . 55 -194 . 47 5 97

1800 106 . 00 463 .61 144 . 63 -213 . 51 -193 . 30 5 61

1900 106 .39 469 .36 155 . 24 -213 .47 -192 . 17 5 28

2000 106 . 74 474 . 82 165 . 89 -213 .43 -191

.

13 4 99

2100 107 . 04 480 . 03 176 . 57 -213 .43 -190 . 12 4 73

2200 107 .29 485 . 02 187 . 28 -213 . 38 -189 . 16 4 49

2300 107 . 52 489 .80 198 . 02 -213 . 34 -188 . 28 4 28

2400 107 . 72 494 . 38 208 . 77 -213 . 26 -187 . 44 4 08

2500 107 . 88 498 . 77 219 . 55 -213 . 22 -186 . 65 3 90

2600 108 . 02 503 . 01 230 . 34 -213 . 17 -185 . 90 3 73

2700 108 . 13 507 . 09 241 . 14 -213 . 13 -185 . 18 3 58

2800 108 . 22 511 . 02 251 . 95 -213 . 05 -184 . 47 3 44

2900 108 . 30 514 . 82 262 . 78 -213 . 01 -183 . 76 3 31

3000 108 .35 518 .49 273 . 60 -212 . 92 -183 . 05 3 19
604 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

-1

C2HF

AH?(0 K) = 117.86 kJ mol"^

AH°(298.15 K) = 117.80 kJ mol"^

S°(298.15 K) = 227.71 J K"^ mol

Moments of Inertia (10"'*^ kg m^)

84.53 84.53

Vibrational Frequencies (cm"^)

441.6 441.6 685.4 685.4 1042.9

2262.5 3293.0

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.177

Cl-Hl 1.055 C2-C1-H1 180.00

C2-F1 1.269 C1-C2-F1 180.00

Enthalpy Reference Temperature = Tj. = 2 98.15 K

Standard State Pressure = p^ = 1 atmosphere

1 _

1 J K'^ mol 1 1

1 1 kJ mol'^ - I

1 T /K lO s'

3 H°-H°(Tj.) 0

f AGf If 150 34 . 'Is 199 . 26 -6 ., 26 ITT Q 0 . 0 A 110 . 42 -38.

2 0 0 3 9 . 94 2 09 . 94 -4 ..40 117 . / 8 107 . 95 - 9 R 1 Q

9 1=; n u ^ ^ -L ^ A 1 -2 ..27 117 . 78 105 . 44 ~ooz z . 300 49. 38 228 . 02 0 .. 09 117 . 82 102 . 93 -17 .. 92

350 52 . 98 235 . 91 2 .. 65 117 . 86 100 . 37 -14 .. 98

400 55 . 94 243 . 18 5 .. 38 117 . 91 97 . 82 -12 .. 77

450 58 . 42 249 . 92 8 .. 24 117 . 95 95 . 27 -11

.

. 06

500 60 . 51 256 . 18 11 ..21 117 . 99 92 . 72 -9 ..68

600 63 . 94 267 . 53 17 .. 44 117 . 95 87 . 57 -7 .. 62

700 66 . 70 277 .60 23 .. 98 117 . 82 82 . 47 -6 .. 15

800 69. 03 286 .67 30 .. 77 117 . 61 77 . 45 -5 .. 06

900 71. 04 294 . 91 37 .. 78 117 .32 72 . 47 -4 ..21

1000 72 . 78 302 .49 44 .. 97 117 . 03 67 . 57 -3 .. 53

1100 74 . 30 309 . 50 52 .. 33 116 .69 62 . 72 -2 .. 98

1200 75 . 64 316 . 03 59 .. 82 116 . 32 57 . 95 -2 .. 52

1300 76 . 81 322 . 13 67 ..45 115 . 90 53 . 26 -2 .. 14

1400 77 . 83 327 . 86 75 .. 18 115 . 52 48 . 58 -1

,

. 81

1500 78 . 72 333 .26 83 ,. 01 115 . 10 43 . 93 -1

.

. 53

1600 79 . 51 338 . 36 90 ,. 92 114 . 68 39 . 33 -1 ..28

1700 80 . 20 343 .21 98 .. 91 114 . 31 34 . 69 -1 .. 07

1800 80 . 81 347 . 81 106 ,. 96 113 . 89 30 . 08 -0 .. 87

1900 81

.

35 352 . 19 115 ,. 06 113 .47 25 . 44 -0 .. 70

2000 81

.

83 356 . 38 123 ,.23 113 . 09 20 . 75 -0 .. 54

2100 82 . 26 360 . 38 131

,

.43 112 .68 16 . 07 -0 ..40

2200 82 . 65 364 .22 139 ,. 68 112 .30 11

.

30 -0 ,.27

2300 82 . 99 367 . 90 147 ,. 96 111 . 92 6 . 53 -0 ,. 15

2400 83 . 30 371 .43 156 ,.27 111 . 55 1

.

72 -0 .. 04

2500 83 . 58 374 . 84 164 . 62 111 . 17 -3 . 14 0 ,. 07

2600 83 . 83 378 . 12 172 . 99 110 . 79 -8 . 03 0 ,. 16

2700 84 . 05 381 .29 181 . 38 110 .42 -12 . 89 0 .25

2800 84 . 26 384 .35 189 .80 110 . 08 -17 . 78 0 ,.33

2900 84 . 45 387 .31 198 . 23 109 . 75 -22 . 68 0 ,.41

3000 84 . 62 390 . 18 206 .69 109 .41 -27 . 49 0 ,.48
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 605

C2F2

AH°(0 K)

AHf(298.15 K)

S°(298.15 K)

Moments o£ Inertia

231.40 231.40

29 . 96 kJ mol"^

31.82 kJ mol"^

= 245 .88 J K'^ mol"^

dO'^"^ kg m^)

Vibrational Frequencies (cm"-'-) 291.1 291.1 407.1 407.1 771.4

1321.2 2502.4

Bond Distances (10"-^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.166

Cl-Fl 1.275 C2-C1-F1 180.00

C2-F2 1.275 C1-C2-F2 180.00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -j I kJ mol"^

T/K ( L

S° H°-H AH0

f AGf Log Kf

150 43 .72 210 . 49 - 7 . 76 30 . 42 26 . 69 - 9 .. 30

200 50 .36 224 . 02 -5 . 40 30 . 84 25 . 40 -6 .. 63

250 55 .28 235 . 81 -2 . 75 31 . 34 23 . 93 -5 .. 00

300 59 . 05 246 . 24 0 . 11 31 . 84 22 . 38 -3 ,.89

350 62 . 05 255 . 58 3 . 14 32 .30 20 . 71 -3 .. 09

400 64 . 51 264 . 03 6 . 31 32 . 72 18 . 95 -2 ..48

450 66 . 59 271 . 75 9 . 59 33 . 10 17 . 20 -1 .. 99

500 68 .37 278 . 86 12 . 96 33 .39 15 . 36 -1

.

. 60

600 71 . 31 291

.

60 19 . 95 33 . 85 11

.

63 -1 .. 01

700 73 . 66 302 . 78 27 . 20 34 . 14 7 . 87 -0 .. 59

800 75 . 58 312 . 74 34 . 67 34 . 31 4 . 10 -0 ..27

900 77 . 16 321 . 74 42 . 30 34 .35 0 . 33 -0 ,. 02

1000 78 .48 329 . 93 50 . 09 34 .31 -3 . 35 0 .. 18

1100 79 . 58 337 . 47 57 . 99 34 . 18 -7 . 03 0 ,.33

1200 80 . 51 344 . 44 66 . 00 34 . 02 -10 . 63 0 .. 46

1300 81 .30 350 . 91 74 . 09 33 . 81 -14 . 23 0 ,. 57

1400 81 . 96 356 . 96 82 . 26 33 . 56 -17 . 78 0 ..66

1500 82 . 53 362 . 64 90 . 48 33 .26 -21

.

30 0 .. 74

1600 83 . 01 367 . 97 98 . 76 32 . 97 -24 . 81 0 ,. 81

1700 83 .43 373 . 02 107 . 08 32 . 64 -28 . 37 0 ,.87

1800 83 . 80 377 . 80 115 . 44 32 .34 -31

.

88 0 ,. 93

1900 84 . 12 382 . 34 123 . 84 32 . 01 -35 . 44 0 .. 97

2000 84 .39 386 . 66 132 . 26 31 .67 -39 . 04 1

,

. 02

2100 84 . 63 390 . 79 140 . 72 31 . 34 -42 . 63 1

,

. 06

2200 84 . 85 394 . 73 149 . 19 31 . 00 -46 . 32 1

,

. 10

2300 85 . 04 398 . 51 157 . 69 30 .67 -50 . 04 1

.

, 14

2400 85 .21 402 . 13 166 . 20 30 . 33 -53 . 76 1

.

. 17

2500 85 . 36 405 . 61 174 . 73 30 . 00 -57 . 53 1

,

. 20

2600 85 . 50 408 . 96 183 . 27 29 .66 -61 . 34 1

,

. 23

2700 85 . 62 412 . 19 191

.

83 29 . 33 -65 . 19 1

.

. 26

2800 85 . 73 415 . 30 200 . 39 29 . 00 -68 . 99 1

,

.29

2900 85 . 83 418 . 32 208 . 97 28 . 70 -72 . 80 1

.

.31

3000 85 . 93 421

.

23 217 . 56 28 .37 -76 . 57 1

,

. 33
606 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

•C2F

AH2(0 K) = 448.94 kJ mol"^

Ah|(298.15 K) = 453.86 kJ mol"^

S°(298.15 K) = 218.02 J K"^ mol"^

Moments of Inertia (10'^^ kg m^)

75.82 75.82

Vibrational Frequencies (cm"-"-) 217.8 1003.0 2071.7

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.196

Cl-Fl 1.272 C2-C1-F1 179.35

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

j - J K"^ mol"^ - |

j kJ mol"^

T/K 1 S° H°-H°{Tj.) AH0

f AG° Log

150 39 Is 190 . 07 - (ZD ,. U b 451 . 91 4 3/. D b - 152 ,. 3 9 200 40 . 38 201

.

51 - 4 ,. U 1 452 . 71 /I "3 0 4 J Z . 1 b - 113 ,. 01

2 50 4

1

. 44 210 . 64 - 2 ,. 02 453 . 38 42 7. 6 0 -89,. 34

300 42 . 4 9 218 . 2 8 0 ,. 08 4 53 . 88 422 . 37 -73 54

350 43 . 53 224 . 91 2 ,.23 454 .26 417 . 06 -62 ,. 24

400 44 . 53 230 . 79 4 ,.43 454 . 51 411 . 66 -53 ,. 75

450 45 .49 236 . 09 6 ,. 68 454 .68 406 . 27 -47 ,. 15

500 46 .40 240 . 93 8 ,. 98 454 . 76 400 . 83 -41

.

. 87

600 48 . 08 249 . 54 13 .. 70 454 . 68 389 . 99 -33 ,. 95

700 49 . 53 257 . 06 18 ,. 59 454 . 34 379 . 15 -28 ,.29

800 50 . 77 263 . 76 23 ,. 61 453 . 80 368 . 48 -24 ,. 06

900 51 . 81 269 . 80 28 ,. 74 453 . 13 357 . 86 -20 ., 77

1000 52 . 68 275 . 31 33 ,. 96 452 . 33 347 . 40 -18 .. 15

1100 53 .40 280 . 36 39 ,.27 451 .45 337 . 02 -16 .. 00

1200 54 . 00 285 . 04 44 .. 63 450 . 45 326 . 81 -14 ..22

1300 54 . 50 289 . 38 50 ,. 06 449 .40 316 . 69 -12 .. 72

1400 54 . 93 293 . 43 55 .. 53 448 . 32 306 . 65 -11

,

.44

1500 55 .29 297 . 24 61 .. 04 447 . 14 296 . 69 -10 ,. 33

1600 55 .60 300 . 81 66 .. 59 445 . 97 286 . 81 -9 .. 36

1700 55 . 86 304 . 19 72 .. 16 444 . 76 276 . 98 -8 .. 51

1800 56 . 08 307 . 39 77 ., 76 443 . 55 267 . 19 -7 .. 75

1900 56 . 28 310 . 43 83 .. 38 442 .29 257 . 40 -7 .. 08

2000 56 .45 313 . 32 89 ,, 01 440 . 99 247 . 65 -6 ..47

2100 56 . 60 316 . 08 94 ..67 439 . 74 237 . 94 -5 .. 92

2200 56 . 73 318 . 72 100 ., 34 438 . 44 228 . 20 -5 ..42

2300 56 . 85 321

.

24 106 .. 01 437 . 14 218 . 49 -4 ., 96

2400 56 . 95 323 . 66 Ill .. 70 435 . 85 208 . 74 -4 ,. 54

2500 57 . 04 325 . 99 117 ..40 434 . 55 199 . 03 -4 ., 16

2600 57 . 13 328 . 23 123 .. 11 433 .25 189 . 33 -3 ., 80

2700 57 .20 330 . 39 128 ..83 431 . 96 179 . 66 -3 .,48

2800 57 .27 332 . 46 134 .. 55 430 . 62 170 . 00 -3 ., 17

2900 57 .33 334 . 48 140 ..28 429 . 32 160 . 41 -2 .,89

3000 57 .38 336 . 42 146 .. 02 428 . 02 150 . 88 -2 ., 63
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 607

CHF=C=0

AH2(0 K)

AHf(298.15 K)

S°(298.15 K)

Moments of Inertia

14.45 180.24

= -144 . 68 kJ mol"^

= -147.19 kJ mol'^

= 270.51 J K"^ mol"

(10"^^ kg m^; 194 .69

Vibrational Frequencies (cm' ) 252.4 479.7 615.2 681.3 1027

1207.3 1425.2 2107.9 3073.7

Bond Distances {10

C1-C2 1.308

Cl-Fl 1.339

Cl-Hl 1.067

C2-01 1.147

-10 m) Bond Angles (°) C2-C1-F1 120.10

C2-C1-H1 122.43

C1-C2-01 180.00

Dihedral Angles

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I - J K"^ mol"^ -I

j kJ mol"^

T/K 1 L

« O s H°-H° (Tj.) AHo

f AG° Log Kf

150 41 . 64 237 . 24 - 7 . 30 -146 . 06 -145 . 35 50 ,. 61

200 46 . 89 249 . 94 -5 . 09 -146 .48 -145 . 06 37 ,. 88

250 52 . 00 260 . 95 -2 . 62 -146 . 86 -144 . 68 30 ,.23

300 56 . 74 270 . 86 0 . 10 -147 . 19 -144 . 26 25 .. 12

350 61 . 06 279 . 93 3 . 05 -147 .49 -143 . 80 21.46

400 64 . 95 288 . 34 6 . 20 -147 . 74 -143 . 30 18 ,, 71

450 68 .45 296 . 20 9 . 54 -147 . 99 -142 . 76 16 .. 57

500 71 . 58 303 . 58 13 . 04 -148 .20 -142 . 21 14 .. 86

600 76 . 94 317 . 12 20 . 48 -148 . 53 -141 . 08 12 .. 28

700 81 .30 329 . 32 28 . 40 -148 .87 -139 . 83 10 ,.43

800 84 . 89 340 . 41 36 . 71 -149 .20 -138 . 53 9 .. 04

900 87 . 88 350 . 59 45 . 35 -149 .49 -137 . 15 7 .. 96

1000 90 .40 359 . 99 54 . 27 -149 . 75 -135 . 69 7 ,. 09

1100 92 . 51 368 . 70 63 . 42 -150 . 04 -134 . 18 6 ..37

1200 94 .30 376 . 83 72 . 76 -150 .29 -132 . 59 5 .. 77

1300 95 . 83 384 . 44 82 . 27 -150 . 54 -130 . 96 5 .. 26

1400 97 . 13 391

59 91 . 92 -150 . 79 -129 . 33 4 .. 83

1500 98 .25 398 . 33 101 . 69 -151 . 04 -127 . 65 4 ..45

1600 99 . 22 404 . 71 Ill57 -151 .25 -125 . 98 4 .. 11

1700 100 . 06 410 . 75 121

53 -151 . 50 -124 . 31 3 .. 82

1800 100 . 79 416 . 49 131

57 -151 . 71 -122 . 67 3 .. 56

1900 101 . 44 421 . 96 141

69 -151 . 92 -121

04 3 ..33

2000 102 . 00 427 . 17 151 . 86 -152 . 13 -119. 45 3 .. 12

2100 102 . 50 432 . 16 162 . 08 -152 . 34 -117 . 95 2 .. 93

2200 102 . 94 436 . 94 172 . 36 -152 . 51 -116 . 44 2 .. 76

2300 103 .33 441 . 52 182 . 67 -152 . 72 -114 . 98 2 ,.61

2400 103 .68 445 . 93 193 . 02 -152 . 88 -113 . 60 2 ,.47

2500 104 . 00 450 . 17 203 . 41 -153 . 05 -112 . 21 2 ,. 34

2600 104 .28 454 . 25 213 . 82 -153 . 18 -110 . 92 2 ,. 23

2700 104 . 54 458 . 19 224 . 26 -153 . 34 -109 . 62 2 ,. 12

2800 104 . 77 462 . 00 234 . 73 -153 .47 -108 . 32 2 . 02

2900 104 . 98 465 . 68 245 . 22 -153 . 64 -107 . 03 1 . 93

3000 105 . 17 469 . 24 255 . 72 -153 . 76 -105 . 69 1 . 84
608 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF2=C=0

AH?(0 K) = -289.91 kJ mol"^

AH|(298.15 K) = -290.43 kJ mol"^

S°(298.15 K) = 288.81 J K"^ mol"^

Moments of Inertia (10"^^ kg m^)

76.39 209.01 285.39

Vibrational Frequencies (cm"-'^) 206.6 273.3 378.2 442.2 672.8

793.3 1314.7 1450.3 2208.3

Bond Distances (10"-'^° m) Bond Angles (°) Dihedral Angles (' C1-C2 1.300

Cl-Fl 1.312 C2-C1-F1 122.97

C1-F2 1.312 C2-C1-F2 122.97

C2-01 1.151 C1-C2-01 180.00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K'^ mol"^ -I I kJ mol'^

t/k C

o S° H°-H°(Tj.) AHf AG° Log Kf

150 50 . 12 248 . 43 -8 . 86 -290 79 -286 . 35 99 71

200 57 . 28 263 . 88 -6 . 16 -290 75 -284 . 89 74 40

250 63 . 05 277 . 30 -3 . 15 -290 62 -283 . 47 59 22

300 67 . 84 289 . 23 0 . 13 -290 41 -282 . 09 49 12

350 71

.

94 300 . 01 3 . 62 -290 20 -280 . 79 41 90

400 75 . 49 309 . 85 7 . 31 -290 03 -279 . 49 36 50

450 78 . 62 318 . 93 11 . 16 -289 83 -278 . 24 32 29

500 81

.

39 327 . 3 6 15 . 17 -289 66 -276 . 98 28 94

600 86 . 04 342 . 62 23 . 55 -289 32 -274 . 55 23 90

700 89 . 73 356 . 17 32 . 34 -289 07 -272 . 17 20 31

800 92 . 67 368 . 36 41

.

47 -288 91 -269 . 74 17 61

900 95 . 03 379 . 41 50 . 86 -288 78 -267 . 36 15 52

1000 96 . 93 389 . 52 60 . 46 -288 70 -264 . 89 13 84

1100 98 . 47 398 . 84 70 . 23 -288 65 -262 . 42 12 46

1200 99 . 73 407 . 46 80 . 14 -288 65 -259 . 91 11 31

1300 100 . 77 415 . 48 90 . 17 -288 70 -257 . 40 10 34

1400 101

.

64 422 . 99 100 . 29 -288 74 -254 . 85 9 51

1500 102 . 37 430 . 02 110. 49 -288 82 -252 . 30 8 79

1600 102 . 98 436 . 65 120 . 76 -288 91 -249 . 78 8 15

1700 103 . 51 442 . 91 131

.

08 -289 03 -247 . 23 7 60

1800 103 . 96 448 . 84 141

.

46 -289 11 -244 . 72 7 10

1900 104 . 35 454 . 47 151

.

88 -289 24 -242 . 25 6 66

2000 104 . 69 459 . 83 162 . 33 -289 37 -239 . 83 6 26

2100 104 . 98 464 . 95 172 . 81 -289 49 -237 . 44 5 91

2200 105 . 24 469 . 84 183 . 32 -289 62 -235 . 10 5 58

2300 105 . 47 474 . 52 193 . 86 -289 74 -232 . 84 5 29

2400 105 . 67 479 . 01 204 . 42 -289 87 -230 . 62 5 02

2500 105 . 85 483 . 33 214 . 99 -289 99 -228 . 40 4 77

2600 106 . 01 487 . 49 225 . 58 -290 12 -226 . 27 4 55

2700 106 . 16 491 . 49 236 . 20 -290 20 -224 . 14 4 34

2800 106 . 29 495 . 35 246 . 82 -290 33 -222 . 00 4 14

2900 106 . 40 499 . 08 257 . 45 -290 45 -219 . 87 3 96

3000 106 . 51 502 . 70 268 . 10 -290 54 -217 . 69 3 79
SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH 609

•CF=C=0

AH|(0 K) = 66.65 kJ mol"^

AHf(298.15 K) = 69.03 kJ mol"^

S°(298.15 K) = 276.17 J K"^ mol"^

Moments of Inertia (10 kg m^)

5.47 195.70 201.18

Vibrational Frequencies (cm"-'-) 250.8 279.2 496.8 764.5 1332.9

1712 . 2 Bond Distances (10'^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.398

Cl-Fl 1.296 C2-C1-F1 125.03

C2-01 1.163 X1-C2-01 128.40 C1-C2-01 180.00

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I

] kJ mol'^

r° n n

° (T ) AHo

f AG? Log

X o u . ± D Z ± . Z .J / . D J D / 59 . 54 z u

1 A

o n n A Q Q Q cD . Z D C O b o

1 c 56 . 74 - X 4

Q O O Z

o c n C "3 D J . b onZoo£Z C . - z . D bC. Q D O . b z 53 . 81 - X XO A z 4 300 56 . 84 276 . 52 0 . 10 69 . 04 50 . 75 -8 84

350 59 .56 285 . 50 3 . 02 69 . 45 47 . 66 -7 11

400 61 . 96 293 . 61 6 . 05 69 . 83 44 . 43 -5 80

450 64 . 10 301

.

03 9 . 21 70 . 12 41

.

21 -4 78

500 66 . 02 307 . 89 12 . 46 70 .42 37 . 95 -3 96

600 69 .23 320 . 22 19 . 23 70 . 84 31

.

34 -2 73

700 71 . 75 331

.

09 26 . 28 71 . 09 24 . 69 -1 84

800 73 . 72 340 . 80 33 . 56 71 .21 18 . 03 -1 18

900 75 .27 349 . 58 41

.

02 71 . 17 11 . 42 -0 66

1000 76 .49 357 . 58 48 . 61 71 . 04 4 . 90 -0 25

1100 77 . 46 364 . 91 56 . 30 70 . 84 -1 . 63 0 08

1200 78 . 24 371 . 69 64 . 09 70 . 54 -8 . 08 0 35

1300 78 . 89 377 . 98 71

.

95 70 . 17 -14 . 52 0 58

1400 79 . 41 383 . 84 79 . 86 69 . 75 -20 . 88 0 78

1500 79 .85 389 . 34 87 . 83 69 .29 -27 . 20 0 95

1600 80 .22 394 . 50 95 . 83 68 . 78 -33 . 51 1 09

1700 80 . 53 399 . 38 103 . 87 68 .24 -39 . 79 1 22

1800 80 . 79 403 . 99 Ill

.

93 67 . 66 -46 . 11 1 34

1900 81 . 02 408 . 36 120 . 03 67 . 07 -52 . 38 1 44

2000 81 .22 412 . 52 128 . 14 66 .44 -58 . 70 1 53

2100 81 .39 416 . 49 136 . 27 65 . 81 -65 . 06 1 62

2200 81 . 53 420 . 28 144 . 41 65 . 19 -71 . 42 1 70

2300 81 .67 423 . 91 152 . 58 64 . 56 -77 . 82 1 77

2400 81 . 78 427 . 38 160 . 75 63 .89 -84 . 22 1 83

2500 81 . 89 430 . 72 168 . 93 63 .26 -90 . 67 1 89

2600 81 . 98 433 . 94 177 . 13 62 . 59 -97 . 11 1 95

2700 82 . 06 437 . 03 185 . 33 61 . 92 -103 . 60 2 00

2800 82 . 14 440 . 02 193 . 54 61 .30 -110 . 04 2 05

2900 82 .20 442 . 90 201 . 75 60 .63 -116 . 44 2 10

3000 82 .26 445 . 69 209 . 98 59 . 96 -122 . 80 2 14
610 SECTION 5. FLAME INHffilTION CHEMISTRY AND THE SEARCH

CHjF-CHO fcco trans

AH2(0 K)

AHf(298.15 K)

S°(298.15 K)

319.20 kJ mol"^

•328 . 72 kJ mol"^

286.63 J K"^ mol"

Moments of Inertia

20.71 190.45

(10"^'^ kg m^)

205 . 95

Vibrational Frequencies (cm"^)

84.2 315.4 514.9 721.1 1019.8

1078.4 1105.5 1222.5 1328.4 1391.8

1461.2 1821.7 2857.6 2910.4 2959.4

Bond Distances (10 -10 m)

C1-C2

C2-01

C2-H1

Cl-Fl

C1-H2

C1-H3

511

186

093

364

082

082

Bond Angles

C1-C2-01

01-C2-H1

C2-C1-F1

C2-C1-H2

C2-C1-H3

122 . 12

122 . 54

109 .29

109 . 91

109 . 91

Dihedral Angles (°)

C1-C2-H1

01-Cl-Fl

F1-C1-H2

F1-C1-H3

180 . 00

180 . 00

119 . 87

-119 . 87

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

- J K"^ mol"^ -

I I kJ mol"^

T/K S° H°-H°(T^) AHo

f AG? Log Kf

150 48 .26 248 . 84 -8 . 27 -324 . 22 -308 . 74 107 51

200 53 .38 263 . 40 -5 . 74 -325 . 81 -303 . 38 79 23

250 58 .64 275 . 86 -2 . 95 -327 .31 -297 . 65 62 19

300 64 . 14 287 . 03 0 . 12 -328 . 78 -291

.

62 50 77

350 69 . 75 297 . 33 3 . 46 -330 . 12 -285 . 39 42 59

400 75 .28 307 . 00 7 . 10 -331 . 33 -278 . 95 36 42

450 80 . 60 316 . 17 11 . 00 -332 . 46 -272 . 34 31 61

500 85 .60 324 . 92 15 . 17 -333 . 46 -265 . 64 27 75

600 94 . 55 341

.

33 24 . 20 -335 . 14 -251 . 96 21 94

700 102 . 19 356 . 50 34 . 07 -336 . 44 -238 . 03 17 76

800 108 .68 370 . 58 44 . 64 -337 . 44 -223 . 89 14 62

900 114 . 22 383 . 75 55 . 81 -338 . 15 -209 . 66 12 17

1000 118 . 94 396 . 35 67 . 48 -338 .65 -195 . 56 10 21

1100 122 . 98 408 . 18 79 . 60 -338 . 95 -181

.

46 8 62

1200 126 .43 419 . 32 92 . 09 -339 . 11 -167 . 36 7 28

1300 129 .40 429 . 84 104 . 89 -339 . 11 -153 . 26 6 16

1400 131 . 95 439 . 79 117 . 97 -339 . 03 -139 . 20 5 19

1500 134 . 16 449 . 22 131 . 29 -338 . 90 -125 . 14 4 36

1600 136 . 07 458 . 18 144 . 80 -338 .65 -Ill

.

17 3 63

1700 137 . 73 466 . 70 158 . 50 -338 .36 -97 . 28 2 99

1800 139 . 18 474 . 83 172 . 36 -338 . 03 -83 . 43 2 42

1900 140 .46 482 . 60 186 . 35 -337 .65 -69 . 71 1 92

2000 141 . 59 490 . 03 200 . 46 -337 .23 -56 . 02 1 46

2100 142 . 58 497 . 16 214 . 67 -336 . 77 -42 . 47 1 06

2200 143 .46 503 . 99 228 . 98 -336 . 31 -29 . 04 0 69

2300 144 .25 510 . 56 243 . 37 -335 .77 -15 . 65 0 36

2400 144 . 95 516 . 88 257 . 83 -335 . 26 -2 . 38 0 05

2500 145 . 58 522 . 97 272 . 36 -334 . 68 10 . 79 -0 23

2600 146 . 15 528 . 85 286 . 95 -334 . 13 23 . 85 -0 48

2700 146 . 67 534 . 53 301 . 60 -333 . 55 36 . 90 -0 71

2800 147 . 13 540 . 02 316 . 29 -332 . 92 49 . 87 -0 93

2900 147 . 55 545 . 33 331

.

03 -332 . 34 62 . 80 -1 13

3000 147 . 93 550 . 47 345 . 80 -331 . 67 75 . 77 -1 32
SECTION 5. FLAME INHIBrnON CHEMISTRY AND THE SEARCH 611

CHjF-CHO fcco gauche

AH|(0 K)

AHf(298.15 K)

S°(298.15 K)

Moments of Inertia (10

•312 .67 kJ mol"^

•322 . 75 kJ mol"^

281.30 J K"^ mol"^

-47 kg m^)

41 . 73 135 . 31 171 . 87

Vibrational Frequencies (cm"-'-) 164.6 251.0 722.4 726.2 845.9

1111.1 1124.2 1239.4 1380.4 1405.7

1464.4 1835.0 2815.4 2881.6 2918.0

Bond Distances (10' m) Bond Angles (°) C1-C2

C2-01

C2-H1

Cl-Fl

C1-H2

C1-H3

1 . 507

1 . 183

1 . 096

1 .351

1 . 085

1 . 085

C1-C2-01

01-C2-H1

C2-C1-F1

C2-C1-H2

C2-C1-H3

124 .47

121 . 76

111 . 18

109.23

109.21

Dihedral Angles C

C1-C2-H1 180.03

01-Cl-Fl 0.10

F1-C1-H2 120.95

F1-C1-H3 -120.96

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

]- J K"^ mol"^ -

I I kJ mol"^

T/K C

lO S° H°-H°(T^) AHo

f AG° Log Kf

150 46 . 245 . 13 -7 . 96 -317 . 94 -30192 105 .. 13

200 50 . 97 259 . 05 -5 . 53 -319 . 62 -296 . 35 77 ..39

250 56 . 38 270 . 93 -2 . 85 -321 .25 -290 . 37 60 .. 66

300 62 . 45 281

68 0 . 12 -322 . 80 -284 . 05 49 .. 46

350 68 . 80 291

71 3 . 39 -324 . 22 -277 . 52 41

. 41

400 75 . 10 301

24 6 . 99 -325 .47 -270 . 79 35 ,. 36

450 81 . 15 310 . 35 10 . 89 -326 .60 -263 . 88 30.,63

500 86 . 83 319 . 18 15 . 08 -327 . 57 -256 . 90 26 .. 84

600 96 . 93 335 . 93 24 . 27 -329 . 11 -242 . 71 21 .. 13

700 105 . 41 35153 34 . 39 -330 . 16 -228 . 28 17 .. 03

800 112 . 46 366 . 07 45 . 28 -330 . 83 -213 . 68 13 ., 95

900 118 . 31 379 . 67 56 . 82 -331 . 16 -198 . 99 11 .. 55

1000 123 . 16 392 . 39 68 . 90 -331 .25 -184 . 22 9 .. 62

1100 127 . 19 404 . 32 81

42 -331 . 16 -169 . 41 8 ,. 05

1200 130 . 55 415 . 53 94 . 31 -330 . 91 -154 . 60 6 .. 73

1300 133 . 37 426 . 10 107 . 51 -330 . 54 -139 . 79 5 ..62

1400 135 . 73 436 . 07 120 . 98 -330 . 08 -125 . 02 4 ..66

1500 137 . 74 445 . 50 134 . 66 -329 . 53 -110 . 25 3 ,. 84

1600 139 . 44 454 . 45 148 . 53 -328 . 99 -95 . 52 3 .. 12

1700 140 . 90 462 . 95 162 . 56 -328 . 36 -80 . 88 2 ,.48

1800 142 . 16 471 . 04 176 . 72 -327 .69 -66 . 27 1. 92

1900 143 . 26 478 . 75 19100 -327 . 02 -51

76 1

.42

2000 144 . 20 486 . 13 205 . 38 -326 .35 -37 . 32 0 ,. 97

2100 145 . 03 493 . 18 219 . 86 -325 . 60 -22 . 97 0 ,. 57

2200 145 . 76 499 . 95 234 . 41 -324 .89 -8 . 74 0 .21

2300 146 . 41 506 . 44 249 . 03 -324 . 18 5 . 44 -0 ,. 12

2400 146 . 98 512 . 68 263 . 71 -323 .42 19 . 54 -0 . 42

2500 147 . 49 518 . 69 278 . 44 -322 . 67 33 . 51 -0 . 70

2600 147 . 95 524 . 49 293 . 22 -321 . 88 47 . 45 -0 . 95

2700 148 . 35 530 . 08 308 . 05 -321 . 12 6134 -1 . 19

2800 148 . 72 535 . 48 322 . 91 -320 .33 75 . 14 -1 .40

2900 149 . 06 540 . 70 337 . 81 -319 . 57 88 . 99 -1 .60

3000 149 . 36 545 . 76 352 . 74 -318 . 78 102 . 80 -1 . 79
612 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHFj-CHO hcco cis

AH?(0 K)

AHf(298.15 K)

S°(298.15 K)

•530 . 70 kJ mol"^

•538.95 kJ mol"^

304 . 75 J K"^ mol

Moments of Inertia

87.24 211.96

(lO"'*'^ kg m^)

272 .48

Vibrational Frequencies (cm"-'-) 76.3 315.5 368.5 416.3 592.2

977.7 1081.8 1127.5 1131.2 1312.2

1368.9 1395.3 1841.3 2875.1 3003.1

Bond Distances (10

C1-C2 1.516

C2-01 1.183

C2-H1 1.091

C1-H2 1.077

Cl-Fl 1.339

C1-F2 1.339

10 m) Bond Angles C

C1-C2-01

01-C2-H1

C2-C1-H2

C2-C1-F1

C2-C1-F2

122 . 03

123 . 13

112 . 55

108 .25

108 .25

•Dihedral Angles (°)

C1-C2-H1

01-C1-H2

H2-C1-F1

H2-C1-F2

180 . 00

0 . 00

121 . 72

-121 . 72

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

|- J K"^ mol'^ -| I kJ mol"^

T/K S° H°-H AHo

f AG° Log Kf

150 53 261

.

96 -9 . 38 -535 . 51 -516 . 60 179 89

200 60 .40 278 . 25 -6 . 55 -536 .81 -510 . 11 133 22

250 67 . 02 292 . 44 -3 . 37 -537 . 94 -503 . 38 105 17

300 73 .30 305 . 21 0 . 13 -538 . 98 -496 . 39 86 43

350 79 . 34 316 . 95 3 . 95 -539 . 90 -489 . 28 73 02

400 85 . 08 327 . 93 8 . 07 -540 . 74 -482 . 04 62 94

450 90 . 42 338 . 26 12 . 46 -541 .45 -474 . 67 55 10

500 95 . 34 348 . 04 17 . 11 -542 . 08 -467 . 27 48 81

600 103 . 90 366 . 20 27 . 10 -543 . 08 -452 . 25 39 37

700 110 . 95 382 . 76 37 . 87 -543 . 79 -437 . 06 32 61

800 116 . 78 397 . 97 49 . 28 -544 .21 -421 . 79 27 54

900 121 . 62 412 . 01 61

.

22 -544 . 51 -406 . 43 23 59

1000 125 . 66 425 . 19 73 . 60 -544 . 59 -391

.

12 20 43

1100 129 . 06 437 . 62 86 . 36 -544 . 59 -376 . 02 17 85

1200 131 . 94 449 . 25 99 . 42 -544 . 46 -360 . 87 15 71

1300 134 .38 460 . 18 112 . 75 -544 .30 -345 . 81 13 89

1400 136 .47 470 . 47 126 . 31 -544 . 00 -330 . 75 12 34

1500 138 . 26 480 . 20 140 . 05 -543 . 71 -315 . 77 11 00

1600 139 . 80 489 . 41 153 . 96 -543 .33 -300 . 83 9 82

1700 141 . 14 498 . 14 168 . 02 -542 . 92 -285 . 98 8 79

1800 142 .30 506 . 46 182 . 20 -542 . 50 -271

.

21 7 87

1900 143 . 32 514 . 39 196 . 48 -542 . 00 -256 . 56 7 05

2000 144 .21 521

.

95 210 . 87 -541 . 54 -241

.

96 6 32

2100 145 . 00 529 . 20 225 . 33 -540 . 99 -227 . 48 5 66

2200 145 . 70 536 . 14 239 . 87 -540 .45 -213 . 13 5 06

2300 146 . 32 542 . 80 254 . 48 -539 . 90 -198 . 82 4 52

2400 146 . 88 549 . 21 269 . 14 -539 . 32 -184 . 68 4 02

2500 147 . 38 555 . 38 283 . 86 -538 . 73 -170 . 58 3 56

2600 147 . 82 561 . 32 298 . 62 -538 . 15 -156 . 57 3 15

2700 148 . 23 567 . 06 313 . 43 -537 . 52 -142 . 63 2 76

2800 148 . 59 572 . 60 328 . 27 -536 . 89 -128 . 74 2 40

2900 148 . 93 577 . 96 343 . 15 -536 . 22 -114 . 89 2 07

3000 149 .23 583 . 15 358 . 06 -535 . 59 -101 . 09 1 76

i
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 613

CHFj-CHO hcco trans

AH2(0 K) = -514.55 kJ mol"^

AHf(298.15 K) = -525.02 kJ mol"^

S°(298.15 K) = 287.06 J K"^ mol"^

Moments of Inertia (10"*^ kg m^)

99.38 183.42 232.28

Vibrational Frequencies (cm"-'-) -109.3 266.0 308.3 496.2 782.7

871.4 960.5 1151.3 1157.2 1379.3

1385.4 1413.6 1842.9 2853.1 2964.4

Bond Distances (10"-^° m) Bond Angles {°) Dihedral Angles (°) C1-C2 1.529

C2-01 1.181 C1-C2-01 121.33

C2-H1 1.092 01-C2-H1 122.22 C1-C2-H1 180.00

C1-H2 1.079 C2-C1-H2 112.87 01-C1-H2 180.00

Cl-Fl 1.335 C2-C1-F1 108.92 H2-C1-F1 120.90

C1-F2 1.335 C2-C1-F2 108.92 H2-C1-F2 -120.90

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

- J K"^ mol"^ -! j kJ mol"^

T/K S° H°-H°{T_) AHo

f AG? Log Kf

150 44 250 . 93 -7 95 -520 . 15 -499 . 57 173 96

200 50 . 58 264 . 57 -5 57 -521 .87 -492 . 46 128 62

250 56 . 80 276 . 52 -2 88 -523 . 54 -484 . 97 101 32

300 63 .27 287 . 44 0 12 -525 . 09 -477 . 14 83 08

350 69 . 76 297 . 69 3 44 -526 . 51 -469 . 11 70 01

400 76 . 02 307 . 42 7 09 -527 . 81 -460 . 87 60 18

450 81 .89 316 . 71 11 04 -528 . 94 -452 . 50 52 52

500 87 .28 325 . 62 15 27 -529 . 99 -443 . 96 46 38

600 96 . 66 342 . 39 24 48 -531 . 79 -426 . 64 37 14

700 104 . 34 357 . 89 34 54 -533 . 17 -409 . 07 30 52

800 110 . 65 372 . 25 45 30 -534 .30 -391

.

25 25 55

900 115 . 86 385 . 59 56 64 -535 . 13 -373 . 30 21 66

1000 120 . 19 398 . 03 68 45 -535 . 84 -355 . 22 18 55

1100 123 . 82 409 . 66 80 65 -536 .35 -337 . 02 16 00

1200 126 . 87 420 . 57 93 19 -536 . 77 -318 . 78 13 87

1300 129 .45 430 . 83 106 01 -537 . 10 -300 . 45 12 07

1400 131 . 64 440 . 50 119 07 -537 . 31 -282 . 09 10 52

1500 133 . 52 449 . 65 132 33 -537 .48 -263 . 72 9 18

1600 135 . 14 458 . 32 145 77 -537 .60 -245 . 39 8 01

1700 136 . 54 466 . 56 159 35 -537 . 64 -227 . 02 6 98

1800 137 . 75 474 . 40 173 07 -537 .69 -208 . 70 6 06

1900 138 . 81 481

.

88 186 90 -537 .69 -190 . 46 5 24

2000 139 . 74 489 . 02 200 82 -537 . 64 -172 . 21 4 50

2100 140 . 56 495 . 86 214 84 -537 . 56 -154 . 05 3 83

2200 141 .29 502 . 41 228 93 -537 .48 -135 . 94 3 23

2300 141 . 93 508 . 71 243 09 -537 . 35 -117 . 86 2 68

2400 142 . 51 514 . 76 257 32 -537 .23 -99. 87 2 17

2500 143 . 02 520 . 59 271 59 -537 . 06 -81

.

96 1 71

2600 143 .49 526 . 21 285 92 -536 . 89 -64 . 06 1 29

2700 143 . 90 531. 63 300 29 -536 . 72 -46 . 19 0 89

2800 144 .28 536 . 87 314 70 -536 . 51 -28 . 37 0 53

2900 144 . 62 541

.

94 329 14 -536 .31 -10 . 54 0 19

3000 - 144 . 93 546 . 85 343 62 -536 . 10 7 . 32 -0 13
614 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF3-CHO fcco cis

AH|(0 K) = -767.26 kJ mol"^

AHf(298.15 K) = -774.30 kJ mol"^

S°(298.15 K) = 318.28 J K'^ mol"^

Moments of Inertia (10"*^ kg m^)

148.91 273.72 278.81

Vibrational Frequencies (cm"^)

73.1 251.9 309.1 423.1 514.7

515.0 688.4 834.9 981.5 1225.3

1226.0 1329.8 1382.7 1857.1 2893.7

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.524

C2-01 1.178 C1-C2-01 122.16

C2-H1 1.089 01-C2-H1 123.85 C1-C2-H1 180.00

Cl-Fl 1.306 C2-C1-F1 112.37 01-Cl-Fl 0.00

C1-F2 1.321 C2-C1-F2 109.29 F1-C1-F2 121.11

C1-F3 1.321 C2-C1-F3 109.29 F1-C1-F3 -121.11

Enthalpy Reference Temperature = Tj. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

I- J K"^ mol"^ -I I kJ mol"^

T/K %S° H°-H°(T) AH0

f AG? Log Kf

150 57 271 . 58 -10 32 -771 . 74 -748 . 85 260 76

200 65 .83 289 . 25 -7 22 -772 . 78 -741 . 07 193 54

250 73 .49 304 . 75 -3 72 -773 . 62 -733 . 08 153 16

300 80 . 68 318 . 78 0 15 -774 .33 -724 . 92 126 21

350 87 .39 331

.

72 4 37 -774 . 92 -716 . 68 106 95

400 93 . 55 343 . 80 8 91 -775 .42 -708 . 35 92 50

450 99 . 14 355 . 41 13 74 -775 .80 -700 . 11 81 26

500 104 . 15 366 . 44 18 84 -776 . 09 -691

.

87 72 27

600 112 .60 386 . 82 29 71 -776 .42 -675 . 38 58 80

700 119 .27 405 . 25 41 33 -776 . 59 -658 . 94 49 17

800 124 . 59 422 . 03 53 55 -776 . 55 -642 . 54 41 95

900 128 . 85 437 . 41 66 24 -776 . 38 -626 . 14 36 34

1000 132 . 31 451

.

58 79 31 -776 . 13 -609. 82 31 85

1100 135 . 14 464 . 70 92 70 -775 . 80 -593 . 50 28 18

1200 137 .48 476 . 91 106 34 -775 .42 -577 . 22 25 12

1300 139 .43 488 . 31 120 19 -774 . 96 -560 . 99 22 54

1400 141 . 07 499 . 00 134 22 -774 . 54 -544 . 84 20 33

1500 142 .45 509 . 07 148 41 -774 . 04 -528 . 73 18 41

1600 143 . 63 518 . 56 162 72 -773 . 54 -512 . 71 16 74

1700 144 .65 527 . 54 177 13 -772 . 99 -496 . 77 15 26

1800 145 . 52 536 . 07 191 65 -772 .45 -480 . 91 13 95

1900 146 . 28 544 . 18 206 24 -771 . 91 -465 . 14 12 79

2000 146 . 94 551 . 91 220 91 -771 . 32 -449 . 45 11 74

2100 147 . 52 559 . 29 235 63 -770 . 73 -433 . 92 10 79

2200 148 . 04 566 . 36 250 41 -770 . 11 -418 . 44 9 93

2300 148 .49 573 . 13 265 24 -769 .48 -403 . 09 9 15

2400 148 . 89 579 . 63 280 11 -768 .85 -387 . 81 8 44

2500 149 . 26 585 . 89 295 02 -768 .22 -372 . 67 7 79

2600 149 . 58 591

.

91 309 96 -767 . 60 -357 . 61 7 18

2700 149 .87 597 . 72 324 94 -766 . 93 -342 . 59 6 63

2800 150 . 13 603 . 32 339 94 -766 .26 -327 . 61 6 11

2900 150 . 37 608 . 74 354 97 -765 . 59 -312 . 71 5 63

3000 150 . 58 613 . 98 370 02 -764 . 92 -297 . 78 5 18
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 615

CHjF-CO* fcco trans

AH?{0 K) = -166.94 kJ mol"^

AHf{298.15 K) = -172.61 kJ mol"^

S°(298.15 K) = 286.12 J K'^ mol"^

Moments of Inertia (10""^^ kg m^)

16.81 192.13 203.68

Vibrational Frequencies (cm"-'-) 137.9 312.2 496.2 841.8 895.3

1100.8 1215.2 1346.3 1450.6 1924.3

2907.3 2963.7

Bond Distances (10'-'-° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.518

C2-01 1.163 C1-C2-01 126.50

Cl-Fl 1.354 C2-C1-F1 110.32 01-Cl-Fl 180.00

Cl-Hl 1.082 C2-C1-H1 108.76 Fl-Cl-Hl 120.13

C1-H2 1.082 C2-C1-H2 108.76 F1-C1-H2 -120.13

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^^ = 1 atmosphere

|- J K'^ mol"^ -| I kJ mol"^

T/K Co S° H°-H AHo

f AG? Log Kf

150 46 . 'Is 249 . 59 -8 . 03 -170 . 04 -162 . 92 56 .. 73

200 51

.

81 263 . 63 -5 . 57 -170 . 96 -160 . 41 41 ., 89

250 56 . 95 275 . 65 -2 . 86 -171 . 84 -157 . 69 32 ,. 95

300 62 . 26 286 . 50 0 . 11 -172 .63 -154 . 85 26 ,, 96

350 67 . 50 296 . 49 3 . 35 -173 . 34 -151

.

88 22 ..67

400 72 . 46 305 . 83 6 . 84 -174 . 01 -148 . 82 19 ..44

450 77 . 02 314 . 63 10 . 57 -174 . 56 -145 . 73 16 ,. 91

500 81 . 14 322 . 95 14 . 52 -175 . 06 -142 . 55 14 .. 89

600 88 . 15 338 . 39 23 . 00 -175 . 85 -136 . 02 11 .. 84

700 93 . 82 352 . 41 32 . 10 -176 .44 -129 . 37 9 ..65

800 98 . 47 365 . 25 41 . 74 -176 . 90 -122 . 59 8 .. 00

900 102 . 33 377 . 07 51

.

79 -177 .28 -115 . 73 6 .. 72

1000 105 . 57 388 . 02 62 . 21 -177 . 53 -108 . 83 5 ..68

1100 108 . 32 398 . 22 72 . 92 -177 . 74 -101

.

80 4 .. 83

1200 110 . 65 407 . 74 83 . 89 -177 . 90 -94 . 77 4 .. 12

1300 112 . 64 416 . 68 95 . 07 -177 . 99 -87 . 70 3 .. 52

1400 114 . 34 425 . 09 106 . 43 -178 . 07 -80 . 58 3 ,. 01

1500 115 . 81 433 . 03 117 . 96 -178 . 15 -73 . 47 2 .. 56

1600 117 . 09 440 . 54 129 . 62 -178 .20 -66 . 36 2 .. 17

1700 118 . 20 447 . 68 141

.

40 -178 .20 -59 . 29 1

.

. 82

1800 119 . 16 454 . 55 153 . 28 -178 .20 -52 . 38 1 ,. 52

1900 120 . 01 461 . 18 165 . 25 -178 .20 -45 . 69 1

,

. 26

2000 120 . 76 467 . 51 177 . 30 -178 .20 -39 . 08 1

,

. 02

2100 121 . 42 473 . 57 189 . 42 -178 . 15 -32 . 55 0 . 81

2200 122 . 01 479 . 39 201 . 60 -178 . 11 -26 . 07 0 . 62

2300 122 . 53 484 . 97 213 . 84 -178 . 07 -19 . 71 0 .45

2400 122 . 99 490 . 34 226 . 12 -178 . 03 -13 . 39 0 .29

2500 123 . 41 495 . 51 238 . 45 -177 . 95 -7 . 15 0 . 15

2600 123 . 79 500 . 49 250 . 82 -177 . 86 -1 . 00 0 . 02

2700 124 . 13 505 . 31 263 . 22 -177 . 78 5 . 15 -0 . 10

2800 124 . 44 509 . 96 275 . 66 -177 . 69 11

.

21 -0 .21

2900 124 . 72 514 . 46 288 . 13 -177 . 61 17 . 28 -0 .31

3000 124 . 97 518 . 81 300 . 62 -177 . 53 23 . 39 -0 .41
616 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CHjF-CO* fcco gauche

AH?(0 K) = -164.01 kJ mol"^

AHf(298.15 K) = -169.85 kJ mol"^

S°(298.15 K) = 285.88 J K"^ mol"^

Moments of Inertia (10"'*^ kg m^)

36.25 139.26 170.32

Vibrational Frequencies (cm"-'-) 180.3 228.0 671.5

1081.2 1240.7 1370.1

2914.0 2961.7

Bond Distances (10'^° m)

C1-C2 1.513

C2-01 1.160

Cl-Fl 1.356

Cl-Hl 1.082

C1-H2 1.082

809.3 894.7

1453.3 1943.7

Bond Angles (°) Dihedral Angles

C1-C2-01 129.10

C2-C1-F1 111.84 01-Cl-Fl 0.00

C2-C1-H1 108.40 Fl-Cl-Hl 120.87

C2-C1-H2 108.38 F1-C1-H2 -120.84

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p^ = 1 atmosphere

i- J K"^ mol"^ -

I I kJ mol"^

T/K nO S° H°-H AHo

f AG? Log Kf

150 46 :i7 250 . 10 -7 . 86 -167 . 11 -160 . 08 55 74

200 50 . 74 264 . 00 -5 . 44 -168 . 07 -157 . 61 41 16

250 55 . 51 275 . 76 -2 . 79 -168 . 99 -154 . 89 32 36

300 60 . 67 286 . 26 0 . 11 -169 . 87 -152 . 05 26 47

350 65 . 92 295 . 92 3 . 27 -170 .67 -149 . 03 22 24

400 71 . 02 304 . 98 6 . 69 -171 .42 -145 . 90 19 05

450 75 .82 313 . 6 3 10 . 36 -172 . 05 -142 . 72 16 57

500 80 . 26 321

.

85 14 . 25 -172 .59 -139 . 49 14 57

600 88 . 01 337 . 19 22 . 66 -173 .43 -132 . 88 11 57

700 94 .35 351

.

25 31 . 78 -174 . 01 -126 . 11 9 41

800 99 . 51 364 . 19 41 . 47 -174 .43 -119 . 29 7 79

900 103 . 71 376 . 15 51

.

63 -174 .68 -112 . 34 6 52

1000 107 . 16 387 . 26 62 . 18 -174 . 81 -105 . 31 5 50

1100 110 . 00 397 . 61 73 . 04 -174 .85 -98 . 28 4 67

1200 112 .35 407 . 29 84 . 17 -174 . 85 -91

.

17 3 97

1300 114 . 32 416 . 36 95 . 51 -174 .81 -84 . 06 3 38

1400 115 . 98 424 . 89 107 . 04 -174 . 72 -76 . 94 2 87

1500 117 .38 432 . 94 118 . 72 -174 .64 -69 . 83 2 43

1600 118 . 58 440 . 55 130 . 53 -174 . 51 -62 . 72 2 05

1700 119 .61 447 . 78 142 . 45 -174 .39 -55 . 65 1 71

1800 120 .49 454 . 64 154 . 47 -174 .26 -48 . 62 1 41

1900 121 . 26 461

.

17 166 . 57 -174 . 14 -41 . 63 1 14

2000 121 . 93 467 . 41 178 . 74 -173 . 97 -34 . 69 0 91

2100 122 . 52 473 . 37 190 . 98 -173 . 85 -27 . 78 0 69

2200 123 . 04 479 . 08 203 . 27 -173 .68 -21

.

00 0 50

2300 123 . 50 484 . 56 215 . 61 -173 . 55 -14 . 23 0 32

2400 123 . 90 489 . 83 227 . 99 -173 .38 -7 . 53 0 16

2500 124 .27 494 . 89 240 . 42 -173 . 22 -0 . 92 0 02

2600 124 .60 499 . 77 252 . 87 -173 . 09 5 . 69 -0 11

2700 124 .89 504 . 48 265 . 35 -172 . 92 12 . 26 -0 24

2800 125 . 16 509 . 03 277 . 87 -172 . 76 18 . 79 -0 35

2900 125 .39 513 . 51 290 . 41 -172 . 59 25 . 06 -0 45

3000 125 . 61 517 . 87 302 . 97 -172 .42 31. 30 -0 55
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 617

CHFj-CO* hcco cis

AHf(0 K) = -372.92 kJ mol"^

AHf(298.15 K) = -377.35 kJ mol"^

S°(298.15 K) = 310.33 J K'^ mol"^

Moments of Inertia (10'^^ kg m^)

82.99 209.20 274.09

Vibrational Frequencies (cm"-'-) 53.8 366.7 407.3

946.7 1133.6 1150.5

1940.6 2976.4

Bond Distances (10'-'^° m)

C1-C2 1.539

C2-01 1.158

Cl-Hl 1.078

Cl-Fl 1.330

C1-F2 1.330

410.6 590.0

1319.5 1366.4

Bond Angles (°) Dihedral Angles (°) C1-C2-01 126.14

C2-C1-H1 112.14 01-Cl-Hl 0.00

C2-C1-F1 108.00 Hl-Cl-Fl 121.17

C2-C1-F2 108.00 H1-C1-F2 -121.17

Enthalpy Reference Temperature = T^. = 298.15 K

Standard State Pressure = p = 1 atmosphere

1

1 J K

^ mol" -1- _ 1 1

1 kJ mol"^ 11 T/K (c| S'

^

1 H°-H°(T_) AHo

f AG° Log Kf

150 50 270 . 46 -8 .. 79 -375 .47 -366 . 10 127 ..47

200 56 .46 285 . 83 -6 .. 09 -376 . 14 -362 . 84 94 .. 76

250 61 . 86 299 . 01 -3 .. 11 -376 . 77 -359 . 45 75 .. 10

300 67 . 12 310 . 75 0 ., 13 -377 .35 -355 . 97 61 .. 98

350 72 . 19 321 . 71 3 .. 62 -377 . 90 -352 . 50 52 ..60

400 76 . 96 332 . 08 7 ..36 -378 . 36 -349 . 03 45 .. 58

450 81 .38 341 .80 11

.

,33 -378 . 78 -345 . 56 40 .. 11

500 85 .41 350 . 95 15 .. 51 -379 . 15 -342 . 04 35 .. 73

600 92 . 32 367 . 82 24 ..42 -379 . 78 -335 . 01 29 ,. 16

700 97 . 91 383 . 07 33 .. 95 -380 . 24 -327 . 94 24 ..47

800 102 .42 396 . 97 43 .. 99 -380 . 62 -320 . 87 20 .. 95

900 106 . 11 409 . 72 54 ..43 -380 . 91 -313 . 76 18 ..21

1000 109 . 14 421 .48 65 .,20 -381 . 12 -306 . 65 16 .. 02

1100 111 . 64 432 .39 76 ..25 -381 . 33 -299 . 49 14 ..22

1200 113 . 73 442 .55 87 .. 52 -381 . 50 -292 . 34 12 .. 72

1300 115 .48 452 . 04 98 .. 99 -381 . 62 -285 . 18 11 ..46

1400 116 . 97 460 . 96 110 .. 62 -381 . 71 -278 . 07 10 ,.37

1500 118 .23 469 .36 122 ..38 -381 . 79 -270 . 96 9 ..43

1600 119 . 32 477 .29 134 .. 26 -381 . 87 -263 . 88 8 ,.61

1700 120 .25 484 . 80 146 .. 24 -381 . 96 -256 . 86 7 ,.89

1800 121 . 06 491 . 93 158 ..31 -382 . 00 -249 . 87 7 ,.25

1900 121 . 76 498 . 72 170 ..45 -382 . 00 -242 . 92 6 ,.68

2000 122 . 38 505 . 19 182 .. 66 -382 . 04 -236 . 06 6 ,. 17

2100 122 . 92 511 . 37 194 .. 93 -382 . 04 -229 . 28 5 ,. 70

2200 123 .40 517 .29 207 ..25 -382 . 04 -222 . 59 5 ,. 28

2300 123 . 83 522 . 97 219 .. 61 -382 . 04 -215 . 98 4 ,. 90

2400 124 .20 528 .43 232 ,. 01 -382 . 04 -209 . 41 4 ,. 56

2500 124 . 54 533 .67 244 ,.45 -382 . 00 -202 . 92 4 ,. 24

2600 124 . 85 538 . 72 256 .. 92 -381 . 96 -196 . 48 3 .. 95

2700 125 . 12 543 .60 269 ..42 -381 . 92 -190 . 12 3 . 68

2800 125 .37 548 . 30 281

,

. 95 -381 .87 -183 . 76 3 ,.43

2900 125 . 59 552 . 85 294 ,. 50 -381 . 79 -177 . 40 3 .20

3000 125 .80 557 .25 307 ,. 06 -381 .75 -171 . 04 2 . 98
618 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

CF3-C0* fcco cis

AH?{0 K) = -606.51 kJ mol"^

AH^(298.15 K) = -609.88 kJ mol"^

S°(298.15 K) = 323.59 J K"^ mol"^

Moments of Inertia (10'^^ kg m^)

144.25 275.19 275.38

Vibrational Frequencies (cm"-"^) 58.1 234.3 388.6 411.9 523.5

531.4 667.3 795.4 1203.3 1240.3

1269.0 1959.9

Bond Distances (10"^° m) Bond Angles (°) Dihedral Angles (°) C1-C2 1.541

C2-01 1.155 C1-C2-01 126.31

Cl-Fl 1.310 C2-C1-F1 112.19 01-Cl-Fl 0.00

C1-F2 1.313 C2-C1-F2 108.91 F1-C1-F2 120.64

C1-F3 1.313 C2-C1-F3 108.91 F1-C1-F3 -120.64

Enthalpy Reference Temperature = T^. = 2 98.15 K

Standard State Pressure = p^ = 1 atmosphere

1

1 mo 1

1 1

1

1 _ _

i kJ mol"^ 11 T /V 1/ Is. %

D H°-H°(T) AHo

f AG? IT 150 53 277 . 64 -9 ,. 77 -608 . 94 -595 . 22 207 ,. 26

200 62 . 06 294 .65 -6 ,. 86 -609 .40 -590 . 61 154 .25

250 69 . 84 310 . 08 -3 ,. 54 -609 .69 -586 . 09 122 ,.45

300 76 . 67 324 . 08 0 .. 14 -609 . 86 -581

.

58 101

,

.26

350 82 . 65 336 . 94 4 ,. 14 -609 . 99 -577 . 10 86 ,.12

400 87 .87 348 . 83 8 ..41 -609 . 99 -572 . 66 74 ,. 78

450 92 .42 359 . 91 12 .. 93 -609 . 94 -568 . 23 65 ,. 96

500 96 . 37 370 .28 17 .. 66 -609 . 90 -563 . 84 58 ,. 90

600 102 . 82 389 . 17 27 .. 64 -609 .69 -555 . 13 48 ..33

700 107 . 71 406 . 03 38 .. 19 -609 . 44 -546 . 56 40 ,. 78

800 111 .45 421 .21 49 .. 16 -609 . 19 -538 . 02 35 ,, 13

900 114 . 35 434 . 99 60 .. 46 -608 . 94 -529 . 53 30 .. 73

1000 116 . 63 447 .60 72 .. 02 -608 .69 -521

.

08 27 .. 22

1100 118 .43 459 . 19 83 .. 78 -608 . 52 -512 . 67 24 ,. 34

1200 119 . 88 469 . 92 95 .. 70 -608 .31 -504 . 26 21 .. 95

1300 121 . 06 479 . 90 107 .. 75 -608 . 14 -495 . 89 19 ,. 92

1400 122 . 03 489 .21 119 ., 91 -608 . 02 -487 . 56 18 ,, 19

1500 122 . 83 497 . 94 132 .. 15 -607 .89 -479 . 28 16 ..69

1600 123 . 50 506 . 16 144 ,.47 -607 .81 -470 . 99 15 ,.38

1700 124 . 08 513 . 92 156 ..85 -607 .68 -462 . 79 14 .. 22

1800 124 . 56 521 . 26 169 .. 28 -607 .60 -454 . 68 13 ,. 19

1900 124 . 98 528 .23 181

.

. 77 -607 . 52 -446 . 60 12 ..28

2000 125 . 34 534 . 86 194 ,.28 -607 .43 -438 . 61 11

.

.45

2100 125 .65 541 . 19 206 .. 83 -607 . 35 -430 . 70 10 .. 71

2200 125 . 93 547 .23 219 ,,41 -607 .27 -422 . 84 10 .. 04

2300 126 . 16 553 . 02 232 ,, 02 -607 . 18 -415 . 09 9 ..43

2400 126 . 38 558 . 57 244 ,.65 -607 . 06 -407 . 40 8 .. 87

2500 126 . 57 563 . 90 257 ..29 -606 . 97 -399 . 78 8 ..35

2600 126 . 73 569 . 03 269 .. 96 -606 .89 -392 . 25 7 .. 88

2700 126 .89 573 . 97 282 .. 64 -606 . 81 -384 . 76 7 ,.44

2800 127 . 02 578 . 74 295 .. 34 -606 .68 -377 . 27 7 .. 04

2900 127 . 14 583 .35 308 .. 05 -606 .60 -369 . 82 6 ..66

3000 127 . 26 587 .80 320 .. 77 -606 .47 -362 . 33 6 ,.31
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 619

CFj-CO" fcco trans

AH?(0 K)

AHf{298.15 K)

S°(298.15 K)

•606.43 kJ mol"^

611.24 kJ mol"^

306 . 16 J K"^ mol"^

Moments of Inertia (10

142.55 273.76 275.81

kg

Vibrational Frequencies (cm"-'-) -78.5 225.9 320.7

527.3 664.5 792.6

1270.4 1953.8

417 . 8 1199 .

3

518 . 9 1240 .

1

Bond Distances (10

C1-C2 1.546

C2-01 1.155

Cl-Fl 1.311

C1-F2 1.312

C1-F3 1.312

-10 m) Bond Angles (°) C1-C2-01

C2-C1-F1

C2-C1-F2

C2-C1-F3

124 . 97

110 .39

109 . 83

109 . 83

Dihedral Angles ('

01-Cl-Fl 180.00

F1-C1-F2 119.86

F1-C1-F3 -119.86

Enthalpy Reference Temperature

Standard State Pressure == P,

1

1 J mol 1 T/K S°

150 49 . 12 264 ., 78

200 58 ., 05 280 ., 15

250 65 . 93 293 ., 97

300 72 ., 78 306 ., 61

350 78 .. 74 318 ,,29

400 83 ., 94 329 ., 15

450 88 .,46 339 .,31

500 92 .,40 348 ,, 83

600 98 .,80 366 .,27

700 103 .,67 381

.

, 89

800 107 .,39 395 ,, 98

900 110 .,28 408 ,, 81

1000 112 ., 54 420 .. 55

1100 114 ., 34 431

,

. 36

1200 115 .. 78 441 ..37

1300 116 .. 95 450 .. 69

1400 117 .. 91 459 ..39

1500 118 .. 71 467 .. 55

1600 119 ..38 475 .. 24

1700 119 .. 95 482 ,.49

1800 120 ..43 489 ,. 36

1900 120 .. 84 495 ,. 88

2000 121

.

.20 502 ,. 09

2100 121 ,. 51 508 ,. 01

2200 121

,

. 78 513 ,.67

2300 122 ,. 02 519 ,. 09

2400 122 ,. 24 524 .29

2500 122 .42 529 .28

2600 122 . 59 534 . 09

2700 122 . 74 538 . 72

2800 122 . 88 543 . 18

2900 123 . 00 547 . 50

3000 123 . 11 551 . 67

= Tj. = 298.15 K

1 atmosphere

kJ mol"^

H°-H°(T^) AHo

f AG? Log

-9 . 13 -609 .65 -594 . 00 206 ., 84

-6 . 44 -610 . 36 -588 . 69 153 ,, 74

-3 . 34 -610 . 86 -583 . 25 121

.

, 86

0 . 13 -611 .24 -577 . 73 100 ., 59

3 . 92 -611 . 53 -572 . 16 85 .,39

8 . 00 -611 . 78 -566 . 56 73 ., 98

12 . 31 -611 . 95 -560 . 95 65 .. 11

16 . 83 -612 . 08 -555 . 34 58 ., 01

26 . 41 -612 .29 -544 . 00 47 .. 36

36 . 54 -612 .45 -532 . 67 39 .. 75

47 . 10 -612 . 62 -521

.

24 34 .. 03

57 . 99 -612 . 79 -509 . 82 29 .. 59

69 . 14 -612 . 96 -498 . 27 26 ,. 03

80 . 48 -613 . 17 -486 . 72 23 ,. 11

91

.

99 -613 .42 -475 . 09 20 ,. 68

103 . 63 -613 .67 -463 . 42 18 .. 62

115 . 37 -613 . 92 -451

.

70 16 ..85

127 . 21 -614 .21 -439 . 99 15 ,. 32

139 . 11 -614 . 50 -428 . 27 13 ,. 98

151

.

08 -614 . 84 -416 . 52 12 .. 80

163 . 10 -615 . 13 -404 . 80 11

,

. 75

175 . 16 -615 .47 -393 . 13 10 ,. 81

187 . 27 -615 .80 -381

.

46 9 ,. 96

199 . 40 -616 . 14 -369 . 82 9 ,.20

211 . 57 -616 .47 -358 . 23 8 ,. 51

223 . 76 -616 . 81 -346 . 69 7 ,.87

235 . 97 -617 . 10 -335 . 18 7 .29

248 . 20 -617 .43 -323 . 72 6 . 76

260 . 45 -617 . 77 -312 . 25 6 . 27

272 . 72 -618 . 06 -300 . 83 5 . 82

285 . 00 -618 .40 -289 . 41 5 .40

297 . 30 -618 .69 -277 . 98 5 . 01

309 . 60 -619 . 02 -266 . 48 4 . 64

SECTION 5. FLAME INfflBITION CHEMISTRY AND THE SEARCH ... 621

Appendix B. Fluorinated Hydrocarbon Reaction Rate Constants

REACTION A b E note/ reference H

H2/O2 System: Combination, Decomposition

H + H +M = H2 + M 1 . 00E18 -1 00 0 . 1MBA13 9 H2/0/ H2O/0/ CO2/0/

H + H +H2 = 2H2 9 . 20E16 -0 60 0 . 1MBA14 0 H + H +H20 = H2 + H20 6 . 00E19 -1 25 0 . 1MBA141

H + H +C02 = H2 + C02 5 . 49E20 -2 00 0 .

1 MBA14 2 0 + 0 +M =02 + M 1 . 89E13 0 00 -1788

! MBA14 5 86TSA/HAM

H + OH +M = H20 + M 1 . 60E22 -2 00 0 . 1MBA14 3 H20/5/

H + 0 +M = OH + M 6 . 20E16 -0 60 0 . 1MBA144

H20/5/

System: Atom Transfers

0 + OH =02 + H 4 . 00E14 -0 50 0 . ! MBAl 3 2 BIHOW/SMI

0 + H2 = OH + H 5 . 06E04 2 6 7 6290 . ! MBAl 3 3 8 8SUT/MIC

2 OH = 0 + H20 6 . 00E08 1 30 0 . ! MBAl 3

H2 +02 = 20H 1 . 70E13 0 00 47780! MBAl 3 0 71JAC/H0U

OH + H2 = H20 + H 1 . 17E09 1 30 3626 . ! MBAl 31 79C0H/WES

H2/O2 System: Peroxyl and Peroxide

H + 02 + M = H02 + M 3 . 61E17 -0 72 0 . ! MBAl 3

H20/18.6/ C02/4. 2/ H2/2.9/ CO/2

1/ N2/1 3/

H + H02 = H2 + 02 1 . 25E13 0 00 0 . IMBA14 6 0 + H02 = 02 + OH 1 . 40E13 0 00 1073 . ! MBAl 3

OH + H02 = H20 + 02 7 . 50E12 0 00 0 . ! MBAl 3

H202 + M = 20H + M 1 . 30E17 0 00 45500

! MBAl4

H + H02 = 20H 1 . 40E14 0 00 1073 . ! MBAl 3 6 2H02 = H202 + 02 2 . 00E12 0 00 0 . ! MBAl 4

REACTION A b E note/ reference H

C]_ Hydrocarbons: Methane

CH4 + H = CH3 + H2 2 20E04 3 00 8750 . 1MBA0 04 7 3CLA/DOV

CH4 + 0 = CH3 + OH 1 02E09 1 50 8604 . 1MBA0 0 5 86SUT/MIC

CH4 + OH = CH3 + H20 1 60E06 2 10 2460 . 1MBA0 0 5 8 3BAU/CRA

CH4 + 02 = CH3 + H02 7 90E13 0 00 56000

1MBA0 0 3 72SKI/LIF

CH4 + HO2 = CH3 + H202 1 80E11 0 00 18700

! MBA0 0 6 72SKI/LIF

Cj^ Hydrocarbons: Methyl

CH3 + H ( +M) CH4 (+M) 6 00E16 -1 . DO 0 . 1MBA002 84WAR

LOW/8 . 00E26 -3.0 0./ ! (89STE/SMI2)

SRI/0.45 797 979./

H2/2 . 0/ CO/2 . 0/ C02/3. 0/ H20/5 0/

CH3 + H CH2 + H2 9 00E13 0 . 00 15100

1MBA013

CH3 + 0 CH20 + H 8 00E13 0 . 00 0 . 1MBA0 0 9 CH3 + OH CH2 + H20 7 50E06 2 . 00 5000 . 1MBA012

CH3 + OH CHS OH 2 24E40 -8 .20 11673

! 8 7DEA/WES

CH3 + OH CH20H + H 2 64E19 -1 . 80 8068 . ! 8 7DEA/WES

CH3 + OH CH30 + H 5 74E12 -0 .23 13931

! 8 7DEA/WES

CH3 + OH CH2SING + H20 8 90E19 -1 .80 8067 . ! 8 7DEA/WES

CH3 + 02 CH30 + 0 2 05E18 -1 . 57 29229

1MBA008 86TSA/HAM

CH3 + H02 CH30 + OH 2 00E13 0 . 00 0 . !MBA007 86TSA/HAM

CH3 + CH3 (+M) = C2H6 (+M) 9 03E16 -1 .20 654 . IMBAOOl 8 8WAG/WAR

LOW/3 . 18E41 -7.0 2762./ ! (8 8WAG/WAR)

TROE/06041 6927. 132./

H2/2 . 0/ CO/2 . 0/ C02/3. 0/ H20/5 0/
622 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

C]^ Hydrocarbons; Methylene (triplet)

u , u u u

1 MTU 7\ n o n ntri "3 U Uii± J u . u u U CH2+C02 = CH20+C0 1 lOEll 0.00 1000 ! MBA042

CH2+0 = C0+H2 3 00E13 0 . 00 0 IMBA044

CH2+02 = C02+2H 1 60E12 0 . 00 1000 1MBA04 5 CH2+02 = CH20+0 5 00E13 0 . 00 9000 1MBA04 6 CH2+02 = C02+H2 6 90E11 0 . 00 500 IMBA04 7 CH2+02 = C0+H20 1 90E10 0 . 00 -1000 1MBA04 8 CH2+02 = CO+OH+H 8 60E10 0 . 00 -500 1MBA04 9 CH2+02 = HCO+OH 4 30E10 0 . 00 -500 ! MBAO 5 0 CH2+CH3 = C2H4+H 3 00E13 0 . 00 0 1MBA0 72

2CH2 = C2H2+H2 4 00E13 0 . 00 0 1MBA114

Hydrocarbons: Methylene (singlet)

CH2SING+M = CH2+M 1 00E13 0 . 00 0 1MBA10 6 H/0 . 0/

CH2SING+H = CH2+H 2 00E14 0 . 00 0 IMBAlll

CH2SING+02 = CO+OH+H 3 00E13 0 . 00 0 ! MBAl 0 9 CH2SING+H2 = CH3+H 7 00E13 0 . 00 0 IMBAllO

CH2SING+CH4 = 2CH3 4 00E13 0 . 00 0 ! MBAl 07

CH2SING+C2H6 = CH3+C2H5 1 20E14 0 . 00 0 ! MBAl 0 8 Hydrocarbons: Methylidyne

CH2+H = CH+H2 1 00E18 -1 . 56 0 ! MBAO 2

4

CH2+0H = CH+H20 1 13E07 2 . 00 3000 ! MBAO 2

5

CH+02 = HCO+0 3 30E13 0 . 00 0 ! MBAO 2

7

82BER/FLE

CH+0 = CO+H 5 70E13 0 . 00 0 ! MBAO 2

8

83MES/FIL

CH+OH = HCO+H 3 00E13 0 . 00 0 ! MBAO 2

9

CH+C02 = HCO+CO 3 40E12 0 . 00 690 ! MBAO 3 0 82BER/FLE

CH+H20 = CH20+H 1 17E15 -0.75 0 ! MBAO 3

2

CH+CH20 = CH2C0+H 9 46E13 0 . 00 -515 ! MBAO 3 3 88ZAB/FLE

CH+CH2 = C2H2+H 4 00E13 0 . 00 0 ! MBAO 3

5

CH+CH3 = C2H3+H 3 00E13 0 . 00 0 ! MBAO 3

6

CH+CH4 = C2H4+H 6 00E13 0 . 00 0 ! MBAO 3

7

80BUT/FLE

C2H3+CH = CH2+C2H2 5 00E13 0 . 00 0 ! MBAO 8

6

HCCO+CH = C2H2+CO 5 00E13 0 . 00 0 ! MBAl 04

REACTION A b E note/reference H

Oxidized C^^ Hydrocarbons: Methanol

CH30H + H = CH30 + H2 8 00E12 0 . 00 6095 . 90NOR

CH30H + OH = CH30 + H20 1 50E13 0 . 00 5960 . 7 5 BOW

CH30H + CH3 = CH30 + CH4 1 45E + 1 3 . 10 6935 . 9 0 NOR

+ 16

CH30H +02 = CH20H+ H02 2 05E13 0 . 00 44717

.

. 9 0NOR

CH30H + H = CH20H+ H2 3 20E13 0 . 00 6095 . 9 0NOR

CH30H +0 = CH20H+ OH 3 88E05 2 . 50 3080 . . 9 0NOR

CH30H + OH = CH20H+ H20 1 50E13 0 . 00 5960 . . 7 5BOW

CH30H + H02 = CH20H+ H202 3 98E13 0 . 00 19400

.

. 9 0NOR

CH30H + CH3 = CH20H+ CH4 3 19E01 3 . 17 7172 . . 90NOR

Oxidized C^^ Hydrocarbons: Methoxy, Hydroxyme thy1 CH30+M = CH20+H+M 1 00E14 0 . 00 25000

.

.MBAO 14

CH30+02 = CH20+H02 6 30E10 0 . 00 2600 . .MBAO 2

2

CH30+H = CH20+H2 2 00E13 0 . 00 0 . .MBAO 16

CH30+0 = CH20+0H 1 00E13 0 . 00 0 . .MBAO 2 0 CH30+0H = CH20+H20 1 00E13 0 . 00 0 . .MBAO 18

CH20H+M = CH20+H+M 1 00E14 0 . 00 25000

.

.MBA015

CH20H+02 = CH20+H02 1 48E13 0 . 00 1500 . .MBAO 2

3

CH20H+H = CH20+H2 2 00E13 0 . 00 0 . MBAO 17

CH20H+0 = CH20+0H 1 00E13 0 . 00 0 . .MBAO 21

CH20H+0H = CH20+H20 1 00E13 0 . 00 0 . .MBAO 19
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 623

Oxidized. C Hydrocarbons : Formaldehyde

CH20+M HCO+H+M 3 . 31E16 0 . 00 81000 IMBA05 3 8 0DEA/ JOH

CH20+H HC0+H2 2 . 19E08 1 . 77 3000 1MBA052 86TSA/HAM

CH20+0 HCO+OH 1

80E13 0 . 00 3080 1MBA0 54 80KLE/SOK

CH20+0H HC0+H20 3 . 43E09 1 . 18 -447 1MBA051 86TSA/HAM

Oxidized Hydrocarbons: Formyl

HCO + M CO + H + M 2 . 50E14 0 . 00 16802 1MBA0 5 6 8 4 WAR

CO/1 . 9/ H2/1 .9/ CH4/2.8/ C02/3. 0/ H20/5

.0/

HCO + H CO + H2 1

19E13 0 . 25 0 ! MBAO 5 7 HCO + 0 CO + OH 3 . 00E13 0 . 00 0 ! MBAO 5 8 8 4 WAR

HCO + 0 C02 + H 3 . 00E13 0 . 00 0 ! MBAO 5 9 8 4 WAR

ur^o _i_ r\xj CO + H20 1

00E14 U . U U U I PitsAU 3 D

0 A T^TIVT /TaT7\ HCO + 02 CO + H02 3 . 30E13 -0 .40 0 IMBA06 0 81VEY/LES

Oxidized C^^ Hydrocarbons : Carbon Monoxide

CO + 0+M C02 + M 6 . 17E14 0 . 00 3000 IMBA061 86TSA/HAM

CO + OH C02 + H 1

51E07 1 . 3 0 -758 1MBA062

CO + 02 C02 + 0 1

60E13 0 . 00 41000 1MBA063 71BRA/BEL

CO + H02 C02 + OH 5 . 80E13 0 . 00 22934 !MBA064 7 7ATR/BAL

REACTION A b E note/reference H

Hydrocarbons: Ethane

C2H6 + H C2H5 + H2 5 . 40E02 3 . 50 5210 , 1MBA066 73CAL/DOV

C2H6 + 0 C2H5 + OH 3 . 00E07 u u 5115 !MBA06 7 8 4 WAR

C2H6 + OH C2H5 + H20 8 . 70E09 1

05 1810 1MBA06 8 8 3TUL/RAV

C2H6 + CH3 C2H5 + CH4 5 . 50E-1 4 . 00 8300 1MBA06 5 7 3CLA/DOV

HydrocarbonsEthyl

C2H5+H CH3+CH3 1

00E14 0 . 00 0 , ! MBAO 74

C2H5 + 0 = CH20 + CH3 1

60E13 0 . 0 , 186TSA/HAM

C2H5+02 C2H4+H02 2 . 56E19 -2 . 77 19 7 7 I90BOZ/DEA

Cj Hydrocarbons

Ethylene

C2H4+M C2H2+H2+M 1

50E15 0 . 00 558001MBA12 8 83KIE/KAP

C2H4+M C2H3+H+M 1

40E16 0 . 00 82360 1MBA12 9 C2H4+H (+M) C2H5 (+M) 8 . 40E08 1

5 990 186TSA/HAM

LOW/ 6

37E27 -2.8 -54./ ! MBAO 7

H2/2.0/ CO/2.0/ CO2/3.0/ H20/5 .0/ ! MBAO 7

C2H4+H C2H3+H2 1

10E14 0 . 00 8500 1MBA06 9 73PEE/MAH

C2H4+0 CH3+HC0 160E09 1 .20 746 , ! MBAO 7 0 84W7m

C2H4+OH C2H3+H20 4 . 50E06 2 . 00 2850 ! fit ref

C2 Hydrocarbons: Vinyl

C2H3+H C2H2+H2 1 . 20E13 0 . 00 0 ! 92BAU/COB

C2H3+OH C2H2+H20 5 . 00E12 0 . 00 0 ! MBAO 8 3 C2H3+CH2 C2H2+CH3 3 . 00E13 0 . 00 0 1MBA084

C2H3+02 CH20+HC0 1 . 05E38 - 8 . 22 7030 ! 92WES .a/s

C2H3+02 CH20+HC0 4 . 48E26 -4 . 55 5480 192WES (direct

C2H3+0 CH2C0+H 3 . 00E13 0 . 00 0 1MBA081 8 4 WAR

Cj Hydrocarbons

Acetylene

C2H2+H (+M) C2H3 (+M) 5 . 54E12 0 . 00 2410 ! MBAO 7

76PAY/STI

LOW/ 2

67E27 -3.5 2410./

H2/2.0/ CO/2.0/ CO2/3.0/ H20/5 .0/

C2H2+OH HCCOH+H 5 . 04E05 2 .30 13500 IMBA088

C2H2+OH CH2C0+H 2 . 18E-4 4 .50 -1000 ! MBAO 8

C2H2+OH CH3+C0 4 . 83E-4 4 . 00 -2000 1MBA090

C2H2+0 CH2+C0 1 . 02E07 2 . 00 1900 1MBA076

C2H2+0 HCCO+H 1 . 02E07 2 . 00 1900 1MBA077

C2H2+02 HCCO+OH 2 . 00E08 1 . 50 30100 1MBA12 6
624 SECTION 5. FLAME INHmiTIGN CHEMISTRY AND THE SEARCH

Oxidized Hydrocarbons: Ketene, HCCOH

HCCOH+H CH2C0+H 1 . 00E13 0 .. 00 0. IMBA091

CH2C0+H CH3+C0 1 . 13E13 0 ..00 3428. IMBA094

CH2C0+H HCC0+H2 5 . 00E13 0 ..00 8000. IMBA095

CH2C0+0 C02+CH2 1 . 75E12 0 ..00 1350. 1MBA093

CH2C0+0 HCCO+OH 1 . 00E13 0 ..00 8 000. 1MBA0 96

CH2C0+0H HCC0+H20 7 . 50E12 0 ..00 2000. IMBA097

CH2C0 (+M) CH2+C0 (+M) 3 . 00E14 0 ..00 7096!0. 1MBA098

LOW/

3

; . 60E15 0.0 59270./

Oxidized Hydrocarbons: Ketyl

HCCO+H CH2SING+C0 1 . 00E14 0 .. 00 0. IMBAlOl

HCCO+0 H+2C0 1 . 00E14 0 .. 00 0. 1MBA102

HCCO+02 2C0+0H 1 . 60E12 0 ,.00 8 54. 1MBA10 3 2HCC0 C2H2+2CO 1 . 00E13 0 ,. 00 0. IMBA10 5 HCC0+CH2 C2H3+CO 3 . 00E13 0 .. 00 0. 1MBA115

REACTION A b E note/ reference H

H/F/0 Chemistry

HF + M H + F 3 .,12E13 0. 00 99320

.

1 81BAU/DUX+134

H2 + F H + HF 2 ..56E12 0. 50 650 . !fit 8 9STE/BRU -31

OH + F 0 + HF 2 ..00E13 0. U U U . H+HN03 -33

H02 + F 02 + HF 2 ..89E12 0. 50 0 . !fit 8 3WAL/WAG -86

H20 + F OH + HF 1 .30E09 1. 50 0 . !fit 8 3WAL/WAG -16

H202 + F H02 + HF 1 .73E12 0. 50 0 . !fit 8 3WAL/WAG -48

Fluoromethanes

:

Thermal Decomposition

CH2SING + HF CH3F 1

,

.91E23 -3. 62 1780 . !fit 73SCH/WAG -88

CHF + H2 CH3F 2 ..25E17 -2. 85 13000

.

!RRKM BAC

CH2F + H CH3F 3 .03E21 -3. 38 3460 . !RRKM BAC -97

CHF + HF CH2F2 3 .64E24 -4. 26 4060 . ! RRKM BAC -71

CF2 + H2 CH2F2 1

.

.70E06 -0. 71 40900

.

IRRKM BAC

CHF2 + H CH2F2 2 .75E06 -0. 32 7690 . ! RRKM BAC -102

CHF3 + M CF2 + HF 2 .,39E30 -4. 00 69050

.

!fit 91HID/NAK + 61

CF3 + F CF4 3 ,.60E15 -1. 00 0 . !fit 8 6PLU/RYA- 130

Fluoromethanes

:

Activated Decomposition

CH2F + H CH2SING + HF 8 .19E15 -0. 63 505 . !RRKM BAC -8

CH2F + H CHF + H2 5 .21E08 1. 16 1000 . !RRKM BAC

CH2SING + HF CHF + H2 2 .08E07 1. 27 8330 . ! RRKM BAC -14

CHF2 + H CHF + HF 1

.

.49E14 -0. 11 101

.

! RRKM BAC

CHF2 + H CF2 + H2 5 .50E03 2. 42 -420 . ! RRKM BAC

CHF + HF CF2 + H2 5 ,.77E06 1. 35 17900

.

! RRKM BAC

CF3 + H CF2 + HF 5 ,.50E13 0. 00 0 . IRRKM 8 9TSA/MCF -46
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 625

FluoromethanesH Atom Abstraction by H, 0, F, OH, H02

CH3F + H - CH2F + H2 2 . 70E03 3 . 00 5300 . !fit 7 SWPc; /nPH' - 6 CH2F2 + H CHF2 + H2 1 . 65E03 3 . 00 5600 . ! fit 7 9P TN /'nQ'\7 - CHF3 + H CF3 + H2 1 . 70E04 3 . 00 8500 . ! fit 7 R ART /RPT.

CH3F + 0 — CH2F + OH 6 . 50E07 1 . 50 7000 . !fit 6 7PAR/AZA

CH2F2 + 0 CHF2 + OH 2 . 25E07 1 . 50 6100 . ! fit 6 8PAR /NATi

CHF3 + 0 CF3 + OH 1 . 00E08 1 . 50 9250 . ! fit 7 RiTOTT/T.FR

CH4 + F — CH3 + HF 5 . 90E12 0 . 50 450 . !fit 71WAG/WAR -31

CH3F + F CH2F + HF 1 . 35E14 0 . 00 1200 . !fit -37

(_rlz r A + cr i_n.r z + rir . 00E13 0 . UU IB b U . [fit -35

CHF3 + F CF3 + HF 4 . 50E13 0 . 00 3700 . !fit -27

CH3F + OH CH2F + H20 2 . 60E08 1 . 50 2940 . 91C0H/WES -20

CH2F2 + OH CHF2 + H20 2 . 80E07 1 . 70 2540 . 91C0H/WES -10

CHF3 + OH CF3 + H20 5 . 77E06 1 . 80 4292 . 87COH/BEN - 10

CH2F + H202 CH3F + HO2 1 . 20E10 0 . 00 -600 . ! ref -11

CHF2 + H202 CH2F2 + HO2 1 . 20E10 0 . 00 -600 . ! ref -12

CF3 + H202 CHF3 + H02 1 . 2 OEIO 0 . 00 -600 . ! ref - 2

1

REACTION A b E note/reference H

FluoromethanesH Atom Abstraction by CH3

,

C2H3, C2H5

CH3F + CH3 CH2F + CH4 1 . 50E11 0 . 00 11400

.

! 65PRI/BRY - 6

CH2F2 + CH3 CHF2 + CH4 8 . 70E10 0 . 00 T n o n n X U z U U . 165PRI/BRY chk - 2 CF3 + CH4 CHF3 + CH3 8 . 34E11 0 . 00 10920

.

! 78ART/BEL - 4 CH3F + C2H3 CH2F + C2H4 1 . 50E11 0 . 00 10300

.

! ref E*0 . 90 - 8 CH2F2 + C2H3 CHF2 + C2H4 9 . OOEIO 0 . 00 9200 . !ref E*0 . 90 - 6 CF3 + C2H4 CHF3 + C2H3 8 . OOEll 0 . 00 12000

.

! ref E*l . 1 - 2 CH3F + C2H5 CH2F + C2H6 1 . 50E11 0 . 00 16300

.

! ref E*l .43 U CH2F2 + C2H5 CHF2 + C2H6 9 . OOEIO 0 . 00 14600

.

! ref E*l . 43 + 2 CF3 + C2H6 CHF3 + C2H5 7 . 59E11 0 . 00 7980 . ! 78ART/BEL -10

FluoromethanesH Atom Abstraction by Fluoromethyls

CH3F + CF3 CH2F + CHF3 1 . 35E12 0 . 00 11200

.

! 67GIL/QUI - 10

CH2F2 + CH2F CHF2 + CH3F 9 . OOEIO 0 .00 14000

.

! ref CH3 E*1.4 + 1 CH2F2 + CF3 CHF2 + CHF3 7 . 20E11 0 . 00 11200

.

! 67GIL/QUI - 8 Fluoromethanes

F Atom Abstraction

CH3F + H CH3 + HF 2 . 75E14 0 . 00 31400

.

!BAC -28

CH2F2 + H CH2F + HF 5 . 50E13 0 . 00 34100

.

!BAC -16

CHF3 + H CHF2 + HF 8 . 00E13 0 . 00 40300

.

!BAC -4

CF4 + H CF3 + HF 1 . 10E15 0 . 00 44600

.

1 71K0C/M0I -4
626 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoromethyls ; H Atom Abstraction from CH20, CH30H, HCO

CH20 + CH2F = HCO + CH3F 5 . 54E03 2 . 81 8300 . ! ref E*l . 4 - 9 CH20 + CHF2 = HCO + CH2F2 5 . 54E03 2 .81 7800 . !ref E*l . 3 -10

CH20 + CF3 = HCO + CHF3 5 . 54E03 2 .81 4600 . !ref E* . 79 -19

CH30H + CH2F = CH30 + CH3F 1 . 44E01 3 . 10 9800 . ! ref E*l . 4 + 5 CH30H + CHF2 = CH30 + CH2F2 1 . 44E01 3 . 10 9000 . ! ref E*l . 3 + 4 CH30H + CF3 = CH30 + CHF3 1 . 44E01 3 . 10 5500 . ! ref E* . 79 - 4 CH30H + CH2F = CH20H + CH3F 3 . 20E01 3 .20 10000

.

! ref E*l . 4 - 3 + (_rir Z = (_rlzUrl + ^riZ r Z -5 . Z\J tjU ± . Z U y J u u . ! ref E* 1 . 3 - 5 CH30H + CF3 = CH20H + CHF3 3 . 20E01 3 . 20 5700 . ! ref E* . 79 - 13

HCO + CH2F = CO + CH3F 0 . 90E14 0 . 00 0 . ! ref

HCO + CHF2 = CO + CH2F2 0 . 90E14 0 . 00 0 . ! ref

HCO + CF3 = CO + CHF3 0 . 90E14 0 . 00 0 . ! ref -91

HCO + CH2F — v_n ^ »w w _|_ 9 7 DPI n n n n ! est

HCO + CHF2 = CHFCO + HF 2 . 70E13 0 . 00 0 . ! est

HCO + CF3 = CF2C0 + HF 2 . 70E13 0 . 00 0 . ! est

REACTION A b E note/ reference H

Fluoromethyls

:

Oxidation

CH2F + 02 =>CHF:0 + 0 +H 2 . 26E09 1

.

14 19500

.

! ref cf3 +dH

CHF2 + 02 =>CF2 :

0

+ 0 +H 2 . 26E0 9 1

.

14 20500 . ! ref cf3 +dH

CF3 + 02 = CF30 + 0 2 . 26E09 1

.

14 21500

.

IRRKM

CF30 + M = CF2 :0 + F 9.03E26 - 3 . 42 21700

.

!RRKM

CH2F + 0 = CHF:0 + H 5 . 70E13 0 . 00 0 . ! ref ch3 cf3 - 90

CHF2 + 0 = CF2 :

0

+ H 3 . 70E13 0 . 00 0 . !ref ch3 cf3 -103

CF3 + 0 = CF2 :

0

+ F 1 . 90E13 0 . 00 0 . 89TSA/BEL - 80

CH2F + OH = CH20 + HF 2 . 50E13 0 . 00 0 . !ref ch3 -95

CHF2 + OH = CHF:0 + HF 2 . 50E13 0 . 00 0 . ! ref ch3 -105

CF3 + OH = CF2 :

0

+ HF 2 . 00E13 0 . 00 0 . !ref ch3 -114

CH2F + HO2 = CH3F + 02 3 . 00E12 0 . 00 0 . ! ref ch3 -48

CHF2 + HO2 = CH2F2 + 02 3 . 00E12 0 . 0 0 0 . ! ref ch3 - 50

CF3 + HO2 = CHF3 + 02 2 . 00E12 0 . 00 0 . ! ref ch3

CH2F + HO2 =>CHF:0 + 0H+ H 1 . 50E13 0 . 00 0 . ! ref

CHF2 + HO2 = >CF2 : 0 + 0H+ H 1 . b Uiil

3

u

.

(J u u . ! ref

CF3 + HO2 = CF30 + OH 1 . 00E13 0 . 00 0 . ! ref

CF30 + M = CF2:0 + F 9 . 03E26 -3 . 42 2170 0 . ! wt kl9

CF30 + H = CF2:0 + HF 1 . 00E14 0 . 00 0. !nistkl9

CH3 + F = CH2SING + HF 1.62E16 0 . 88 -981

.

! ref -17

CH2F + F = CHF + HF 5 . 00E13 0 . 00 0 . ! ref -46

CHF2 + F = CF2 + HF 3 . 00E13 0 . 00 0 . ! ref -68

CH3 + F = CH2F + H 1.36E12 0 . 39 -265 . ! ref - 9 REACTION A b E note /reference H
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 627

Fluoromethylenes : Oxidation

CHF + 02 CHF:0 + 0 2 00E13 0 . 00 16500

.

! ref cf2 E- 10 -60

CF2 + 02 CF2 :0 + 0 2 00E13 0 00 26500

.

1 7 7KEA/MAT -50

CHF + 0 CO + HF 9 00E13 0 00 0 .

1 90TSA/MCF- 181

CF2 + 0 CF:0 + F 5 00E13 0 00 0 . ! ref chf 169

CHF + OH HCO + HF 5 00E12 0 00 0 . ! ref cf2 -94

CF2 + OH CF:0 + HF 5 00E12 0 00 0 . ! ref cf2 -71

CHF + OH CHF:0 + H 5 00E12 0 00 0 . ! ref cf 2 -77

CF2 + OH CF2 :0 + H 5 00E12 0 00 0 . !fit 78BIO/LAZ - 68

Fluoromethylenes : Destruction

CHF + H20 CH2F + OH 5 00E12 0 00 29000

.

! est Hrxn + 29

CF2 + H20 CHF2 + OH 5 00E12 0 00 51000

.

! est Hrxn + 51

CHF + H CH + HF 2 00E14 0 00 0 . 8 9TSA/MCF

CF2 + H CF + HF 2 00E14 0 00 0 .

1 8 9TSA/MCF

CF4 +CH2SING CHF:CF2 + HF 4 00E13 0 00 5000 . ! ref Ch4

CH4 + CHF C2H4 + HF 4 00E13 0 00 29000

.

! BAC

CH4 + CF2 CH2 : CHF + HF 4 00E13 0 00 80000

.

!BAC

CHF + H20 CH2F + OH 5 00E12 0 00 29000

.

! est dH + 29

CF2 + H20 CHF2 + OH 5 00E12 0 00 51000

.

! est dh + 51

CH20 + CHF HCO + CH2F 1 00E13 0 00 29000 . ! ref CH4

CH20 + CHF CH2C0 + HF 1 00E13 0 00 29000

.

! ref CH4

CH20 + CF2 HCO + CHF2 1 00E13 0 00 80000

.

! ref CH4 + 22

CH20 + CF2 CHFCO + HF 1 00E13 0 00 80000

.

! ref CH4

HCO + CHF CO + CH2F 2 00E13 0 00 29000

.

! ref CH4

HCO + CF2 CO + CHF2 2 00E13 0 00 80000

.

! ref CH4

Fluoromethylenes Insertion into Fluorome thanes

CH3F +CH2SING C2H4 + HF 4 00E13 0 00 0 . ! ref

CH2F2+CH2SING CH2 : CHF + HF 4 00E13 0 00 0 . ! ref

CHF3 +CH2SING CH2 :CF2 + HF 4 00E13 0 00 0 . ! ref

CF4 +CH2SING CHF:CF2 + HF 4 00E13 0 00 5000 . ! ref

CH4 + CHF C2H4 + HF 4 00E13 0 00 29000

.

!BAC

CH3F + CHF CH2 :CHF + HF 4 00E13 0 00 29000

.

! ref CH4

CH2F2 + CHF CH2 :CF2 + HF 2 00E13 0 00 29000

.

! ref CH4

CH2F2 + CHF CHF:CHF- Z +HF 1 00E13 0 00 29000

.

! ref CH4

CH2F2 + CHF CHFrCHF- E +HF 1 00E13 0 00 29000

.

! ref CH4

CHF3 + CHF CHF : CF2 + HF 4 00E13 0 00 29000

.

! ref CH4

CF4 + CHF CF2 :CF2 + HF 4 00E13 0 00 29000

.

! ref CH4

CH4 + CF2 CH2 : CHF + HF 4 00E13 0 00 80000

.

!BAC

CH3F + CF2 CH2 :CF2 + HF 2 00E13 0 00 80000

.

! ref CH4

CH3F + CF2 CHF:CHF- Z +HF 1 00E13 0 00 80000

.

! ref CH4

CH3F + CF2 CHF:CHF- E +HF 1 00E13 0 00 80000

.

! ref CH4

CH2F2 + CF2 CHF:CF2 + HF 4 00E13 0 00 80000

.

!ref CH4

CHF3 + CF2 CF2 :CF2 + HF 4 00E13 0 00 80000

.

! ref CH4

CF4 + CF2 CF3-CF3 4 00E13 0 00 80000

.

! ref CH4
628 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoromethylenes : Insertion into Fluoromethyls

CH2F + CH2 CH2 : CHF + H 4 00E13 0 00 0 . ! ref CH3

K^tiZ r + Crlz = C2H4 + r 4 U Ubl i u u u 0 . ! ret CH3

CHF2 + CH2 CH2 : CF2 + H 4 00E13 0 00 0 . !ref CH3

CHF2 + CH2 CH2 : CHF + F 4 00E13 0 00 0 . !ref CH3

CF3 + CH2 CH2 : CF2 + F 4 00E13 0 00 0 . ! ref CH3

CH2F + CH2SING = CH2 :CHF + H 4 00E12 0 00 0 . ! ref CH3

CH2F + CH2SING = C2H4 + F 2 00E12 0 00 0 . ! ref CH3

CHF2 + CH2SING = CH2 :CF2 + H 2 00E12 0 00 0 . ! ref CH3

CHF2 + CH2SING = CH2 :CHF + F 4 00E12 0 00 0 . !ref CH3

+ url^ o UNO = CH2 :CF2 +

trr b U Uti-L^ U U (J A

! ret CH3

CH3 + CHF CH2 : CHF + H 1 00E12 0 00 2000 . ! est -46

CH2F + CHF CHF : CHF-Z + H 1 00E12 0 00 2000 . ! est -46

CH2F + CHF CHF: CHF-E + H 1 00E12 0 00 2000 . ! est -46

CH2F + CHF CH2 : CHF + F 1 00E12 0 00 2000 . ! est -37

CHF2 + CHF CHF: CF2 + H 1 00E12 0 00 2000 . ! est -42

CHF2 + CHF CHF : CHF-Z + F 1 00E12 0 00 2000 . ! est -27

CHF2 + CHF CHF: CHF-E + r U Uhi _L z u u u

O A A A z u u u . ! est -2.1 CF3 + CHF CHF: CF2 + F 1 00E12 0 00 5000 . ! est -20

CH3 + CF2 CH2 : CF2 +

TJri ±

-1 n T? T T u

A A u u

C A A b z U U . \ ret - 2

1

CH2F + CF2 CHF: CF2 + H 1 ± Ubl 1 u

A A u ucr A A A

b z 0 U . ! ref - 2 0 CH2F + CF2 CH2 : CF2 + F 1 lOEll 0 0 0 5200 . ! re f - 12

CHF2 + CF2 CF2 : CF2 + H 1 lOEll 0 00 5200 . ! ref - 6 CHF2 + CF2 CHF : CF2 + F 1 lOEll 0 00 5200 . ! ref - 1 Fluoromethylidyne : Oxidation

CF + 02 CF:0 + 0 3 00E13 0 00 0 . ! ref ch

CF + H20 >CHF: 0 + H 3 00E13 0 00 0 . ! ref ch

CF + 0 >C0 + F 7 00E12 0 00 0 . !fit CF + OH >C0 + HF 3 00E13 0 00 0 . ! est

CF + H02 >CF:0 + OH 1 00E13 0 00 0 . ! est

Fluoromethylidyne: Destruction

CH + HF = CF + H2 3 00E13 0 00 0 . ! ref ch -17

CF + CH3 =>CH2:CF + H 3 00E13 0 00 0 . ! est

CF + C2H5 =>CH2:CF + CH3 3 00E13 0 00 0 . ! est

CF + C2H3 =>C2HF + CH2 3 00E13 0 00 0 . ! est

CF + CH2 =>C2HF + H 3 00E13 0 00 0 . ! est

CF + CH2SING=>C2HF + H 3 00E12 0 00 0 . lest

CF + CH4 =>CH2:CHF + H 1 00E13 0 00 0 . ! est

CF + C2H4 =>C2H2 + CH2F 1 00E13 0 00 0 . ! est

CF + CH20 =>CHF + HCO 1 00E13 0 00 8000 . ! est + 7 CF + HCO =>CHF + CO 1 00E13 0 00 0 . ! est

CF + F = CF2 6 00E13 0 00 0 . !fit
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 629

Carbonyl Fluorides: CHF=0, CF2=0

CHF: 0 + M CO + HF 2 48E25 - 3 00 43000

.

!fit 8 5SAI/KUR

CF2 : 0 + H20 - >C02 +2HF 3 92E00 3 08 27600

.

!BAC

CF2 : 0 + ri K^r : U + HF ^ U U J -5 z y X u . CF2 : 0 + H - CF:0 + HF 1 02E02 2 89 32980

.

iQRRK BAC Cadd

CHF: 0 + H HCO + HF 2 91E13 - 0 03 32910

.

! ref cf2o Oadd

CHF: 0 + H - HCO + HF 1 02E02 2 89 32980

.

! ref cf2o Cadd

CHF: 0

TJri + H2 ± -L UIj u o _L 1 1 o n n n J U U U . I ire r / z CHF: 0 + 0 CF:0 + OH 9 00E12 0 00 3080 . ! ref ch2o -2

CHF: 0 + OH = CF:0 + H20 1 72E09 1 18 -447 . ! ref ch2o 18

CF:0 + H202 CHF:0 + H02 1 OOEll 0 00 3900 . ! ref E*0.56 13

CHF: 0 + CH3 CF:0 + CH4 2 00E12 0 00 9000 . ! ref uplim 5 CHF: 0 + CH2F CF:0 + CH3F 2 00E12 0 00 9000 . ! ref uplim + 1 CHF: 0 + CHF2 CF:0 + CH2F2 2 00E12 0 00 9000 . ! ref 0 CHF: 0 + CF3 CF:0 + CHF3 2 00E12 0 00 9000 . ! ref 8

CHF: 0 + C2H3 CF:0 + C2H4 2 . 00E12 0 00 5000 !ref E*0.9

CHF: 0 + C2H5 CF:0 + C2H6 2 . 00E12 0 00 13000 !ref E*1.43

Carbonyl Fluorides: CF=0

CF 0 + M = CO + F 1 89E18 -0 865 39000 RRKM + 32

CF 0 + H = CO + HF 1 20E14 0 00 0 ref hco

CF 0 + 0 = C02 + F 3 00E13 0 00 0 ref hco

CF 0 + OH = C02 + HF 3 00E13 0 00 0 ref hco

CF 0 + H02 = >C02 + F+OH 3 00E13 0 00 0 ref

CF 0 + CH3 = CO + CH3F 1 00E13 0 00 0 est

CF 0 + CH3 = CH2C0 + HF 2 70E13 0 00 0 est

CF 0 + F = CF2:0 1 00E12 0 00 0 wtsangl -131

CF 0 + 02 = >C02 + F + 0 2 00E13 0 00 24000 nist k20 +24

CF 0 + CH2F = CO + CH2F2 1 . 00E13 0 . 00 0 lest, kl9

CF 0 + CHF2 = CO + CHF3 1 . 00E13 0 . 00 0 lest, kl9

CF 0 + CF3 = CO + CF4 1 . 00E13 0 . 00 0 lest, kl9

CF 0 + CH2F = CHFCO + HF 2 . 70E13 0 . 00 0 lest, kl9

CF .0 + CHF2 = CF2C0 + HF 2 . 70E13 0 . 00 0 lest, kl9

CF :0 + CF:0 = CO + CF2 :0 2 . 23E13 0 . 00 318 192MAR/SZE2

Fluoroethanes : Thermal Decomposition

CH3-CH2F C2H4 + HF 2 00E13 0 00 58800. 1 7 0CAD/DAY + 11

CH3-CHF2 CH2 : CHF + HF 7 94E13 0 00 61900. 1 70TSC/QUI + 22

CH3-CF3 CH2 : CF2 + HF 1 00E14 0 00 69000. ! 71TSC/QUI

CH2F-CH2F CH2 : CHF + HF 2 50E13 0 00 61800. ! 71KER/TIM + 5 CH2F-CHF2 CHF: CHF- Z + HF 1 20E14 0 00 69600. 1 74SEK/TSC

CH2F-CHF2 CHF: CHF- E + HF 3 00E13 0 00 65800. 1 74SEK/TSC

CH2F-CHF2 CH2 : CF2 + HF 1 00E13 0 00 65800. 1 74SEK/TSC + 8 CH2F-CF3 CHF: CF2 + HF 2 50E13 0 00 63200. 1 72MIL/TSC + 37

CHF2-CHF2 CHF: CF2 + HF 2 00E13 0 00 70000. 1 71MIL/HAR + 32

CHF2-CF3 CF2 : CF2 + HF 4 00E13 0 00 72000. ! 71TSC/MIL + 41
630 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoroethanes

:

Activated Decomposition, Ethyl + H (HF elimination)

CH2F-CH2 + H = C2H4 + HF 1 44E20 z . ± z 1730 . IRRKM

CH3-CHF + H - C2H4 + HF 2 27E20 z , O 1 1950 . IRRKM

CH2F-CHF + H = CH2 : CHF + HF 2 06E23 - J . O "2 Z J 2280 . IRRKM

CHF2-CH2 + H CH2 : CHF + HF 5 24E16 - U . Q T "3 y J J 880 . IRRKM

CH3 -CF2 + H CH2 : CHF + HF 2 09E16 - U . O D 4 848 . ! RRKM

CHF2-CHF + H CHF: CHF- Z+ HF 1 86E20 -2 . 2 9 1 '7 c; n ; KKrJ^l CHF2-CHF + H CHF:CHF-E+ HF 1 07E19 -1 . 84 1520 . CHF2-CHF + H CH2 :CF2 + HF 9 95E19 -2 . 34 17 8 0. 1 RRKM

CH2F-CF2 + H CHF: CHF- Z+ HF 6 32E19 -2 . 21 16 3 0. ! RRKM

CH2F-CF2 + H CHF:CHF-E+ HF 3 95E18 -1 . 77 13 90. ! RRKM

CH2F-CF2 + H CH2 :CF2 + HF 3 36E19 -2 . 26 1660 .

1 RRKM

CF3-CH2 + H CH2 :CF2 + HF 1 12E21 -2 . 27 2240 .

1 RRKM

CHF2-CF2 + H CHF:CF2 + HF 1 81E22 -2 . 92 3 070. 1 RRKM

CF3-CF2 + H CF2 : CF2 + HF 1 41E21 -2 . 40 3 6 3 0. 1 RRKM

Fluoroethanes

:

Activated Decomposition, Ethyl + H

/ «... A. 1.. -T (methyl formation)

CH2F-CH2 + H CH3 + CH2F 3 BOEll D

"5J ^D 633 . ! RRKM

CH3-CHF + H CH3 + CH2F 3 47E10 0 . 90 1370 . IRRKM

CH2F-CHF + H CH2F + CH2F 1 79E14 -0 . 105 1320 . IRRKM

CHF2-CH2 + H CH3 + CHF2 2 02E06 2 . 16 451

.

IRRKM

CH3-CF2 + H CH3 + CHF2 1 62E05 2 . 50 1370 . IRRKM

CHF2-CHF + H CH2F + CHF2 6 36E12 0 . 318 1460 . IRRKM

CH2F-CF2 + H CH2F + CHF2 3 64E12 0 . 329 1180 . IRRKM

CF3-CH2 + H CH3 + CF3 2 48E11 0 . 816 2870 . IRRKM

CHF2-CF2 + H CHF2 + CHF2 3 25E15 -0 . 524 3000 . IRRKM

CF3-CF2 + H CHF2 + CF3 4 37E16 -0 . 746 4360 . IRRKM

Fluoroethanes

:

Activated Decomposition, Ethyl + H (stabilization)

CH2F-CH2 + H CH3-CH2F 1 19E35 D X 8140 . ! RRKM

CH3-CHF + H CH3-CH2F 9 57E38 - 9 . 24 7360 . IRRKM

CH2F-CHF + H CH2F-CH2F 1 56E45 -10 . 80 8070 . IRRKM

CHF2-CH2 + H CH3-CHF2 2 96E37 -9 . 05 7160 . IRRKM

CH3-CF2 + H CH3-CHF2 3 11E40 -9 . 59 7190 . IRRKM

CHF2-CHF + H CH2F-CHF2 1 20E44 -10 . 60 7520 . IRRKM

CH2F-CF2 + H CH2F-CHF2 2 74E43 -10 . 50 7670 . IRRKM

CF3-CH2 + H CH3-CF3 7 27E42 -9 . 86 7360 . IRRKM

CHF2-CF2 + H CHF2-CHF2 3 77E46 -10 . 80 8980 . IRRKM

CF3-CF2 + H CHF2-CF3 1 12E47 -10 . 80 4100 . IRRKM

Fluoroethanes

:

Activated Decomposition, Ethyl + H (H atom elimination)

CH2F-CH2 + H = CH3-CHF + H 1 87E01 3 . 10 137 . IRRKM

CHF2-CH2 + H = CHF2-CH2 + H 1 59E-3 4 . 35 157 . IRRKM

CH2F-CF2 + H = CHF2-CHF + H 0 64E00 3 . 53 1830 . IRRKM

Fluoroethanes

:

Activated Decomposition, Methyl + Methyl (HF elimination)

CH3 + CH2F = C2H4 + HF 2 35E19 -1

.

86 1870 . IRRKM

CH2F + CH2F = CH2 : CHF + HF 7 56E21 -2 . 79 2590 . IRRKM

CH3 + CHF2 = CH2 :CHF + HF 1 90E15 -0 . 586 634 . IRRKM

CH2F + CHF2 = CHF: CHF- Z+ HF 3 88E20 -2 . 35 2888 . IRRKM

CH2F + CHF2 = CHF:CHF-E+ HF 1 09E19 -1

.

83 2670 . IRRKM

CH2F + CHF2 = CH2:CF2 + HF 2 23E20 -2 . 41 2910 . IRRKM

CH3 + CF3 = CH2:CF2 + HF 5 53E19 -1

.

94 2440 . I RRKM

CHF2 + CHF2 = CHF:CF2 + HF 2 20E19 -1

.

95 4100 . I RRKM

CHF2 + CF3 = CF2 :CF2 + HF 7 00E16 -1

.

17 4330 . I RRKM
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 631

Fluoroethanes : Activated Decomposition, Methyl + Methyl (stablization)

CH3 + CH2F CH3-CH2F 1 57E31 -6 27 4440 . ! RRKM

CH2F + CH2F CH2F-CH2F 2 37E24 -3 79 2290 . ! RRKM

CH3 + CHF2 CH3-CHF2 1 93E35 -7 69 5760 . ! RRKM

CH2F + CHF2 CH2F-CHF2 9 61E38 -8 36 6940 . IRRKM

CH3 + CF3 CH3-CF3 1 78E33 -6 64 5020 . IRRKM

CHF2 + CHF2 CHF2-CHF2 2 26E24 -3 50 3360 . ! RRKM

CHF2 + CF3 CHF2-CF3 2 61E26 -4 16 4100 . ! RRKM

CF3 + CF3 CF3-CF3 1 63E36 -7 26 7050 . IRRKM

Fluoroethanes: H Atom Abstraction by H, 0, OH

CH3-CH2F + H = CH2F-CH2 + H2 5 50E08 1 60 9100 . ! ref lOAE+8 + 2 CH3-CH2F + 0 CH2F-CH2 + OH 2 90E08 1 60 6100 . ! ref 5AE + 5 + 3 CH3-CH2F + OH CH2F-CH2 + H20 5 50E07 1 60 1093 . ! ref -13

CH3-CH2F + H = CH3-CHF + H2 3 30E08 1 60 9100 . ! ref lOAE+8

CH3-CH2F + 0 CH3-CHF + OH 1 60E0B 1 60 6100 . ! ref 5AE + 5 CH3-CH2F + OH = CH3-CHF + H20 3 30E07 1 60 1093 . ! ref

CH3-CHF2 + H CHF2-CH2 + H2 1 50E08 1 60 9600 . ! ref lOA E+8

CH3-CHF2 + 0 CHF2-CH2 + OH 7 50E08 1 60 6600 . ! ref A*5 E+5

CH3-CHF2 + OH CHF2-CH2 + H20 1 54E07 1 60 1132 . ! ref

CH3-CHF2 + H = CH3-CF2 + H2 0 44E08 1 60 9300 . ! ref A*10 E+8

CH3-CHF2 + 0 = CH3-CF2 + OH 2 20E07 1 60 6300 . ! ref A*5 E+5

CH3-CHF2 + OH — CH3-CF2 + H20 0 44E07 1 60 1332 . ! ref

CH3-CF3 + H = CF3-CH2 + H2 4 OOEIO 1 10 12700

.

! ref A*10E+8 + 3 CH3-CF3 + 0 = CF3-CH2 + OH 2 OOEIO 1 10 9700 . ! ref A*5E+5 + 4 CH3-CF3 + OH CF3-CH2 + H20 4 08E09 1 10 4670 . ! 87COH/BEN -12

CH2F-CH2F + H = CH2F-CHF + H2 6 00E08 1 70 9600 . ! ref A*10 E+8 -11

CH2F-CH2F + 0 = CH2F-CHF + OH 3 OOEOB 1 70 6600 . ! ref A*5 E+5 -10

CH2F-CH2F + OH CH2F-CHF + H20 6 16E07 1 70 1610 . ! 87COH/BEN -26

CH2F-CHF2 + H CHF2-CHF + H2 2 00E08 1 70 9800 . ! ref A*10 E+8

CH2F-CHF2 + 0 CHF2-CHF + OH 1 00E08 1 70 6800 . ! ref A*5 E+5

CH2F- CHF2 + OH CHF2-CHF + H20 2 05E07 1 / (J 1 O (J U . ! ref

CH2F-CHF2 + H CH2F-CF2 + H2 1 00E08 1 70 11000

.

! ref A*10 E+8

CH2F-CHF2 + 0 CH2F-CF2 + OH 5 00E07 1 70 8000 . !ref A*5 E+5

CH2F-CHF2 + OH CH2F-CF2 + H20 1 06E07 1 70 3000 . ! ref

CH2F-CF3 + H CF3-CHF + H2 2 00E08 1 70 10500

.

! ref A*10 E+8 - 2 CH2F-CF3 + 0 CF3-CHF + OH 1 00E08 1 70 7500 . ! ref A*5 E+5 - 1 CH2F-CF3 + OH CF3-CHF + H20 2 10E07 1 70 2524 . ! 8 7COH/BEN -17

CHF2-CHF2 + H CHF2-CF2 + H2 1 60E07 1 70 10600

.

! ref A*10 E+8

CHF2-CHF2 + 0 CHF2-CF2 + OH 8 00E07 1 70 7600 . ! ref A*5 E+5

CHF2-CHF2 + OH CHF2-CF2 + H20 1 60E07 1 70 2643 . ! 8 7COH/BEN

CHF2-CF3 + H CF3-CF2 + H2 1 40E07 1 60 10200

.

! ref A*10 E+8 + 0 CHF2-CF3 + 0 CF3-CF2 + OH 7 00E07 1 60 7200 . ! ref A*5 E+5 + 1 CHF2-CF3 + OH CF3-CF2 + H20 1 40E07 1 60 2246 . ! 8 7COH/BEN -15

! * * * NOTE F ATOM ABSTRACTION * * * CF3-CF3 + H CF3-CF2 + HF 1 . 00E15 0 00 30000 !ref CF4
632 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoroethanes : H Atom Abstraction by F

C2H6 + F = C2H5 + HF 8 .. 00E12 0 ., 00 300 . ! 60FET/KNO2 -36

PHI - CH? F p -)- HF 9 _ 00E13 0 0 0 8 0 0. . J. c i- p HF 1 . 00E14 0 0 0 8 0 0. . I. c: J_ p CF3 - CH2 HF ]_ ^ 0 0E14 0 n n 4.^ nw n\J n\J , , L ^ L. CH3-CH2F + F CH3-CHF + HF 6 ., 00E13 0 .. 00 200 . ! ref

CH2F-CH2F + F CH2F-CHF + HF 1

.

, 30E14 0 .. 00 800 . ! ref

CH2F-CHF2 + F CHF2-CHF + HF 1

,

, 30E14 0 ., 00 800 . ! ref

CH2F-CF3 + F CF3-CHF + HF 6 ., 00E13 0 ., 00 1200 . ! ref

CH3-CHF2 + F CH3-CF2 + HF 3 ., 00E13 0 .. 00 800 . ! ref

CH2F-CHF2 + F CH2F-CF2 + HF 3 ,. 00E13 0 .. 00 1200 . ! ref

CHF2-CHF2 + F CHF2-CF2 + HF 6 ., 00E13 0 .. 00 1200 . ! ref

CHF2-CF3 + F CF3-CF2 + HF 4 .. 00E13 0 .. 00 1400 . ! ref -33

Fluoroethanes

:

Association with H02

CH3-CHF + H202 = CH3-CH2F + H02 9 00E09 0 . 0 0 10 0 0. . i. C J- CH3-CF2 + H202 = CH3-CHF2 + H02 9 ., 00E09 0 .. 00 1000 . ! ref

CH2F-CH2 + H202 = CH3-CH2F + H02 9 .. 00E09 0 .. 00 1000 . ! ref

CH2F-CHF + H202 = CH2F-CH2F+ H02 9 ., 00E09 0 ,. 00 1000 . ! ref

CH2F-CF2 + H202 = CH2F-CHF2+ H02 9 .. 00E09 0 .. 00 1000 . ! ref

CHF2-CH2 + H202 = CH3-CHF2 + H02 9 .. 00E09 0 ,. 00 1000 . ! ref

CHF2-CHF + H202 = CH2F-CHF2+ H02 9 .. 00E09 0 ,. 00 1000 . ! ref

CHF2-CF2 + H202 = CHF2-CHF2+ H02 9 .. 00E09 0 ,. 00 1000 . ! ref

CF3-CH2 + H202 = CH3-CF3 + H02 9 ., 00E09 0 ,. 00 1000 . ! ref

CF3-CHF + H202 = CH2F-CF3 + H02 9 ., 00E09 0 ,. 00 1000 . ! ref

CF3-CF2 + H202 = CHF2-CF3 + H02 9 .. 00E09 0 ,. 00 1000 . ! ref
SECTION 5. FLAME INHIBrnON CHEMISTRY AND THE SEARCH 633

Fluoroethanes : H Atom Abstraction by CH3 , C2H5, C2H3

CH3-CH2F + CH3 = CH2F-CH2 + CH4 2 OOEll 0 00 13400

.

! ref + 1 CH3-CH2F + C2H5 = CH2F-CH2 + C2H6 2 OOEll 0 00 16000

.

! ref CH3 + 3 + 6 CH3-CH2F + C2H3 = CH2F-CH2 + C2H4 2 OOEll 0 00 10000

.

! ref CH3 -3 + 4 CH3-CH2F + CH3 = CH3-CHF + CH4 1 50E11 0 00 10000

.

! ref

+

TX D U _L XnU u u

1 T n n n X J u u u . + -5 CH3-CH2F + C2H3 = CH3-CHF + C2H4 1 50E11 0 00 7000 . ! ref CH3 -3

CH3-CHF2 + CH3 = CHF2-CK2 + CH4 2 OOEll 0 00 12000

.

! ref

^ ZXl J

pxj-p9 _ PUO nU DTTi U 1 IL_L _L u T± -3c nU nU nU . 1 ire X CH3-CHF2 + C2H3 CHF2-CH2 + C2H4 2 OOEll 0 00 9000 . ! ref CH3 -3

CH3-CHF2 + CH3 = CH3-CF2 + CH4 8 OOEIO 0 00 10000

.

! ref

CH3-CHF2 + C2H3 = CH3-CF2 + C2H4 8 OOEIO 0 00 7000 . ! ref CH3 -3

^ XT ^ PK"? - PF9 C2H6 Qo n OPT n n n 1X J nU nvj nu , i X tr X -1- 7

CH3-CF3 + CH3 - CF3-CH2 + CH4 2 OOEll 0 00 12000

.

! ref + 2 CH3-CF3 + C2H3 = CF3-CH2 + C2H4 2 OOEll 0 00 9000 . ! ref CH3 -3 + 5 CH3-CF3 + C2H5 = CF3-CH2 + C2H6 2 OOEll 0 00 15000

.

! ref CH3 + 3 + 7 CH2F-CH2F + CH3 = CH2F-CHF + CH4 3 OOEll 0 00 10400

.

! ref -12

+ V— z nJ wriz r L^rir + P9 Wzl U U i_j X X nu nu nvj 7 n n n ! X e X

-3 CH2F-CH2F + C2H5 CH2F-CHF + C2H6 3 OOEll 0 00 13000

.

! ref CH3 + 3 - 7 CH2F-CHF2 + CH3 CHF2-CHF + CH4 2 OOEll 0 00 10000

.

! ref

i^riz r L.rlr + L.Z rlo *.-rir z v_nr + z n'± z VJ U Ji ± X u n n 7/ nu nu nu . i X e X CH2F-CHF2 + C2H5 - CHF2-CHF + C2H6 2 OOEll 0 00 13000

.

! ref CH3 + 3 CH2F-CHF2 + CH3 = CH2F-CF2 + CH4 1 OOEll 0 00 9600 . ! ref

CH2F-CHF2 + C2H3 = CH2F-CF2 + C2H4 1 OOEll 0 00 7000 . ! ref CH3 -3

TX nU nPi U 1 Xlj X X 1X -J nu nu nu . ! X X Pin

CH2F-CF3 + CH3 = CF3-CHF + CH4 2 OOEll 0 00 10000

.

! ref - 3 CH2F-CF3 + C2H3 = CF3-CHF + C2H4 2 OOEll 0 00 7000 . ! ref CH3 -3 + 0 CH2F-CF3 + C2H5 CF3-CHF + C2H6 2 OOEll 0 00 13000

.

! ref CH3 + 3 + 2 CHF2-CHF2 + CH3 CHF2-CF2 + CH4 3 OOEll 0 00 10000

.

! ref

CHF2-CHF2 + C2H3 CHF2-CF2 + C2H4 3 OOEll 0 00 7000 . ! ref CH3 -3

CHF2-CHF2 + C2H5 CHF2-CF2 + C2H6 3 OOEll 0 00 13000

.

! ref CH3 + 3 CHF2-CF3 + CH3 CF3-CF2 + CH4 5 70E10 0 00 9500 . ! 64PRI/THO - 1 CHF2-CF3 + C2H3 CF3-CF2 + C2H4 6 OOEIO 0 00 7000 . ! ref CH3 -3 + 2 CHF2-CF3 + C2H5 CF3-CF2 + C2H6 6 OOEIO 0 00 13000 . ! ref CH3 + 3 + 4
634 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoroethanes : H Atom Abstraction by Fluoromethyls

- PHP F CH2F PH9P_PK9 _j_ CH3 F 2 . OOEll n n 1X

TJ nu 0u nu . L-xlJ + b

V_>i,liJ \^ ±1 tLi X CH2 F CWK - PHF CH3 F . 50E11 n n 1X nU nU nU nVJ . i i tr J- P'H"5 L,xl J CH3-CHF2 + CH2F — CHF2-CH2 + CH3F 2 . OOEll 0 00 12000

.

! ref CH3

CH3-CHF2 + CH2F = CH3-CF2 + CH3F 8 . OOEIO 0 00 10000

.

! ref CH3

rH3 V_Xj. _) - cv~^ IT J ^_ CH2 F PF'^ - PH? CH3F 2 n DFi 1 . VJ VJ iZj J. J. n n n 1 ^9 n\j n\j n\j . i X fc; X P"P"^ + 7

_i_ CH2 F pprop-pwp CH3 F

"3

. vj u rj X J. n n n 1X nu nu nu nu . ! ITGf P"H"^ L-xl J 7

_i_ piJO p pupo _ pwp ^ PH"? F 2 n nFi 1 n n n 1X nu nu nu nu , i X t= J_

+ CH2F pup p_ ppo CH3 F ]_ . OOEll Q 0 0 j~1 n\j n\j n\j n\j . . X X V— xl J CH2 F CF3 - CHF -{- CH3 F 2 . OOEll 0 0 0 X1 nVJ nVJ nw nu . i X t: X '_xlo + 2

_{_ CH2 F pupo V XXX £j _ ppp \v x^ ^ CH3 F 3 . OOEll n n n 1X X1 nVJ nVJ nVJ , 1 X c: X V-XX J + 10

Pfjpp - PF^ _i_ CH2 F CF3 - CF2 CH3 F 2 . OOEll n n n 1^ n\j n\j n\j n\j

,

! X X + 4 CH3 -rH2F -(- CHF2 CH2F-CH2 -f CH2F2 2 . OOEll 0 0 0 13000

.

1 f v^xx O CH3-CH2F + CHF2 = CH3-CHF + CH2F2 1 . 50E11 0 00 10000

.

!ref CH3

CH3-CHF2 + CHF2 = CHF2-CH2 + CH2F2 2 . OOEll 0 00 12000

.

!ref CH3

CH3 -CHF2 CHF2 CH3 -CF2 CH2F2 8 . OOEIO 0 00 10 0 0 0

.

! ref CH3

CH3 -CF3 4- CHF2 CF3 -CH2 CH2F2 2 . OOEll 0 00 12000

.

! ref CH3

puop.pfTop \^ r 1 ^ X. v Li. £j c CHF2 puo \_1X^X^p_ pup \_. XX X CH2 F2 3 . OOEll 0 0 0 10 0 0 0

.

! ref CH3

_l_ pupo _ Pfjp pup pp p . nU nFi U I-J X 1-L n n n 1X nVJ nVJ nu nVJ . ! ref CH3

\_11^L V 1 1 X^ ^ _|. CHF2 pup V 1 1 ^ p_X ppo \^ X^ ^ CH2F2 ]_ . OOEll 0 0 0 10 0 0 0

.

!ref CH3

\— il ^ 1. \^ J -I- CHF2 CF3 - CHF CH2F2 2 . OOEll n 0 0 10000

.

! ref CH3

v_n.x^ ^ v^jniT ^ 4_ CHF2 pupp _ ppp CH2 F2 3 . OOEll n 0 0 1J. noon \J \J \J w . ! ref CH3

p-p-po _ PPT _l_ CHF2 CF3 - CF2 CH2 F2 2 . OOEll n n n 1J. n\j n\j n\j n\j . ! ref CH3

0^3 v. XI ^ - rH2XX ^ x^F _{- CF3 ^XX^X \^xi.£j CHF3 9 . 5 OEl

1

0 0 0 8 2 0 0. ! 71QUI/WHI

CH3 -CH2F + CF3 CH3 -CHF + CHF3 9 . 50E11 0 00 8200 . ! 71QUI/WHI

CH3-CHF2 + CF3 CHF2-CH2 + CHF3 3 . lOElO 0 00 7100 . ! 660KS/PRA

CH3-CHF2 + CF3 CH3-CF2 + CHF3 3 . lOElO 0 00 7100 . ! 660KS/PRA

CH3-CF3 + CF3 CF3-CH2 + CHF3 1 .45E12 0 00 13500

.

! 67GIL/QUI

CH2F-CH2F + CF3 CH2F-CHF + CHF3 5 . OOEll 0 00 7600 . !ref

CH2F-CHF2 + CF3 CHF2-CHF + CHF3 3 . OOEll 0 00 7200 . ! ref

CH2F-CHF2 + CF3 CH2F-CF2 + CHF3 2 . OOEll 0 00 8000 . ! ref

CH2F-CF3 + CF3 CF3-CHF + CHF3 3 . OOEll 0 00 6400 . !ref

CHF2-CHF2 + CF3 CHF2-CF2 + CHF3 5 . 70E11 0 00 11900

.

! 67GIL/QUI

CHF2-CF3 + CF3 CF3-CF2 + CHF3 1 . 40E11 0 00 10100 . ! 67GIL/QUI

Fluoroethanes: H Atom Abstraction by HCO

CH3-CHF + CH20 CH3-CH2F + HCO 5 . 2 Js q^ nVJ nVJ , ! ref

CH3-CF2 + CH20 CH3-CHF2 + HCO 5 . 50E03 2 80 5900 . !ref

CH2F-CH2 + CH20 CH3-CH2F + HCO 5 . 50E03 2 80 5900 . ! ref

CH2F-CHF + CH20 CH2F-CH2F+ HCO 5 . 50E03 2 80 5900 . ! ref

CH2F-CF2 + CH20 CH2F-CHF2+ HCO 5 . 50E03 2 80 5900 . •ref

CHF2-CH2 + CH20 CH3-CHF2 + HCO 5 . 50E03 2 80 5900 . ! ref

CHF2-CHF + CH20 CH2F-CHF2+ HCO 5 . 50E03 2 80 5900 . ! ref

CHF2-CF2 + CH20 CHF2-CHF2+ HCO 5 . 50E03 2 80 5900 . !ref

CF3-CH2 + CH20 CH3-CF3 + HCO 5 . 50E03 2 80 5900 . ! ref

CF3-CHF + CH20 CH2F-CF3 + HCO 5 . 50E03 2 80 5900 . ! ref

CF3-CF2 + CH20 CHF2-CF3 + HCO 5 . 50E03 2 80 5900 . !ref
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 635

Fluoroethyls : Association with 02

CH3-CHF + 02 = CH2 CHF + H02 2 56E19 -2 77 1977 . ! ref

= CH2 CV? + 'HO? 2 -J O Hj _L J7 7 7 1 Q 7 7 . I. t3 I. — PH? 2 -J O Hj J. ^ 7 7 1 Q 7 7 i X J_ r^XJO PXJT7 L.rlz r L.rlr + — ^nr v^nir 111 Triw^ T

_L Z o -L 7 7 1 Q 7 7 ! ITS

f

CH2F-CHF + 02 = CHF CHF-Z +H02 1 28E19 -2 77 1977 . ! ref

CH2F-CF2 + 02 = CHF CF2 + H02 2 56E19 -2 77 1977 . !ref

CHF2-CH2 + 02 = CH2 CF2 + H02 2 56E19 -2 77 1977 . !ref

CHF2-CHF + 02 = CHF CF2 + H02 2 56E19 -2 77 1977 . !ref

CHF2-CF2 + 02 = CF2 CF2 + H02 2 56E19 -2 77 1977 . ! ref

CF3-CH2 + 02 = >CF3 + CH20 + 0 1 30E13 0 00 44000

.

! ref CF3+dH

CF3-CHF + 02 = >CF3 + CHF:0+ 0 1 30E13 0 00 23000

.

! ref CF3

CF3-CF2 + 02 = >CF3 + CF2 :0+ 0 1 30E13 0 00 23000

.

! ref CF3

Fluoroethyls: Association with 0

CH2F-CH2 + 0 CH20 + CH2F 6 60E13 0 00 0 . ! ref

CHF2-CH2 + 0 CH20 + CHF2 6 60E13 0 00 0 . ! ref

CF3-CH2 + 0 CH20 + CF3 6 60E13 0 00 0 . ! ref

CH3-CHF + 0 CHF:0 + CH3 4 40E13 0 00 0 . ! ref

CH2F-CHF + 0 CHF:0 + CH2F 4 40E13 0 00 0 . ! ref

CHF2-CHF + 0 CHF:0 + CHF2 4 40E13 0 00 0 . ! ref

CF3-CHF + 0 CHF:0 + CF3 4 40E13 0 00 0 . ! ref

CH3-CF2 + 0 CF2 :0 + CH3 2 20E13 0 00 0 . ! ref

CH2F-CF2 + 0 CF2 :0 + CH2F 2 20E13 0 00 0 . ! ref

CHF2-CF2 + 0 CF2 :0 + CHF2 2 20E13 0 00 0 . ! ref

CF3-CF2 + 0 CF2 :0 + CF3 2 20E13 0 00 0 . ! ref
636 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoroethyls : Association with H02

CH3-CHF + HO2 =>CH3+CHF:0+0H 3 . 00E13 0 00 0 . ! iref

CH3-CF2 + H02 =>CH3+CF2 :0+0H 3 . 00E13 0 00 0 . !ref

CH2F-CH2 + HO2 =>CH2F+CH20+ OH 3 . 00E13 0 0 . Iref

CH2F-CHF + HO2 =>CH2F+CHF:0+0H 3 . 00E13 0 0 . ! ref

CH2F-CF2 + H02 =>CH2F+CF2 :0+0H 3 . 00E13 0 0 . ! ref

CHF2-CH2 + H02 =>CHF2+CH20+ OH 3 . 00E13 0 0 . ! ref

CHF2-CHF + HO2 =>CHF2+CHF:0+0H 3 . 00E13 0 0 . ! ref

CHF2 -CF2 + H02 =>CHF2+CF2 :0+0H 3 . 00E13 0 0 . ! ref

CF3-CH2 + HO2 =>CF3 +CH20+ OH 3 . 00E13 0 0 . ! ref

CF3-CHF + HO2 =>CF3 +CHF:0+OH 3 . 00E13 0 0 . ! ref

CF3-CF2 + HO2 =>CF3 +CF2:0+0H 3 . 00E13 0 0 . ! ref

CH3 -CHF + HO2 = CH2 : CHF + H202 3 . OOEll 0 0 . ! ref

CH3-CF2 + HO2 = CH2 : CF2 + H202 3 . OOEll 0 0 . ! ref

CH2F-CH2 + HO2 = CH2:CHF + H202 2 . OOEll 0 0 . ! ref

CH2F-CHF + HO2 = CHF:CHF-Z+ H202 2 . OOEll 0 0 . ! ref

CH2F-CHF + HO2 = CHF:CHF-E+ H202 2 . OOEll 0 0 . ! ref

CH2F-CF2 + HO2 = CHF:CF2 + H202 2 . OOEll 0 0 . ! ref

CHF2-CH2 + HO2 = CH2:CF2 + H202 1 . OOEll 0 0 . ! ref

CHF2-CHF + HO2 = CHF:CF2 + H202 1 . OOEll 0 0 . ! ref

CHF2-CF2 + HO2 = CF2 :CF2 + H202 1 . OOEll 0 0 . ! ref

CH3-CHF + HO2 = CH3-CH2F + 02 3 . OOEll 0 0 . ! ref

CH3-CF2 + HO2 = CH3-CHF2 + 02 3 . OOEll 0 0 . !ref

CH2F-CH2 + HO2 = CH3-CH2F + 02 3 . OOEll 0 0 . Iref

CH2F-CHF + HO2 = CH2F-CH2F+ 02 3 . OOEll 0 0 . ! ref

CH2F-CF2 + H02 = CH2F-CHF2+ 02 3 . OOEll 0 0 . ! ref

CHF2-CH2 + HO2 = CH3-CHF2 + 02 3 . OOEll 0 0 . ! ref

CHF2-CHF + HO2 = CH2F-CHF2+ 02 3 . OOEll 0 0 . ! ref

CHF2-CF2 + HO2 = CHF2-CHF2+ 02 3 . OOEll 0 0 . ! ref

CF3-CH2 + H02 = CH3-CF3 + 02 3 . OOEll 0 0 . . J- w J- CF3-CHF + H02 = CH2F-CF3 + 02 3 . OOEll 0 0 . ! ref

CF3-CF2 + HO2 = CHF2-CF3 + 02 3 . OOEll 0 0 . ! ref

Fluoroethylenes

:

Activated Decomposition, Methylene comh '

n

(HF elimination)

CH2SING + CHF = C2H2 + HF 17 . 02E19 -2 12 2380 . ! ref

CH2SING + CF2 = C2HF + HF 17 . 02E19 -2 12 2380 . ! ref

CHF + CHF = C2HF + HF 17 . 02E19 -2 12 2 3 8 0. . i- J- CHF + CF2 = C2F2 + HF 8 . 51E19 -2 12 2380 . ! RRKM

Fluoroethylenes

:

Activated Decomposition, Methylene comh '

n

( stabilization)

CH2SING + CHF = CH2:CHF 3 . 10E24 -3 80 2830 . ! ref

CH2SING + CF2 = CH2:CF2 3 . 10E24 -3 80 2830 . ! ref

CHF + CHF = CHF: CHFZ1 . 55E24 -3 80 2830 . ! ref

CHF + CHF = CHF:CHF-E 1 . 55E24 -3 80 2830 . ! ref

CHF + CF2 = CHF:CF2 3 . 10E24 -3 80 2830 . ! RRKM

Fluoroethylenes

:

Activated Decomposition, Methylene comb'

n

(H atom elim'n)

CH2SING + CHF = CH2 :CF + H 1 . 64E07 1 56 5740 . ! ref

CH2SING + CHF = CHF:CH-Z + H 1 . 64E07 1 56 5740 . ! ref

CH2SING + CHF = CHF:CH-E + H 1 . 64E07 1 56 5740 . ! ref

CH2SING + CF2 = CF2 :CH + H 3 . 28E07 1 56 5740 . ! ref

CHF + CHF = CHF:CF-Z + H 1 . 64E07 1 56 5740 . ! ref

CHF + CHF = CHF:CF-E + H 1 . 64E07 1 56 5740 . ! ref

CHF + CF2 = CF2 :CF + H 1 . 64E07 1 56 5740 . IRRKM
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 637

Fluoroethylenes : Activated Decomposition, Vinyl + H (HF elimination)

CH2 : CF + H = C2H2 + HF 5 . 98E20 -2 31 1940 . ! ref

CHF:CH-Z + H = C2H2 + HF 5 . 98E20 -2 31 1940 . ! ref

CHF:CH-E + H = C2H2 + HF 5 . 98E20 -2 31 1940 . ! ref

CF2 : CH + H = C2HF + HF 5 . 98E20 -2 31 1940 . ! ref

CHF:CF-Z + H = C2HF + HF 5 . 98E20 -2 31 1940 . ! ref

CHF:CF-E + H = C2HF + HF 5 . 98E20 -2 31 1940 . ! ref

CF2 :CF + H = C2F2 + HF 5 . 98E20 -2 31 1940 . !RRKM

Fluoroethylenes

Activated Decomposition, Vinyl + H (stabilization)

CH2 :CF + H = CH2 : CHF 2 . 40E34 -7 11 5040 . ! ref

CHF:CH-Z + H = CH2 : CHF 2 . 40E34 -7 11 5040 . ! ref

CHF:CH-E + H = CH2 : CHF 2 .40E34 -7 11 5040 . ! ref

CF2 : CH + H = CH2:CF2 2 .40E34 -7 11 5040 . ! ref

CHF:CF-Z + H = CHF:CHF- Z 2 .40E34 -7 11 5040 . ! ref

CHF :CF-E + H = CHF:CHF- E 2 .40E34 -7 11 5040 . ! ref

CF2 :CF + H = CHF:CF2 2 .40E34 -7 11 5040 . IRRKM

Fluoromethylene Combination

CHF + CF2 = CF2:CF + H 1.64E07 1.56 5740. ! ref

CF2:CF2 + M = CF2 + CF2 + M 7.80E15 0.50 55690. ! 65TSC

Fluoroethylenes: Thermal Decompositions

CH2 : CHF = C2H2 + HF 1 00E14 0 .00 70800

.

! 70TSC

CH2 : CF2 = C2HF + HF 2 50E14 0 . 00 86000

.

! 70TSC

CHF : CHFZ = C2HF + HF 2 50E14 0 . 00 78000

.

! ref +dH 37

CHF : CHF

E = C2HF + HF 2 50E14 0 . 00 79000

.

!ref +dH 38

CF2 :CF2 + M = CF2 + CF2 + M 7 80E15 0 . 50 55690

.

1 65TSC

CHF : CHFZ = CHF: CHE- E 2 63E13 0 . 00 60700

.

! 74JEF

CH2 :CF2 + CF3 = CF2 : CH + CHF3 7 . 24E10 0 . 00 4000 . ! 7 2KER/PAR + 7 C2H4 + F2 = CH2F -CH2 + F 1 . 38E14 0 . 00 13600

.

! 8 8RAF

C2H4 + F2 = CH2 : CHF + HF 1 . 90E13 0 . 00 11040

.

! 88RAF

C2H4 + F = C2H3 + HF 1 . 00E14 0 . 00 2000 . ! ref -34

Fluoroethylenes: H Atom Addition (stabilization)

CH2 CHF + H (+M)

CH2F-CH2 (+M) 4.20E08 1 50 990 . ! ref

CH2 CHF + H (+M)

CH3-CHF (+M) 4.20E08 1 50 990 . !ref

CH2 CF2 + H (+M) = CHF2-CH2 (+M) 4.20E08 1 50 990 . ! ref

CH2 CF2 + H (+M)

CH3-CF2 (+M) 4.20E08 1 50 990 . ! ref

LOW/ 3 . 19E2 7 -2 . 8 -54./ !ref

H2/2.0/ CO/2 0/ C02/3 0/ H2O/5.0/ !ref

CHF CHF-Z+ H (+M)CH2F-CHF (+M) 8.40E08 1 50 990 . !ref

CHF CHF-E+ H (+M)

CH2F-CHF (+M) 8.40E08 1 50 990 . !ref

LOW/6 . 37E27 -2 . 8 -54./ ! ref

H2/2.0/ CO/2 0/ C02/3 0/ H2O/5.0/ ! ref

CHF CF2 + H(+M)CHF2-CHF (+M) 4.20E08 1 50 990 . ! ref

CHF :CF2 + H(+M)= CH2F-CF2 (+M) 4.20E08 1 50 990 . ! ref

LOW/3 . 19E27 -2 . 8 -54./ ! ref

H2/2.0/ CO/2 . 0/ C02/3 0/ H2O/5.0/ !ref

CF2 :CF2 + H(+M)

CHF2-CF2 (+M) 8.40E08 1 . 50 990 . !ref

LOW/6 . 37E27 -2.8 -54./ ! ref

H2/2.0/ CO/2 .0/ C02/3 0/ H2O/5.0/ ! ref

Fluoroethylenes: H Atom Addition (F atom displacement)

C2H4 + F = CH2 CHF + H 2 O0E13 0 00 0 . ! ref

CH2 : CHF + F = CH2 CF2 + H 2 00E12 0 00 0 . 1 ref

CH2 : CHF + F = CHF CHF- Z + H 5 00E12 0 00 0 . ! ref

CH2 : CHF + F = CHF CHF- E + H 5 00E12 0 00 0 . ! ref

CHF : CHF

Z + F = CHF CF2 + H 4 00E12 0 00 0 . ! ref

CHF : CHF

E + F = CHF CF2 + H 4 00E12 0 00 0 . ! ref

CHF : CF2 + F = CF2 CF2 + H 2 00E12 0 00 0 . !ref
638 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

Fluoroethylenes : H Atom Abstraction by H

CH2 ;:CHF + H = CHF;: CH- Z + H2 0 . 33E06 2 . 53 12241 !ref

CH2 :;CHF + H = CHF;;CH- E + H2 0 . 33E06 2 . 53 12241 !ref

CH2 :CHF + H = CH2 ;:CF + H2 0 . 33E06 2 . 53 12241 ! ref

CH2 :CF2 + H = CF2 ;;CH + H2 0 . 67E06 2 . 53 12241 ! ref

CHF;;CHF- Z + U = CHF;;CF- Z + H2 0 . 33E06 2 . 53 12241 ! ref

CHF;:CHF- E + H = CHF;;CF- E + H2 U . ^ J x:^ U b Z . D Ji 1 z z 4 1 ! ref

CHF::CF2 + H = CF2 ;;CF + H2 0 . 33E06 2 . 53 12241 ! ref

Fluoroethylenes

:

Oxidation by 0

CH2 :CHF + 0 = CH2F + HCO 5 .. 30E09 1

.

00 1310 . !fit 8 7CVE

CHF:;CHF- Z + 0 = CH2F + CF:0 7 .. 00E09 1

.

00 1590 . !fit 8 7CVE

CHF:;CHF- E + 0 = CH2F + CF:0 8 .. 00E09 1

.

00 1320 . !fit 8 7CVE

CH2 :CF2 + 0 = CHF2 + HCO 4 .. 30E09 1

.

00 1490 . !fit 8 7CVE

CHF::CF2 + 0 = CHF2 + CF:0 6 .. 00E09 1

.

00 1150 . !fit 8 7CVE

CF2 :;CF2 + 0 = CF2 + CF2 :0 1 .. 90E09 1

.

00 0 . !fit 8 7CVE

CH2 :CHF + 0 = CH3 + CF:0 5 .. 30E09 1

.

00 2300 . !ref+l 7 4 SLA/GUT

Fluoroethylenes: H Atom Abstraction by OH

CH2 ;:CHF + OH = CHF::CH- Z + H20 2 .. 00E06 2 .. 00 2850 . ! ref -6

CH2 :;CHF + OH = CHF:;CH- E + H20 2 .. 00E06 2 ., 00 2850 . ! ref -6

CH2 ;:CHF + OH = CH2 :CF + H20 1 .. 00E06 2 .. 00 2850 . ! ref -10

CHF;; CHF- Z + OH = CHF::CF- Z + H20 2 .. 00E06 2 .. 00 2850 . ! ref - 5 CHF;:CHF- E + OH = CHF:;CF- E + H20 2 .. 00E06 2 .. 00 2850 . ! ref - 5 CH2 :CF2 + OH = CF2 :CH + H20 2 .. 00E06 2 .. 00 2850 . ! ref

CHF::CF2 + OH = CF2 :;CF + H20 1 .. 00E06 2 .. 00 2850 . ! ref + 1 Fluorovinyls : Oxidation by 02

CH2 ;:CF + 02 CH20 + CF:0 4 ,. 48E26 -4 ,, 55 5480 . ! ref direct

CHF;;CH- Z + 02 CHF:0 + HCO 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CHF;:CH- E + 02 CHF:0 + HCO 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CHF ;:CF- E + 02 CHF:0 + CF:0 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CHF;:CF- Z + 02 CHF:0 + CF:0 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CF2 ;:CH + 02 CF2 :0 + HCO 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CF2 :;CF + 02 CF2 :0 + CF:0 4 ,. 48E26 -4 ,. 55 5480 . ! ref direct

CH2 ;;CF + 02 CH20 + CF:0 1

.

. 05E38 -8 ,.22 7030 . ! ref a/s

CHF;;CH- Z + 02 CHF:0 + HCO 1 ,. 05E38 -8 ,.22 7030 . ! ref a/

s

CHF;:CH- E + 02 CHF:0 + HCO 1 ,. 05E38 -8 ,. 22 7030 . ! ref a/s

CHF;:CF- E + 02 CHF:0 + CF:0 1 ,. 05E38 -8 ,.22 7030 . ! ref a/

s

CHF;:CF- Z + 02 CHF:0 + CF:0 1 ,. 05E38 -8 ,. 22 7030 . ! ref a/

s

CF2 ;:CH + 02 CF2 :0 + HCO 1 ,. 05E38 -8 ,. 22 7030 . ! ref a/

s

CF2 ;;CF + 02 CF2 :0 + CF:0 1 ,. 05E38 -8 ,. 22 7030 . ! ref a/

s

Fluorovinyls;

:

Oxidation by 0

CH2 ;:CF + 0 CH2C0 + F 3 ,. 00E13 0 ,. 00 0 . ! ref

CHF;:CF- E + 0 CHFCO + F 3 ,, 00E13 0 ,. 00 0 . ! ref

CHF;:CF- Z + 0 CHFCO + F 3 ,. 00E13 0 ,. 00 0 . ! ref

CF2 ;:CF + 0 CF2C0 + F 3 ,. 00E13 0 ,. 00 0 . ! ref

CHF:CH-E + O

CHF:CH-Z + O

CF2 :CH + O

= CHFCO + H

= CHFCO + H

= CF2C0 + H

3.00E13 0.00

3.00E13 0.00

3.00E13 0.00

0. !ref

0. !ref

0. !ref
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 639

Fluorovinyls : Oxidation by OH

CH2 CF + OH _ CH2C0 + HF 3 00E13 0 00 0 . ! ref

CHF CH-E + OH _ CH2C0 + HF 3 00E13 0 00 0 . ! ref

CHF CH-Z + OH _ CH2C0 + HF 3 00E13 0 00 0 . ! ref

CHF CF-E + OH — CHFCO + HF 2 00E13 0 00 0 . ! ref

CHF CF-Z + OH CHFCO + HF 2 00E13 0 00 0 . ! ref

CF2 CF + OH CF2C0 + HF 1 00E13 0 00 0 . ! ref

CH2 CF + OH CH3 + CF:0 3 00E13 0 00 0 . ! ref

CHF CH-E + OH CH2F + HCO 3 00E13 0 00 0 . ! ref

CHF CH-Z + OH CH2F + HCO 3 00E13 0 00 0 . ! ref

CHF CF-E + OH CH2F + CF:0 4 00E13 0 00 0 . ! ref

CHF CF-Z + OH CH2F + CF:0 4 00E13 0 00 0 . ! ref

CF2 CF + OH CHF2 + CF:0 5 00E13 0 00 0 . ! ref

Fluoroethynes : H Atom Addition/Stabilization

C2HF +H (+M) CH2:CF (+M) 2 80E12 0 00 2410 . ! ref

C2HF +H (+M) CHF: CH-Z (+M) 1 40E12 0 00 2410 . ! ref

C2HF +H (+M) CHF: CH-E (+M) 1 40E12 0 00 2410 . ! ref

LOW/0 . 67E27 -3.5 2410 /

H2/2./ CO/2. / C02/3 . / H20/5 /

C2F2 +H (+M) CHF: CF-Z (+M) 2 80E12 0 00 2410 . ! ref

C2F2 +H (+M) CHF: CF-E (+M) 2 80E12 0 00 2410 . ! ref

LOW/1 . 33E27 -3.5 2410 /

H2/2./ CO/2. / C02/3 . / H20/5 /

Fluoroethynes: Oxidation

C2HF + 0 = FCCO-E + H 1 00E07 2 00 1900 . !ref*2 MB A077

C2F2 + 0 = FCCO-E + F 1 00E07 2 00 1900 . !ref*2 MB A077

C2HF + OH = CHFCO + H 2 18E-4 4 50 -1000

.

!ref MB A089

C2HF + OH = CH2F + CO 2 50E-4 4 00 -2000

.

!ref/2 MB A090

C2HF + OH = HCCO + HF 2 50E-4 4 00 -2000

.

!ref/2 MB A090

C2F2 + OH = CF2C0 + H 2 18E-4 4 50 -1000

.

!ref MB A089

C2F2 + OH = FCCO-E + HF 2 50E-4 4 00 -2000

.

!ref MB A090

Fluoroketenes , Fluoroketyl

CHFCO + H CH2F + CO 1 13E13 0 00 3428 . ! ref MB A094

CF2C0 + H CHF2 + CO 1 13E13 0 00 3428 . ! ref MB A094

CHFCO + H FCCO-E + H2 5 00E13 0 00 8000 . !ref MB A095

CHFCO + 0 CHF:0 + CO 1 00E13 0 00 8000 . ! est

CF2C0 + 0 CF2 :0 + CO 1 00E13 0 00 8000 . ! est

CHFCO + OH FCCO-E + H20 7 50E12 0 00 2000 . ! ref MB A097

HCCO + F CHF + CO 3 00E13 0 00 0 . ! ref MB AlOl

FCCO-E + F CF2 + CO 3 00E13 0 00 0 . ! ref MB AlOl

FCCO-E + H CHF + CO 1 00E14 0 00 0 . ! ref MB AlOl

FCCO-E + 0 CF:0 + CO 1 00E14 0 00 0 . ! ref MB A102
640 SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH

F Atom Reactions

CH30H + F = CH30 + HF 2 62E09 1 44 -205 . ! 91GLA/K0S

CH30H + F = CH20H + HF 4 62E07 1 97 -300 . ! 91GLA/K0S

CH20 + F = HCO + HF 6 00E13 0 00 2000 . ! ref -45

CHF : 0 + F = CF:0 + HF 2 65E13 0 00 1800 . ! 90FRA/ZHA -35

CH30 + F CH20 -1- HF 3 00E13 0 0 0 0 . HCO + F = CO + HF 1 00E13 0 00 0 . ! est -120

C2H3 -1- F = C2H2 + HF 2 00E13 0 00 0 .

1 yezi-F

. J- o J- CH2 :CF + F = C2HF + HF 2 00E13 0 00 0 . ! ref

CHF : CH-Z + F = C2HF + HF 2 00E13 0 00 0 . . J- ^ X. CHF:CH-E + F = C2HF + HF 2 00E13 0 00 0 . ! ref

CHF:CF-Z + F = C2F2 + HF 2 00E13 0 00 0 . ! ref

CHF:CF-E + F = C2F2 + HF 2 00E13 0 00 0 . ! ref

CF2 :CH + F = C2F2 + HF 2 00E13 0 00 0 . ! ref

CHF:CF-Z + F = CHF + CF2 1 00E13 0 00 0 . ! ref

CHF:CF-E + F = CHF + CF2 1 00E13 0 00 0 . ! ref

CF2 :CF + F = CF2 + CF2 2 00E13 0 00 0 . ! ref

CF2:CF2 + F = CF3 + CF2 3 00E13 0 00 0 . ! ref -18

NOTATION

-^^-E/RT

prefactor (s , etc.)

temperature dependence (dimensionless

)

activation energy (J/mol/4.18)

heat of reaction (kJ/mol/4 . 18

)

REFERENCES

65PRI/BRY Pritchard et al .

, 1965

65TSC Tschuikow-Roux, 1965

67PAR/AZA Parsamyan et al . , 1970

68PAR/NAL Parsamyan and Nalbanddyan, 1968

70CAD/DAY Cadman et al .

, 1970

70TSC/QUI Tschuikow-Roux and Quiring, 1970

71KER/TIM Kerr and Timlin, 1971

71K0C/M0I Kochubei and Moin, 1971

71MIL/HAR Millward et al .

, 1971

71TSC/MIL Tschuikow-Roux et al . , 1971a

71TSC/QUI Tschuikow-Roux et al . , 1971b

71WAG/WAR Wagner et al . , 1971

72KER/PAR Kerr and Parsonage, 1972

72MIL/TSC Millward and Tschuikow-Roux, 1972

72RID/DAV Ridley et al .

, 1972

73SCH/WAG Schug and Wagner, 1973

74SEK/TSC Sekhar and Tschuikow-Roux, 1974

74SLA/GUT Slagle et al . , 1974

75WES/DEH Westenberg and deHaas, 1975

76PAY/STI Payne and Stief, 1976

77KEA/MAT Keating and Matula, 1977

78ART/BEL Arthur and Bell, 1978

78BIO/LAZ Biordi et al ., 1978

78JOU/LEB Jourdain et al . , 1978

79SCH/WAG Schug et al . , 1979
SECTION 5. FLAME INHIBITION CHEMISTRY AND THE SEARCH 641

REFERENCES (continued)

81BAU/DUX Baulch et al . , 1981

83WAL/WAG Walther and Wagner, 1983

84RYA/PLU Ryan and Plumb, 1984

84WAR Warnatz, 1984

85SAI/KUR Saito et al ., 1985

86PLU/RYA Plumb and Ryan, 1986a, 1986b

86TSA/HAM Tsang and Hampson, 1986

87COH/BEN Cohen and Benson, 1987a, 1987b

87CVE Cvetanovic, 1987

87DEA/WES Dean and Westmoreland, 1987

89STE/BRU Stevens et al . , 1989

89TSA/BEL Tsai et al . , 1989

89TSA/MCF Tsai and McFadden, 1989

90BOZ/DEA Bozzelli and Dean, 1990

90TSA/MCF Tsai and McFadden, 1990a, 1990b

91C0H/WES Cohen and Westberg, 1991

91HID/NAK Hidaka et al . , 1991

92BAU/COB Baulch et al . , 1992

92MAR/SZE Maricq and Szente, 1992

RRKM this work, RRKM calculations

QRRK this work, QRRK calculations

BAC this work, BAC-MP4 & RRKM calculations

ref this work, by analogy to hydrocarbon reaction

643

6. AGENT STABILITY UNDER STORAGE AND

DISCHARGE RESIDUE

Richard D. Peacock, Thomas G. Cleary, and Richard H. Harris Jr. Building and Fire Research Laboratory

6.1 Agent Stability Under Storage

Halon 1301 is known to be stable in metal containers for many years. Any by-products do not affect

its fire suppression effectiveness or result in an unacceptable residue. For candidate replacement

chemicals, comparable data are needed, reflecting the storage conditions of elevated temperature and

pressure. Significant losses in fire suppression effectiveness and increases in toxicity are possible if the extinguishing agent degrades during storage. Thus, stability during the multi-year storage

environment is an important concern. The storage environment fosters conditions which may have an adverse effect on the stability of

halon replacements. Stored chemicals may engage in oxidation-reduction reactions, hydrolysis, and - other corrosive interactions with metal cylinders. They are also subject to unimolecular decomposition and attack by reactive impurities in the agent. Water and oxygen, for example, will sorb to surfaces of cylinders and transfer lines and can never be completely excluded. These sources of

instability, along with the possibility of catalytic interactions with the cylinder walls, can promote the

evolution of undesirable products and a concomitant loss of fire suppression effectiveness. Toxicity

and corrosiveness are particularly important concerns with respect to halogenated compounds, due to the tendency to liberate hydrogen halide in the process of degradation.

NIST Technical Note 1278 (Gann et ai, 1990) details a screening test for the stability of

chemicals, with the purity of the chemical determined by infrared spectral analysis. This test involves

storing the candidate in a carbon-steel cylinder for two days at a temperature of 170 °C (338 °F).

The infrared spectra of the original and "aged" samples are compared. Although appropriate for a screening procedure, promising candidate agents need a more rigorous examination.

In this project, samples of each of the 12 candidate agents were evaluated in pressurized

cylinders. It was presumed that NaHC03 is stable under the likely storage temperatures and pressures. In order to allow for potential interactions analogous to actual storage conditions, a measured

amount of metal (with separate tests for each candidate cylinder metal) was introduced into the

containers prior to the experiments. The vessel and its contents were stored in an oven at elevated

temperature for 28 days. After cooling to ambient conditions, an infrared spectrum of the aged

sample was compared to a spectrum of the original sample. Degradation of the sample would be

indicated by a systematic decrease in the absorbance of peaks attributable to the agent and/or the

appearance of new peaks in the IR spectrum of the aged agent.

6.1.1 Selection of Experimental Technique. The most reliable way to assess storage stability is to determine the shelf life of each chemical by monitoring its degradation in the actual storage environment. Existing ASTM test methods for assessing the storage stability of aircraft cleaning compounds

are available (ASTM, 1987a and 1987b). After storage at specified conditions for 12 months,

samples are inspected for signs of corrosion and sediment formation, and their performance is tested
644 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

against contract specifications. For evaluation of potential halon replacements, such long term storage

is only practical for evaluating a near-final selection of agents. For the current project, an alternative

approach must be taken.

The temperature dependence of potential reactions can be dramatic. For example, the rate of a chemical reaction with an activation energy of 126 kJ/mole, which is reasonable for catalytic

degradation, increases by almost a factor of 10^ over the temperature range from 25 to 225 °C. The

ASTM test method for distillate fuels uses this phenomenon to accelerate aging in samples stored

from 4 to 24 weeks at 43 °C. The elevated temperatures used in these test accelerate aging and make

it possible to complete long-term stability assessments in a time frame of weeks rather than years. The strategy of using temperature to study stability is also the basis of differential thermal

analysis (DTA). This technique is frequently used to study phase transitions and assess thermal

stability in liquids and solids (Wendlandt, 1964) In DTA, thermocouples are used to monitor the

temperature difference between the sample and an inert reference material as a function of increasing

temperature. A thermal event is signaled by a peak in the trace of voltage versus temperature and is indicative of a reaction in the sample. DTA is useful for measuring transition temperatures and for determining the signs of observed enthalpy changes. An ASTM standard test method for assessing

thermal stability using DTA (E 537-86) is performed by observing the temperature changes which

result from heating the sample from room temperature to about 500 °C. In performing this test, the

investigator makes note of all enthalpy changes, the temperatures at which these transitions occur, whether they are exothermic or endothermic, and the effects of changing the atmosphere in the test

cell. The criteria for thermal stability is the absence of a reaction.

Neither DTA (or the related differential scanning calorimetry) are appropriate for the current

study. These techniques are typically applied to detect chemical reactions and phase transitions in

solids. They would be difficult, if not impossible, to apply to potential halon replacements, which

are, almost exclusively, gases and volatile liquids. More importantly, they measure thermal, as opposed to storage stability. The presence of reactive impurities or heterogenous catalysis in the

storage environment cannot be determined by studying pure samples of the agent alone.

Actual storage at elevated temperatures for some period of time is thus desirable. Degradation can be monitored via conventional analytical techniques, such as gas chromatography (GC) (Grob,

1977) or infrared spectroscopy (IR) (Willis, et al., 1987).

A GC consists of a heated injection port connected to a column which is coated with an

adsorbent material called the stationary phase. Mixtures are transported through the column by an

inert gas. The components are separated on the basis of the amount of time they require to traverse the length of the column. Retention times in the column increase with the affinity of the component

to the stationary phase. A detector is positioned so that it can respond to the individual components

as they are sequentially eluted from the column. Ideally, the chromatogram, which is a plot of

detector response versus retention time, consists of a series of well-resolved peaks, each of which

corresponds to a pure compound. In practice, however, co-elution of components is an important concern which must be addressed by temperature programming and careful selection of the stationary

phase. Accurate measurements of the concentrations of identified mixture components are frequently

made using a GC equipped by an ionization detection or electron capture detector. To make

definitive compound identifications, the GC is often paired with a mass spectrometer.

Chemical analysis by IR spectroscopy is based on the assumption that the IR spectrum of a compound is sufficiently unique to identify it. An IR spectrum is obtained by measuring the ratio of

the intensity of IR radiation which passes through the sample, /, to the intensity of the incident

radiation, 1^, as a function of frequency. Concentrations of individual components are quantified by

applications of Beer's law
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 645

~Ajy)

A{y) = -log

/(V) (1)

where the subscript r indicates known values obtained from a reference sample. IR spectroscopy is extremely versatile in the sense that almost all compounds, with the notable exception of single atomic

and homonuclear diatomics, are IR active. In principle, IR analysis makes it possible to monitor the

degradation of each candidate and identify the corresponding products from a simple before and after comparison of the spectra. In practice, IR analysis can suffer from overlapping peaks in a spectrum

similar to problems with gas chromatography. In addition, quantitative determination of degradation

products require either pre-existing spectra for comparison (libraries of spectra for common materials

are commercially available) or available reference samples.

GC is much more sensitive than IR. The major disadvantage is the considerable effort and

expense involved in column preparation and selection. This situation is exacerbated by the need to use different columns to resolve a wide variety of alternative chemicals from their degradation

products. In addition, since it was expected that the additional sensitivity afforded by GC would not be needed in most cases (most of the compounds were expected to be stable), the speed, cost, and

acceptable resolution of IR spectroscopy made this approach the best alternative for the screening tests conducted. R3r longer-term studies, to be performed once the list of acceptable chemicals has been

reduced to a few alternatives, IR studies can be supplemented with GC analysis as appropriate.

Since the agents may interact with metal storage cylinders during long term storage, it is important to evaluate the agents in the presence of a variety of metals which may be encountered

during typical use. Table 1 shows the metals selected for the current smdy Although the primary

interest is in storage c>iinder materials, additional materials which may come in contact with the

agents for extended periods of time were included. Ideally, containers constructed of each material

would be used for agent storage during the smdy. However, cost and availability of containers made

from the eight metals necessitated the use of metal coupons inside an inert container. In this approach, an amount of metal with surface area roughly equal to the storage c>'linder was introduced

into teflon-lined stainless steel cylinders prior to agent filling. With new metal samples for each

experiment, this also eliminated potential longer-term degradation of the metals over multiple

experiments.

6.1.2 Experimental. A total of thirteen agents were screened for their long-term storage stability

(twelve candidate replacement agents and halon 1301). Each agent was exposed to fixed amounts of

eight different metals at elevated temperature and pressure. Testing conditions dictated that the

exposure temperature be fixed at 149 °C, and that the room temperature fill pressure be 4.13 MPa

(600 psia). In order to assure that the storage cylinder pressure not exceed the pressure rating on the

c>'linder or valve when exposed to the elevated temperature, a conservatively low amount of agent was placed in each cylinder. The cylinders were filled with agent up to a maximum pressure of the

saturation vapor pressure of the agent at room temperamre (in the case of the gaseous agent, up to a

fixed pressure). Nitrogen was then added to achieve the final pressure. Since small amounts of

additional liquid would vaporize at the elevated temperamre only a single phase would exist anyway,

though a significant pressure increase would be observed. The effect of the system pressure on

degradation was not explored, but is thought to be small.

Eight different metals were chosen to be exposed to the agents for the screening. The metals were introduced into the cylinders as coupons 10.2 cm long, 0.8 cm wide, and 0.2 cm thick. The

number of coupons was fixed at 30 for each cylinder because that amount is approximately the
646 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Table 1 . Metals used in Agent Stability Study

Metal Major Constituents Typical Use

Nitronic-40 (21-6-9 Stainless) Cr Mn Ni Fe Pressure vessel

304 Stainless Cr Ni Fe Pressure vessel, pressure

vessel fitting, gage fitting,

and diagnostic devices

AM 355 Stainless Cr Mo Ni Fe Outlet closure

4130 Alloy Steel Cr Mn Si Fe Pressure vessel

13-8 Mo Alloy Steel Al Cr Mo Ni Fe

Inconel 625 Cr Fe Mo Ni Bourdon tube

6061-T6 Aluminum Alloy Al Si Fe Mg Compression gasket, fill fitting

CDA 172 Cu-Be Alloy Cu Be Bourdon tube, outlet closure

surface area of the interior of the cylinder. Overlapping of the coupons in the cylinder effectively

decreased the exposed surface area, but this decrease, though not calculated, should be similar for all

tests.

6.1.2.1 Test Preparation. The storage cylinders were constructed of stainless steel, lined with

polytetrafluoroethylene (PTFE), and had a 1000 ml capacity. All new cylinders were assembled with

high temperature/high pressure stainless steel valves and end plugs, then heated to 149 °C (300 °F)

for at least 48 hours with the valves completely open. For re-used cylinders, the heating time was cut

to 24 hours.

The metal coupons were immersed in dichloromethane solvent and stirred around to remove any

oil residue and markings on the surfaces. The coupons were removed from the solvent, stripped with

clean paper towels, and dried at room temperature. Care was taken to handle the cleaned coupons

with gloved hands only. Thirty pieces of each metal were weighed and placed into their respective

cylinders. The end plugs threads were wrapped with PTFE tape and tightened. The cylinder valves were opened and the cylinders were heated for at least 12 hours at 149 °C. After the cylinders were removed from the ovens, the valves were closed.

6.1.2.2 Filling Procedure. When the cylinders were cool enough to be handled with gloves,

they were evacuated to an absolute pressure of less than 20 Pa (0.15 torr). Each evacuated cylinder was weighed. The agents were added and the cylinder was pressurized to the final pressure with

nitrogen. The detailed filling procedure is given below.

The filling procedure was designed to insure low levels of contaminant (including room air) and

a reasonably accurate measurement of the amounts of material introduced into the cylinder. Highly

accurate measurement of amounts of agent and/or nitrogen are not required in this screening method
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 647

Regulated

Nitrogen

Storage Cylinder

Relief

Valve

Pressure

Gage

High

FYessure

I

Agent

Supply

House

Vacuum

Figure 1. Experimental apparatus for agent filling study.
648 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

since it is the change in spectra that was desired. Figure 1 is a schematic of the filling apparatus. It consisted of a three valve manifold, an absolute pressure gage, and filling lines, all constructed from

stainless steel and capable of handling high pressures. The lines to the agent tank and the house vacuum were not high pressure rated and were isolated by valves. The following steps outline the

filing procedure.

1. The evacuated cylinder was attached to the filling apparatus.

2. The apparatus was evacuated with the house vacuum which was capable of pulling a vacuum down to 34 kPa. Then the valve to the house vacuum was closed.

3. The apparatus was pressurized with nitrogen in excess of 1.4 MPa. The pressure was released

through the relief valve.

4. The apparatus was then evacuated again with the house vacuum. Steps 2-4 purge the filling

apparatus. After purging the amount of room air in the apparatus was less than 1 % of the

original amount. The volume of the filling apparatus is much less than a storage cylinder, so the

percentage of air in any filled cylinder was very small (on the order of 0.01 %).

5 . The agent inlet valve was opened slowly and the pressure in the apparatus checked (the agent

inlet line is pre-charged with agent). At this point if the agent cylinder included a regulator, the

delivery pressure could be adjusted. The valve to the cylinder is slowly opened and agent flowed

in. Both the inlet valve and the cylinder valve must be opened slowly to avoid the possibility of

drawing liquid agent into the apparatus. The system was allowed to equilibrate for one to two

minutes, the pressure recorded, and the agent inlet valve closed.

Here we note that the delivery pressure of the agent was, in most cases, the saturation vapor

pressure at room temperature of that particular agent. One agent (FC-116) was available as a high pressure gas, and two other agents (HFC- 125 and the azeotrope of HFC-32 and HFC- 125)

had saturation vapor pressures somewhat higher than the other agents, therefore the delivery

pressure for these agents was specified to yield a mass of agent delivered to the cylinder close to the mass delivered for most of the other agents.

6. Nitrogen was added to increase the total pressure to nominally 4.20 MPa. This pressure was

slightly higher than the prescribed target pressure of 4.13 MPa, but because a small amount of

the agent/N2 mixture was removed from the cylinder for the initial FTIR analysis, a slight

decrease in the initial (cold) storage pressure was observed. The total pressure was recorded and

the valve to the storage cylinder closed.

7. The pressure was bled from the apparatus and the cylinder disconnected and weighed. At this point, a new cylinder can be connected and filled starting with Step 1 above.

6.1.2.3 Conduct of the Tests. Each cylinder was connected to the inlet of the gas cell on the

FTIR system. A vacuum of less than 20 Pa was drawn on the gas cell and inlet line up to the

cylinder valve. A needle valve was placed in line between the cylinder valve and the gas cell to accurately control the flow of the agent into the gas cell. The agent/N2 was introduced into the gas

cell to an absolute pressure of 1330 ± 10 Pa (9.95 + .05 torr). Three spectra were taken for each

sample. The cylinders were then re-weighed.
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 649

The filled cylinders were placed in the ovens at 149 °C and the time and date were recorded.

Periodically, the cylinders were removed and weighed while still hot, then placed back in the oven to check for leakage. It was observed that the valves had to be re-tightened (while hot) after being in the ovens for a day or two. In very few cases was leakage significant, though it appeared that the

valves used in this study degraded after some re-use leading to some leaking. In any long term study

the valves should be capped to prevent leakage. After 28 days the cylinders were removed from the ovens and allowed to cool before the spectra were taken.

The final spectra were obtained in the same manner as the initial spectra. Comparisons between

the initial and final spectra were made.

6.1.2.4 Exposed metals. The agent/N2 mixture was released from each cylinder in an exhaust

hood. The end plugs were removed and the metal coupons shaken from the cylinders. The metals were examined visually for changes, wrapped in aluminum foil and retained for reference. The

cylinders were then ready for reuse.

6.1.3 Results. Table 2 lists the metal mass, agent pressure, calculated agent mass, and initial agent/N2 pressure and mass before and after storage. In this table, the agent mass was calculated

from the agent density at the recorded pressure and ambient temperature (along with the fixed 1000

ml volume of the cylinders). The error in this calculated value is the combination of the uncertainty

in the agent pressure reading which is estimated to be + 0.01 MPa (± 2 psia), deviations of the

agent temperature from ambient temperature (at least two minutes were allowed for the system to equilibrate) and the error associated with the equation of state. Here, generalized compressibility

charts were used (Balzhiser et al., 1972). The combined uncertainties in the calculated values were

not estimated. Uncertainty in the recorded masses depends on the weighing apparatus (agent/N2 masses were obtained by subtracting the initial mass of the evacuated cylinder with metals in place).

Neglecting the cases where obvious leaking occurred, the masses of the agent/N2 were within ±5%

for any given agent with the exception of FC-31-10. Inspection of the calculated agent masses, and

measured agent/N, masses for the FC-31-10 indicates that some liquid agent was probably drawn into the cylinder during filling (approximxately 40 extra grams of agent). It is remarkable that the filling

procedure still yielded final agent/N2 masses within ±10%. The extra liquid vaporized at the

elevated temperature.

6.1.3.1 FTIR Spectra. The spectra taken before and after the storage period were examined to identify changes in the gas phase. Any new peaks observed in the spectra taken after storage would

indicate that some degradation of the agent had occurred during the storage time. In addition a decrease in the area under specific peaks of the final spectra would indicate degradation. Equation 1 would yield quantitative concentrations.

Figures 2-4 show the initial and final overlaid IR spectra for three of the agents. In some of the

spectra, baseline perturbations can be seen. However, these perturbations generally occur above 2500

cm'^ outside the area of interest for fluorinated compounds and are likely due to detector sensitivity

fall ofi" or gas cell window contamination.

For all thirteen agents examined, no new peaks were observed in any of the spectra. Peaks of new compounds could be masked by the peaks formed by the original compound, so examination of

the peak heights or areas under the peaks would provide further validation of whether or not the

agents were stable. During the course of the storage experiments, the FTIR system experienced a catastrophic failure of the computer. The FTIR was updated with new computer hardware, but several electronic files were lost which included the initial and final spectra for the agents HCFC-22, HCFC-124, HFC- 125,
650 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Table 2. Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: HCFC-22

Agent Agent/N2

Metal

Metal

Mass

(±0.01 g) Pressure (+ 0.02 MPa)

Mass

(g)

Pressure (± 0.02 MPa)

Mass

Before

Storage

(± 0.1 g)

Mass

After

(+ 0.1 g)

304 stainless steel 272.07 0.97 40 4.13 75.6 74.8

Nitronic-40 (21-6-9) 325.45 0.97 40 4.13 72.8 73.0

Inconel 625 320.29 0.97 40 4.13 75.4 75.2

6061-T6 aluminum 96.55 0.97 40 4.13 76.8 75.5

4130 alloy steel 304.45 0.95 39 4.17 76.5 75.0

AM 350 stainless 285.68 0.94 39 4.13 __a 77.1

Cu/Be C82500 311.20 0.96 39 4.17 77.8 76.6

13-8 Mo stainless 280.13 0.95 39 4.13 74.6 73.5

Agent: HCFC-124

Agent Agent/N2

Metal

Metal

Mass

(+ 0.01 g) Pressure (± 0.02 MPa)

Mass

(g)

Pressure (± 0.02 MPa)

Mass

Before

Storage

(± 0.1 g)

Mass

After Storage

(+ 0.1 g)

304 stainless steel 272.61 0.31 19 4.13

Nitronic-40 (21-6-9) 323.56 0.31 19 4,20 63.1 63.1

Inconel 625 319.78 0.33 20 4.15 62.9 63.3

6061-T6 aluminum 96.95 0.31 19 4.13 64.8 63.8

4130 alloy steel 303.52 0.33 20 4.13 63.5 63.6 AM 350 stainless 283.72 0.33 20 4.17 62.8 51.3

Cu/Be C82500 311.48 0.33 20 4.17 65.3 63.2

13-8 Mo stainless 285.73 0.31 19 4.17 64.2 64.6

^ - data not available
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 651

Table 2. (continued) Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: HFC- 125

Metal

Metal

Mass

(± 0.01 g)

Agent Agent/N2

Prf^ ecu rA (± 0.02 MPa)\A Q C C

(g)

r IcSSUlC (± 0.02 MPa)

Mass

Before

Storage

(+ 0.1 g)

Mass

After Storage

(+ 0.1 g)

304 stainless steel 273. Us 0.69 38 4.20 72.1 72.7

Nitronic-40 (21-6-9) 323.72 0.69 38 4.20 74.4 74.2

Inconel 625 320.11 0.69 38 4.20 74.1 74.4

6061-T6 aluminum 96.87 0.69 38 4.20 76.5 76.7

4130 alloy steel 303.57 0.69 38 4.20 73.4 73.6

AM 350 stainless 286.23 0.69 38 4.20 76.0 45.5 Cu/Be C82500 311.06 0.72 39 4.20 76.9 77.2

13-8 Mo stainless 283.45 0.69 38 4.20 72.6 73.2

Agent: HFC-32/HFC-125

Agent Agent/N2

Metal

Metal

Mass

(± O.OIg)

Pressure (+ 0.02 MPa)Mass

(g)

Pressure (± 0.02 MPa)

Mass

Before

Storage

(± 0.1 g)

Mass

After Storage

(± 0.1 g)

304 stainless steel 273.49 0.69 24 4.20 55.2 55.4

Nitronic-40 (21-6-9) 323.03 0.69 24 4.20 55.2 55.7

Inconel 625 320.34 0.69 24 4.20 53.9 54.7

6061-T6 aluminum 96.72 0.69 24 4.20 55.2 55.4

4130 alloy steel 303.44 0.69 24 4.20 55.7 56.1 AM 350 stainless 286.22 0.69 24 4.20 55.6 55.7

Cu/Be C82500 311.02 0.69 24 4.20 58.1 58.3

13-8 Mo stainless 284.32 0,69 24 4.20 55.2 56.0
652 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Table 2 (continued). Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: HFC- 134a

Agent Agent/N2

Metal

ivi& lal Mass

(± 0.01 g) Pressure Mass Pressure Mass

Before Mass

After (± 0.02 MPa) (g) (± 0.02 MPa) Storage

(± 0.1 g)

Storage

(± 0.01 g)

304 stainless steel 273.60 0.59 27 4.20 66.8 66.6

Nitronic-40 (21-6-9) 324.52 0.59 27 4.20 62.3 60.1

Inconel 625 319.85 0.59 27 4.20 64.2 63.6

6061-T6 aluminum 96.83 0.59 27 4.41 68.3 67.9

4130 alloy steel 303.77 0.59 27 4.20 64.1 63.9

AM 350 stainless 284.73 0.59 27 4.20 66.7 66.9

Cu/Be C82500 311.65 0.59 27 4.20 62.2 61.6

13-8 Mo stainless 285.24 0.59 27 4.20 64.2 64.3

Agent: FC-218

Agent Agent/N2

Metal

Metal Mass

(± 0.01 g) Pressure Mass Pressure Mass

Before MassAfter (± 0.02 MPa) (g) (+ 0.02 MPa) Storage

(± 0.1 g)

Storage

(± 0.1 g)

304 stainless steel 274.10 0.81 80.1 4.24 107.7 107.3

Nitronic-40 (21-6-9) 324.78 0.79 78.7 4.24 104.7 104.2

Inconel 625 317.72 0.79 78.7 4.24 107.1 107.1

6061-T6 aluminum 97.01 0.79 78.7 4,24 107.2 107.2

4130 alloy steel 303.34 0.79 78.7 4.24 101.

1

102.1 AM 350 stainless 284.92 0.79 78.7 4.24 101.9 105.9 Cu/Be C82500 311.66 0.79 78.7 4.24 103.7 103.6

13-8 Mo stainless 282.99 0.76 75.3 4.28 83.4
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 653

Table 2 (continued). Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: HFC-227

Metal

IViCull Mass

(± 0.01 g)

Agent Agent/N2

Pressure (+ 0.02 MPa)

Mass

(g)

Pressure (+ 0.02 MPa)

Mass

Before

Storage

(+ 0.1 g)

Mass

After Storage

(± 0.1 g)

304 stainless steel 272.8 0.40 31 4.26 72.8 73.0

Nitronic-40 (21-6-9) 323.29 0.40 31 4.20 71.6 71.5

Inconel 625 318.02 0.41 32 4.20 69.6 70.3

6061-T6 aluminum 97.02 0.41 32 4.20 69.7 69.7

4130 alloy steel 303.18 0.40 31 4.20 66.4 66.5 AM 350 stainless 285.37 0.40 31 4.20 70.0 69.9

Cu/Be C82500 310.51 0.41 32 4.24 70.2 70.0

13-8 Mo stainless 282.26 0.40 31 4.24 68.5 68.5

Agent: FC-31-10

Agent Agent/N2

Metal

Metal

Mass

(+ 0.01)

Pressure (+ 0.02 MPa)Mass

(g)

Pressure (+ 0.02 MPa)

Mass

Before

Storage

(± 0.1 g)

Mass

After Storage

(+ 0.1 g)

304 stainless steel 273.79 0.23 24 4.20 97.7 97.8

Nitronic-40 (21-6-9) 325.27 0.21 22 4.20 103.4 103.3

Inconel 625 320.01 0.21 22 4.20 111.9 112.4

6061-T6 aluminum 97.03 0.21 22 4.20 117.1 100.4

4130 alloy steel 302.73 0.21 22 4.20 113.6 111.9 AM 350 stainless 287.56 0.21 22 4.20 106.4 106.4

Cu/Be C82500 311.62 0.21 22 4.20 110.9 110.9

13-8 Mo stainless 284.99 0.21 22 4.20 110.8 110.3
654 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Table 2. (continued) Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: FC-116

Metal

Metal

Mass

(± 0.01 g)

Agent Agent/N2

Pressure

(-1- 0 02 MPa)

Mass

(s)

Pressure (+ Q 02 MPal

Mass

Before O LUl due (+ 0.1 g)

Mass

After

(± 0.1 g)

634 stainless steel 273.23 0.55 32.4 4.20 66.4 66.0

Nitronic-40 (21-6-9) 324.53 0.55 32.4 4.20 68.7 68.6

Inconel 625 320.85 0.56 33.2 4.20 71.2 71.4

6061-T6 aluminum 97.01 0.56 33.2 4.20 71.8 71.6

4130 alloy steel 304.05 0.56 33.2 4.27 72.2 71.8

AM 350 stainless 286.24 0.56 32.8 4.20 72.7 72.6 Cu/Be C82500 311.97 0.57 33.6 4.22 47.5 47.2

13-8 Mo stainless 281.79 0.57 33.6 4.21 70.8 70.6

Agent: FC-318

Metal

Metal

Mass

(+ 0.01 g)

Agent Agent/N2

Pressure (± 0.02 MPa)Mass

(g)

Pressure (+ 0.02 MPa)

Mass

Before

Storage

(+ 0.1 g)

Mass

After Storage

(± 0.1 g)

304 stainless steel 272.47 0.26 23.4 4.20 66.3 66.3

Nitronic-40 (21-6-9) 324.44 0.26 23.4 4.20 65.6 65.4

Inconel 625 319.30 0.25 22.7 4.20 66.5 66.3

6061-T6 aluminum 69.55 0.25 22.7 4.22 - 67.0 66.9

4130 alloy steel 303.38 0.26 23.4 4.20 60.5 60.6

AM 350 stainless 284.62 0.25 22.7 4.20 64.5 64.5

Cu/Be C82500 313.05 0.26 23.4 4.20 64.2 63.9

13-8 Mo stainless 281.84 0.25 22.7 4.22 65.6 65.4
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 655

Table 2 (continued). Agent and Metal Amounts in Storage Cylinders for Stability Under Storage Tests

Agent: HFC-236

Agent Agent/N2

Metal

Metal

Mass

(+ 0.01 g) (± 0.02 MPa) (g) (± 0.02 MPa)

Mass

Storage

(+ 0.1 g)

Mass

/\l LCI Storage

(± 0.1 g)

304 stainless steel I i'\.iL 0.26 18 4.20 57.1 57.1

Nitronic-40 (21-6-9) 325.85 0.25 17 4.20 58.0 58.0

Inconel 625 320.19 0.26 18 4.21 58.2 58.3

6061-T6 aluminum 97.18 0.25 17 4.20 57.8 57.8

4130 alloy steel 302.58 0.25 17 4.22 57.0 57.2 AM 350 stainless 283.61 0.25 17 4.22 58.4 58.7

Cu/Be C82500 311.32 0.24 16 4.20 57.6 57.8

13-8 Mo stainless 283.96 0.24 16 4.21 56.6 56.6

Agent: halon 1301

Metal

Metal

Mass

(± 0.01 g)

Agent Agent/N2

Pressure f± 0.02 MPa)Mass

(g)

Pressure (± 0.02 MPa)

Mass

Before

Storage

(+ 0.1 g)

Mass

After Storage

(± 0.1 g)

304 stainless steel 272.72 1.46 113 4.22 128.3

Nitronic-40 (21-6-9) 324.55 1.46 113 4.24 142.6 142.7

Inconel 625 318.41 1.45 112 4.27 143.3 143.6

6061-T6 aluminum 97.73 1.41 109 4.24 141.0 141.2

4130 alloy steel 301.56 1.36 102 4.13 141.3 141.3 AM 350 stainless 284.83 1.34 101 4.19 136.6 136.8

Cu/Be C82500 323.26 1.32 99 4.15 137.1 137.4

13-8 Mo stainless 289.68 1.32 99 4.19 134.3 134.5
656 SECTION 6. AGENT SHABILm' UNDER STDR.\GE .\NT) DISCKARGE RESIDUE

Table 2 (continued). Agent and Meml .\niount5 in Storage Cylinders for Siabilit}' Under Storage Tests

.\gent: CF3I

Metal

Metal

Mass

(= 0.01 g)

Agent Agent NPresstire (= 0.02 MPa)

Mass

(g)

Pressure (= 0.02 MPa)

Mass

Before

Storage

(- o.Tg)

Mass

.After Storage

(= O.Tg)

304 stainless steel 271.69 0.34 30 4.20 71.5 71.2

Nitromc-40 (21-6-9) 324.22 0.34 30 4.20 70.4 70.4

Inconel 625 318.68 0.34 30 4.20 66.5 66.5

6061-T6 aluminum 98.99 0.34 30 4.20 69.3 67.0

4130 allo>- steel 300.22 0.34 30 4.20 71.7 71.8

.\M 350 stainless 286.06 0.34 30 4.20 71.3 52.4

Cu.Be C82500 323.63 0.34 30 4.20 73.8 72.9

13-8 Mo stainless 29". 59 0.34 30 4.20 "1.2 "1.6
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 657

3.0

2.5

Z 2.0

-r 1.5

< 1.0

Average of Pcst-Exposure Spectra

0.5

0.0

4000 3500 3000 2500 2000 1500 1000 500

Wavenumbers (cm )

Figure 2. Pre- and post-storage spectra for halon 1301.
658 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Average of Pre — Exposure Spectra

U

Average of Post -Exposure Spectra

4000 3500 3000 2500 2000 1500 1000

Wavenumbers (cm )

500

Figure 3. Pre- and post-storage spectra for HFC-236.
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 659

4000 3500 3000 2500 2000 1500 1000 500

Wavenumbers (cm )

Figure 4. Pre- and Post-storage spectra for CF3I.
660 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

and HFC-32/HFC-125. Therefore no peak areas are available for those agents. Judgment for these

agents were based on inspection of the hard-copy plots of the spectra.

Selected peaks in the spectra before and after the 28 day storage were integrated for the

remaining agents. Table 3 lists the initial and final integrated areas of selected peaks for halon 1301,

HFC-236, and CF3I. The integrated peaks were selected on the basis of the following criteria to ensure possible absorbance linearity:

• an absorbance of < 1 • acceptable baseline resolution

• wave number greater than 700 cm"^

For halon 1301, there was a tendency for the peak areas to increase, whereas for CF3I there was

a tendency for the peak areas to decrease. The experiments are very repeatable within a short time

span (as indicated by the small uncertainties for each agent-metal combination). Over longer time

periods, alignment of the FTIR becomes critical and the comparison of initial and final spectra can be

expected to have a larger uncertainty. Thus, the key indicator of degradation should be taken as the

formation of new compounds evidenced in the spectra. No new compounds were identified in any of

the spectra for any of the agents. For CF3I, an initial spectra and two final spectra were available

(one taken after eight days of storage and one after the full 28 day exposure). The peak areas

systematically decreased over the four week exposure, but the change may be within the experimental

measurement error. Since no new compounds were observed in the CF3I spectra, the possible

degradation of CF3I deserves further examination and will be included in the forthcoming longer term

studies.

6.1.3.2 Examination of Metals. The metals were removed from the cylinders, examined, and

then wrapped in aluminum foil for later reference. Changes in the appearance of the metals are given

in Table 4. When the metals from the CF3I cylinders were removed, a dark solid was apparent on

the coupons, then disappeared within seconds. The solid may have been which then sublimed to gaseous I2. For the other metals, visual changes may be due to the prolonged heating of the metal or by interaction with the agent. For the former, eff'ects would be observed for all of the agents (for

example, the Cu/Be C82500 coupons). Interaction with the agents is far more common, with some

metals alTected by only a few agents and others effected by nearly all of the agents. HCFC-22 and

CF3I aff'ected most of the metals, while for others, only the Cu/Be C82500 coupons were aff"ected.

6.2 Agent Residue Study

The objective of this phase of the project was to determine if the agents as received contained

measurable contamination that might have an eff'ect on the results of our experiments. An attempt was made to determine high volatile residue contaminants as well as low volatile residue contaminants. If a residue was discovered some effort was made to determine the source of the contamination.

6.2.1 Experimental Method. The apparatus used for determining residue is shown in Figure 5. The cylinders as received were fitted with hardware that allowed them to be connected to the copper

cooling coil with a commercially available flexible hose commonly used in the refrigerant industry.

Graduated tubes of ca. 2 ml capacity (the bottoms of larger graduated centrifuge tubes) were used to
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 661

Table 3. Initial and Final Integrated Areas of Selected Peaks in the IR Spectra of Halon 1301

Metal

Band

(cm"^)

Integrated Area CflCX 1

Initial Final

304 stainless steel 1351.85-1259.28 15.45+0.01 17.49±0.01 1.13

Nitronic 40 (21-6-9) n 15.56+0.02 18.32+0.02 1.18

Inconel 625 15.72+0.02 18.55±0.04 1.18

6061-T6 aluminum 15.90+0.02 19.80+0.66 1.25 4130 alloy steel 15.32±0.04 18.46+0.04 1.20

AM 350 stainless 15.14+0.03 18.44+0.04 1.22

Cu/Be C82500 15.24+0.03 18.61+0.03 1.22

13-8 Mo stainless 14.98+0.05 17.51+0.11 1.17

304 stainless steel 1984.39-1928.46 1.88 + 0.01 2.08+0.02 1.11

Nitronic 40 (21-6-9)

II 1.90+0.01 2.16±0.02 1.14

Inconel 625 1.95+0.01 2.21+0.01 1.13

6061-T6 aluminum 1.95+0.02 2.35+0.07 1.21

4130 alloy steel 1.88+0.02 2.18+0.01 1.16 AM 350 stainless 1.87+0.01 2.18±0.01 1.17 Cu/Be C82500 1.89+0.01 2.23+0.02 1.18

13-8 Mo stainless 1.85+0.02 2.11-0.01 1.14

304 stainless steel 2302.58-2250.52 7.90±0.03 8.20+0.08 1.04

Nitronic 40 (21-6-9)

n 7.92+0.10 8.63+0.10 1.09

Inconel 625 n 8.10+0.05 8.80+0.02 1.09

6061-16 aluminum II 8.10+0.03 9.30+0.25 1.15

4130 alloy steel

fi 7.78+0.01 8.64 + 0.03 1.11 AM 350 stainless

n 7.77+0.04 8.74+0.06 1.12

Cu/Be C82500 II 7.82+0.01 8.88+0.12 1.14

13-8 Mo stainless

n 7.65+0.04 8.41 ±0.05 1.10
662 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Table 3, (continued). Initial and Final Integrated Areas of Selected Peaks in the IR Spectra of

CF3I.

Band

(cm ')

Integrated Area Cf/C:

1 1

Metal Initial Final

Blank 1312.32 + 1225.54 3.55+0.02 3.30+0.01 0.93

304 stainless steel

tl 3.77+0.02 3.60+0.03 0.95

Nitronic 40 (21-6-9) 3.55+0.03 3.47+0.05 0.98

Inconel 625 3.63+0.01 3.26+0.01 0.90

6061-T6 aluminum 3.68+0.01 3.56+0.01 0.97

4130 alloy steel 3.70±0.02 3.46+0.03 0.94

AM 350 stainless 3.27±0.02^ 3.35+0.04 1.02 Cu/Be C82500 3.69+0.01 3.69+0.04 1.00

13-8 Mo stainless 3.48+0.03 3.39+0.02 0.97

Blank 762.70-717.38 27.02±0.12 22.63±0.04 0.84

304 stainless steel 29.03+0.08 24.38+0.13 0.84

Nitronic 40 (21-6-9)

fi 27.18+0.23 23.75+0.04 0.87

Inconel 625 M 27.58±0.03 22.54+0.04 0.82

6061-T6 aluminum fl 28.21+0.13 24.07+0.05 0.85

4130 alloy steel

II 27.99+0.12 23.47+0.03 0.84

AM 350 stainless

II 24.69+0.12^ 22.95+0.10 0.93

Cu/Be C82500 II 27.64+0.27 24.66+0.01 0.89

13-8 Mo stainless

n 26.13+0.27 22.99+0.02 0.88

^ - Outliers
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 663

Table 3, (continued). Initial and Final Integrated Areas for Selected Peaks in the IR Spectra of

- HFC-236

IVICLCU

Band

Integrated Area Cf/Ci

Initial Final

304 stainless steel IDUo.UO- 14 / J . j4 3.71^ 4.29+0.09 1 1 1 . 16 Mitmnir 40 f91-fS-Q^ 3.68±0.23 4.65^ 1 Of\ 1 . L\J

\x\ccw\fA f\0^ „ 3.78±0.24 4.77^ 1i . lf\ Z.U

UVU i 1 U dlUilliilUili 3.66+0.14 3.37+0.01 n Q9

tuu cuiuy oLCwi 3.78±0.33 3.15+0.01 \J . O J) 3.65+0.14 3.27+0.01 0 QD

V^U/ DC \^O^J\J\J

„ 3.62+0.10 3.28+0.01 nW . Qi-7 1 kjO iviu alallilCoo 3.70+0.26 3.28+0.01

juh- siajniebi) sieci 1 84"^ f^^ 1 7QQ 'JS 2.16^ 2.30+0.04 1 .UD

Nitronic 40 (21-6-9) 2.09+0.14 2.42^ 1.16

Inconel 625 2.11±0.15 2.45=» 1.16

6061-76 aluminum 2.03+0.08 1.93+0.01 0.95

4130 alloy steel 2.09+0.20 1.82+0.06 0.87

AM 350 stainless 2.00+0.09 1.82+0.01 0.91

Cu/Be C82500 1.97+0.06 1.84+0.01 0.93

13-8 Mo stainless 2.02+0.15 1.83+0.03 0.91

^ - Only two usable spectra
664 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Charging Hose

(later replaced w/ Cu tubing)

Cylinder

Valve

iiiiiiiiiiiiiiiiiiiii Agent

Cylinder

Copper Cooling

Coil

Metering Valve

Sample Draw Off

Dewar

ice or

dry ice

Figure 5. Apparatus for determining agent residue.
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 665

collect 1 ml of liquid agent. All tubes were dried at 105 °C for one hour, kept in a desiccator, then

weighed with an analytical balance with 0.1 mg accuracy before filling. Three tubes were used to collect each agent. After the 1 ml samples were collected, the tubes were kept in an ice or dry ice

Table 4. Visible Changes in Metals After 28 Day Exposure at 149°C (300°F)

Ocn Agent

HCFC-22 /3 / / / y

HCFC-124 /

HFC-125 y

HFC-32/HFC-125

HFC- 134a /

FC-218 / / /

HFC-227 / /

FC-31-10 / / / / y

FC-116 / /

FC-318 / / /

HFC-236 y

halon 1301

CF3I / / / / / y

a - / indicates visible change in metal appearance after exposure b - data not available

bath (depending on the boiling point of the agent) to allow controlled evaporation. When all of the

agent had evaporated, the tubes were then placed in a desiccator for at least one hour and weighed on

the previously described balance.

Agents with a boiling point greater than -26 °C were processed with the cooling coil immersed

in ice and the tubes evaporated in ice. Those agents with a boiling point less than -39 °C required

dry ice and acetone in the dewar holding the cooling coil and controlled evaporation of the tubes in a

tray containing dry ice. All agents with the exception of FC-116, which is a gas at room temperature, were removed from the cylinders as a liquid and collected in the graduated tubes as a liquid.

The FC-116 condensed in the cooling coil and was collected as a liquid.

W

O

o

'Sp

CM

I—,o

1—1 o

o

on OmCO

o

in

(N00U

CQ

U

O

00

I
666 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

An effort was made to determine the presence of highly volatile residues when the graduated

tubes were at low temperature after filling and during the controlled evaporation. The formation of

frost prevented any clear observation of the presence of residue.

If measurable residue was present in the graduated tubes after the agents had boiled away and a weighing at ambient was done, the tubes were then heated to 150 °C for one hour, cooled to ambient

in a desiccator,, and reweighed.

6.1.2 Results of residue analyses. Initially, the residue analysis was performed on several of the

agents contained in 1000 ml cylinders. These agents were prepared in small laboratory quantities and no measurable residues were detected. A representative of the manufacturer of these agents indicated

that no low volatile residues should be present because they are distilled prior to filling and are of

high purity.

As production and pilot quantities were received, the residue analysis was performed. Initially,

the HCFC-22 and HCFC-124 samples were found to have as much as 1.3 and 0.3%, respectively, of

a yellow, oily residue. The residue was still visible when the tubes were heated to 150 °C, however,

the amount was reduced to 0.9 and less than 0.1 %, respectively. The manufacturer of these agents was contacted with a lot number from the HCFC-22. One of their research chemists contacted the

plant where this agent had been filled and found that the 13.6 kg non-returnable cylinders do contain

a small amount of oil to protect the inner walls. Since many of the applications of HCFC-22 require

the addition of oil to the agent, no effort was made to remove the oil. In later discussions with

another representative of the manufacturer, this statement was said to be incorrect.

If left in the high pressure flexible filling lines commonly used by the refrigeration industry (and

used to connect the cylinders to the cooling coil in our experiments), interaction between HCFC-22

and HCFC-124 and the black, rubber lining resulted in a yellow, oily residue. If the connection from

the cylinder to the cooling coil was made with a metal tube such as copper, no significant residue was

collected. This dissolution or extraction phenomenon was only seen with HCFC-22 and HCFC-124,

which also contain chlorine.

Table 5 summarizes the results of the low volatile residue analyses for all agents, except FC-318

and CF3I. There was some uncertainty in the weighing of the tubes after they were conditioned to ambient temperature in the desiccator and weighed. In some cases, the weight after collection was

actually slightly less than the initial weight. In the table, a report of a trace amount indicates that the

tubes showed less than 1 mg of residue. Only the HFC- 134a and HFC-236 contained low volatile

residues greater than 1 mg and the weight increase was consistent for the three tubes.

6.3 Conclusions

• No new compounds, observable above the background, were evident after 28 day, 150 °C

exposure of any of the candidate agents.

• A possible decrease in integrated areas for selected spectral peaks was observed only for CF3I. This observation may be within the experimental error, and was without accompanying

formation of new compounds. An observation of a dark solid on metal coupons postexposure for this agent may have been I2 which could result from either degradation or impurity in the original agent. Longer term study is warranted for this agent.
SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE 667

Table 5. Low Volatile Residue Results

Agent Weight % Residue

HCFC-22 Trace

HCFC-124 Trace

HFC- 134a <0.1

HFC- 125 Trace

HFC-32/HFC-125 Trace

HFC-227 Trace

FC-31-10 Trace

HFC-236 0.2

FC-116 Trace

FC-218 Trace

FC-318 Not Measured

CF3I Not Measured

• All agents except HFC- 134a and HFC-236 appeared to be relatively free of low volatile residues. The small amounts detected for the two agents appeared to be oily and might have come from the container walls.

• For the chemicals studied, stability in long-term storage and agent residue should not be

major deciding factors in determining selection of appropriate agents for further study. As

noted above, CF3I could be an exception.

6.4 References

Standard Test Method for Preparing Aircraft Cleaning Compounds, Liquid Type, Witer Basedfor

Storage Stability Testing, ASTM F-1104, Annual Book of ASTM Standard, Volume 15.03, American

Society for Testing and Materials, 1987.

Standard Test Method for Preparing Aircraft Cleaning Compounds, Liquid Type, Solvent Based

for Storage Stability Testing, ASTM F-1105, Annual Book of ASTM Standard, Volume 15.03,

American Society for Testing and Materials, 1987.
668 SECTION 6. AGENT STABILITY UNDER STORAGE AND DISCHARGE RESIDUE

Gann, R.G., Barnes, ID., Davis, S., Harris, J.S., Harris, R.H., Jr., Herron, J.T., Levin, B.C.,

Mopsik, F.I., Notarianni, K.A., Nyden, M.R., Paabo, M., and Ricker, R. E., Preliminary Screening

Procedures and Criteria for Replacements for Halons 1211 and 1301, Natl. Inst. Stand. Technol.,

Technical Note 1278, 1990.

Grob, R.L. (Ed.), Modern Practice of Gas Chromatography, John Wiley & Sons, New York,

1977.

Wendlandt, W.W., Thermal Methods of Analysis, John Wiley & Sons, New York, 1964.

Willis, H.A., Van der Maas, J.H., and Miller, R.G.J. (Eds.), Laboratory Methods in Vibrational

Spectroscopy, John Wiley & Sons, New York, 1987.
669

7. CORROSION OF METALS

Richard E. Ricker, Mark R. Stoudt, James F. Dante, James L. Fink,

Carlos R. Beauchamp and Thomas R Moffat

Materials Science and Engineering Laboratory

7.1 Introduction

Since a fire suppressant which is corrosive to either the materials used for storage and distribution of

the agent or to aircraft materials after deployment would be undesirable, experiments were conducted

to assess the relative corrosivity of the potential fire suppressants to diff'erent metallic alloys used on

aircraft. These experiments were designed to evaluate the propensity of the different agents to cause

failure by any of the possible corrosion failure modes during service or after deployment. In addition

to these experiments, experiments were conducted to develop an electrochemical technique that would

rapidly, and less expensively, evaluate the corrosivity of an agent and the influence of changes in composition, temperature and contaminant levels on the corrosivity of an agent.

Corrosion is usually defined as degradation of the properties of a material as a result of chemical

reaction with the environment. Corrosion can be classified into different categories based on the

material, environment, or the morphology of the corrosion damage. Based on the morphology of the

damage, corrosion can be classified into eight different modes of attack:

1. General Corrosion - General corrosion is the result of chemical or electrochemical

reactions which proceed over the entire exposed surface at about the same rate. General corrosion results in the metal becoming thinner and usually alters the

appearance of the surface. General corrosion could result in failure through lowering

the mechanical strength of components or by reducing wall thickness until leaking

results.

2. Pitting Corrosion - Pitting corrosion is corrosion at a high rate in a small spot on the

surface of a material. It commonly occurs at flaws in coatings applied to prevent

corrosion of the underlying metal, but pitting corrosion also occurs in metals that resist corrosion through the formation of a native oxide, hydroxide or salt film ("passive"

metals or alloys). To prevent corrosion, these films must be continuous and they must be

able to reform rapidly by reaction between the material and the envirormient when the

film cracks, ruptures or becomes damaged through any mechanical or chemical process. Apparently, halide ions either stimulate the breakdown of these films or inhibit their

reformation (repassivation) as pitting is commonly observed in environments that contain

ions of these elements. The time required to initiate pitting depends on temperature and

the composition of the environment typically increasing in a logarithmic fashion as the

halide ion concentration (activity) decreases. Once initiated, the corrosion rate at the

bottom of a pit can be several orders of magnitude greater than that outside the pit and

pitting could cause perforation of storage bottles, etc. or it could initiate cracks that propagate to failure by other mechanisms.
670 SECTION 7. CORROSION OF METALS

3. Crevice Corrosion - When a material is exposed to an environment, there are usually

occluded regions, such as crevices, where the environment in this region does not freely

mix with the bulk environment. The chemistry of the environment in these occluded

regions can differ significantly from that of the bulk environment and this difference can

result in corrosive attack. Crevice corrosion is frequently observed in passivated metals

and alloys when they are exposed to environments that contain halide ions (especially

chloride). Similar to pitting corrosion, a long incubation period may be required to

initiate this attack, but once initiated, this attack can occur at rates that are orders of

magnitude greater than rates outside the crevice. Crevice corrosion could result in failure through leaking of joints, mechanical failure of joints, freezing of joints, or it could lead

to the initiation of cracks that propagate to failure through other mechanisms.

4. Intergranular Corrosion - Virtually all engineering alloys are fabricated by solidifying a melt composed of a mixture of elements. During this solidification process, crystals

nucleate and grow together; forming the solid. As a result of this process, the chemical

composition and the properties of the region between the crystals can differ significantly

from those inside the crystals or grains. Alloy manufacturers have a variety of techniques

that they employ to homogenize alloys and reduce this difference, but complete elimination of this difference is virtually impossible. As a result, certain alloy environment

combinations can result in rapid corrosion of the region between the crystals. This is referred to as intergranular corrosion. When intergranular corrosion occurs, the surface

of the material can appear unattacked, but the mechanical strength of the alloy can

deteriorate slowly or rapidly until the materials break under any applied load or until indi-- vidual grains begin to fall out and the material falls apart.

5. Environmentally Induced Fracture - When a material is exposed to a chemically reactive

environment, the application of a mechanical stress to this material can result in the

formation and propagation of a crack at stresses well below those required for cracking

by purely mechanical processes in the absence of the corrosive environment. Exposure to aqueous solutions, organic solvents, liquid metals, solid metals, and gases have been

found to cause failure in this manner. Depending on the environment and the type of

loading, this type of failure has many different names such as stress corrosion cracking

(SCC), chloride stress cracking, hydrogen embrittlement, liquid metal embrittlement,

solid metal embrittlement, sulfide stress cracking, and corrosion fatigue, but they all consist of crack propagation induced by exposure to an environment. The environments

that cause this type of failure are not always very aggressive and this mechanism of

failure could cause sudden loss of agent or prevent the proper operation of valves.

6. Dealloying - Because different elements have different chemical activities, one element in an alloy may be selectively removed from the surface. This selective leaching of an

alloying element can result in the surface becoming dealloyed and in some cases this dealloyed layer is not limited to the surface and it grows into the solid causing a serious

loss of mechanical strength. In addition to lowering the strength of any component,

dealloyed layers are porous and dealloying could cause loss of agent from a storage

container.

7. Galvanic Corrosion - Because different elements have different chemical activities,

different alloys with different chemical compositions will have different chemical and
SECTION 7. CORROSION OF METALS 671

electrochemical properties in the same environment. Joining of dissimilar metals in a manner such that an electrical circuit is completed through the joint and the environment

will result in increasing the rate of corrosion of the more active metal and reducing the

rate of corrosion of the more noble metal. The extent of the change depends on a variety

of factors including the ratio of the areas of each alloy exposed to the environment, the

potential difference between the alloys, the electrical resistivity of the joint, and the

conductivity of the environment. In the case of metals and alloys exposed to fire suppressants, the conductivity of these suppressants is usually extremely low (unless they become

contaminated with impurities) and this should minimize the damage that galvanic coupling

of dissimilar alloys can cause.

8. Erosion Corrosion - Mixing or stirring of a solution can accelerate the rate of corrosion

either by simply increasing the rate of transport to the surface or by mechanically

damaging or removing surface films that may protect the underlying metal from attack. An increase in the rate of corrosion as a result of relative motion of the environment is termed "erosion corrosion." In general, the relative motion of the agents in the storage

and distribution systems will be quite low; except during deployment, where erosion

corrosion may occur in the nozzles.

Of these eight corrosion failure modes, six are of potential concern for the storage, distribution,

and post deployment corrosivity of fire suppressant agents on aircraft: general corrosion, pitting

corrosion, crevice corrosion, intergranular corrosion, environmental induced fracture and dealloying.

Erosion corrosion may be a concern for nozzles but, the duration of the flow through the nozzle is very short and damage by this mechanisms should be noted in the fire suppressant eflRciency tests if it occurs. Galvanic coupling may influence corrosion damage accumulation, but due to the relatively

low conductivity of these agents, galvanic eff'ects should be at a minimum. As a result, experiments were designed to evaluate the propensity of the diff^erent agents to cause failure by these six corrosion

failure modes.

On board aircraft, metallic materials will be exposed to fire suppressant agents during storage

(storage containers, valves etc.), deployment (distribution tubes, nozzles, etc.) and post-deployment

(structural and engine components exposed to residuals and combustion products) but, the exposures

during deployment are very short and should not be of concern. Of greatest concern for the safe operation of aircraft is the possibility that the fire suppressant might cause attack of the storage

container resulting in loss of agent or preventing the proper operation of the release valve. These

storage containers hold the agent for up to five years at high pressures ( = 5.8 MPa) and temperatures

as high as 150 °C. Very little is known about the corrosion of metals in these environments and

there is virtually no information in the literature on the corrosivity of most of these agents to the

metals used for aircraft storage containers. As a result, corrosion experiments that emulate the

suppressant storage conditions were determined to be a necessary part of the evaluation of new fire suppressant agents.

A second concern for the safe operation of aircraft, is the potential for significant corrosion

damage to the structural metals of the aircraft and aircraft components by residual suppressant or by

suppressant combustion products. Of particular concern would be the formation of halide acids or

salts as a result of the combustion of the agent during fire suppression. For suppressants that are not

a gas at normal temperatures and pressures, residual suppressant left on the surface of metals after suppressant deployment could result in corrosion damage before the agent is cleaned off the surface

(30 days maximum exposure). As a result of this concern, experiments were designed to evaluate the

potential for corrosion damage to metals by suppressant residual and combustion products.
672 SECTION 7. CORROSION OF METALS

7.2 Experimental

There were two main objectives of this initial phase of the study: 1) to rank the corrosivity of the

alternative fire suppressants with respect to storage vessel materials and 2) to rank the corrosion

resistance of the storage vessel materials when exposed to alternative fire suppressants. This allows

the elimination of the most corrosive fire suppressants and the least corrosion resistant storage vessel

material. To achieve these objectives, experiments were designed to evaluate the propensity of the

difi'erent agents to cause failure by the possible corrosion failure modes during service or post-deployment. In addition to these experiments, experiments were also conducted to develop an electro- chemical technique that would rapidly, and less expensively, evaluate the corrosivity of an agent and

the influence of changes in composition, temperature and contaminant levels on the corrosivity of an

agent.

7.2.1 Aircraft Storage System Corrosion. Experiments were designed and conducted to evaluate

the corrosivity of the different potential fire suppressants during normal service in aircraft storage

containers. Since no electrochemical techniques have been developed for the rapid evaluation of the

corrosivity of low conductivity media such as fire suppressants, these experiments had to consist of

exposures that emulate the conditions under which corrosion is expected to occur followed by careful examination and evaluation of the resulting corrosion damage. To evaluate the potential for failure of

the aircraft storage containers by the six potential corrosion failure modes during normal service,

three diff'erent types of exposures tests were conducted to evaluate the potential for failure of the

aircraft storage containers by the six possible corrosion failure modes expected during normal service.

Exposure tests were conducted to determine the changes in mass during the exposures which would,

in turn, assess the rate of formation of corrosion scales or the rate of removal of metallic species by

corrosion. Also, visual and optical microscope examination of these samples after exposure allowed

for the evaluation of the occurrence of pitting, intergranular corrosion and dealloying. Using the mass change and the exposure time data, the average mass change was estimated allowing for a comparison between the agents. The weld/crevice coupon exposures were conducted to allow for the

evaluation of localized corrosion failure modes such as crevice corrosion, weld zone attack, and

intergranular attack in the heat affected zone. The environmental induced fracture resistance of the

alloys in the agents was evaluated by conducting exposure tests on cylindrical samples of the alloys that are loaded in tension by slowly increasing the strain on the sample until failure occurs by either

normal mechanical means or by an environmentally assisted or induced means (the slow strain rate tensile test). By comparing the strain or load required to cause failure in an inert environment to that required to cause failure in the agents, the propensity of each agent to cause environmentally induced

fracture in each alloy can be evaluated.

7.2.1.1 Materials. The following materials were chosen for this study: 304 stainless steel, 13-8

Mo stainless steel, AM 355 stainless steel, stainless steel alloy 21-6-9 (Nitronic 40), 4130 alloy steel,

Inconel alloy 625, CDA-172 copper/beryllium alloy and 6061-T6 aluminum alloy. (The compositions

of these alloys are given in Table 1.) These alloys were selected for these experiments based on

discussions with the Air Force on the materials typically used for aircraft storage containers and

distribution system components.

7.2.1.2 Mass Change (Smooth Coupon) Exposure Experiments. The samples used for the mass change experiments were flat smooth surface coupons of the geometry shown in Figure 1(a).
SECTION 7. CORROSION OF METALS

Table 1 . Composition of the alloys in weight percent

673

Nit Al In 304 CDA 13-8 AM AISI Element 40 6061 625 SS 172 Steel 355 4130

Ni 7.1 61.39 8.26 0.06 8.4 4.23 0.08

Cr 19.75 0.04 21.71 18.11 0.01 12.65 15.28 0.98 Mn 9.4 0.15 0.08 1.41 0.02 0.8 0.51 Mg 1

Si 0.5 0.4 0.09 0.49 0.08 0.04 0.16 0.23 Mo 8.82 0.17 2.18 2.6 0.16 Nb 3.41 N 0.29 0.03 0 0.12 C 0.02 0.02 0.06 0.03 0.12 0.32

Be 1.9

Co 0.11 0.2

Zn 0.25

Cu 0.15 97.9

Fe bal 0.7 3.97 bal 0.06 bal bal bal

Al bal 0.23 0.04 1.11 0.04

7.83 2.70 8.44 7.94 8.23 7.76 7.91 7.85

7,830 2,700 8,440 7,940 8,230 7,760 7,910 7,850

^Nominal Density

After machining, the surfaces were glass bead blasted to remove any remaining surface oxides or

scale and to provide a consistent smooth surface finish (120 grit nominal). Additional preparation

prior to the start of these tests consisted of cleaning and three initial weight measurements for each

individual sample. Each sample was ultrasonically cleaned, first in acetone and then in alcohol, dried

with warm air and immediately weighed. Three separate weighings were taken at approximately 30

second intervals and averaged. This average value was then referred to as the initial weight of the

coupon. The balance used for these weight measurements was self calibrating on start-up to maintain an accuracy to within +10x10"^ g with a reproducibility of no less than +1.51x10"^ g. Representative photographs of the surface of each alloy were also taken prior to the start of testing.

At the start of each exposure test, three coupons of each of the eight alloys were mounted on a

polytetraflouroethylene (PTFE) rod with PTFE spacers between the samples as shown in Figure 2. This was done 1) to separate and 2) to isolate the samples electrically; thus eliminating galvanic

coupling effects. Three samples of each alloy type were tested and PTFE shields were placed

between the different alloys sets to protect the samples from contact with any corrosion products that may form. The samples were next placed in a custom designed PTFE liner and charged into a two
674 SECTION 7. CORROSION OF METALS

W= 25.4 mm (1.0 in) L= 50.8 mm (2.0 in)

t= 1 .6 mm (0.0625 in) d= 6.7 mm (0.265 in)

(a) General corrosion coupon design (ASTM G-01).

X = 50.8 mm (2.0 in)

y = 50.8 mm (2.0 in) R = 4.75 mm (0.1875 in) C = 31.75 mm (1.25 in)

h = 3.2 mm (0.125 in) = 3.2 mm (0.125 in)

(b) Weld/crevice corrosion coupon design.

D = 19 mm (0.25 in)

L = 178 mm (7 in)

I > 1 1 1 ll|l 1 g = 25.4 mm (1 in)

\ I- h -1 h = 57 mm (2.25 in)

s = 19 mm (0.75 in) M6x20 Threads /

(1/4 X 20)

(c) Slow strain rate tensile sample design (ASTM E-8, G-49).

Figure 1. Design of samples used in immersion and slow strain rate tensile tests.
SECTION 7. CORROSION OF METALS

General Corrosion

Samples

1 =AI Alloy AA 6061-16

2 = Cu Alloy CDA-172

3 = AM 355

4 = 13-8 Mo Steel

Weld/Crevice Corrosion

Samples

5 = In 625

6 = 304 Stainless Steel

7 = AISI 4130 Steel

8 = Nitronic 40

Figure 2. Immersion testing chamber used for general corrosion coupon tests and weld/crevice

corrosion coupon tests.
676 SECTION 7. CORROSION OF METALS

liter pressure vessel for testing. The vessel was then attached to a mechanical vacuum pump and

evacuated for a minimum of 30 minutes.

The mass of agent required to produce a pressure of approximately 5.86 MPa at 150 °C was

determined by using the mass balance shown in Equation 1

:

Mr - P,V^ + pV (1)

where Mj is the total mass of the system, p, is the density of the liquid agent at 20 °C, V, is the

volume of the liquid agent at 20 °C, Pg is the density of the vapor phase at 20 °C, and Vg is the

volume of the vapor at 20 °C. A computer program based on Equation 1, the available thermodynamic data, and the ideal gas law was developed to calculate the required mass. To facilitate charging of the agents, the test vessels were chilled in a bath of either ice and water, or dry ice and

alcohol, depending on the temperature necessary to maintain the liquid phase. The vessels were then

placed on a balance that allowed weighing of the entire vessel and determination of the mass of the

agent added continuously during filling. Upon completion of the charging step, the vessels were

placed in proportionally controlled, calrod-type heaters that kept the temperature constant at 150+1 °C for the 25 day exposure period.

At the conclusion of the testing period, the heaters were turned off and the vessels were allowed

to cool naturally to ambient temperature. After the agent was released, the coupons were extracted

and immediately re-weighed using the same procedure that was used for the initial weight measurements. The average of those three measurements was then referred to as the final weight of the

coupon. Representative photographs of the surfaces were again taken and compared to those of the

initial condition.

7.2.1.3 Localized Corrosion (Weld/Crevice Coupon) Exposure Experiments. The samples

used for these experiments were designed to provide specific information about changes in the nature

of the corrosion in a crevice and in a weld heat afi'ected zone. The geometry of these is shown in Figure 1(b). The testing procedure used for these experiments was identical to the procedure used for the smooth coupon mass change experiments except that the samples were not weighed before and

after the exposures.

At the conclusion of the 25 day test period, the samples were removed from the test chambers

and the samples were split along the fusion boundary using a water cooled slitting saw. The internal

surfaces of the crevice were then analyzed by eye and in an optical microscope. The resulting

appearance within the crevice region of the three samples for each agent/alloy combination was then

given a numerical rating based on the appearance of the corrosion damage within the crevice region. In addition, the relative performance of each alloy in a given agent was determined by placing all of

the samples tested in that agent on a table and making direct (optical) comparisons. Then, the relative corrosivity of the different agents for a given alloy was determined by placing all of the samples of

that type of alloy on a table and making direct (optical) comparisons between the damage that resulted

in each agent.

7.2.1.4 Environmental Induced Fracture (Slow Strain Rate Tensile Tests). The stress corrosion cracking susceptibilities of the eight alloys in the replacement candidates were evaluated

using the slow strain rate (SSR) tensile test technique. This technique was selected because it generates mechanical properties data that reveals any interactions that may have occurred between that alloy and the testing environment within a relatively short time frame.
SECTION 7. CORROSION OF METALS 677

All of the samples used for these experiments were machined with the tensile axis parallel to the

rolling direction of the plate stock (Figure 1(c)) and tested in the "as received" condition. The sample

preparation consisted of m.easurement followed by degreasing in acetone and alcohol. The vessels

used for these experiments, Figure 3, were 250 ml volume autoclaves with a similar design to those

used for the exposure testing, except that these vessels were modified so that load could be applied to the tensile specimen in situ under constantly maintained environmental conditions (5.86 MPa at 150+1 °C). The test vessels were evacuated and charged with agent in the same manner as the

exposure test vessels, but due to the number of mechanical tests required per agent, and the reduced

capacity of the test vessels, the appropriate mass was based solely on the ideal gas law. This

approach was selected over the mass balance because it produced a slightly lower final pressure and

the number of moles of the agent was held constant for each test; regardless of the agent used.

The mechanical tests were conducted using a computer controlled slow strain rate testing system

which operated at a constant crosshead speed of 0.0254 ^m/sec. The computer was configured to sample and record the applied load, the crosshead displacement and the elapsed time at 90 second

intervals. After failure, the agent was released, the vessels were allowed to cool to ambient

temperature, and the samples were removed from the vessel and stored in a desiccator until analyzed.

The fracmre surfaces were cut from the broken SSR samples for analysis. The influence of the

agents on the ductility of the alloys was determined from reduction in area (RA) measurements

performed on the fracture surfaces with a optical measuring microscope with a ± 0.5x10'^ m

resolution. Scanning electron microscopy was performed on selected samples to allow evaluation of

the mechanism of crack propagation based on the fracture morphology. The results of these

experiments were used to formulate a ranking of the potential for failure by stress corrosion cracking

for each of the alloys in a replacement candidate.

7.2.2 Post Deployment Experiments. When a fire suppressant is applied to a fire, the metals in the

aircraft engine nacelle or dry bays may become covered with deposits of the fire suppressant or fire suppressant combustion products. As a result, an evaluation of the relative corrosivity of the expected

combustion product layers was included in this investigation.

7.2.2.1 Materials and Sample Preparation. The alloys used for these experiments were the same as those used for the aircraft storage and distribution system experiments. These alloys were

used because samples were readily available and because this matrix included representative alloys of

the different types that might be exposed to fire suppressant combustion products. The samples used were identical to those given in Figure 1(a) except that these samples did not contain the hole in the

center. As before, the samples were glass bead blasted to the equivalent of a 120 grit finish. The

samples were cleaned in acetone in an ultrasonic cleaner and then treated with acid solutions in an

ultrasonic cleaner to remove existing corrosion products and surface films. The acid solutions were

rinsed off of the metal surfaces using distilled water. Following is a list of the acid mixtures used for surface cleaning or pretreatment:

• 6061-T6. The surface was treated by exposing it to concentrated HNO3 for three minute

intervals until the mass loss for this exposure time became constant.

• CDA-172. The surface was treated by exposing it to 250 ml of a solution made by adding

25 ml of H2SO4 to double distilled water for three minute intervals until the mass loss for

this exposure became constant.
678 SECTION 7. CORROSION OF METALS

Figure 3. Slow strain rate tensile test chamber.
SECTION 7. CORROSION OF METALS 679

• IN 625. The surface was treated by exposing it to 250 ml of a solution made by adding

37.5 ml of HCl to double distilled water for three minute intervals until the mass loss for

this exposure became constant.

• AISI 4130, AM 355 and 13-8 Mo Steel. A solution was prepared by adding 80 ml of a 35

weight % solution of 2 butyne 1 ,4 diol to 60 ml of HCl and then adding double distilled water to make one liter of solution. The AM355 and 13-8 Mo surfaces were exposed to this mixmre for three minute intervals until the mass loss for this exposure became constant. The 4130 surface was exposed to this mixmre for one minute intervals until the mass loss

for this exposure became constant. The 4130 steel became tarnished by the water rinse used

to clean the surface.

• Nitronic 40 and 304 Stainless Steel. One liter of solution was prepared by adding 100 ml

of HNO3 to 20 ml to double distilled water. The samples were exposed to this mixture for

five minute intervals until the mass loss for this exposure became constant.

7.2.2.2 Combustion Product Surface Layer Emulation. The fire suppressant candidates can be classified into two categories: those that contain fluorine and/or other halides and those that contain sodium bicarbonate. Fire suppressants that contain halides, of which fluorine is usually the

worst, will produce halide ions during combustion that will deposit on metallic surfaces and form

halide acids when they react with water. Sodium bicarbonate (NaHC03) will decompose on heating

to form sodium carbonate (Na2C03), and on combustion to form sodium hydroxide (NaOH). As a

result, there are essentially three diff'erent combustion product surface films of interest: (1) fluoride

ions, (2) NaHC03/Na2C03 and (3) NaOH.

To produce a surface film rich in fluoride ions, the surface of the samples were first sprayed

with ASTM artificial seawater then sprinkled with sodium fluoride (NaF) powder while the surface was still wet. While the films produced by combustion will be much more complex than this simple

analog, this surface film should contain the most critical corrosive species that can be expected

following deployment: water, chloride and fluoride. The ASTM seawater step was included to simulate the conditions in a marine atmosphere, which should be the worst case situation, as well as

to provide water for reaction with the sodium fluoride. To simulate the sodium bicarbonate and

sodium carbonate (NaHC03/Na2C03) surface film, the samples were sprayed with ASTM artificial seawater and sprinkled with a 50/50 mixmre of sodium bicarbonate/sodium carbonate. The 50/50

mixmre was created by heating 100 g of sodium bicarbonate + 1.5 weight % Si02 overnight at 150 °C to form sodium carbonate and then mixing this with another 100 g of sodium bicarbonate +

1.5 weight % SiOj- To simulate the NaOH rich surface film that may result from the deposition of

NaHC03 combustion products, the samples were sprayed with ASTM artificial seawater and then

sprayed with a 0.1 M solution of NaOH. The samples were placed into their exposure chambers

immediately following these surface treatments.

7.2.2.3 Exposure Environment. After deployment of a fire suppressant, a layer of combustion

products may remain on the exposed surfaces for up to one month before the surfaces are cleaned to remove the surface film. During this time, the environment which will cause corrosion will be the

ambient air environment and the hydration of the surface film that may result from absorption of water from the ambient environment (corrosion in condensate droplets as a result of thermal cycling was not included in this smdy). As a result, the exposures were conducted in three different glass

desiccators each with the a different fixed relative humidity (RH). To keep the relative humidity

constant, the desiccant in each of the desiccators was replaced with a saturated salt solution which has
680 SECTION 7. CORROSION OF METALS

a different equilibrium water vapor partial pressure or relative humidity. For a low relative humidity

exposure, a saturated solution of potassium acetate (KC2H3O2) was included in the desiccator for which the equilibrium relative humidity is 20%. For a moderate relative humidity environment, a

saturated sodium dichromate dihydrate (Na2Cr207-2H20) solution was used in the desiccator for which the equilibrium relative humidity is 52%. And for a high relative humidity, a saturated

solution of sodium sulfate decahydrate (Na2SO4 l0H2O) for which the equilibrium relative humidity is 93% was used.

7.2.2.4 Testing Procedure. The samples of the different alloys were exposed under nine

diflFerent conditions: three different pretreatments and three different humidities. The samples were weighed between the cleaning and the surface pre-treatment steps. After the application of the ASTM

artificial seawater, the samples were weighed again in order to determine the mass present on the

surface. Then, after the NaF, 50/50 NaHC03/Na2C03 mixture, or the NaOH was applied, the

samples were weighed again. Then, the specimens were exposed for 30 days with one of each alloy

and pretreatment in each humidity. Following the exposure, the coupons were weighed, then rinsed

in double distilled water, dried and reweighed. Finally, the corrosion product films were chemically

removed by using the same cleaning procedure used during the sample preparation and the samples were weighed again to determine the total mass loss due to corrosion.

7.2.3 Development of an Electrochemical Measurement Technique. Electrochemical techniques can be used to measure the interfacial properties directly associated with the rate of corrosion. These

measurements allow for a relatively fast determination of the corrosion resistance of a metal or alloy

in solution as compared to the time required for acquisition of similar data from exposure tests in the same environment. The development of such a technique would provide a more thorough assessment

of the corrosion rates of the storage vessel materials in the alternative fire suppressants.

Figure 4 is a schematic representation of the most simple equivalent electrical circuit that can be

used to describe the processes that occur in an electrochemical cell. In this figure, represents the

resistance between the sample and the measurement point that arises due to the resistivity of the

electrolyte, C^j] is the capacitance of the sample electrode solution interface and R^^ is the charge

transfer resistance that arises from the kinetics of the electrochemical reactions occurring at the

electrode surface (e.g., corrosion). If is high, as in the case of the extremely high resistivity

associated with fire suppressants, significant error or artifacts in the measured interfacial properties can be expected. Therefore, a technique must be developed to either 1) reduce the solution resistivity

(without compromising the relative order of corrosivity of the agents) or 2) allow for the correction of

the solution resistivity and eliminate the observed artifacts. Experimentally, the best approach for obtaining useful data is done by applying a low amplitude sine wave to the sample and measuring the

cell impedance as the frequency of the sine wave is varied. By examining Figure 4, it can be seen

that for a low frequency sine wave the impedance of this circuit becomes R^ + R^.^ while for a high

frequency signal the impedance of the circuit becomes just R^. This technique is known as ac impedance or electrochemical impedance spectroscopy (EIS).

For EIS, one normally uses a potentiostat and three electrode cell. Figure 5. The sample, a

reference electrode and a counter electrode make up the three electrodes in this system. This allows

for the measurement of the potential of the sample with respect to a reference electrode with a fixed

and well known potential on the hydrogen electrode (thermodynamic) scale while applying current with a counter electrode whose potential can vary with current without creating measurement errors. However, as the resistivity of the media increases, problems arise due to the appearance of artifacts which result in considerable measurement errors. Investigators have shown that a two electrode

system is more suited for EIS measurements in high resistivity media than a three electrode system
SECTION 7. CORROSION OF METALS

Sample Potential,

E (vs. RE) ~1 I Reference

Electrode

Capillary

to Reference

CEIectrode

Cell Current, I

Cell Potential, Vceij

X=0 X=d

Interface

F-7,--^

i

I

' —I

I

I Equivalent Circuit for Electrode Measurements

Counter

(Auxilliary)

Electrode

External

Measurement

Instrument

A/WV

External Current

Source

X=D

Figure 4. Simple equivalent circuit for an electrochemical interface.
682 SECTION 7. CORROSION OF METALS

Computer

Frequency

Response

Analyzer

Potentiostat

Electrochemical

Cell

Measured

Z = Z + Z" Imp = Real + Imaginary

Measured

E

Desirred

E

-WW

A^

, ^ Reference Counter

Electrode * ^ i-i * ^

/c^ , X Electrode Electrode

(Sample)

Figure 5. Simple diagram of a three electrode cell used for electrode impedance measurements.
SECTION 7. CORROSION OF METALS 683

(Halsal, 1992, and DeSouza, 1986). As a result, a two electrode cell capable of placing two

relatively large area samples in close proximity and conducting EIS measurements at different

temperatures and pressures was designed and fabricated. This cell design is shown schematically in Figure 6 and consists essentially of a autoclave similar to that used for the exposure tests modified by

the addition of electrical feed troughs in the head along with a micrometer for adjusting the distance

between the samples. Inside the chamber, two identical disk samples are held in place by PTFE

holders one of which is moved by the micrometer. To evaluate the functionality of this chamber,

measurements were made on samples in an environment similar to that used by the previous

investigators (Halsal, 1992, and DeSouza, 1986).

7.2.3.1 Materials. Disk shaped samples 38 mm in diameter and 12.7 mm thick were machined

out of 304 stainless steel and polished to a 0.05 fim finish. Before placing the samples into the cell,

they were rinsed with reagent anhydrous ethyl alcohol. Anhydrous ethyl alcohol with approximately

0.005 % water and a conductivity on the order of 10'^ ohms" ^ cm"^ was used initially for the testing

environment because this is similar to the environment used by the previous investigators. It should

be remembered that the conductivity of the halon alternatives is orders of magnitude lower than this value. However, results of tests performed in ethyl alcohol should reveal the effectiveness of the cell in measuring electrochemical parameters in low conductivity media. The addition of nonaggressive

charge carriers to the fire suppressants may increase their conductivity up to or above this level.

7.2.3.2 Procedure. After the cell was assembled, ethyl alcohol was introduced. The initial separation distance of the electrodes was measured. The impedance of this cell at different frequencies from 65,000 Hz to 0.1 Hz was determine with a 10 mV amplitude signal. From this data, the

solution resistance was determined. The separation of the electrodes was increased by 1 mm and

another scan was taken. This procedure was repeated until the electrodes were 4 mm from there

original position. The electrodes were then reset to their original position and the entire procedure was repeated.

7.3 Results

7.3.1 Aircraft Storage and Distribution System Corrosion.

7.3.1.1 Mass Change (Smooth Sample) Exposure Experiments. The mean mass changes

observed for three samples of each of the alloys on exposure to each of the agents for 25 days at 150 °C are given in Table 2, the standard deviation determined for these measurements are given in

Table 3 and the original measurements are given in Appendix A. On examination of these tables, it can be seen that all of the mass changes are relatively small and most are mass increases (positive

values in Table 2). The mass change rate was estimated from these mass change measurements from

the relationship

(2)
SECTION 7. CORROSION OF METALS

ure 6. Electrochemical cell for two electrode impedance measurements in low conductivity

electrolytes at elevated temperatures and pressures.
SECTION 7. CORROSION OF METALS 685

Table 2. Mean Mass Change for the three samples of Each Alloy tested in Each Agent After 25

Day Immersion at 150 °C [mg]

Environmcnt Nit Al In 304 CDA 11 J3-8o AM AISI Agent OUj . 40 6061 625 SS 172 Steel 355 4130 Avg. Dev.

HCFC-22 0.176 0 458 -0.329 0.168 3.064 -0.033 -0 087 2 847 0 783 1.362 HCFC-124 0.092 0 219 -0.267 -0.037 1.474 -0.086 -0 206 0 319 0 189 0.557

FC-31-10 0.247 0 533 -0.037 0.583 2.663 0.357 0 219 1 102 0 708 0.858

HFC-227 0.009 0 128 -0.056 0.141 1.973 0.069 0 007 0 297 0 321 0.676

HFC-125 -0.011 0 199 0.031 0.136 0.890 0.151 0 029 0 673 0 262 0.333

FC-116 0.006 0 240 0.026 0.147 1.509 0.210 0 139 0 403 0 335 0.491

HFC-134a 0.013 0 131 -0.103 0.069 0.229 0.088 -0 014 0 262 0 084 0.122

HFC-236 0.024 2 479 -0.063 0.078 0.866 0.107 0 007 0 343 0 480 0.861

FC-C318 0.008 1 183 0.018 0.199 0.431 0.169 0 006 0 310 0 290 0.392

FC-218 0.153 0 373 0.051 0.090 0.768 0.279 0 150 0 381 0 281 0.232

HFC-32/125 0.300 0 587 0.146 0.241 2.263 0.492 0 311 0 227 1 960 3.851 NaHC03 0.396 1 038 0.079 0.262 2.182 0.800 0 259 0 890 0 738 0.677

Alloy Avg. 0.118 0 631 -0.042 0.173 1.526 0.217 0 068 1 597

Std. Dev. 0.137 0 675 0.138 0.152 0.911 0.243 0 150 3 152

Positive Values = Scaling

Negative Values = Mass Loss

where aM is the mass change, A is the total exposed area of the sample and t is the exposure time in

days and are given in Table 4. This relationship estimates the average mass change rate over the

exposure period and assumes that the mass change rate was essentially constant over the entire

exposure period.

The magnitudes of these mass changes (Table 4) are small but, they are, for the most part, larger

than that required to be statistically significant for the measurement equipment and measurement

techniques employed. Since each sample was weighed three times before and after each exposure, the

standard deviation for the mass measurement process can be estimated and it was found to be less than 1.4x10"^ g. If one conducts a t-test for the significance of the change in mass based on this estimated standard deviation, one will find t values that indicate significant mass change at the 99%

confidence level in all but 20 of the 96 agent/alloy combinations. However, when the mass change

measurements over the three samples of each alloy tested in each agent are averaged, a larger

standard deviation is determined because this standard deviation includes scatter from the natural

corrosion processes and sample handling during preparation and removal from the exposure chamber

in addition to the normal measurement scatter. In Table 5, a minimum and maximum mass change

rate at the 90% confidence level is given for each alloy/agent combination. The values in this table were determined from the standard deviation estimated from the three samples of each alloy tested in each agent and estimated standard deviations for the surface area and exposure time using standard
686 SECTION 7. CORROSION OF METALS

Table 3 . Standard Deviation of mass change measurements for the three samples of each alloy

tested in each environment (mg)

Environment Nit 40

A! 6061In 625

304 SS CDA

172

13-8

Steel AM

355

AISI 4130

Agent

Avg. HCFC-22 0 079 0.113 0 088 0 057 0 331 0.043 0.050 1 280 0.255 HCFC-124 0 045 0.091 0 048 0 040 0 294 0.072 0.103 0 099 0.099

FC-31-10 0 040 0.079 0 028 0 035 0 107 0.070 0.046 0 173 0.072 HFC-227 0 034 0.043 0 018 0 025 0 159 0.063 0.020 0 047 0.051 HFC- 125 0 121 0.117 0 044 0 013 0 018 0.014 0.007 0 091 0.053

FC-116 0 124 0.087 0 025 0 033 0 157 0.012 0.063 0 013 0.064 HFC- 134a 0 021 0.025 0 024 0 002 0 039 0.061 0.048 0 091 0.039 HFC-236 0 024 1.895 0 030 0 030 0 028 0.015 0.090 0 043 0.269

FC-C318 0 040 0.602 0 065 0 054 0 072 0.014 0.109 0 035 0.124

FC-218 0 083 0.048 0 Oil 0 050 0 032 0.021 0.044 0 080 0.046

HFC-32/125 0 033 0.324 0 022 0 037 0 217 0.140 0.035 0 122 0.116 NaHC03 0 297 0.501 0 022 0 047 0 235 0.263 0.042 0 168 0.197

Alloy Avg. 0 078 0.327 0 035 0 035 0 141 0.066 0.055 0 187

error propagation techniques (Bevington, 1969). Many of these ranges include zero which indicates

that we cannot conclude at this level of confidence that the corrosion rate is a value other than zero. On the other hand, scale formation and spalling off of this scale will increase measurement scatter which could result in an estimated mass change rate range that includes zero. As a result, each

agent/alloy combination was given a numerical rating based on the absolute value of the maximum or minimum indicated mass change rate in Table 4 rather than on the mean mass change rate. Table 6

gives the absolute mass change rate range for each numerical rating and the equivalent penetration

rate for each alloy for mass loss at the maximum rate for the range, while Table 7 gives the score for each agent/alloy combination. Because on cannot determine the amount of metal reacting for mass

increases without information on the chemical composition of the scale, mass increases and decreases were given the same weighting for this analysis. It is believed that this rating basis will result in conservative scores for each agent/alloy combination especially since scale formation usually results in a corrosion rate that decreases with time as the scale inhibits mass transport.

Visual examination confirmed the presence of surface films on many of the samples. These films were very thin and could be detected by a color change in the samples. For these exposures, the

films may be the result of reaction with residual gasses, decomposition products or the agents themselves. The formation of salt films when metal are exposed to organic solvents has been observed by

other investigators. Mass increases are not uncommon during immersion testing and usually

descalling techniques are employed to removed these scales for the determination of the remaining

metal. Considerable research has been conducted into the development of techniques for the removal

of the scales which form in aqueous solutions without attacking the underlying metal. Even with this
SECTION 7. CORROSION OF METALS 687

Table 4. Mean Mass Change Rate for Each Alloy in Each Agent for 25 Day Immersion at 150 °C

[mg/(square meter"day)

Environment Nit Al In 304 CDA 13-8 AM AISI Agent Std. 40 6061 625 SS 172 Steel 355 4130 Avg. Dev. HCFC-22 2.512 6.550 -4 706 2 401 43 847 -0.477 -1 240 40.731 11.202 19.481 HCFC-124 1.320 3.132 -3 816 -0 525 21 097 -1.224 -2 941 4.563 2.701 7.965 FC-31-10 3.529 7.631 -0 525 8 346 38 108 5.103 3 132 15.771 10.137 12.278 HFC-227 0.127 1.828 -0 795 2 Z J J 0.986 AU 4.245 4.593 9.677 HFC- 125 -0.159 2.846 0 445 1 940 12 734 2.162 0 413 9.634 3.752 4.770 rL-1 Id U.U /y 0 366 2 099 21 589 3.005 1 987 5.771 4.791 7.021 HFC- 134a 0.191 1.876 -1 479 0 986 3 275 1.256 -0 207 3.752 1.206 1.752 HFC-236 0.350 35.468 -0 906 1 113 12 385 1.526 0 095 4.912 6.868 12.321 FC-C318 0.111 16.931 0 254 2 846 6 168 2.416 0 079 4.436 4.155 5.610 FC-218 2.194 5.342 0 731 1 288 10 986 3.990 2 146 5.453 4.016 3.326 HFC-32/125 4.292 8.394 2 083 3 450 32 384 7.043 4 451 3.243 28.040 55.097 NaHC03 5.660 14.849 1 129 3 752 31 224 11.447 3 704 12.734 10.562 9.690

Alloy Avg. 1.684 9.023 -0 601 2 476 21 836 3.103 0 976 22.852

Std. Dev. 1.953 9.652 1 973 2 181 13 039 3.480 2 151 45.098

Positive Values = Scaling

Negative Values = Mass Loss

research, these techniques usually remove some unreacted metal or leave scales of their own which

induce an error in this mass measurement. Normally, this mass error is small and insignificant

compared to the mass of the scale being removed but, the films on the samples in this study are very

thin and the composition could be very different from that formed in aqueous solutions. As a result,

descalling was not attempted as a means of estimating the thickness of the scales. Evaluation of the relative performance of the alloy/agent combinations was hampered by the

comparison of mass loss and mass gain measurements. Typically, mass loss measurements are evaluated by assuming that all of the lost mass was the result of corrosion and that no corrosion

products were left on the surface to create errors in this determination. Then, the quantity of metal

reacting and rate of the reactions can be calculated directly from the mass loss. Similarly, mass gain

measurements can be evaluated if the reaction and the reaction products are known and it is assumed

that none of the scale spalls off and is lost to the environment. However, when these are not known,

evaluation contains greater uncertainties, but it is clear that the relative magnitude of mass change is

still related to the reaction rate between the metallic species and the environment.

7.3.1.2 Localized Corrosion (Crevice/Weld Coupon) Exposure Experiments. The

weld/crevice coupons were cut open and the corrosion damage in the crevice region was examined.

Based on these examinations, the results of these experiments were evaluated in three different ways.

First, the type of damage observed for each alloy/environment combinations was categorized into one

of five different categories based on the appearance of the overlap region: (1) no visible evidence of

attack, (2) light discoloration of surface, (3) moderate discoloration with slight attack visible, (4)

substantial discoloration and some attack visible, and (5) attack throughout the overlap region.
688 SECTION 7. CORROSION OF METALS

Table 5. Maximum and minimum estimated mass change rate at 90% confidence level based on the

three measurements for each alloy/agent combination and standard statistical error estimation techniques (mg((m^*day))

Environment Nit 40Al 6061In 625

304 SS CDA

172

13-8

Steel AM

355

AISI 4130

HCFC-22

4.418

0.606

9.290

3.810

-2.587

-6.825

3.770

1.031

51.856

35.837

0.570

-1.524

-0.037

-2.443

71.660

9.801

HCFC-124

2.416

0.223

5,325

0.939

-2.654

-4.977

0.452

-1.501

28.206

13.987

0.511

-2.959

-0.444

-5.438

6.955

2.171

FC-31-10

4.506

2.553

9.530

5.732

0.163

-1.213

9.199

7.494

40.691

35.524

6.795

3.412

4.231

2.033

19.957

11.584

HFC-227

0.942

-0.687

2.867

0.790

-0.351

-1.238

2.613

1.425

32.087

24.383

2.506

-0.535

0.585

-0.394

5.387

3.103

HFC- 125 2.758

-3.076

5.661

0.030

1.518

-0.627

2.265

1.614

13.162

12.306

2.497

1.827

0.581

0.246

11.824

7.445

FC-116

3.068

-2.909

5.532

1.335

0.959

-0.228

2.892

1.305

25.389

17.790

3.295

2.714

3.499

0.476

6.094

5.448

HFC-134a

0.694

-0.312

2.480

1.271

-0.898

-2.059

1.032

0.939

4.219

2.331

2.724

-0.212

0.944

-1.358

5.945

1.559

HFC-236

0.932

-0.233

81.249

-10.312

-0.190

-1.622

1.835

0.391

13.070

11.699

1.877

1.175

2.276

-2.085

5.960

3.865

FC-C318

1.072

-0.849

31.476

2.387

1.835

-1.326

4.155

1.536

7.900

4.437

2.752

2.081

2.709

-2.550

5.288

3.583

FC-218

4.188

0.200

6.503

4.180

0.990

0.472

2.490

0.085

11.752

10.219

4.508

3.473

3.202

1.090

7.375

3.531

HFC-32/125

5.098

3.487

16.217

0.571

2.619

1.546

4.348

2.552

37.627

27.141

10.416

3.670

5.305

3.598

6.183

0.303

NaHC03

12.843

-1.524

26.941

2.757

1.653

0.605

4.880

2.623

36.903

25.544

17.800

5.094

4.711

2.697

16.800

8.669

Table 8 gives the results of this analysis, and the only alloy to show category five attack was the 4130

steel which received a five rating in HCFC-22 and HCFC-124. This table also gives the average

rating for each agent and ranking of the agents based on this average rating. The lowest rated agents on this basis were HCFC-22 (agent average = 3. 143), sodium bicarbonate (agent average = 3. 125), and
SECTION 7. CORROSION OF METALS 689

Table 6. Mass Change Rate Values Used in Rating Alloy/Agent Behavior in Mass Change

Measurements and equivalent penetration rate based on alloy nominal density

Score

Mass Chg Rate mg/(m^»d)Equivalent Penetration Rate for Upper Boundary

/im/year

Lower Bound

Upper

Bound

Nit 40Al 6061

In 625

304 SS CDA

172

13-8

Steel AM355

AISI 4130

1 0 0.1487 0.0069 0.0201 0.0064 0.0068 0.0066 0.0070 0.0069 0.0069

2 0.1487 0.923 0.0431 0.1249 0.0399 0.0425 0.0410 0.0434 0.0426 0.0429

3 0.923 5.00 0.2332 0.6764 0.2164 0.2300 0.2219 0.2353 0.2309 0.2326

4 5.00 25.00 1.166 3.382 1.082 1.150 1.109 1.177 1.154 1.163

5 25.00 125 5.831 16.909 5.409 5.750 5.547 5.883 5.772 5.816

6 125 600 27.99 81.16 25.96 27.60 26.63 28.24 27.70 27.92

7 600 3,000 139.9 405.8 129.8 138.0 133.1 141.2 138.5 139.6

8 3,000 15,000 699.7 2029.1 649.1 690.0 665.7 706.0 692.6 697.9

9 15,000 75,000 3498 10146 3246 3450 3328 3530 3463 3490

10 75,000 OO OO OO OO 00 00 OO OO 00

(Rating is based on absolute value of mass change rate and penetration rate is only valid for mass loss.)

HFC-227 (agent average = 2. 750). The best rated agents by this technique were HFC-236 (agent

average = 1.750), FC-31-10 (agent average = 1.875), and HFC-134a (agent average = 2.0). However, one should keep in mind that agent averages can be adversely influenced by a single rating, when in

service, that specific alloy/agent combination could be avoided and the agent would serve satisfactorily. On this basis, only five of the twelve agents did not have an alloy where no evidence of attack was observed (a one rating) and all of the agents had at least three alloys where only slight discoloration was observed (a two rating). As a result, the crevice/weld immersion experiments show that while there is a clear difference in the corrosivity of the agents, it should be possible to find container

alloys for each of the candidates.

While a visual rating of the appearance is a fair estimate of the relative corrosivity of each agent,

there are numerous cases where several alloy/agent combinations received the same rating and there is no indication of which agent was more or less corrosive for a specific alloy. As a result, all of the

weld/crevice samples of each alloy were compared to each other and the agents were ranked based on

the appearance of the corrosion damage in the crevice region for each alloy. Each agent was given a

score between one and twelve for each alloy where one showed less evidence of corrosion damage

than two and so forth. This analysis allows the identification of the least corrosive agent for each

alloy and the results are given in Table 9. The data in this table indicates the same general trend as Table 8 with the low ranking agents in Table 8 receiving most of the lower rankings; but since there

is no relationship between the scores from one agent to another {e.g., an eight for one agent/alloy

combination may have less attack than a much lower rating for another combination), this table should
690

Table 7.

SECTION 7. CORROSION OF METALS

Mass Loss Rate Evaluation of Agents and Alloys based on 25 Day Immersion at 150 °C

Environment Nit 40

Al 6061In 625

304 SS CDA

172

13-8

Steel AM

355

AISI 4130

Agent Ave

Std. Dev.

Agent

Rank

HCFC-22 3 4 3 3 5 2 3 5 3.50 1 .07 9 HCFC-124 3 3 3 2 4 3 3 3 3.00 0.53 7

FC-31-10 3 4 2 4 5 4 3 4 3.63 0.92 11

HFC-227 1 3 2 3 5 3 1 3 2.63 1 .30 1

HFC-125 2 3 2 3 4 3 2 4 2.88 0.83 4

FC-116 1 3 2 3 4 3 3 4 2.88 0.99 4

HFC- 134a 2 3 3 3 3 3 2 3 2.75 0.46 3

HFC-236 2 5 2 3 4 3 1 3 2.88 1.25 4

PP-P'<18 \^ J L o 1 4 2 3 4 J 1 -J 1 1 Q 1

FC-218 3 4 2 3 4 3 3 4 3.25 0.71 8

HFC-32/125 3 4 3 3 5 4 3 3 3.50 0.76 9 NaHC03 4 4 3 3 5 4 3 4 3.75 0.71 12

Alloy Avg. 2.33 3.67 2.42 3.00 4.33 3.17 2.33 3.58

Std. Dev. 0.98 0.65 0.51 0.43 0.65 0.58 0.89 0.67

Alloy Rank 1 7 3 4 8 5 1 6

Score

12

34

5

67

89

10

Freq

6

14 47

22

70

0

0

0

0

No evidence of attack (t<0.741, 50% CI). Some inconclusive evidence of attack (0.741(50% CI) < t < 4.6 (99% CI)). Conclusive evidence of attack (t>4.6, 99% CI). Corrosion of little concern, (less than = 1 /xm/year, 3^n\ for Al alloy). Corrosion may be a concern

Corrosion rate should be considered in design Corrosion about 1 to 5 mils per year Rapid Corrosion

Very rapid corrosion Extremely rapid corrosion

only be used for identifying the relative corrosivity of each agent to a specific alloy. However, it is worthy of note that sodium bicarbonate (NaHC03) received a twelve rating for half of the alloys and

a ten or higher rating for six of the eight alloys.

Perhaps a more valuable comparison would be a comparison of the performance of the alloy in a given agent. Table 10. As with Table 9, this ranking was developed by a side by side comparison of

the crevice/weld region of the samples only this time the different alloys tested in the same agent were compared. This ranking allows for identification of the alloy which is more resistant to attack

in the crevice and weld heat aff'ected zone by each agent. Again, a low rating in one

alloy/environment combination may be considerably better than a higher rating in another combination
SECTION 7. CORROSION OF METALS 691

Table 8. Weld/Crevice Coupon Visual Classification of Attack After 25 Day Immersion at 150 °C

Environment

Nit Al In 304 CDA 13-8 AM AISI Agent Std. Agent 40 6061 625 ss 172 Steel 355 4130 Avg. Dev. Rank

HCFC-22 2 2 2 3 4 3 3 5 3.14 1.07 12 HCFC-124 2 2 1 T 3 3 3 5 2.63 1.19 8 FC-31-10 1 1 2 2 2 2 2 3 1.88 0.64 2 HFC-227 2 2 2 3 3 3 3 4 2.75 0.71 10 HFC- 125 2 2 2 2 3 2 3 2 2.25 0.46 5 FC-116 2 2 1 2 3 2 3 2.13 0.64 4 HFC- 134a 1 1 1 3 4 2 2.00 1.07 3 HFC-236 2 2 1 1 2 2 2 2 1.75 0.46 1 FC-C318 2 3 2 2 2 2 2 4 2.38 0.74 6 FC-218 3 4 2 2 2 2 2 4 2.63 0.92 8 HFC-32/125 3 2 1 2 2 2 3 4 2.38 0.92 6 NaHC03 2 3 2 2 4 4 4 4 3.13 0.99 11 Alloy Avg. 2.00 2.17 1.58 2.08 2.58 2.58 2.75 3.50 2.42

Std. Dev. 0.60 0.83 0.51 0.51 0.79 0.67 0.75 1.09 0.45

Alloy Rank 1 4 1 3 5 5 7 8 0.45

Classification Criteria

1 =No visible staining 2 = Light Staining in overlap region

3=Moderate staining with slight attack visible in overlap region. 4 = Substantial staining with moderate attack in overlap region. 5=Heavy attack throughout overlap region.

so comparisons between environments should not be done with this data, but this tabulates the results of a direct visual comparison of the different alloy's performance in each agent.

7.3.1.3 Environmental Induced Fracture Experiments (Slow Strain Rate Tensile Tests).

Slow strain rate tensile tests result in three parameters that can be used as an indicator of environmental induced fracmre. First, the ultimate tensile strength (UTS) is a measure of the fracture strength of

the sample and it is determined from the maximum load observed during the tensile test according to the relationship:

UTS=^ (3)

where P^^^^ is the maximum load supported by the sample during the test and A^ is the initial crosssectional area of the gage section of the sample. Environmental interactions that lower the strength of

the sample either by corrosion, cracking or other interaction will result in a change in this parameter.

Environments can also interact with deformation occurring at surfaces and at crack tips in the sample

to limit ductility and cause cracking. So, a second indicator of environmental induced fracture is total strain to cause failure (e^) which is determined from the relationship

(4)
692 SECTION 7. CORROSION OF METALS

Table 9. Weld/Crevice Coupon Visual Ranking of Agents for Each Alloy After 25 Day Immersion

at 150 °C

Environment INlt In jU4 LUA too 13-0 AM AISI Agent Std. 40 6061 625 SS 172 Steel 355 4130 Avg. Dev.

HCFC-22 4 4 7 12 11 9 9 11 8.38 3.11 HCFC-124 3 3 3 8 10 7 8 12 6.75 3.45

FC-31-10 1 1 6 4 1 2 3 5 2.88 1.96

HFC-227 8 7 8 11 9 8 5 7 7.88 1.73

HFC-125 5 5 9 7 8 5 10 4 6.63 2.20

FC-116 7 6 4 9 4 11 4 2 5.88 3.00 HFC- 134a 2 2 2 10 7 10 11 1 5.63 4.31

HFC-236 6 9 5 1 2 6 6 3 4.75 2.60

FC-C318 9 11 10 3 6 3 1 9 6.50 3.78

FC-218 11 12 11 2 5 1 2 8 6.50 4.57

HFC-32/125 12 8 1 5 3 4 7 10 6.25 3.69 NaHC03 10 10 12 6 12 12 12 6 10.00 2.62

1 = Best Performance

where If is the total change in sample length during the experiment and 1^ is the initial gage length.

This value includes the elastic and plastic strains required to cause failure. Another means for measuring the ductility of the sample is the reduction in area (RA) which is a third means for indicating environmentally induced fracture. The RA is determined by measuring the area of the

fracture surface after the test is completed and the relationship

RA

--^

(5)

where Af is the cross sectional area of the fracture surface and A^ is the initial gage section cross

sectional area. This measurement includes only the plastic deformation required to cause failure and

for some alloys is a better measure of environmental interactions than the strain to failure. Typically,

these parameters are analyzed by forming a ratio of the value observed in the environment to observed in a reference environment (usually an inert environment). Another approach to analysis is to estimate the statistical significance of the difference between the mean determined for the parameter

in the inert reference environment and the mean determined for the agent. The significance of this difference is determined by dividing the difference in the means by the standard deviation estimated

for that parameter for that specific alloy and temperature in this type of test.
SECTION 7. CORROSION OF METALS 693

Table 10. Weld/Crevice Coupon Visual Ranking of Alloys for Each Agent After 25 Day Immersion

at 150 °C

Environment Nit Al In 304 CDA 13-8 AM AISl 40 6061 625 SS 172 Steel 355 4130

HCFC-22 1 2 3 4 7 5 6 8 HCFC-124 3 2 1 4 7 5 6 8

FC-31-10 2 1 3 4 5 6 7 8

HFC-227 3 1 2 4 6 7 5 8

HFC-125 3 4 2 1 7 5 8 6

FC-116 2 3 1 4 6 8 5 7

HFC- 134a 2 3 1 4 6 7 8 5

HFC-236 4 3 1 2 5 7 6 8

FC-C318 2 7 3 1 6 5 4 8

FC-218 4 7 2 1 6 3 5 8

HFC-32/125 6 3 1 2 4 5 7 8 NaHC03 2 4 3 1 8 5 6 7

Alloy Avg. 2.83 3.33 1.92 2.67 6.08 5.67 6.08 7.42

Std. Dev. 1.34 1.97 0.90 1.44 1.08 1.37 1.24 1,00

1 = Best Performance

Tables 1 1 and 12 gives the average UTS determined for each alloy in air at room temperature,

argon at 150 °C, and each agent at 150 °C. The ratio of the average UTS in each agent to the

average UTS in Ar at the same temperature is given in Table 13. In this table, it can be seen that for 87 of the 96 agent alloy combinations the UTS ratio is 1.00+.05. For the nine outside this range, two have a higher average UTS in the agent than in Ar. Of the seven with a lower average UTS in the agent, only two is more than 10% below the value determined in Ar: 304 and 4130 steel in FC-C318. Table 14 gives the results of an estimate of the significance of the differences. A

significant decrease in the average UTS may be an indication of cracking, but it could also be due to corrosion reactions reducing effective cross-section or a flaw in the sample. An increase in the

average UTS is unusual and this indicates that sample/environment interactions are inhibiting

deformation and fracture, but it could also indicate a interference between the corrosion products

being generated on the sample and the seal of the autoclave through which the sample must slide. The most important point is that no agent caused significant changes in all alloys indicating that materials for the containment and reliable distribution of any of these candidates can be identified. The ductility results for the SSR tests are given in Tables 15-20. Table 15 gives the average

strain to failure for each alloy/environment combination, laboratory air at room temperature and Ar at 150 °C. Table 16 gives the ratio of the average strain to failure for each alloy in each environment

to that observed in the same alloy in Ar at the same temperature and Table 17 gives the statistical
694 SECTION 7. CORROSION OF METALS

Table 11. Average Ultimate Tensile Strength in Each Agent at 150 °C as Compared to Ar at 150

°C and Air at 25 °C (MPa)

Environment

Nit 40

Al 6061In 625

304 SS CDA

172

13-8

Steel AM

355

AISI 4130

Air, 25 °C 727 341 957 773 763 1117 1005 673

Ar, 150 °C 610 240 927 667 874 1136 969 647 HCFC-22 602 232 892 674 855 1155 938 624 HCFC-124 624 235 937 648 860 1027 945 665

FC-31-10 619 235 910 649 841 1176 935 673

HFC-227 610 235 911 667 865 1156 928 650

HFC-125 611 235 924 650 847 1154 917 669

FC-116 599 221 880 638 830 1163 922 629

HFC- 134a 618 225 932 653 851 1144 929 688

HFC-236 612 245 930 633 861 1168 935 648

FC-C318 597 289 924 550 851 1182 950 558

FC-218 624 235 944 681 842 1170 928 663

HFC-32/125 624 225 886 622 843 1138 919 631 NaHC03 623 252 947 643 852 1217 953 674

Alloy Avg. 614 239 918 642 850 1154 933 648

Std. Dev*. 16 16 22 37 22 44 15 35

Alloy Avg. Average of the means determine for each of the agents.

Std. Dev.* jtie "Alloy Standand Deviation" was estimated from the Ar at 150 °C tests

analysis of the significance of this difi^erence. In Table 16, it can be seen that measured strain to failure increased for most of the alloy/environment combinations indicating the deformation was

easier in these environments than in Ar. Environmentally induced cracking, or stress corrosion

cracking (SCC), is usually indicated by a reduction in ductility. Of the 96 alloy environment

combinations, the strains to failure of three of the alloy/environment combinations were more than

10% less than that observed in Ar: Nit 40 in HCFC-124 (0.87), Nit 40 in FC-31-10 (0.88), and Nit

40 in HFC-236 (0.88). However, these strain to failure measurements are based on load frame

displacement measurements taken by a transducer outside of the autoclave during the experiment. As

a result, the reduction in area measurements which are based on measurement of the fracture surface

are in an optical microscope after the experiment are a more reliable indication of changes in

ductility. Table 18 presents this data for each alloy in laboratory air at 25 °C, Ar gas at 150 °C and

each agent at 150 °C. Table 19 gives the ratio of the average reduction in area measured for each al- loy/agent combination to Ar at 150 °C and the statistical significance of this diff'erence is given in Table 20. In Table 19, it can be seen that the reduction in area is reduced by more than 5% in 11 of
SECTION 7. CORROSION OF METALS 695

Table 12. Average Ultimate Tensile Strength in Each Agent at 150 °C as Compared to Ar at 150

°C and Air at 25 °C (ksi)

Environment Nit Al In 304 CDA 13-8 AM AISI 40 6061 625 SS 172 Steel 355 4130

Air. 25 = C 105.4 49.4 138.8 112.1 110.7 162.0 145.7 97.6

At, 150°C 88.5 34.8 134.4 96.7 126.7 164.7 140.5 93.8 HCFC-22 87.3 33.6 129.4 97.7 123.9 167.5 136.0 90.4 HCFC-124 90.5 34.1 135.9 93.9 124.8 149.0 137.1 96.4

FC-31-10 89.7 34.1 131.9 94.2 122.0 170.6 135.7 97.7

HFC-227 88.5 34.0 132.2 96.7 125.4 167.6 134.6 94.3 HFC- 125 88.6 34.1 134.0 94.3 122.9 167.4 133.0 97.0

FC-116 86.9 32.1 127.6 92.5 120.3 168.6 133.7 91.2

HFC-134a 89.7 32.6 135.2 94.7 123.4 165.9 134.7 99.8

HFC-236 88.8 35.5 134.9 91.8 124.8 169.5 135.6 94.0

FC-C318 86.6 41.9 134.1 79.8 123.4 171.4 137.8 80.9

FC-218 90.5 34.1 136.9 98.8 122.1 169.6 134.5 96.1

HFC-32/125 90.5 32.6 128.4 90.2 122.2 165.0 133.2 91.5 NaHC03 90.3 36.5 137.4 93.3 123.6 176.5 138.1 97.8

Alloy Avg. 89.0 34.6 133.2 93.2 123.2 167.4 135.3 93.9

Std. Dev.* 2.3 2.3 3.2 5.3 3.2 6.4 2.2 5.1

Alloy Avg. = Average of the means determine for each of the agents.

Std. Dev.* = The "Alloy Standard Deviation" was estimated from the Ar at 150°C tests

the 96 alloy/agent combinations and is reduced by more than 10% in only four alloy/agent combinations: Al 6061 in HCFC-124 (0.85), Al 6061 in NaHCOj (0.83), SS 304 in NaHCOj (0.83), and

AISI 4130 in NaHC03 (0.87). Since three of these five are with the same agent, NaHC03, this agent

may have problems with this failure mode.

To summarize the results of the SSR tensile tests, Table 21 was assembled by taking the highest

(worst) score received by each agent/allo>' combination in Tables 14, 17, and 20 (UTS, strain to

failure and RA) and assigning a number to each agent/alloy combination. Since the agent/alloy scores given in these tables are the absolute value of the difference between the mean for the alloy in the

agent and Ar divided by the standard deviation for this alloy and parameter at 150 °C, this rating is a measure of the significance of any postulated environmental influence on the deformation and

fracmre of the agent/alloy combinations. If the maximum deviation was less than one standard

deviation, the agent/alloy combination was assigned a value of one, if it was greater than one standard

deviation, but less than two, it was assigned a two and so on. In Table 21, it can be seen that 65 of

the 96 agent/alloy combinations received a one or two rating. Every agent had at least one alloy with

a one rating and only two agents had only one alloy with a rating of one.
696

Table 13.

SECTION 7. CORROSION OF METALS

Average Ultimate Tensile Strength Ratios [(Mean for Envir.)/(Mean for Hot Ar)]

Environment Nit Al In 304 CDA 13-8 AM355 AISI Agent

40 6061 625 SS Steel SS 4130 Avg.

HCFC-22 0 99 0 97 0.96 1 01 0 98 1 02 0 97 0 96 0.96 HCFC-124 1 02 0 98 1.01 0 97 0 98 0 90 0 98 1 03 0.98

FC-31-10 1 01 0 98 0.98 0 97 0 96 1 04 0 97 1 04 0.99

HFC-227 1 00 0 98 0.98 1 00 0 99 1 02 0 96 1 00 0.98

HFC- 125 1 00 0 98 1.00 0 97 0 97 1 02 0 95 1 03 0.98

FC-116 0 98 0 92 0.95 0 96 0 95 1 02 0 95 0 97 0.98 HFC- 134a 1 00 0 94 1.01 0 98 0 97 1 01 0 96 1 06 0.98

HFC-236 1 00 1 02 1.00 0 95 0 99 1 03 0 97 1 00 0.99

FC-C318 0 98 1 20 1.00 0 83 0 97 1 04 0 98 0 86 0.96

FC-218 1 02 0 98 1.02 1 02 0 96 1 03 0 96 1 02 0.99

HFC-32/125 1 02 0 94 0.96 0 93 0 96 1 00 0 95 0 97 0.99 NaHC03 1 02 1 05 1.02 0 96 0 98 1 07 0 98 1 04 1.01

Alloy Avg. 1 00 1 00 0.98 0 97 0 97 1 01 0 96 0 98

On the negative side, no alloy received only one's and two's in all of the agents and no agent

received only one's and two's for all of the alloys. The best performance by an alloy was demonstrated by the 304 stainless steel alloy which received a one in all except for three of the environ- ments. The best performing agents were HFC-236 and HFC-125 with FC-31-10 coming in third. However, each of these agents received a three or a four rating on at least one alloy. Since the alloys

have diflferent chemical compositions, surface films and corrosion susceptibilities, this is not

surprising, but it indicates the importance of conducting experiments of this type.

7.3.1.4 Summary of Aircraft Storage and Distribution System Corrosion Experiments. To

combine the performance of the different agent/alloy performances on the three different types of tests that comprise this part of the investigation, each agent/alloy combination was given a rating between one and ten based on its performance in that particular test. Table 22 reviews the results for each

agent/alloy combination in the three types of experiments that comprise this part of this investigation:

Mass Change Exposure Experiments (Exp.), Localized Corrosion Exposure Experiments (Crevice)

and Environmental Induced Fracture (SSR). To allow for ready comparison between these results, a

similar rating scheme was adopted for each of the three different types of tests used in this part of this study. Basically, this scheme assigned a numerical score between one and ten to each agent/alloy

combination investigated in that particular test. Whenever possible, a numerical rule was used for the

rating scheme with the magnitude of the standard deviation used to distinguish between one, two, and

three while order of magnitude steps were used from six to 10. The interpretation of the meaning of

the numerical ratings is given at the bottom of Table 22. Basically, a one indicates no discernable

evidence of attack, a five indicates attack that may be a problem, and a ten indicates extremely rapid
SECTION 7. CORROSION OF METALS 697

Table 14. Score for the Average Ultimate Tensile Strength j [(Mean for Envir.)-(Mean for Hot Ar)]/(

Alloy Std. Dev.) |

Environment Nit Al In 304 CDA 13-8 AM355 AISI Agent

40 6061 625 SS 172 Steel SS 4130 Avg.

HCFC-22 0 49 0 53 1 57 0 18 0 88 0.43 2 06 0 67 0.85 HCFC-124 0 89 0 30 0 47 0 52 0 61 2.45 1 55 0 50 0.91

FC-31-10 0 55 0 34 0 79 0 48 1 49 0.91 2 20 0 76 0.92

HFC-227 0 00 0 35 0 70 0 00 0 41 0.45 2 70 0 09 0.59

HFC-125 Au U4 U J

1

n

U

1 11 J UA A 40 11 0.41 J 41 0 62 0.82

FC-116 0 68 1 22 2 13 0 79 2 01 0.60 3 11 0 51 1.38

HFC- 134a 0 52 0 97 0 23 0 38 1 03 0.17 2 64 1 17 0.89

HFC-236 0 14 0 29 0 13 0 93 0 59 0.73 2 23 0 02 0.63

FC-C318 0 80 3 14 0 12 3 19 1 04 1.04 1 24 2 55 1.64

FC-218 0 89 0 33 0 76 0 39 1 46 0.76 2 73 0 45 0.97

HFC-32/125 0 87 1 00 1 87 1 23 1 42 0.05 3 33 0 46 1.28 NaHC03 0 79 0 74 0 92 0 65 0 99 1.83 1 07 0 78 0.97

Alloy Avg. 0 56 0 79 0 82 0 77 1 09 0.82 2 36 0 71

attack. A rating above four indicates more than just superficial corrosion and a better understanding

of the attack responsible for this score should be developed before putting these combinations in service.

From Table 22, it can be seen that some agent/alloy combinations that did well in one test did

poorly in another. For example, AISI 4130 steel performed well in the SSR tests in all of the agents,

but did poorly in most of the other tests. On the other hand, some of the alloys which apparently

retain passivity in the agents, had some indications of problems in the SSR tests such as 304 stainless

steel in FC-C318. The results of these tests can be combined into a single overall rating number for each agent/alloy combination by either averaging the results of the tests, Table 23, or by taking the worst rating achieved for the three tests as the overall rating. Table 24. Since a good rating on two

of the tests could bring an unacceptable rating on one of the tests into the acceptable range for the

overall rating based on averaging, it was decided that an overall rating which is the worst rating

received on the three different tests was the best way to evaluate the overall performance of an

agent/alloy combination. Table 24. By examining this table, it can be seen that by this rating scheme

the best rated agents are HFC- 134a with HFC-227, HFC-125 and HFC-236 not far behind. None of

the agents received an agent average rating for the eight alloys higher than five, but NaHC03 was

over four. Also, all of the agents had at least two alloys with a three or lower rating. That is, none

of the agents were so corrosive that an alloy could not be identified for aircraft storage and distribution systems, but clearly some agents were less aggressive than others and some agent/alloy

combinations may be undesirable. The origin of the differences in the corrosivity of the agents could

be due to reactions between the metals and the agents or it could be due to different contamination or
698 SECTION 7. CORROSION OF METALS

Table 15. Average Strain to Failure in Each Agent at 150 °C (Except lab air which was at room

temperature) (%)

Environment Nit Al In 304 CDA 13-8 AM355 AISI 40 6061 625 SS 172 Steel SS 4130

Lab Air 49.9 13.0 44.4 30.3 13.0 5.9 10.7 8.3

Hot Ar 41.5 7.9 41.0 13.6 10.1 6.2 9.2 7.9 HCFC-22 37.8 8.9 37.7 12.5 12.4 7.5 9.8 8.9 HCFC-124 36.0 10.5 43.7 14.1 12.4 7,9 9.1 8.2

FC-31-10 36.7 10.1 37.2 12.4 12.9 6,5 7.8 8.0 HFC-227 37.8 10.7 42.6 13,6 11.7 6.9 9,6 7.7

HFC-125 40.4 9.8 43.4 15.1 11.6 6.7 9,8 8.1

FC-116 40.5 9.6 46.5 17.1 12.2 6.9 9.1 8.8 HFC- 134a 37.7 9.4 46.5 12.7 10.1 7.7 9.2 7.8

HFC-236 36.3 9.7 41.5 16.9 9.7 7.0 9.7 8.2

FC-C318 41.7 8.5 43.1 36.0 12.8 8.1 9.1 9.6

FC-218 41.0 9.1 45.0 15.0 11.7 6.6 9.0 9.5

HFC-32/125 39.4 9.2 46.2 17.4 11.6 7.2 9.6 8.5 NaHC03 42.5 7.5 46.0 17.4 11.6 10.5 9.5 7.5

CF3I 40.6 8.6 47,8 21.3 11,0 7.4 8.0 6.8

Alloy Avg. 39.3 9.3 43.4 16.8 11,6 7.4 9.2 8.3

Std. Dev.* 3.0 1.2 2.6 5.5 1.6 0.9 0.6 0.7

Alloy Avg. Average of the means determine for each of the agents.

Std. Dev.* The "Alloy StandanJ Deviation" was estimated from the Ar at 150 °C tests

residual levels in the agents. If the corrosivity is due to contaminants and residuals, then changes in processing or storage that alters these levels would also alter the corrosivity of the agents. Additional

testing will be required to determine the cause of the observed corrosion behaviors and the role of

contaminants and residuals in determining the corrosivity.

7.3.2 Post-Deployment Corrosion.

7.3.2.1 Literature Review of Corrosion of Metals in Halide Ions. All of the fire suppressants

investigated in this study except for NaHC03, contain fluoride and produce fluoride ions and HE

during combustion. Since other halogens are also incorporated in these, and other potential fire suppressants, other halogen acids will be produced by combustion of these suppressants. Therefore,

the post-deployment corrosion damage to aircraft materials will depend on the corrosivity of surface

films containing these species on the metals of the aircraft frame, engine components, etc.
SECTION 7. CORROSION OF METALS 699

Table 16. Average Strain to Failure Ratio in Each Agent at 150 °C [(Agent)/(Hot Ar)]

Environment INK in 304 CDA 13-8 AM355 AISI Agent 40 6061 625 SS 172 Steel SS 4130 Avg. HCFC-22 0.91 1,12 0.92 0 92 1 23 1 20 1.06 1 13 1.01 HCFC-124 0.87 1.32 1.06 1 04 1 23 1 28 0.99 1 03 1.10

FC-31-10 0.88 1.27 0.91 0 91 1 28 1 05 0.85 1 01 0.96 HFC-227 0.91 1.34 1.04 1 00 1 17 1 11 1.05 0 97 1.13 HFC- 125 0.97 1.24 1.06 1 11 1 16 1 08 1.06 1 02 1.14

FC-116 0.98 1.21 1.13 1 26 1 21 1 11 0.99 1 12 1.06

HFC-134a 0.91 1.19 1.13 0 93 1 00 1 24 1.00 0 99 1.09

HFC-236 0.88 1.23 1.01 1 25 0 97 1 13 1.05 1 03 1.13

FC-C318 1.00 1.07 1.05 2 65 1 27 1 30 0.99 1 22 1.13

FC-218 0.99 1.14 1.10 1 11 1 17 1 06 0.97 1 20 1.14

HFC-32/125 0.95 1.16 1.13 1 28 1 15 1 16 1.04 1 08 1.11 NaHC03 1.03 0.94 1.12 1 29 1 15 1 70 1.03 0 95 1.21

Alloy Avg. 0.94 1.19 1.05 1 23 1 17 1 20 1.01 1 06

Most engineering alloys that exhibit good corrosion resistance do so due to the formation of a "passivating" surface film. Aluminum alloys and stainless steels are excellent examples of alloys

based on active elements that behave in a relatively noble manner due to the formation of continuous

protective corrosion product films. These films are usually oxides, hydroxides or oxide-hydroxide

mixtures. These surface films may be crystalline and they may be precipitated out of the solution

adjacent to the metals surface, but it is generally believed that the more rapidly grown, amorphous

films are more protective than precipitated crystalline films. No matter what the nature of the surface

film, it is generally found that halogen ions in the environment and on these films tend to destabilize

the protective passivating films. Presumably, this is because the soluble metal-halide salt is thermodynamically preferred over the relatively insoluble metal-oxide or hydroxide compounds in the

passive film at the passive film to environment interface. This is a very simplified view which

ignores many kinetic issues in passive film growth and breakdown, but many of the essential details

in the relationship between halogen ions and passivity breakdown can be understood based on this simple model: (1) increasing halide ion activity reduces passive film stability, (2) the more aggressive

halide is the more electronegative (F" > Cr> Br"> !"> At"), (3) pH influences stability, (4)

electrode potential influences stability and (5) metal-halide salt solubility should exhibit an influence.

Aluminum alloys are the most common metallic alloy type used in airframe construction and the

most susceptible to pitting by halide ions of the metals of concern so this review will focus on the

effect of halide ions on the corrosion behavior of these alloys. Aluminum is a relatively active

element and aluminum alloys corrode as rapidly as mass transport to the surface will allow if the

passive film becomes unprotective for any reason. As a result, aluminum alloys are not used in environments where passivity is not maintained (e.g., pH<4 or pH> 10). However, pitting, crevice
700 SECTION 7. CORROSION OF METALS

Table 17. Score for the Average Strain to Failure | [(Mean for Envir.)--(Mean for Hot Ar)]/( Alloy

Std. Dev.) I

Environment Nit Al In 304.00 CDA 13-8 AM355 AISI Agent 40 6061 625 SS 172 Steel SS 4130 Avg.

HCFC-22 1 22 0 78 1 29 0 20 1 44 1 47 0 97 1 51 1.11 HCFC-124 1 80 2 10 1 02 0 10 1 46 2 04 0 17 0 42 1.14

FC-31-10 1 57 1 79 1 47 0 22 1 76 0 34 2 26 0 10 1.19

HFC-227 1 21 2 26 0 61 0 01 1 03 0 81 0 69 0 31 0.87 HFC- 125 0 36 1 55 0 90 0 28 0 97 0 60 0 97 0 23 0.73

FC-116 0 32 1 35 2 10 0 64 1 32 0 81 0 18 1 39 1.01

HFC- 134a 1 23 1 22 2 11 0 16 0 02 1 78 0 00 0 14 0.83

HFC-236 1 69 1 49 0 18 0 61 0 21 0 96 0 79 0 41 0.79

FC-C318 0 07 0 49 0 78 4 11 1 68 2 19 0 19 2 61 1.51

FC-218 0 17 0 95 1 51 0 27 1 03 0 44 0 44 2 45 0.91

HFC-32/125 0 66 1 06 1 98 0 69 0 95 1 15 0 59 0 90 1.00 NaHC03 0 35 0 40 1 90 0 71 0 94 5 09 0 48 0 66 1.32

Alloy Avg. 0 89 1 29 1 32 0 67 1 07 1 47 0 64 0 93

corrosion or stress corrosion cracking may result in environments where aluminum is normally

passive if halide ion concentrations exceed critical values. As discussed above, when aluminum is exposed to solutions containing different types of halide

ions, passive film breakdown and pitting should vary with the electronegativity of the halide ion.

Chloride, bromide and iodide follow this trend, but fluoride was found to be less aggressive to aluminum alloys than chloride and bromide (Davies and Prigmore 1984). Since aluminum salts formed with fluoride are less soluble than other aluminum halide salts, several investigators have

postulated that this difference is due to the formation of different types of salt films (Zotikov,

Bakhmutova et al. 1974; Davies and Prigmore 1984). In general, aluminum alloys have excellent

atmospheric corrosion resistance and, in most atmospheres, the corrosion rate of aluminum decreases

with time as the passive film become thicker (Fink and Boyd 1970; Bailey, Porter et al. 1976).

However, cyclic wetting and drying and the presence of aggressive species in the atmosphere can

accelerate attack by damaging the passivating film. Marine and coastal environments with their high

humidity and particulates rich in chloride ions can be particularly aggressive. Also, gases such as sulfur dioxide (SO2) chlorine (CI2) and hydrogen chloride gas have been found to be aggressive for aluminum alloys (Bailey, 1976; Scully, 1990(a)).

7.3.2.2 Post-deployment Corrosion Test Results. The results of the post-deployment

corrosion tests are shown in Figures 7 through 14 for Alloy 6061-T6, AISI 4130, 13-8 Mo steel, AM

355, 304 stainless steel, Nitronic 40, Inconel 625, and CDA 172 (Cu-Be) respectively. The data in these figures are for 30 days of exposure at three different humidities (20, 52 and 93 % RH) following
SECTION 7. CORROSION OF METALS 701

Table 18. Average Reduction in Area in Each Agent at 150 °C (Except lab air which was at room

temperature) (%)

Environment Nit Al In 304 CDA 13-8 AM355

40 6061 625 SS 172 Steel SS 4130

Lab Air 78 1 50.9 73 7 76.7 66.3 69.4 48 6 52.1

Hot Ar 79 2 42.3 69 5 67.8 27.5 60.0 48 6 50.8 HCFC-22 79 3 40.2 65 7 67.9 38.7 61.5 49 3 50.4

HCFC-124 79 6 35.8 65 7 7L7 36.6 56.2 48 3 48.7

FC-31-10 80 2 40.0 64 2 70.2 37.8 60.5 51 4 47.6

HFC-227 79 5 39.1 68 9 69.8 43.3 61.1 48 7 48.2

HFC-125 80 4 39.4 67 3 68.7 40.0 61.9 48 6 49.2

FC-116 79 6 4L1 69 6 71.9 40.5 62.5 48 5 49.2

HFC-134a 80 3 39.0 69 9 69.0 31.1 59.6 50 4 49.6

HFC-236 79 1 40.7 65 7 72.6 32.9 60.1 48 6 49.1

FC-C318 79 5 40.9 70 3 68.6 36.3 61.1 47 5 49.5

FC-218 86 2 42.2 69 3 70.8 31.4 59.5 47 2 47.5

HFC-32/125 79 5 40.9 67 0 72.5 37.9 62.3 48 5 50.1 NaHC03 80 1 35.2 68 4 56.3 33,8 58.1 51 5 44.3

Alloy Avg. 80 2 39.8 67 8 69.1 36.0 60.3 49 1 48.8

Std. Dev.* 2.3 2.3 3.2 5.3 3.2 6.4 2.2 5.1

Alloy Avg. Average of the means determine for each of the agents. Std. Dev.* The "Alloy Standard Deviation" was estimated from the Ar at 150°C tests

three different surface pretreatments: (1) artificial seawater plus NaF, (2) artificial seawater plus

NaHC03/Na2C03 and (3) artificial seawater plus NaOH. The mass changes (after washing with

distilled water) for this exposure period are plotted in these figures as a function of the relative humidity of the exposure environment. In these figures, it can be seen that for most of the alloys and

pretreatments, no particular surface pretreatment was consistently worse than the others, that there is no clear trend in the magnitude of the mass change with relative humidity, and that, except for three

alloys, the mass changes were almost always small mass losses. A mass loss indicates the formation of corrosion products that were removed by washing in

distilled water prior to weighing of the samples. These products may be either ions in thin electrolyte

layers on the surface or solid corrosion products that are dissolved or physically removed by the

washing process. In either case, these measurements indicate corrosion of the underlying alloy.

However, for all of these alloys the mass changes were < 2 mg for these exposures which corre- sponds to a corrosion rate of < 0.024 grams per meter squared per day (0.24 MDD) which is a

relatively low corrosion rate. The three exceptions were: aluminum alloy 6061-T6, AISI 4130 steel,

and the Cu-Be alloy CDA- 172. These alloys exhibited mass increases of a much larger magnitude
702

Table 19.

SECTION 7. CORROSION OF METALS

Average Reduction in Area Ratio in Each Agent at 150 °C [(Agent)/(Hot Ar)]

Environment Nit AI In 304 CDA 13-8 AM355 AISI Agent 40 6061 11 79/ z oo A 1 TA 41 jU Avg.

1 .uu u.yj (\ OS u.yJ

1 c\c\

1 .uu 1/11 1.41 1 .Uj 1 .00 0.99 1.04

1 .Ul U.oJ r> OS u.yJ 1 .uo 11 .ii n u.y4 c\ no 0.98 0.96 1.01 PP T 1 in 1 .uz n OS C\u.yz09 1 .Uj 1 T7 1 . J / I .Uz 1 AC 1 .05 0.94 1 .04 HCP T)l u.yz n oQ u.yy 1 .Uj 1 SO 1 .JO 1 .UJ A AA 0.99 0.95 1 .06 HCP rir\^- M'\ 1 ZD 1 .Ul u.yj r\ 07 u.y /

11 .Um

1

1 A S

i .4j 1 C\A1 .U4 A AA u.yy

A AT U.97 1 .05 CP 1 1 ^ 1 .U

1

u.y / 1 .uu 1 HA 1 .Uo 1 /I Q 1 .48 1 AS 1 .Uj A AA U.yy A Arf 0.9 / 1 .06

HFC-134a 1.0! 0.92 1.01 1.02 1.13 0.99 1.04 0.98 1.01

HFC-236 1.00 0.96 0.95 1.07 1.20 1.01 0.99 0.97 1.02

FC-C318 1.00 0.97 1.01 1.01 1.32 1.02 0.97 0.98 1.04

FC-218 1.09 1.00 1.00 1.04 1.14 1.00 0.96 0.94 1.02

HFC-32/125 1.00 0.97 0.97 1.07 1.38 1.04 0.99 0.99 1.05 NaHC03 1.01 0.83 0.99 0.83 1.23 0.97 1.05 0.87 0.97

Alloy Avg. 1.02 0.94 1.00 1.00 1.23 1.05 0.99 1.01

than the small mass losses of the other five alloys and required the use of a much larger scale for plotting the results for these alloys. For aluminum alloy 6061 (Figure 7), the mass changes ranged

from a small mass loss of about 2 mg at 20% RH for the NaOH pretreatment to a mass increase of

about 15 mg at 93% RH for the carbonate/bicarbonate pretreatment. The NaF pretreatment showed mass increases between 2 and 6 mg for these exposures. Apparently, adding NaF to the surface

electrolyte film was not as detrimental to this alloy as the carbonate/bicarbonate treatment. Since the

films were prepared by first spraying with artificial seawater, all of the surface films contained

chloride ions which are typically found to be more detrimental than fluoride ions to aluminum alloys.

While the NaOH pretreatment resulted in only small mass changes, it should be noted that this pretreatment resulted in the formation of small pits on the surface for the low humidity exposure

(20% RH). For the AISI 4130 steel alloy (Figure 8), the carbonate/bicarbonate and the NaOH

surface pretreatments resulted in mass changes between -1 mg and 2 mg. However, the NaF

pretreatment resulted in the growth of an insoluble corrosion product film as indicated by a mass

increase of over 17 mg. Apparently, unlike the aluminum alloy, adding NaF to the thin electrolyte

film on the surface of this alloy resulted in significant changes in the corrosivity of this film. For the

Cu-Be alloy (Figure 14), the mass changes for the high humidity exposure (93% RH) were so large

that a much larger scale had to be used for plotting the results than used for any other alloy. For the

carbonate/bicarbonate pretreatment, the mass increased more than 100 mg for the 30 day exposure at 93% RH. This was almost an order of magnitude greater than that observed for aluminum alloy 6061

and 4130 steel and almost two orders of magnitude greater than the magnitude of the mass losses

observed in the other five alloys. For the NaF pretreatment, the mass increased more than 21 mg for the 93% RH exposure (the second largest mass change observed). Clearly, in the high humidity
SECTION 7. CORROSION OF METALS 703

Table 20. Score for the Average Reduction in Area for Each Agent at 150 °C [(Mean for Envir.)-(Mean for Hot Ar)]/( Alloy Std. Dev.)

Environnienl Nit Al In 304 CDA 13-8 AM355 AISI Agent 40 6061 625 SS 172 Steel SS 4130 Avg.

HCFC-22 0.01 0.93 1.16 0 02 3 53 0 29 0 06 0 06 0.76 HCFC-124 0.18 2.87 1.17 0 72 2 88 0 53 0 37 0 40 1.14

FC-31-10 0.63 1.02 1.63 0 44 3 22 0 14 1 04 0 62 1.09

HFC-227 0.13 1.44 0.16 0 37 4 98 0 24 0 21 0 51 1.00 HFC- 125 0.50 1.30 0.67 0 16 3 91 0 35 0 23 0 32 0.93

FC-116 0.18 0.52 0.05 0 77 4 10 0 45 0 31 0 32 0.84 HFC- 134a 0.45 1.46 0.13 0 22 1 14 0 01 0 58 0 23 0.53

HFC-236 0.07 0.70 1.16 0 91 1 70 0 07 0 24 0 32 0.65

FC-C318 0.11 0.63 0.26 0 15 2 76 0 23 0 75 0 24 0.64

FC-218 2.99 0.03 0.04 0 56 1 24 0 01 0 87 0 64 0.80

HFC-32/125 0.11 0.60 0.75 0 88 3 26 0 42 0 27 0 13 0.80 NaHC03 0.37 3.16 0.32 2 17 1 97 0 24 1 07 1 26 1.32

Alloy Avg. 0.48 1.22 0.62 0 61 2 89 0 25 0 50 0 42

environment water vapor was absorbed onto the surface to form insoluble corrosion products at a much higher rate than at the lower humidities. Also, it appears that the addition of NaOH to the thin

surface electrolyte film resulted in altering the surface films in a manner that prevented this process. A mass increase indicates the formation of insoluble corrosion products. These corrosion

products could be a protective surface film or it could be a non-protective scale. For a continuous

protective film, the corrosion rate will decrease as the film grows thicker and, at least theoretically,

the corrosion rate for this situation will exponentially approach zero. However, films can become

non-protecting if they crack or spall as they grow allowing the environment access to bare metal

through the fissures in the surface film. A non-protective scale could form as the metal corrodes to form ions in a thin electrolyte film on the surface which migrate through the electrolyte film to nucleate and grow precipitates. These precipitates grow to form relatively loose aggregate layers

which do not prevent wetting of the metal by the electrolyte film nor do they prevent atmospheric

gases from interacting with the thin electrolyte film, thereby allowing corrosion to continue.

The results for the aluminum alloy were a little surprising in that research has found that the

hydroxide ion is the most detrimental ion with respect to the corrosion of aluminum alloys (Foley and

Trazaskoma 1977). This is a direct result of the amphoteric nature of the passivating films that form

on aluminum alloys and the formation of the aluminate ion (AIO2"). Typically, if the passive film is removed for any reason, aluminum alloys will corrode as rapidly as mass transport can provide

reactants to the surface. In aqueous solutions, the passive film is stable in a pH range between 4 and

9.5. If for any reason the pH goes below or above these values the corrosion rate increases very

rapidly. Apparently, adding NaOH to the artificial seawater sprayed on the surface did not increase
704

Table 21.

SECTION 7. CORROSION OF METALS

Slow Strain Rate Tensile Tests at 150°C Rating for Each Alloy/Agent Combination

Environment Nit Al In 304 CDA 13-8 AM355 AISI Agent Std. Agent 40 6061 625 SS 172 Steel SS 4130 Avg. Dev. Rank

HCFC-22 2 1 2 1 4 2 3 2 2.13 0.99 5 HCFC-124 2 3 2 1 3 3 2 1 2.13 0.83 5

FC-31-10 2 2 2 1 4 1 3 1 2.00 1.07 3

HFC-227 2 3 1 1 5 1 3 1 2.13 1.46 5 HFC- 125 1 2 1 1 4 1 4 1 1.88 1.36 2

FC-116 1 2 3 1 5 1 4 2 2.38 1.51 10

HFC-134a 2 2 3 1 2 2 3 2 2.13 0.64 5

HFC-236 2 2 2 1 2 1 3 1 1.75 0.71 1

FC-C318 1 4 1 5 3 3 2 3 2.75 1.39 11

FC-218 3 1 2 1 2 1 3 3 2.00 0.93 3

HFC-32/125 1 2 2 2 4 2 4 1 2.25 1.16 9 NaHCOS 1 4 2 3 2 6 2 2 2.75 1.58 10

Alloy Avg. 1.67 2.33 1.92 1.58 3.33 2.00 3.00 1.67

Std. Dev. 0.65 0.98 0.67 1.24 1.15 1.48 0.74 0.78

Alloy Rank 2 6 4 1 8 5 7 2

Rating Freq Interpretation

1 31 No evidence of environment influence on deformation and fracture in this test. 2 34 Slight indication of possible environmental influence 3 18 Some evidence 4 9 Strong evidence 5 3 Nature of interaction should be evaluated. 6 1 Nature of interaction should be evaluated. 7 0 Nature of interaction should be evaluated. 8 0 Nature of interaction should be evaluated. 9 0 Nature of interaction should be evaluated. 10 0 Nature of interaction should be evaluated.

the pH of the surface electrolyte enough to destabilize the passive film. Also, it appears that increasing the humidity promoted mixing of this electrolyte film keeping the pH below critical levels and preventing the small pits which were observed at lower humidities. However, the quantity of

NaOH produced by the application of a fire suppressant may exceed that applied in this experiment

and pitting under droplets or particles similar to that observed in these experiments, or much worse, could occur. As a result, even though the samples treated with NaOH did not exhibit significant

attack we believe that it is very risky to use a fire suppressant that produces NaOH during combustion

around aluminum alloys.

The other two alloys which showed mass increases, AISI 4130 and CDA- 172, are not used as commonly on aircraft as aluminum alloys. Both of these alloys showed a significant increase in the
SECTION 7. CORROSION OF METALS

Table 22 . Summary Table of Mass Change Exposure Test Results, Crevice Corrosion Visual

Examination Results and Slow Strain Rate Tensile Test Results

Env 1 ronmcnt Nit Al In 304 CDA 1 JO AIVl J3 J

A I CI Agent 40 6061 625 SS 172 Steel SS 4130 Avg.

Exp. 3 4 3 3 5 2 3 5 3.50 Hcrc-22 Crevice 2 2 2 3 4 3 3 5 3.14 SSR 2 1 2 1 4 2 3 2 2.13 Exp 3 3 3 2 4

-3 J"1J

1J

1 AA

hcr:-i24 Crevice 2 2 1 2 3 3 3 5 2.63 SSR 2 3 1 3 3 2 1 2.13

Exp. 4 J 4 5 4 3 4 3.63 FC-31-10 Crevice 1 1 2 2 2 2 2 3 1.88 SSR 2 2 2 1 4 1 3 1 2.00

Exp. 1 3 2 3 5 3 1 3 2.63 HFC-227 Crevice 2 2 2 3 3 3 3 4 2.75 SSR " 3 1 1 5 1 3 1 2.13

Exp. 2 3 3 4 3 2 4 2.88 HFC-125 Crevice 2 2 2 2 3 2 3 2 2.25 SSR 1 1 1 4 1 4 1 1 .88

Exp. 1 3 2 3 4 3 3 4 2.88 FC-1!6 Crevice 2 2 1 2 2 3 2 3 2.13 SSR 1 2 3 1 5 1 4 2 2.38

Exp. 2 3 3 3 3 3 2 3 2.75 HFC- 134a Crevice 1 1 1 2 2 3 4 2 2.00 SSR 2 2 3 1 2 2 3 2 2.13

Exp. 2 5 2 3 4 3 1 3 2.88 HFC-236 Crevice 2 2 1 1 2 2 2 2 1.75 SSR 2 2 2 1 2 1 3 1 1.75

Exp. 1 4 2 3 4 3 1 3 2.63 FC-C3I8 Crevice 2 3 2 2 2 2 2 4 2.38 SSR 1 4 1 5 3 3 2 3 2.75

Exp. 3 4 2 3 4 3 3 4 3.25 FC-218 Crevice 3 4 2 2 2 2 2 4 2.63 SSR 3 1 2 1 2 1 3 3 2.00

Exp. 3 4 3 3 5 4 3 3 3.50 HFC Crevice 3 2 1 2 2 2 3 4 2.38 32/125 SSR 1 2 2 2 4 2 4 1 2.25

Exp. 4 4 3 3 5 4 3 4 3.75 NaHC03 Crevice 2 3 2 2 4 4 4 4 3.13 SSR 1 4 2 3 2 6 2 2 2.75

Exp. 2.33 3.67 2.42 3.00 4.33 3.17 2.33 3.58

Alloy Avg. Crevice 2.00 2.17 1.58 2.08 2.58 2.58 2.75 3.50 SSR 1.67 2.33 1.92 1.58 3.33 2.00 3.00 1.67

Rating Freg Interpretation

1 47 No evidence of environmental attack 2 100 Some evidence of attack. 3 86 Evidence of superficial attack. 4 42 Superficial attack (more study be required) 5 12 Attack could be a problem. 6 1 Attack needs to be consideied. 7 0 Attack can be managed

8 0 Rapid attack. 9 0 Very rapid attack. 10 0 Extremely rapid attack.
706 SECTION 7. CORROSION OF METALS

Table 23. Average Rating for Mass Change Exposure Tests, Crevice Visual Examination, and Slow

Strain Rate Tensile Test

Env ironment IN It Al in CDA 13-8 AM355 AISI Agent Std. Agent 40 6061 625 SS 172 Steel SS 4130 Avg. Dev. Rank

HCFC-22 2.33 2 33 2.33 2.33 4 33 2 33 3.00 4 00 2 88 0 83 11 HCFC-124 2.33 2 67 2.00 1.67 3 33 3 00 2.67 3 00 2 58 0 56 8

FC-31-10 2.33 2 33 2.00 2.33 3 67 2 33 2.67 2 67 2 50 0 53 5 HFC-227 1.67 2 67 1.67 2.33 4 33 2 33 2.33 2 67 2 50 0 84 5 HFC- 125 1.67 2 33 1.67 2.00 3 67 2 00 3.00 2 33 2 33 0 69 3

FC-116 1.33 2 33 2.00 2.00 3 67 2 33 3.00 3 00 2 46 0 73 4

HFC-134a 1.67 2 00 2.33 2.00 2 33 2 67 3.00 2 33 2 29 0 42 2

HFC-236 2.00 3 00 1.67 1.67 2 67 2 00 2.00 2 00 2 13 0 47 1

FC-C318 1.33 3 67 1.67 3.33 3 00 2 67 1.67 3 33 2 58 0 90 7

FC-218 3.00 3 00 2.00 2.00 2 67 2 00 2.67 3 67 2 63 0 60 9

HFC-32/125 2.33 2 67 2.00 2.33 3 67 2 67 3.33 2 67 2 71 0 55 10 NaHC03 2.33 3 67 2.33 2.67 3 67 4 67 3.00 3 33 3 21 0 80 12

Alloy Avg. 2.00 2 72 1.97 2.22 3 42 2 58 2.69 2 92

Std. Dev. 0.49 0 53 0.26 0.46 0 64 0 73 0.48 0 59

Alloy Rank 2 6 1 3 8 4 5 7

atmospheric corrosion rate as a result of a pretreatment with NaF. As a result, cleaning the surfaces

of these alloys after exposure to fire suppressant combustion products is a very good practice and

substitution of another alloy which is not attack to this extent should be considered in locations where

this type of exposure is expected.

7.3.3 Results of Electrochemical Measurements. The objective of these experiments was 1) to determine whether or not electrochemical measurements could be made in low conductivity electrolytes at elevated temperatures and pressures and 2) to define the lower conductivity limit for these

types of measurements. As a result, a special cell was designed and constructed for this type of

measurement. Figure 6. Then, experiments were conducted in an electrolyte with conductivities near and below that used by previous investigators (Halsal, 1992, deSouza, 1987). Figure 15 shows a typical impedance scan from 1 mHz to 65 kHz for 304 stainless steel in a low conductivity electrolyte

(ethyl alcohol). This figure includes a Bode amplitude and phase angle plot (Figure 15(a)) and a Nyquist plot (Figure 15(b)). The resistance of the solution between the two parallel faces on the two

identical samples can be determined from the Nyquist plot as the high frequency real axis (Z')

intercept. By repeating these measurements at difi'erent electrode separations, a plot of the solution

resistance as a function of electrode separation can be developed. Figure 16. The line in this figure is a linear regression fit through all of the resistance values determined at each separation setting. A
SECTION 7. CORROSION OF METALS 707

Table 24. Maximum (Worst) Rating for Mass Change Exposure Tests, Crevice Visual Examination,

and Slow Strain Rate Tensile Test

Environment Nit Al In 304 CDA 13-8 AM AISI Agent Std. Agent 40 6061 625 ss 172 Steel 355 4130 Avg. Dev. Rank

HCFC-22 3 4 3 3 5 3 3 5 3.63 0.92 9 HCFC-124 3 3 3 2 4 3 3 5 3.25 0.89 5

FC-31-10 3 4 2 4 5 4 3 4 3.63 0.92 9

HFC-227 2 3 2 3 5 3 3 4 3.13 0.99 2 HFC- 125 2 3 2 3 4 3 4 4 3.13 0.83 2

FC-116 2 3 3 3 5 3 4 4 3.38 0.92 8

HFC- 134a 2 3 3 3 3 3 4 3 3.00 0.53 1

HFC-236 2 5 2 3 4 3 3 3 3.13 0.99 2

FC-C318 2 4 2 5 4 3 2 4 3.25 1.16 5

FC-218 3 4 2 3 4 3 3 4 3.25 0.71 5

HFC-32/125 3 4 3 3 5 4 4 4 3.75 0.71 11 NaHC03 4 4 3 3 5 6 4 4 4.13 0.99 12

Alloy Avg. 2.58 3.67 2.50 3.17 4.42 3,42 3.33 4.00

Std. Dev. 0.67 0.65 0.52 0.72 0.67 0.90 0.65 0.60

Alloy Rank 2 6 1 3 8 5 4 7

correlation coefficient of 0.978, indicating a good fit, was determined for this line. As seen from the

figure, the data follow a linear relationship between the electrode separation and the resistance. Since

R = p (d/A), (6)

where R is resistance, p is resistivity, d is the electrode separation and A is the electrode surface area.

If the cell is constructed properly, the instrumentation is stable, and the experimental technique is applied properly, then one should get a linear relationship between the resistance of the electrolyte

and the electrode separation as shown in this figure. However, the corrosion rate must be determined

by subtracting the solution resistance, R^, the high frequency real axis intercept in a Nyquist plot,

from the cell resistance (R3 + Rj,() which is the low frequency real axis intercept in the Nyquist plot.

By examining Figure 16, it can be seen the line in the Nyquist plot has not started approaching the

real axis even for the lowest frequency (0.001 Hz). Since the corrosion rate is inversely proportional

to the charge transfer resistance, very low frequencies will be required for these measurements even

with estimating the real axis intercept from the curvature of the Nyquist plot. In addition, one can

calculate the resistivity of the ethyl alcohol electrolyte from the experimental results. It is equal to
708 SECTION 7. CORROSION OF METALS

Figure 7. Mass change during atmospheric exposures at different relative humidities for aluminum alloy 6061-T6 after surface pretreatment with artificial seawater and either

NaF, NaHCOg/NajCOj or NaOH.
SECTION 7. CORROSION OF METALS 799

Figure 8. Mass change during atmospheric exposures at different relative humidities for AISI

4130 steel after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
710 SECTION 7. CORROSION OF METALS

Figure 9. Mass change during atmospheric exposures at different relative humidities for 13-8

Mo steel after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
Figure 10. Mass change during atmospheric exposures at different relative humidities for AM

355 alloy after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
712 SECTION 7. CORROSION OF METALS

Figure 11. Mass change during atmospheric exposures at different relative humidities for 304

stainless steel after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
SECTION 7. CORROSION OF METALS 71

3

c

03 O

in

CO

CO

-0.0006

-0.0008

-0.001

-0.0012

-0.0014

Surface Pre-treatment

• NaF

NaHC03/Na2C03 —NaOH

-0.0016

Relative Humidity (%)

Figure 12. Mass change during atmospheric exposures at different relative humidities for Nitronic 40 alloy after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
714 SECTION 7. CORROSION OF METALS

Figure 13. Mass change during atmospheric exposures at different relative humidities for Inconel 625 after surface pretreatment with artificial seawater and either NaF,

NaHC03/Na2C03 or NaOH.
SECTION 7. CORROSION OF METALS 7I5

Figure 14. Mass change during atmospheric exposures at different relative humidities for the

CDA 172 alloy after surface pretreatment with artificial.
716 SECTION 7. CORROSION OF METALS

N

0) D

Q.

E

<

o

10-1 lOl 103

Frequency (Hz)

(a) Bode Amplitude and Phase Angle Plot

-500 -

-600

(b) NyquistPlot

100 200 300

Z' (kli)

400 500 600

Figure 15. Bode and Nyquist plots of the impedance of 304 stainless steel in high purity

ethanol.
SECTION 7. CORROSION OF METALS 717

17 18 19 20 21 22 23

Sample Separation (mm)

Figure 16. Solution resistance determined from two electrode impedance measurements for

different electrode separations.
718 SECTION 7. CORROSION OF METALS

the slope of the curve in Figure 16 times the electrode area. Since A = 1,140 mm^ and the slope in Figure 16 is 31,820 ohm/mm, the resistivity of the electrolyte is 3.63 x 10^ ohm m. This resistivity

is still several orders of magnitude lower than the fire suppressants. Lowering the resistivity of the

electrolyte will also make the required testing frequency range lower. As a result, other techniques

for making these measurements are being explored.

7.4 Conclusions

The primary objective of this investigation was to determine if there were any of the candidates that should be eliminated from consideration because they were unacceptably corrosive to metals. This

study did not identify any agent for which it would not be possible to identify a suitable alloy for the

fabrication of aircraft storage and distributions systems. However, while no serious concerns were

observed for all of the alloys in any of the agents, several agents consistently performed poorer than

the others or yielded results that indicate further testing is warranted before they should be placed into service. Similarly, the post-deployment corrosion test results did not indicate any serious problems

with any of the agents. However, the use of an agent which produces deposits of mixtures of

bicarbonate, carbonate and sodium hydroxide may cause corrosion problems especially with aluminum

alloys. A review of the literature on the corrosion of aluminum alloys indicates the hydroxide ion is very aggressive to these metals and very high corrosion rates can result if the pH of a surface film

becomes too high. As a result, sodium bicarbonate is not considered a desirable candidate from a corrosion prevention point of view.

The secondary objective was to identify alloys that were suitable for use in each of the agents. • Nitronic 40, AM 355, and 304 stainless steel consistently performed better than the other alloys in the

agents, but examination of the behavior in the specific agent should be consulted before making decisions. For example, the slow strain rate tensile tests on 304 stainless steel indicated that there may be

a potential environmentally induced cracking problem with this material in FC-C318, but additional

testing would be required before this can be concluded.

The third objective was to develop a rapid electrochemical technique for the assessment of the

corrosivity of fire suppressant agents. The technique examined enables the measurement of corrosion

rates in media of much lower electrical conductivity, higher temperatures and pressures than normally

examined, but very low testing frequencies and long testing times will be required for these low

corrosion rate and very low conductivities. As a result, other techniques for making these measure- ments are being explored.

7.5 References

ASTM, Standard Guide for Applying Statistics to Analysis of Corrosion Data, Annual Book of

ASTM Standards Section 3, Metals Test Methods and Analytical Procedures, Phil., PA, ASTM, 1993.

ASTM, Standard Practice for Laboratory Immersion Corrosion Testing of Metals, Annual Book

of ASTM Standards Section 3, Metals Test Methods and Analytical Procedures, Phil., PA, 1993.

Bailey, J. C, Porter, F. C, and Pearson, A. W., Aluminum and Aluminum Alloys, Corrosion.

Chapter 4.1, 4:3-4:33, Newnes - Butterworths, 1976.

Baumert, B. A., A Study of the Corrosion Behavior of Aluminum Alloys, Univ. of Notre Dame,

Notre Dame, IN, 1986.
SECTION 7. CORROSION OF METALS 719

Bevington, P. R., Data Reduction and Error Analysis for the Physical Sciences, McGraw-Hill,

New York, 1969.

Boyer, H. E., and Gall, T. L., Metals Handbook— Desk Edition, Metals Park, OH, ASM

International, 1985.

Brett, C. M. A., "On the Electrochemical Behaviour of Aluminium in Acidic Chloride

Solution," Corrosion Science 33, No. 2, 203-210, 1992.

Cabot, P. L., Centellas, F. A., Garrido, J. A., Perez, E., and Vidal, H., "Electrochemical

Study of Aluminum Corrosion in Acid Chloride Solutions," Electrochim. Acta 36, 1, 179-187, 1991.

Davies, D. E., and Prigmore, R. M., "The Effect of Sodium Fluoride on the Localized

Corrosion of Aluminum in Distilled Water and 50% Ethanediol Solution," International Congress on

Metallic Corrosion, Toronto, Canada, National Research Council of Canads, 1984.

deSouza, J. P., Mattos, O. R., Sather, L., and Takenouti, H., "Impedance Measurements of

Corroding Mild Steel in an Automotive Fuel Ethanol With and Without Inhibitor in a Two and Three

Electrode Cell." Corro. Sci. 27, 1351-1364, 1987.

Fink, F. W., and Boyd, W. K., The Corrosion of Metals in Marine Environments, Columbus,

Bayer and Company Inc., Columbus, OH, 1970.

Flarsheim, W M., Tsou, Y. M., Trachtenberg, I., Johnston, K. P., and Bard, A. J., "Electrochemistry in Near-Critical and Supercritical Fluids, 3, Studies of Br-, I-, and Hydroquionone in Aqueous Solutions", 7. Phys. Chem., 90, 3857-3862, 1986.

Foley, R. T, and Trazaskoma, P. P., "The Chemical Nature of the Anion Dependency of the

Corrosion of Aluminum Alloy 7075-T6," Corro., 33, 12, 435, 1977.

Halsall, R., Corrosion of Metals in Ethanol and Methanol Containing Fuels, General Motors,

Warren, MI, 1992.

McDonald, A. C, Fan, F.-R. F, and Bard, A. J., "Electrochemistry in Near-Critical and

Supercritical Fluids, Water, Experimental Techniques and the Copper (II) System," J. Phys. Chem.

90, 2, 196-202, 1986.

Mendenhall, W., and Sincich, T., Statistics for Engineering and the Sciences, San Francisco,

Dellen Publ. Co., 1992.

Pourbaix, M., Atlas of Electrochemical Equilibria in Aqueous Solutions. Houston, TX, National

Association of Corrosion Engineers, 1974.

Ricker, R. E., and Duquette, D. J., Potentiodynamic Polarization Studies of an Al-Mg-Li Alloy,

Aluminum Lithium Alloys II, Warrendale, PA, The Metallurgy Society of AIME, 581-596, 1984.

Romans, H. B., and H. L. C, Jr., "Atmospheric Stress Corrosion Testing of Aluminum

Alloys," Metal Corrosion in the Atmosphere, Boston, MA, ASTM, 1967.

Scully, J. R., Frankenthal, R. P., Hanson, K. J., Siconolfi, D. J., and Sinclair, J. D., "Localized Corrosion of Sputtered Aluminum and Al-0.5% Cu Alloy Thin Films in Aqueous HF Solution: 1 Corrosion Phenomena," J. Electrochem. Soc, 137, 5, 1365-1372, 1990a.

Scully, J. R., Frankenthal, R. P., Hanson, K. J., Siconolfi, D. J., and Sinclair, J. D., "Localized Corrosion of Sputtered Aluminum and Al-0.5% Cu Aloy Thin Films in Aqueous HF Solution:

II. Inhibition by COj," / Electrochem. Soc, 137, 1373-1377, 1990b.

Scully, J. R., Peebles, A. D., Romig, A. D., Frear, D. R., and Hills, C. R., "Metallurgical

Factors Influencing the Corrosion of Aluminum, Al-Cu, and Al-Si Alloy Thin Films in Dilute

Hydrofluoric Solution," Metallurgical Transactions A 23A, 2641-2655, September 1992.

Smith, W. F, Structure and Properties of Engineering Alloys, McGraw-Hill Book Company,

New York, 1981.
720 SECTION 7. CORROSION OF METALS

Stoudt, M. R., Vasudevan, A. K., and Ricker, R. E., "Examination of the Influence of Lithium on the Repassivation Rate of Aluminum Alloys," Corrosion Testing of Aluminum Alloys, San

Francisco, CA, ASTM, Phila., R\., 1990.

Walpole, R. E., and Myers, R. H., Probability and Statistics of Engineers and Scientists, The

Macmillan Co., New York, 1972.

Wheeler, K. R., A. B. J. Jr., and May, R. R, "Aluminum Alloy Performance in Industrial AirCooled Applications," Atmospheric Corrosion of Metals, Denver, CO, ASTM, 1980A.

Zotikov, V. S., Bakhmutova, G. B., Bocharova, N. A., and Semenyuk, E. Y., "Corrosion of Al

and its Alloys in Hydrofluoric Acid," Prot. Met., 10, 2, 154-156, 1974.
SECTION 7. CORROSION OF METALS 721

Appendix A. Corrosion Rate Exposure Test Results

Alloy Sample No. Test Average Initial Average A Wt. Average Standard

Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013 NIT40 02 HCFC-22 14.31539 14.31552 0.00013 -0.00018 0.00008 NIT40 03 HCFC-22 13.78000 13.78027 0.00027 NIT40 04 HCFC-124 14.30209 14.30217 0.00008 NIT40 05 HCFC-124 14.11364 14.11370 0.00006 -0.00009 0.00005 NIT40 06 HCFC-124 14.14462 14.14476 0.00014 NIT40 07 FC-31-10 14.01056 14.01079 0.00022 NIT40 08 FC-31-10 14.07978 14.08000 0.00022 -0.00025 0.00004 NIT40 09 FC-31-10 14.03391 14.03421 0.00029 NIT40 10 HFC-227 14.26208 14.26211 0.00003 NIT40 11 HFC-227 14.21813 14.21816 0.00003 -0.00001 0.00003 NIT40 12 HFC-227 14.15156 14.15153 -0.00003 NIT40 13 HFC-125 14.29082 14.29093 0.00012 NIT40 14 HFC-125 14.21129 14.21117 -0.00012 0,00001 0.00012 NIT40 15 HFC-125 13.91204 13.91201 -0.00003 NIT40 16 FC-116 14.09377 14.09379 0.00002 NIT40 17 FC-116 13.88219 13.88232 0.00012 -0,00001 0.00012 NIT40 18 FC-116 14.06229 14.06216 -0.00012 NIT40 19 HFC-134a 13.94796 13.94795 -0.00001 NIT40 20 HFC- 134a 14.24534 14.24537 0.00003 -0.00001 0.00002 NIT40 21 HFC- 134a 14.21014 14.21016 0.00002 NIT40 22 HFC-236 14.12528 14.12532 0.00004 NIT40 23 HFC-236 14.11809 14.11812 0.00004 -0.00002 0.00002 NIT40 24 HFC-236 14.26927 14.26926 0.00000 NIT40 25 FC-C318 14,35759 14.35764 0.00005 NIT40 26 FC-C318 14.33708 14.33707 -0.00001 -0.00001 0.00004 NIT40 27 FC-C318 14.18774 14.18772 -0.00002 NIT40 28 FC-218 14.30586 14.30608 0.00022 NIT40 29 FC-218 13.95517 13.95535 0.00018 -0.00015 0.00008 NIT40 30 FC-218 14.27204 14.27210 0.00006 NIT40 la HFC-32/125 14.85542 14.85569 0.00027 NIT40 2a HFC-32/125 14.79800 14.79830 0.00030 -0.00030 0.00003 NIT40 3a HFC-32/125 14.79587 14.79621 0.00033 NIT40 4a NaHC03 14.82673 14.82690 0.00017 NIT40 5a NaHC03 14.79937 14.79965 0.00028 -0.00040 0.00030 NIT40 6a NaHC03 14,80710 14.80783 0.00073

6061 T6 01 HCFC-22 5.15897 5.15956 0.00059

6061 T6 02 HCFC-22 5.14160 5.14202 0.00041 -0.00046 0.00011

6061-T6 03 HCFC-22 5.16151 5.16188 0.00037

6061-T6 04 HCFC-124 5.15992 5.16021 0.00029
722 SECTION 7. CORROSION OF METALS

Alloy Sample No. Test Average Initial Average A Wt. Average Standard

Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013

6061-T6 05 HCFC-124 5.15997 5.16009 0.00012 -0.00022 0.00009

6061-T6 06 HCFC-124 5.16249 5.16274 0.00025

6061-T6 07 FC-31-10 5.15805 5.15867 0.00062

6061-T6 08 FC-31-10 5.15762 5.15808 0.00047 -0.00053 0.00008

6061-T6 09 FC-31-10 5.15774 5.15825 0.00051

6061-T6 10 HFC-227 5.14240 5.14254 0.00014

6061-T6 11 HFC-227 5.15458 5.15475 0.00016 -0.00013 0.00004

6061-T6 12 HFC-227 5.16393 5.16401 0.00008

6061-T6 13 HFC- 125 5.14067 5.14096 0.00029

6061-T6 14 HFC- 125 5.14888 5.14912 0.00024 -0.00020 0.00012

6061-T6 15 HFC- 125 5.16261 5.16267 0.00007

6061-T6 16 FC-116 5.14425 5.14459 0.00034

6061-T6 17 FC-116 5.16136 5.16155 0.00018 -0.00024 0.00009

6061-T6 18 FC-116 5.15905 5.15924 0.00020

6061-T6 19 HFC- 134a 5.15888 5.15901 0.00013 6061-T6 20 HFC- 134a 5.15054 5.15069 0.00016 -0.00013 0.00003

6061-T6 21 HFC-134a 5.13945 5.13955 0.00011

6061-T6 22 HFC-236 5.15121 5.15562 0.00442 6061-T6 23 HFC-236 5.14617 5.14856 0.00239 -0.00248 0.00189

6061-T6 24 HFC-236 5.15225 5.15288 0.00063

6061-T6 25 FC-C318 5.15709 5.15884 0.00174

6061-T6 26 FC-C318 5.15927 5.16053 0.00126 -0.00118 0.00060

6061-T6 27 FC-C318 5.14241 5.14295 0.00055

6061-T6 Ola FC-218 5.10076 5.10108 0.00032

6061-T6 02a FC-218 5.11183 5.11221 0.00039 -0.00037 0.00005

6061-T6 03a FC-218 5.11843 5.11884 0.00041

6061-T6 10a HFC-32/125 5.11563 5.11658 0.00095

6061-T6 11a HFC-32/125 5.11169 5.11215 0.00047 -0.00059 0.00032

6061-T6 12a HFC-32/125 5.11423 5.11457 0.00034

6061-T6 13a NaHC03 5.10236 5.10328 0,00092 6061-T6 14a NaHC03 5.11181 5.11242 0.00061 -0.00104 0.00050

6061-T6 15a NaHC03 5.11315 5.11474 0.00159

1625 01 HCFC-22 16.71024 16.70993 -0.00031

1625 02 HCFC-22 16.73534 16.73509 -0.00025 0.00033 0.00009

1625 03 HCFC-22 16.74050 16.74008 -0.00042

1625 04 HCFC-124 16.72578 16.72552 -0.00025

1625 05 HCFC-124 16.71996 16.71964 -0.00032 0.00027 0.00005

1625 06 HCFC-124 16.78778 16.78755 -0.00023

1625 07 FC-31-10 16.73050 16.73044 -0.00007

1625 08 FC-31-10 16.71615 16.71614 -0.00001 0.00004 0.00003
SECTION 7. CORROSION OF METALS 723

Alloy Sample No. Test Average Initial Average A Wt. Average Standard

Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013

1625 09 FC-31-10 16.43237 16.43234 -0.00003

1625 10 HFC-227 16.73460 16.73455 -0.00005

1625 1

1

HFC-227 16.73019 16.73015 -0.00004 0.00006 0.00002

1625 12 HFC-227 16.78825 16.78818 -0.00008

1625 13 HFC-125 16.72097 16.72095 -0,00002

1625 14 HFC-125 16.70044 16.70050 0.00006 -0.00003 0.00004

1625 15 HFC-125 16.72375 16.72380 0.00005

1625 16 FC-116 16.79899 16.79900 0.00002

1625 17 FC-116 16.76286 16.76287 0.00001 -0.00003 0.00002

1625 18 FC-116 16.80056 16.80061 0.00005

1625 19 HFC-134a 16.72989 16.72981 -0.00008

1625 20 HFC-134a 16.40575 16.40562 -0.00013 0.00010 0.00002

1625 21 HFC-134a 16.72400 16.72390 -0.00010

1625 22 HFC-236 16.38116 16.38111 -0.00005

1625 23 HFC-236 16.71 103 16.71099 -0.00004 0.00006 0.00003

1625 24 HFC-236 16.73446 16.73437 -0,00010

1625 25 FC-C318 16.79867 16.79876 0.00009

1625 26 FC-C318 16.68154 16.68152 -0.00002 -0.00002 0.00007

1625 27 FC-C318 16.73452 16.73450 -0.00002

1625 FC-218 17.04525 17.04529 0.00004

1625 02a FC-218 17.04924 17.04931 0.00006 -0.00005 0.00001

1625 03a FC-218 17.03955 17.03959 0.00005

1625 10a HFC-32/125 17.04597 17.04613 0.00016

1625 11a HFC-32/125 17.02094 17,02106 0.00012 -0.00015 0.00002

1625 12a HFC-32/125 17.03789 17.03805 0.00016

1625 13a NaHC03 17.00623 17.00631 0.00008

1625 14a NaHC03 16.98212 16.98218 0.00006 -0.00008 0.00002

1625 15a NaHC03 17.00205 17.00215 0.00010

304 01 HCFC-22 14.50690 14.50701 0.00011 304 02 HCFC-22 14.54831 14.54854 0.00022 -0.00017 0.00006, 304 03 HCFC-22 14,54721 14,54738 0.00017

304 04 HCFC-124 14.55392 14.55389 -0.00003 304 05 HCFC-124 14.52197 14.52197 0.00000 0.00004 0.00004 304 06 HCFC-124 14.55766 14.55758 -0,00008

304 07 FC-31-10 14.57636 14.57693 0,00057 304 08 FC-31-10 14.59461 14.59524 0.00062 -0.00058 0.00004

304 09 FC-31-10 14.53066 14.53122 0.00056

304 10 HFC-227 14.50282 14.50297 0.00015 304 11 HFC-227 14.48668 14.48684 0.00016 -0.00014 0.00002

304 12 HFC-227 14.49441 14.49453 0.00011
724 SECTION 7. CORROSION OF METALS

Alloy Sample No. Test Average Initial Average A Wt. Average Standaixl Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013 304 13 HFC-125 14.53445 14.53459 0.00014

304 14 HFC-125 14.58747 14.58759 0.00012 -0.00014 0.00001 304 15 HFC-125 14.51819 14.51834 0.00014

304 16 FC-116 14.60180 14.60199 0.00018 304 17 FC-116 14.58646 14.58660 0.00014 -0.00015 0.00003 304 18 FC-116 14.56847 14.56859 0.00012 304 19 HFC-134a 14.54856 14.54863 0.00007

304 20 HFC- 134a 14.57147 14.57154 0.00007 -0.00007 0.00000

304 21 HFC-134a 14.51002 14.51008 0.00007

304 22 HFC-236 14.50289 14.50299 0.00009

304 23 HFC-236 14.56563 14.56573 0.00010 -0.00008 0.00003 304 24 HFC-236 14.49324 14.49329 0.00004

304 25 FC-C318 14.55940 14.55956 0.00016

304 26 FC-C318 14.48874 14.48892 0.00018 -0.00020 0.00005 304 27 FC-C318 14.52980 14.53006 0.00026

304 Ola FC-218 14.70650 14.70657 0.00007

304 02a FC-218 14.59916 14.59922 0.00005 -0.00009 0.00005 304 03a FC-218 14.67582 14.67597 0.00015 304 10a HFC-32/125 14.62390 14.62411 0.00021 304 11a HFC-32/125 14.72853 14.72875 0.00023 -0.00024 0.00004 304 12a HFC-32/125 14.66874 14.66902 0.00028

304 13a NaHC03 14.66832 14.66853 0.00021 304 14a NaHC03 14.71106 14.71136 0.00030 -0.00026 0.00005 304 15a NaHC03 14.57512 14.57540 0.00028

172 01 HCFC-22 16.08959 16.09228 0.00269

172 02 HCFC-22 16.60931 16.61261 0.00331 -0.00306 0.00033

172 03 HCFC-22 16.35164 16.35484 0.00320

172 04 HCFC-124 16.37510 16.37691 0.00181

172 05 HCFC-124 16.38782 16.38919 0.00137 -0.00147 0.00029

172 06 HCFC-124 16.31289 16.31414 0.00125

172 07 FC-31-10 16.47673 16.47927 0.00254

172 08 FC-31-10 16.47398 16.47671 0.00273 -0.00266 0.00011

172 09 FC-31-10 16.15224 16.15496 0.00272

172 10 HFC-227 16.42018 16.42233 0.00215

172 11 HFC-227 16.20014 16.20199 0.00185 -0.00197 0.00016

172 12 HFC-227 16.52838 16.53030 0.00192

172 13 HFC-125 16.57400 16.57489 0.00088

172 14 HFC-125 16.60772 16.60860 0.00088 -0.00089 0.00002

172 15 HFC-125 16.48618 16.48709 0.00091

172 16 FC-116 16.53656 16.53792 0.00136
SECTION 7. CORROSION OF METALS 725

Alloy Sample No. Test Average Initial Average A Wt. Average Standaid EnV 1 r *T L . , ^1 alllA grams Wt. LOSS Deviation

niO 11 A 1 9303 0 0001

T

1 79 I 7 Pr.i 1 1 f\ 41 OS? 1 A 41 lAQ 0 001 fi7 -U.UUI Jl n Ann i

O.OUUlo

179 1 1 A. i1 8o Fr 1 1 f\ 1

6

1 fi S9^;71 0 001 40

1 79 1 0 HFP 1 \f\ 1 O . 1S4Q^ 1 0-+" J

1 A 1 8S1

4

0 0009

1

U.VJUUZ 1 179 20 HFr-1 "^43 jri 1 v„ i J'-rd 16 48381 1 48401 0 00090 0 00093 A f»ririA4 U . UUUU'l- 179 21 16.37599 11 fiL; . J37fi9fi / uzu 0 00097

1 79 zz HFP 7'^^>

1 o . jzS9 11 09uz 11

fi0 . JZS9 11 8007 0 00087

179 93Z J 111 16 49087 16.49177 0 00080 -0 00087 0 00003

17? 24 HFC-236 1 6 87507 16 89590 0 00084

1 79 Z J Vr P31 8 1 ^ '>74S'> 1 fi 97S9S 0 0004^1

1 79 zo vc n1

8

11 fiO . 47784 H / Z04 0yj .\j\j\jjv OOOSO 0 00043 u.uuuu /

1 79 z97/ FP r3 1 8 1 fi S7734 0 00016

11 79/Z Uia FPr^^-z9 11 8o \f. 1 7"^ 1 fi fiS 1 QQ 0 00077

179 1 /Z 09UZda FP-? 1 8 16 61 "^QO 1 fi fi 1 470 0 00080 -0 00077 0 00003

I1 79/ z f)3a FP-91 8 16 71 7ns 16 71278 0 00074

179 1iUdfla nr^.^-jz/ izj 1 6 64*^47 1 fi fi4S04 0 00747

179 11a HFP-32/1 25 16 6778S 16 67989 0 00204 -0\J . 00776 \J\J z, z u 0 00022

179 1 / z 11 9aZd riFv_ HFP-39/1 jz/ izj95 16 56978 16 56506 0W . 00778 \JWZ z

o

1 79 1 /Z 1 ja rNaric MaHPn^ 1 A AQn64 1 6 6Q787 0 007 1

8

U .uUz 1 o 179 I4a NaHP03 1 ^ al 1 v„ \J ~j 1 6 698'^9 16 70081 0.00242 -0 0071 8 0 00024

11 79/ z 11 JdSa NaHPn3 1 fiOOQQ 16 60704 0 001 OS

n 1 HPFP 9? I S 047 1 A 1 S 0471

S

1 3 .U'+ / 1 3 0 0000

1

U.UUUU 1

1 T 8 09uz HPFP nv^rv^ zz99 IS 1 fil 77 IS13. 11 61U 1 6QU" 0 00008 0U .UUUVJ 00001J 0 00004

1i JT-Ro 03 HCFC-22 15 07 1 96 15.07193 -0 00003

1 1 8 HPFP 1 94 1 S 41 1 S 41 S'^'^ 1 3 1 3 J J 0 00017

1uT 8o OS 15.21775 1I JS . 7^ 1I 760 / Kjy -0 00006 0 00000 0 00007

1 1 8 WO HPFP 1 9d IS 01 S 1

0

i J .U IJ lu 11 3S .u011 JuS07/ 0 00001

1 1 Q U / FP T 1 in ID .J / o 1

o

13.3/ o3 / nu.uuujy ooniQ

1

8

08uo FP 31J V 10\.\J 1 S 7604Q 1 S 76077 0 00078 -0 00016 0 00007

OQ FP 31 10 1 S 1 SQin 13.1 1 S 1 JySQ711 1 0 00041 ion lU HrL-zz /

1 ^ 1 QOQQ r\ c\r\r\f\Q

1 1 8 1 J-O 1 1 1 1 HFP rirL--zz 997/

113.1 S 1 4S43 1 S1 3 11 3S . 11H-3Z 4S77/ 0 000 1

7

U . \j\J\J 1 z 0 00007 0 00006

1 7 8 1 91 Z HFP 997 1 3 . JO / 3Z 1 3 . 30 / 3Z 0 00000

13-8 13 HFC-125 1 J . jzy4z 1 3. 3/^30 U.UUU14

1 T 8 1 A 14 HrC- 1 z J 1 3 . zOzzV 1 3 . Z0Z4

J

0 0001

S

U .UuU 1 3 0 0001

S

-U .UUU 1 3 0 00001

13-8 15 HFC-125 15.13633 15.13649 0.00017

13-8 16 FC-116 15.05156 15.05179 0.00022

13-8 17 FC-116 14.98056 14.98077 0.00021 -0.00021 0.00001

13-8 18 FC-116 15.21466 15.21486 0.00020

13-8 19 HFC-134a 15.42397 15.42412 0.00015

13-8 20 HFC-134a 15.22669 15.22673 0.00003 -0.00009 0.00006
726 SECTION 7. CORROSION OF METALS

Alloy Sample No. Test Average Initial Average A Wt. Average Standard

Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013

13-8 21 HFC- 134a 15.15183 15.15190 0.00008

13-8 22 HFC-236 15.12782 15.12793 0.00012

13-8 23 HFC-236 15.01663 15.01672 0.00009 -0.00011 0.00001

13-8 24 HFC-236 15.34009 15.34021 0.00011

13-8 25 FC-C318 15.09380 15.09395 0.00015

13-8 26 FC-C318 15.37339 15.37357 0.00018 -0.00017 0.00001

13-8 27 FC-C318 15.44205 15.44223 0.00017

13-8 Ola FC-218 15.29468 15.29494 0.00026

13-8 02a FC-218 15.10975 15.11002 0.00027 -0.00028 0.00002

13-8 03a FC-218 15.23719 15.23749 0.00030

13-9 10a HFC-32/125 15.11540 15.11581 0.00042

13-8 11a HFC-32/125 15.34063 15.34103 0.00041 -0.00049 0.00014

13-8 12a HFC-32/125 15.03324 15.03389 0.00065

13-8 13a NaHC03 14.85465 14.85575 0.00110

13-8 14a NaHC03 15.10975 15.11041 0.00066 -0.00080 0.00026

13-8 15a NaHC03 15.08130 15.08194 0.00064

355 01 HCFC-22 15.29150 15.29145 -0.00005 355 02 HCFC-22 15.00926 15.00912 -0.00014 0.00009 0.00005 355 03 HCFC-22 15.32792 15.32785 -0.00007

355 04 HCFC-124 15.01662 15.01629 -0.00032

355 05 HCFC-124 15.42594 15.42581 -0.00013 0.00021 0.00010

355 06 HCFC-124 14.92252 14.92236 -0.00016

355 07 FC-31-10 14.98140 14.98163 0.00023

355 08 FC-31-10 15.39432 15.39458 0.00026 -0.00022 0.00005

355 09 FC-31-10 15.36341 15.36358 0.00017

355 10 HFC-227 15.51429 15.51431 0.00002

355 11 HFC-227 15.39858 15.39860 0.00002 -0.00001 0.00002

355 12 HFC-227 14.84048 14.84047 -0.00002

355 13 HFC-125 14.97517 14.97520 0.00003

355 14 HFC-125 14.75849 14.75853 0.00004 -0.00003 0.00001

355 15 HFC-125 15.34143 15.34145 0.00002

355 16 FC-116 15.23439 15.23460 0.00021

355 17 FC-116 14.88056 14.88069 0.00013 -0.00014 0.00006

355 18 FC-116 14.87070 14.87078 0.00008

355 19 HFC-134a 14.89687 14.89690 0.00003

355 20 HFC-134a 15.51394 15.51393 0.00000 0.00001 0.00005

355 21 HFC- 134a 14.97787 14.97780 -0.00007

355 22 HFC-236 15.26546 15.26536 -0.00010

355 23 HFC-236 14.88965 14.88972 0.00007 -0.00001 0.00009

355 24 HFC-236 15.50902 15.50906 0.00005
SECTION 7. CORROSION OF METALS 727

Alloy Sample No. Test Average Initial Average A Wt. Average Standard Envir g rams Wit T r\cf wl. loss Deviation MTTiin niU i 1 H . i Z JOU 114.4 1izjyj 71Q1 u.uuui ^ JJJ ZJ nr* 1 fi rL.-L, J 1 o 14. /41ZZ 1 A 7/1 178 14. /41Zo A AAAA/C

JJJ 1 j.ZZj /o 1 ^ 77^8^ 1 j.ZZjoj A AAAA"T A AAAA

-0.00001 0.00011 JJJ Z /

1 4 74S18 1/1 7/I^7fi 14. /4jZ0 A AAA 1

JJJ n 1 '1 ula 1 c 1 1 1 1 1 J .J J 101 1 c 11171 1 J. JJl /

A AAA 1 A U.UUUlU

JJJ n^T uza 1 ^ 71187 1 J . / 1 1 oZ 1^711 QC 1j. / 1 lyo A AAA 1

n nnn 1

-U.UUU15 A AAAA A 0.00004

nil uja 1 ^ 777^1 1 J.Z / / Jj 1 ^ 77777 I J .£ 1 1 1

A AAA 1

Q

JJJ 1 n-i lua flrC-jZ/ IZj 1 ^ iQ7nn 1 J. jyzuu 1 c 1Q7 1 /I 1 J. jyzj4 A AAAT/I

JJJ Ilid1 Q T4PP -39/1 9^ 11 JS .H^U 44071/ J

1 s 44 1 n

1 J.441U1 A nnn97 n nnnii -u.uuuj 1 n nnnn/1 U.UUUU4

JJJ iza 1 J . JOOOV 1 ^ ifionI J . joyu

nU.UUUJZ nnmi

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A 1 in 41 JU 1UCP 19^ rlrC- 1 Zj 1 c /: 1 /I /1

I J .0 144

1 fi 1 ^ni 1 J .0 1 jUJ

41JU 16 FC-l 16 15.671 18 15.67160 A AAA/1 1 0.00042

A 1 in 41 JU 1 11CP 1 1 1 j.ojyyo 1 c ^:Ani7 1 J .DoUJ / n nnnm u.uuujy n nnn/in -U.UUU4U n nnnn

U . UUUU

A 1 in 41 JU C/^ 1 1 A rL-l lo 1 ^ fiion^ 1 j.6jyUj 1 ^ /^10/1 ^

1 J .ojy4j n nnn/in U.UUU4U 4130 19 HFC-134a 15.65736 15.65765 0.00029

A 1 in 41 JU zU IT C/^ 1 1/1 ^ HrC-l J4a 1 ^ /;7C7i

1 J.O/jzJ 1 J.6 /jjy n nnn 1 a U.UUUlo n nnnifi -U.UUUzo n nnnno U.uuuuy

/1 1 in 41 JU 1

z 1 HrC-l J4a ID. 00 / /4 1 J.666U0 n nnnii U.UUUJJ 4130 22 HFC-236 15.64327 15.64356 0.00029 4130 23 HFr-23fi 1 5 70320 15.70357 0 00037 -0.00034 0.00004 4130 24 HFC-236 15.72882 15.72919 0.00037

4130 25 FC-C318 15.67104 15.67137 0.00034 4130 26 FC-C318 15.69345 15.69372 0.00027 -0.00031 0.00004 4130 27 FC-C318 15.67568 15.67600 0.00032

4130 Ola FC-218 15.63867 15.6389C 0.00029
728 SECTION 7. CORROSION OF METALS

Alloy Sample No. Test Average Initial Average A Wt. Average Standard

Envir. Wt., grams grams Wt. Loss Deviation NIT40 01 HCFC-22 14.12380 14.12393 0.00013 4130 02a FC-218 15.66631 15.66673 0.00042 -0.00038 0,00008 4130 03a FC-218 15.57494 15.57537 0.00044 4130 10a HFC-32/125 15.68322 15.68357 0.00035 4130 11a HFC-32/125 15.72594 15.72617 0.00023 -0.01134 0.01914 4130 12a HFC-32/125 15,65444 15.68788 0.03344 4130 13a NaHC03 15,67555 15.67662 0.00107 4130 14a NaHC03 15.63064 15.63150 0.00086 -0.00089 0.00017 4130 15a NaHC03 15.69020 15.69093 0.00074
729

8. ELASTOMER SEAL COMPATIBILITY

Gregory B. McKenna, William K. Waldron Jr. and Ferenc Horkay

Materials Science and Engineering Laboratory

8.1 Introduction

Excessive swelling or deterioration of the elastomer seal (o-ring and its lubricant) in the fire suppressant storage container could lead to leakage of the agent leaving the system unready to respond

in case of fire. Short term exposure experiments have been conducted and data have been generated on the proclivity of the eleven fluid agents to alter the properties of various elastomers and greases. The compatibilities of the fire suppressant agents with commonly used elastomers and greases have

been characterized using two types of measurements.

The first type was based on solution thermodynamics and characterizes swelling due to sorption

of the agent into the elastomer or grease. The degree of swelling was determined by measuring the

solvent (agent) weight fraction in the polymer (elastomer or grease)/solvent (agent) system at various temperamres and vapor pressures. Using these weight fractions and the Flory-Huggins theory

(Flory, 1942; Huggins, 1943), a single polymer-solvent interaction parameter x was calculated for each polymer/agent system at various temperatures. The x values were used to characterize the

compatibility of the solvents (agents) with the polymers (elastomers and greases). Small x values

correspond to good solubility or, for present purposes, bad compatibility. Above x - 0.5, the

solubility of polymers gradually decreases with increasing x- A rating system was defined based on

the values of x parameters obtained from swelling measurements at 35 °C. Good compatibility (x >

1.2, < 0.22) implies that an elastomer or lubricant is acceptable for use in the fire suppressant

system. Bad compatibility (x < 0.9, vvj > 0.38) corresponds to excessive swelling. For values of

0.9 < X < 1-2, the agent was considered to have fair compatibility with the elastomer or grease and

represents a marginally acceptable system.

The second type, durability measurements, produced data on residual mechanical (rheological)

properties of the elastomers (greases) after exposure to the fluids at elevated temperature and pressure

(150 °C, 5.86 MPa). Compatibility ratings were based on the results of compression set and tensile

test measurements of the elastomers and viscosity measurements of the greases. These tests provide

direct information on the physical and chemical damage to the samples at extreme conditions. While

the following procedures were used to define the fire suppressant agent/elastomer or lubricant

compatibilities, it is important to note that the long term exposure response of these materials cannot

be extracted from the tests as performed. A major conclusion from the results of the durability

testing is that the 150 °C condition is too severe, i.e., property changes for the most-part were

extreme. Further testing at lower temperatures will be required to provide better estimates of

elastomer and lubricant durabilities in the final down-selected candidate fluids. In compression set measurements, good compatibility was defined as a condition when the

compression set (percent of original deflection retained after release from compression) was less than

90% after a 2 week exposure. Elastomers have bad compatibility if the compression set exceeded

90% after 4 weeks. If the compression set was less that 90% after 2 weeks but exceeded 90% after a 4 week exposure, the agent was considered to have fair compatibility with the elastomer and

represents a marginally acceptable system. If a specimen was split or broken, the agent was

considered to have bad compatibility with the elastomer.
730 SECTION 8. ELASTOMER SEAL COMPATmiLITY

In the tensile test measurements, good compatibility was defined as a condition when the

decrease in ultimate elongation (percent increase of inner circumference at rupture) was less than 65%

after a 4 week exposure. Elastomers have bad compatibility if the decrease in ultimate elongation

exceeded 65 % after a 2 week exposure. If the decrease in ultimate elongation was less than 65 %

after 2 weeks but exceeded 65 % after 4 weeks, the agent was considered to have fair compatibility

with the elastomer and represents a marginally acceptable system.

The viscosity of the greases, when measurable, did not show systematic variation with exposure

time, indicating that no significant chemical degradation occurred. It was found, however, that mobile

substances or fractions were extracted by some of the candidate fire suppressant fluids which resulted

in the greases becoming powder-like, i.e., their viscosities were not measurable. Therefore, the

ratings for the lubricant compatibility were based on the following criteria: Good compatibility was

defined as a condition when the grease did not become powder-like after a 6 week exposure.

Lubricants have bad compatibility if the grease became powder-like after a 4 week exposure. If the

grease's viscosity was measurable after 4 weeks, but became powder-like after 6 weeks, the lubricant was considered to have fair compatibility with the agent.

The following describes the section organization: Section 8.2 describes the characterization of the

agents in terms of thermodynamic swelling measurements. After briefly describing the theoretical

basis of the thermodynamic investigations (Section 8.2.1), the vapor sorption experiments are described (Section (8.2.2) and the results of swelling measurements are presented (Section 8.2.3).

Section 8.3 describes the characterization of the agents in terms of the durability measurements. After

describing the high temperature exposure method (Section 8.3.1) and each durability measurement

technique, the results of mechanical property measurements (compression set, tensile test, and

viscosity) performed on the elastomers and greases are described (Sections 8.3.2 and 8.3.3). In Section 8.4, the results of the durability measurements in combination with the thermodynamic

(swelling) investigations are used to characterize the compatibility of fire suppressants with elastomer

seals.

8.2 Thermodynamic Compatibility

8.2.1 Theoretical Considerations. Swelling is a powerful method for characterizing the affinity of a polymer (elastomer or grease) to a solvent (fire suppressant agent). The driving force for swelling is the change in the free energy of mixing. The Flory-Huggins lattice theory of polymer solutions

(Flory, 1942; Huggins, 1943) describes the free energy of mixing under subcritical conditions and

characterizes the affinity of the polymer to the solvent by a single polymer-solvent interaction

parameter x- Existing network theories allow the evaluation of the relevant thermodynamic

parameters from equilibrium swelling measurements when the polymer is crosslinked (James and

Guth, 1943; Flory and Rehner, 1943; Flory and Erman, 1982). Proposed improvements of the

theories lead to more complicated expressions containing more adjustable parameters. The corre- sponding states theory of polymers (Flory et ai, 1964a and 1964b.; Flory, 1965) provides a more

general framework to analyze the thermodynamic properties of polymer systems both under subcritical and supercritical conditions.

In swelling measurements, the solvent activity is varied to swell the polymer by different

amounts. In vapor sorption experiments, first conducted by Gee et al. (1965), the solvent uptake of

the polymer is controlled by varying the partial pressure (activity) of the swelling agent in the

equilibrium phase. Similar experiments have been performed by Brotzman and Eichinger (1982 and

1983) and McKenna and Crissmann (1993). The same thermodynamic information can also be
SECTION 8. ELASTOMER SEAL COMPATIBILITY 731

obtained using other methods to vary the solvent activity around a swollen network (Boyer, 1945;

Pennings and Prins, 1961; Horkay and Zrinyi, 1982; McKenna et al, 1989; Horkay et ai, 1989;

McKenna et ai, 1990). However, vapor sorption experiments were used to measure the degree of

swelling because of their straightforward interpretation and many polymer/solvent systems can be

analyzed simultaneously.

8,2.1.1 Flory-Huggins Theory. The thermodynamics of mixing is governed by the change in free energy

AG„ = AH^ - rA5„ , (1) m m m ^ ^ '

where AG^ is the change in Gibbs free energy, T is the absolute temperature, AH^ is the heat of

mixing, and AS^ is the entropy of mixing. A negative value of AG^ indicates that the mixing process

occurs spontaneously. The term TAS^ is always positive because the entropy increases upon mixing.

The thermodynamic properties of polymer solutions both in the semi-dilute and concentrated

regimes can be described in a straightforward way using the Flory-Huggins lattice theory (Flory,

1953). This theory calculates the entropy of mixing from the possible arrangements of component

molecules (i.e., solvent molecules and monomeric units of the polymer) with respect to each other,

e. , AS^ is given by

A5^ = k (NjhiVj + ^^v^) , (2)

where and A'2 are the number of molecules of the solvent and polymer, respectively, Vj and V2 are

the corresponding volume fractions and k is the Boltzmann factor. The Flory-Huggins theory

calculates the heat of mixing by introducing the dimensionless interaction parameter

X

= (3)

(kTN^v^)

which when combined with Equation 2 leads to the total free energy of mixing

AG^ - kT (A^,hiv, + N^\nv^ + x^jVj) . (4)

The value of x depends on the thermodynamic quality of the solvent: in good solvent conditions x <

0.5 and in "theta" conditions x = 0.5. For poor solvents x > 0-5.

In general, x depends on the polymer volume fraction (Flory, 1970), i.e..

Xo ^ X1V2 (5)

The chemical potential of the solvent can be obtained by differentiating the Gibbs free energy of

mixing with respect to the number of solvent moles and

^l

- til = Ati, = RT [bid - V2) + (1 - 1/P)V2 + xv'] ,

where is the chemical potential of the pure solvent and P is the degree of polymerization. A[Xi is related to measurable macroscopic quantities, such as the vapor pressure of the solvent p and the

osmotic pressure of the solution 11. From general thermodynamic considerations.
732 SECTION 8. ELASTOMER SEAL COMPATIBILITY

1

' (7)

where is the saturation pressure of the solvent at temperature T, Vy is the partial molar volume of

the solvent and R is the gas constant. Combining Equations 6 and 7 yields

which can be fit to experimental data using adjustable parameters xq. Xi. • frorn Equation 5. Thus,

chemical potentials obtained from vapor pressure or osmotic pressure measurements yield empirical

values of x- These values allow the estimation of the solubility of polymers. Small x values (x <

0.5) correspond to good solubility In good solvents, crosslinked polymers exhibit a large amount of

swelling which can lead to seal failure in elastomeric o-ring based joints. High x values (x > 0.5)

are characteristic of polymer/solvent systems with limited miscibility. Crosslinked polymers swell

only a small amount in thermodynamically poor solvents. The simple Flory-Huggins theory discussed above is based on a series of arbitrary assumptions

(e.g., lattice sites are the same size for polymer segments and solvent molecules, uniform distribution

of the polymer segments in the lattice). The x parameter is, by definition (see Equation 3), governed

by the interaction energy. In reality, x is a measure of Gibbs free energy, i.e., it also contains an

entropy contribution. Because of several assumptions involved in the Flory-Huggins theory, x can

only be considered an empirical fitting parameter, without precisely defined physical significance,

which accurately describes the thermodynamics of the system in question.

Although the Flory-Huggins theory is formulated in terms of volume fractions (Vj and V2), the

theory can also be expressed in terms of weight fractions (iVj and W2). This substitution is arbitrary,

but the functional form of Equation 8 remains the same

where x = Xo + Xi ^2 been substituted and the term {I - \ / P) has become unity since P is large for polymeric materials and infinite in crosslinked networks. Strictly speaking the use of the

volume fraction in the Flory-Huggins theory is also a simplification because the surface area fraction

is the relevant quantity in deriving the enthalpy of mixing on a lattice. In this case the x values difi'er from those calculated from the equilibrium volume fractions (Vj and V2). Throughout the present

report weight fractions are used instead of volume fractions because they are directly measured

quantities.

8.2.1.2 Flory's Equation-of-State Theory. The principle of the corresponding state theories

rests on the assumption that the intermolecular potentials for the polymer and solvent have equivalent

forms when expressed as functions of the distance between molecular (or segment) centers. If the

thermodynamic properties are known for one reference fluid, those for any corresponding fluids are determined by two scale factors: one for the separation distance of molecular centers and the other for the magnitude of intermolecular potential. These factors are embodied in a characteristic temperature

7^ and a characteristic pressure p . In the case of polymers the third important parameter is the

number of intermolecular degrees of freedom.

hxiplp") = ln(l - v^) + (1-1/P)V2 + , (8)

^(pIp") = ln(l - Wj) + + Xo^2 + X1W2, (9)
SECTION 8. ELASTOMER SEAL COMPATIBILITY 733

Flory formulated the partition function of liquids by combining a rudimentary factor for hard

spheres with an intermolecular energy of the van der Waals form (Flory et ai, 1964a; Flory et al.,

1964b; Flory 1965). The reduced equation of state of a pure liquid is

IT V

T

>l/3

,1/3. J_

D T (10)

where v - v / v* , t = T / T* , it = p / p* are the reduced variables and v* is the characteristic (hard

core) volume. A pure component is completely characterized by the three parameters v*, 7^, and p"

The values of these parameters can be evaluated from pvT (pressure-volume-temperature) data by

nonlinear regression fit to the theoretical equation of state (Equation 10). These parameters are tabulated in the literature for many small molecules as well as for several polymers (Shih and Flory,

1972).

The application of the equation-of-state theory for a two component mixture requires the

knowledge of the parameters of the pure components as well as their interactions in the mixture

defined by an interaction enthalpy parameter and interaction entropy parameter Qi2- The chemical

potential of mixing is given by

RT

= hi(l - (p) + (1 - 1/P)(p

Pi Vi RT

3 7, hi

/ 1/3

1 >V3 - + p,(u - Uj)

I)

(11)

RTv R

where <p is the segment fraction and 62 is the site fraction of the polymer.

For the mixture the reduction parameters can be calculated from

p' - p;a - (p) ^ p;^ - n - ip)Q,x^, (12)

and

r =p'K(i - <p)p;/t; + ^p;/t;) (13)

Knowing p* and T* , u can be obtained from Equation 10. The only remaining unknown parameters

are X12 and 2i2- Since the interaction terms usually have very little effect, X12 and (2i2 are usually

treated as free variables (adjustable parameters).

Equation 1 1 allows the calculation of the chemical potential both below and above the critical temperature of the solvent and therefore can be used to describe the solubility of a supercritical gas in a polymer. (The simple Flory-Huggins lattice theory discussed in the previous section cannot be used

above the critical temperature of the fluid because the saturated pressure of the vapor p^ is not defined

under supercritical conditions.)

In Section 8.2.3 some examples are shown which use Flory's equation-of-state theory to describe

the experimental results both below and above T^. This analysis has not been performed for all of the
734 SECTION 8. ELASTOMER SEAL COMPATIBILITY

polymer/solvent systems because the solvent uptake was relatively small above at the pressures

tested.

8.2.2 Swelling Measurements

8.2.2.1 Experimental Setup. The degrees of swelling of three greases and six sets of

crosslinked and uncrosslinked elastomers were determined by measuring the displacement of each

quartz spring using a cathetometer as in the experimental arrangement illustrated in Figure 1 . The

elastomer's or grease's mass uptake of the agent (solvent) was calculated from the relative displacement of the springs. The cathetometer had a vernier scale permitting readings to 0.01 mm but

measurements were recorded to the nearest 0.05 mm because of difficulties in viewing the spring and

pan assemblies.

Six pressure vessels were designed and built for this purpose. The pressure vessels were made

of type 304 stainless steel and designed for a maximum working pressure of 5.86 MPa with two view

ports 180° apart for viewing and backlighting purposes. The view ports were made with 10.2 cm

diameter and 25.4 mm thick circular plates of tempered Corning Glass Pyrex (trademark for borosilicate) glass. The cylindrical vessel had 2.25 liters of inner volume (17.8 cm height by 12.7

cm diameter) and was large enough to admit a 7.6 cm square by 15.2 cm high stainless steel spring

stand with sixteen fused quartz spring and pan assemblies. The quartz springs had a coil diameter of

6.4 mm and either 1 mg/mm or 1.25 mg/mm sensitivity (spring rate). The quartz pans were slightly

curved with diameters of 7-9 mm and weighed 10-25 mg when empty.

Each vessel was placed in a 41 cm diameter by 30 cm high cylindrical Pyrex jar and completely

immersed in Dow Corning silicone bath oil, clear 200 type, with a viscosity of 0.05 Pa • s. For

insulation, a 10 mm thick aluminum lid was placed over the jar and 76 mm thick fiberglass sheet was wrapped around the circumference of the Pyrex jar. Two rectangular holes were made so that the

view ports were visible. The bath was heated using a 1000 W Vycor (Corning Incorporated

trademark) immersion heater and stirred to maintain temperature uniformity using a 37 W ColeParmer Stir-Pak mixer with a 51 mm diameter propeller. The temperature was monitored using a Fluke thermometer with a type K thermocouple and 0. 1 °C resolution.

A pressure charging and recovery system was designed and built to allow the admission of the

agent into the previously evacuated vessel and the recovery of the agent after each measurement. The

system was also capable of overpressurizing the vessels with nitrogen but this capability was not employed. The pressure was monitored using a Druck DPI 602 digital pressure indicator and PDCR

910 transducer with an absolute pressure range of 0-6.21 MPa, combined nonlinearity, hysteresis and

repeatability of + 0.1% using the best straight line method (BSL).

8.2.2.2 Experimental Materials. The elastomers and lubricants used in this study are shown in Tables 1 and 2, respectively. Both crosslinked and uncrosslinked elastomers were studied. The 85%

butadiene-15% acrylonitrile crosslinked copolymer was prepared by first dissolving approximately 2 g

of polymer in 50 ml of benzene. Next, one part dicumyl peroxide per hundred parts butadiene rubber

(1 phr) was added and dissolved. The benzene was allowed to evaporate and the specimen was

placed in an oven at 150 °C for a period of 2 hours in order to crosslink the peroxide containing

specimen. The 55% butadiene-45 % acrylonitrile copolymer was crosslinked in a similar manner

using pyridine as the solvent. The silicone, fluorosilicone, and neoprene polymers were supplied

crosslinked and uncrosslinked by the vendor. The fluorocarbon polymers were cured at NIST for 1 hour at 150 °C and post-cured at 240 °C for 24 hours.

8.2.2.3 Experimental Methods. Samples of the three greases, six crosslinked and six uncrosslinked polymers weighing 25-50 mg were placed on 15 quartz pans. A glass bead with an
SECTION 8. ELASTOMER SEAL COMPATIBILITY 735

Quartz Spring and Pan Assembly

Figure 1. Schematic of the experimental apparatus for isopiestic swelling measurements.
736 SECTION 8. ELASTOMER SEAL COMPATmiLITY

Table 1 . Elastomers used in swelling experiments

Elastomer Vendor Designation

Silicone Colonial Rubber Si

55% Butadiene-45 % Acrylonitrile Goodyear N206

Fluorosilicone Colonial Rubber FSi

Viton E-60 Fluorocarbon Du Pont FKM

Neoprene Colonial Rubber CR

85% Butadiene-15% Acrylonitrile Goodyear N926

Table 2. Lubricants used in swelling experiments

Lubricant Vendor Designation

Krytox 240AC Fluorinated Grease Du Pont 240AC

Braycote 600 Perfluoropolyether Grease, Low Volatility Castrol 600

Braycote 807 Aircraft Grease MIL-G-27617, Type IV Castrol 807

approximate mass of 50 mg and volume of 0.02 cm was placed on a sixteenth pan to correct for buoyancy effects. The 16 pans were suspended from quartz springs which were hung on the spring

stand and placed in the pressure vessel described above and shown in Figure 1 . After aligning the

spring stand so that all 16 springs were visible through the view port, the lid of the vessel was bolted

in place and the vessel was completely immersed in silicone bath oil. The aluminum lid was placed

over the Pyrex jar and the mixer, heater and thermoregulator were adjusted to maintain the maximum

test temperature. After evacuation and temperature equilibrium was achieved, the displacements of

the calibrated springs were measured using a cathetometer. Next, the pressure was increased to the

maximum pressure considered.

Following a swelling equilibration time, which was determined independently and checked, the

displacements of the springs were measured again for each sample. The vapor pressure was then

decreased to the next lower value, allowed to equilibrate and then measurements were taken. The

process was repeated for the remaining pressures considered. The thermoregulator was then reset for the next lower temperature and the system was allowed to equilibrate overnight. After measurements

were obtained for all temperatures considered, the vessel was evacuated and the spring displacements were checked for reversibility.

This method of increasing the temperature and pressure to the maximum value considered and

then reducing the pressure in increments (instead of increasing the pressure incrementally from a vacuum) was used in order to minimize the disruption of the system which occurs when pumping at
SECTION 8. ELASTOMER SEAL COMPATIBILITY 737

high pressures. The undesired saturation condition and associated liquid condensation, which could

ruin the measurements, was also more easily avoided.

In this work, swelling measurements at four vapor pressures were taken for each temperature and

were usually done in the morning after equilibration overnight and in the late afternoon after equilibration during the working hours. Measurements were taken at 35, 70, 105, and 150 °C. In

addition, FC-116 (hexafluoroethane) was also tested at 5 °C (/?° = 2.133 MPa) because this fluid has a low critical temperature {T^ = 19.7 °C) and little swelling was observed at temperatures above T^. The maximum pressure considered was 5.86 MPa. For test temperatures above the critical temperature of the agent, vapor pressures were chosen to cover the range from the maximum pressure (5.86

MPa) to the saturation pressure of the highest test temperature for which saturation occurs. For

temperatures below T^, vapor pressures were chosen to cover the range from saturation for the test temperature to the saturation pressure of the next lower temperature considered. Saturation pressures

for each temperature were determined using thermodynamic properties software (Gallagher et al. , 1991; Allied Signal, Inc., 1989) or vendor supplied data (E.I. du Pont de Nemours, Inc., 1968 and

1993; Robin, 1992; Wilson et al., 1992) and are shown in Table 3.

Table 3. Saturation data for agents

Trade name

Formula Chemical name P°@

35°C 70 °C 105°C (°C)

Pc (MPa)

HFC-236fe CF3CH2CF3 hexafluoropropane 0.381 1.07 2.11 130,6 3,18

HFC-32/125 CH2F2/CHF2CF3 azeotrope 2.16 4.74 __a 73,2 5,06

HFC-227ea C3HF, heptafluoropropane 0.615 1.49 101,7 2,91

HCFC-22 CHF2CI chlorodifluoromethane 1.35 3.00 96,0 4.99

HFC-134a CH2FCF3 tetrafluoroethane 0.887 2.12 101,1 4.07

FC-116 C2F6 hexafluoroethane 19.7 3.03

HCFC-124 CHFCICF3 chlorotetrafluoroethane 0.516 1.26 2,63 118.9 4.60

HFC-125 CHF2CF3 pentafluoroethane 1.78 66.3 3,62

FC-218 C3F8 octafluoropropane 1.15 2.41 71,9 2,68

FC-31-10 C4F10 decafluonDbutane 0.364 0.921 1.97 113.2 2,30

FC-318 cyclo-C4F8 octafluorocyclobutane 0.428 1,08 2,25 115,2 2,78

*not defined

8.2.3 Results and Discussion. In Figures 2-4 typical plots of solvent weight fraction vs. vapor

pressure p are shown for different elastomer/agent systems. The solvent uptake of the polymers

strongly increases with the vapor pressure. The vertical arrows in the figures show the saturation

vapor pressures of the solvent p°j at different temperatures. In Figure 4 it is demonstrated that the

swelling degree of silicone in decafluorobutane (FC-31-10) significantly decreases with increasing
738 SECTION 8. ELASTOMER SEAL COMPATIBILITY

35

0.4

0.3

0.2

0.1

0.0

1 1 FLUOROCARBON (FKM)

1

# STTJCONF CSi)

V FLUOROSILICONE (FSi)

NEOPRENE (CR)

•

• V

T •

V

1

1 1 0.20 0.25 0.30 0.35 0.40

p (MPa)

Figure 2. Weight fraction of FC-31-10 as a function of pressure for various polymers at 35 °C. The arrow shows the saturation vapor pressure = 0.36 MPa of the

agent.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 739

o o P DFB npR P HFP CFM

0.4

0.3

0.2

0.1

0.0

\ /' \

1 1 1 1 1 DECAFLUOROBUTANE (FC-31-10)

•

V HEPTAFLUOROPROPANE

(HFC-227)

CHLORODIFLUOROMETHANE

(HCFC-22)

• V

T

•

T

V

V

1 1 1

T

1 1 1 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

p (MPa)

Figure 3. Swelling of uncrosslinked silicone in various agents as a function of pressure at 35 °C. The arrows show the saturation vapor pressures of the respective agents.
740 SECTION 8. ELASTOMER SEAL COMPATIBILITY

0.4

0.3

0.2

0.1

0.0

o o P 35 P 70

105

1 1 T =

c

1 113.2 °C •

1

•

-

V

T

70 °C

1 Uo C

150 °C •

V

• V

T

•

V

T

V T

1 n 1 1 1 0 0.5 1.0 1.5 2.0 2.5 3.

p (MPa)

Figure 4. Weight fraction of FC-31-10 as a function of vapor pressure in uncrosslinked

silicone at various temperatures. The arrows show the saturation vapor pressures at the respective temperatures.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 741

temperature. At 150 °C, i.e., above the critical temperature of decafluorobutane (T^ = 113.2 °C), the

fluid uptake was approximately one order of magnitude smaller than at 105 °C.

In Figures 5-6, the same experimental data shown in Figures 2-4 are shown according to the

Flor>'-Huggins representation. The continuous curves were calculated using least squares fits of

Equation 9 to the data points. The experimental values of In (p / jf) vs. were satisfactorily

described using a temperature and concentration dependent interaction parameter. The values of the

Flory-Huggins interaction parameters (x = Xo + Xi) calculated from solubility data at difi'erent temperatures are given in Tables 4 and 5 for all the polymer/solvent systems investigated. The x

values are measures of the solubilities (compatibilities) of the subcritical agents in various polymers.

(Small x values correspond to good solubility or, for present purposes, bad compatibility. Above x = 0.5, solubility of the polymers gradually decreases with increasing x-) In Figure 7 the swelling of crosslinked and uncrosslinked silicone are compared. The difl'erence between the swelling behavior of the two samples is apparent: the crosslinked polymer exhibits

considerably lower affinity to the solvent than the uncrosslinked one. The values of the interaction

parameters calculated from solubility data at 35 °C for the crosslinked and uncrosslinked polymers

are x — 0.99 and x = 104, respectively. Strictly speaking Equation 9 can only be applied for the

uncrosslinked polymer. The appropriate expression for the crosslinked polymer contains an additional

term arising from the change in Gibbs free energy due to the elastic deformation of the network on

swelling. For lightly crosslinked networks at high polymer concentration (wj > 0.7), the contribution

of the elastic term is negligible in comparison to the mixing term. Above the critical temperature of the fluid the swelling behavior of the polymers can be analyzed on the basis of Equation 11 (Flory's equation-of-state theory). In Figure 8, the fits of Equation 11 to the swelling data for the silicone/decafluorobutane system at four temperatures between 35 °C and

150 °C are shown. Values of the equation-of-state parameters used in these calculations were

obtained from the literature (Shih and Flory, 1972; Wilson et ai, 1992). The fit of Equation 9

(Flory-Huggins theory) to the data are also shown in Figure 8 for the three temperatures below the

critical temperature = 113.2 °C of FC-31-10 (decafluorobutane). Both theories yield reasonable

fits to the experimental data below the critical temperature of the solvent. In addition, Flory's

equation-of-state theory yields a reasonable fit to the experimental data above T^. In order to rank the agents in terms of their compatibility with the sealing materials, a rating

system was defined for the data at 35 °C. This temperature was chosen for several reasons. First, as shown in Figure 4, swelling is greatest at 35 °C and decreases with increasing temperature. Second,

it is near the initial storage temperature of 25 °C and third, it is below the critical temperature for every agent considered except FC-116 (hexafluoroethane). FC-116 was rated based on its swelling

data at 5 °C because it has a critical temperature of 19.7 °C. A good, bad, or fair rating is given for each agent/crosslinked elastomer and agent/lubricant system in Table 6. Bad compatibility was

defined as a condition of excessive swelling and corresponding to x < 0.9, where x is the FloryHuggins polymer-solvent interaction parameter. Good compatibility implies that the elastomer or

lubricant is acceptable for use in the fire suppressant system and corresponds to x > 1-2. For values

of X between 0.9 and 1.2, the agent was considered to have fair compatibility with the elastomer or grease and represents a marginally acceptable system. Under saturation conditions, x values of 0.9

and 1.2 represent agent weight fractions of 0.38 and 0.22, respectively.

The X values shown in Tables 4 and 5 were calculated using the curve fitter in the scientific graphing software SigmaPlot for DOS (Jandel, 1992). The curve fitter uses the Marquardt-Levenberg

algorithm (Press et ai, 1986) to find the parameters which minimize the variance s^ (In p I p°), i.e., the squared differences between the observed and predicted values of the dependent variable. The

standard error or experimental standard deviation of the x parameter, s ( x ), and the coeflficient ol variation, CV = s ( x ) / X x 100, were also calculated for each x value. Thus, the standard
742 SECTION 8. ELASTOMER SEAL COMPATIBILITY

Table 4. Flory-Huggins interaction parameters x for the subcritical temperatures studied for each

agent/lubricant system considered

Agent (T °C) Krytox 240AC Braycote 600 Braycote 807

236fa (35) 1.54 1,60 1,72

236fe (70) 1.53 1,61 1.75

236fe (105) 1.65 1,73 1.88

32/125 (35) 1.98'' 1,63 1,86''

32/125 (70) 2.08 1,66 1.96

227ea (35) 1.14" 0,93 0,93''

227ea (70) 1.23 0,97 1.02

22 (35) 1.08 1,30 1.20

22 (70) 1.10 1,31 1.23

134a (35) 1.64 1,21 1.76

134a (70) 1.68 1,24 1.65

116 (5) 1.08 1,13 1.05

124 (35) 1.31'' 1.37'' 1.31

124 (70) 1.28 1.37 1.30

124 (105) 1.30 1.39 1.33

125 (35) 1,18 1.37 1.35

218 (35) 1.83^ 1.45 1.76

218 (70) 1,86 1,47 1.79

31-10 (35) 0.93 0,73 0.82

31-10 (70) 0.95 0,75 0.83

31-10 (105) 1.01 0,80 0.88

318 (35) 1,03 1,00 1.00

318 (70) 1.04 1,03 1.01

318 (105) 1,10 1,10 1.10

"12.5 < CV < 20% and 0.64 < x < 1-5 (35 °C)

^CV > 20% (35 °C)

uncertainty of the parameter x is « ( x ) = s ( x ) and the percentage of relative standard uncertainty

isM(x)/xxlOO= CV. The expanded uncertainty is U - k u { x ), where k = 2, and the

percentage of relative expanded uncertainty is f/ / x x 100 = k CV.

Unless otherwise indicated in Tables 4, 5, and 6, the coefficient of variation was less than 12.5%

for the X parameters at 35 °C which gives a percentage of relative expanded uncertainty of 25%.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 743

Figure 5. Solvent activity in various polymers in FC-31-10 as a function of polymer weight

fraction at 35 °C. The continuous curves show the least squares fits according to Equation 9.
Figure 6. Solvent activity in uncrosslinked silicone as a function of the polymer weight

fraction H'2 at 35 °C in various agents. The continuous curves show the fits according to Equation 9.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 745

Table 5. Flory-Huggins interaction parameters x for subcritical temperatures studied for each

agent/elastomer system considered

Agent (T u X u X u X u X u X

Si Si N206 N206 FSi FSi FKM FKM CR CR N926 i^y

1.09 1.10 1 .02 1.05 0.66 0.68 0.76 0.77 n 81 n 84 u.ou U.OH

1 .08 1 .12 1 .03 1 .07 0.68 0.71 0.77 0.79 n 81 fl 85 n (^1 U.Oj

1.12 1 15 1.19 1 .20 0.76 0.79 0.81 0.83 0\j . QOy\j fl Q7 nW . 84oH flU.oO86

1 .UJ 1 48 1 57*' 1 . ^z 1 .24*' 11 . zo78 1 1 4''

1 1

0

1 1 7 1 . J /

1 /lib

1 .41

1 \I\J ) 1 .08 1.13 1 .51 1 .56 1 .27 1 .34 1 .18 1 .23 1.16 I

1

. 1L J5 11.4-1 41 1 47

0.89 0.93^ 1.34*' 1 .40*' 0.97 1.04^ 1 .01 1 .07'' 1,25^ 1 .33^ 0.80 0.84

227ea HOI 0.94 0.96 1 .48 1 .50 1.10 1 .13 1 .09 1 .13 1 .31 1 .38 0.85 0 87

0.98 1 .00 0.72 0.75 1 .02 1 .05 0.97 1.02 1.20'' 1 .26 0.76 n 81

22 f70^ 0.99 1 .00 0.74 0.75 1 .05 1 .07 0.99 1 .06 1 .26 1.28 0.78 0.82

134a f35) 1.16 1.20 1.35 1.36 1.00 1 .02 1.25 1 .30 1 .96 2.00'' 1 .42 1 .44

1.19 1 .22 1 .37 1 .38 1 .02 1 .06 1 .28 1 .31 2.03 2.04 1 .44 1 .46

116 CSI 1 .4P 1 .47^ 1 .94^ 1 .92*' 1.35^ 1.42 1.33^ 1 .40 1.60 1 .68'' 2. 1 1'' 2.27^

124 (35) 0.88*' 0.91*' 0.88*' 0.92'' 0.93^ 0.95'' 0.96'' 0.97'' 0.97'' 0.98'' 0.91*' 0.94''

124 (70) 0.85 0.88 0.86 0.89 0.90 0.93 0.95 0.96 0.96 0.98 0.89 0.92

124 n05^ 0.86 0.91 0.91 0.94 0.93 0.95 0.97 0.97 0.98 1 .01 0.92 0.96

125 (35) 1 .29 1 .30 1 .69^ 1.72 1 .33 1 .35 1.18 1 .21 4.10'' 1 .99 1 .23 1 .28

1 55 1 11^ 1 84 1 5fi^ 1 . JO 1 f\A L . OH 1 fiS 1 . DJ 1 .OJ 1 ^;8 1 .00 1 5^;''

1 . JO 1 7

1

1 . /

1

218 (70) 1.64 1.69 2.04 1.99 1.67 1.74 1.66 1.67 1.75 1.77 1.69 1.74

31-10 (35) 0.99 1.03 1.18 1.24 1.32 1.39 1.09 1.15 1.55 1.59 1.10 1.13

31-10 (70) 1.01 1.06 1.20 1.25 1.35 1.44 1.13 1.19 1.58 1.66 1.13 1.15

31-10 (105) 1.09 1.09 1.30 1.31 1.40 1.45 1.18 1.25 1.60 1.64 1.19 1.23

318 (35) 1.04*' 1.27 1.41 1.46 1.27 1.32^ 1.32 1.34 1.32'' 1.35 1.10 1.14

318 (70) 1.26 1.29 1.46 1.47 1.27 1.33 1.33 1.36 1.32 1.38 1.10 1.16

318 (105) 1.29 1.37 1.52 1.56 1.33 1.37 1.41 1.44 1.34 1.39 1.19 1.23

M2.5 < CV < 20% and 0.64 < x < 1-5 (35 °C)

^CV > 20% (35 °C)

"Hincrosslinked polymer

''crosslinked polymer
746 SECTION 8. ELASTOMER SEAL COMPATIBILITY

0.0

0.

1

• V

o -0.2

-0.3

V

0.4 ® uncrosslinked

V crosslinked

0.5

0.6 0.7 0.8 0.9

Figure 7. Comparison between the swelling of crosslinked and uncrosslinked silicone in

FC-31-10 at 35 °C.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 747

0.4

0.3 '

0.2

0.1

0.0

o o P 35 P 70

105

\1/

0 -35 °C

T =113.2 °C

c

Flory-Huggins Equation

Flory Equation of State

70 °C

105 °C

150 °C

0.0 0.5 1.0 1.5 2.0 2.5 3.0

p (MPa)

Figure 8. Comparison between measured and calculated values of the weight fraction of

FC-31-10 in silicone.
748 SECTION 8. ELASTOMER SEAL COMPATIBILITY

Hence, for an estimated x value of 1.2, the measurand (true value of x) is 1.2 + 0.3. This level of

confidence suggests that an agent which was assigned a good rating with a x value slightly greater

than 1.2 may actually fall into the fair range (0.9 < x < 1-2) but a bad rating (x < 0.9) is statistically unlikely. A similar argument can be made for an agent with a bad rating and an

estimated x value slightly less than 0.9. The agent/polymer systems for which 0.64 < x < 1.5 and 12.5 < CV < 20% at 35 °C are shown in Tables 4 and 5 with superscripts. If x = 1-5 and CV = 20%, the value of the measurand

is 1.5 ± 0.6 and the agent may fall into the fair range but a bad rating (x < 0.9) is statistically

unlikely. Similarly, if x = 0-64 and CV — 20%, the value of the measurand is 0.64 + 0.26 and the

agent may fall into the fair range but a good rating (x > 1.2) is statistically unlikely.

Estimated values of x for which CV > 20% are also shown with superscripts in Tables 4 and 5. In some instances, the uncertainty represented by large coefficients of variation may be attributed to the small degree of swelling (W2 = 1.0) in poor solvents. Using the idealization that the mixing

contribution to the chemical potential is zero gives the expression

ln(l - W-) + w,

X

- 1 (14)

where the slope of x goes to infinity as W2 1.0. Hence, small errors in due to uncertainty in the

cathetometer readings (± 0.05 mm) produce large differences in the estimate of x as W2 -* 1.0

because -ln(l - 00.

8.3 Durability

8.3.1 High Temperature Exposure Methods. As an independent method of characterizing the

compatibility of fire suppressant agents, mechanical property measurements of the elastomers and

greases were obtained after short term exposures to the agents at extreme conditions of 150 °C and

5.86 MPa. For this purpose, two separate pressure vessels without view ports were designed and

built. These pressure vessels were also made of stainless steel and designed for a maximum working

pressure of 5.86 MPa. The cylindrical vessel had 10 liters of inner volume (25 cm height by 23 cm

diameter) and was large enough to fit an 18 cm diameter compression set fixture, (20) samples for each of the six elastomers, and a 2 ml vial for each of the three greases. After placing the test samples in the vessel, it was placed in a Fisher Scientific Isotemp forced air lab oven with an inside capacity of 51 x 61 x 51 cm, temperature range of 30 to 200 °C with 1 °C

resolution, ± 3 °C accuracy and ± 3 °C uniformity. After evacuating the vessel, the charging and

recovery system described previously was used to pump a predetermined mass of agent into the

chamber at room temperature (24 + 1 °C). The value of mass necessary to obtain 5.86 MPa at 150 °C was determined using thermodynamic properties software (Gallagher et al., 1991; Allied

Signal, Inc., 1989) or vendor supplied data (E.I. du Pont de Nemours, Inc., 1968 and 1993; Robin,

1992; Wilson era/., 1992).

The oven was turned on, set to 150 °C, and monitored using a Fluke thermometer with a type K

thermocouple and 0.1 °C resolution. The pressure was monitored using a Druck DPI 602 digital

pressure indicator and PDCR 910 transducer with an absolute pressure range of 0-6.21 MPa,

combined nonlinearity, hysteresis and repeatability of ± 0.1% using the best straight line method

(BSL). After attaining thermal equilibrium, the pressure was reduced until the desired pressure of
SECTION 8. ELASTOMER SEAL COMPATIBILITY 749

Table 6. Compatibility of lubricants and crosslinked elastomers based on swelling measurements in various fluorocarbon agents at 35 °C

Agent OVJ 807/ ol rol r IS-M LK Mm£ Nyzo

good^ good good lair fair aa bad bad bad

good^ good good'^ fair good^ good good good good

HFL-227ea tair lair lair lair good*^ tair fair good*^ bad HCFC-22 fair good fair fair bad fair fair good bad

HFC-134a good feir good fair good fair good good*^ good

FC-116^ fair fair fair good^ good*^ good good good*^ good*^

HCFC-124 good*^ good*^ good badf fairf fairf fairf fairf fairf

HFC-125 fair good good good good good fair good good

FC-218 good*^ good good good good good good good good

FC-31-10 feir bad bad fair fair good fair good fair

FC-318 fair fair fair good good good^ good good fair \ > 1.2 •^.9 < X ^ 1-2 =X < 0.9

''measured at 5 °C

'12.5 < CV < 20% and 0.64 < x < 15

'CV > 20%

5.86 MPa was achieved. The pressurized vessel was then left undisturbed until the exposure time was completed. Short term exposure times of one, 2, 4, and 6 weeks were considered for the initial screening of the agents. After the completion of each exposure, the oven was turned off and the

agent was recovered into an evacuated storage cylinder immersed in a water bath.

To maintain the project time schedule, two of the small pressure vessels used for swelling

measurements were used as containers for additional high temperature exposures. The procedure

described above was followed except a silicone oil bath setup similar to that used for the swelling

measurements was employed for heating. However, in order to expedite the sample placement and

removal process, the pressure vessels were only partially (approximately half) immersed in the

silicone oil. This allowed for the removal of the vessel's top without waiting for the silicone oil to cool and eliminated subsequent pumping of oil to access the samples. The compression set fixtures

and lubricant samples were placed near the bottom of the vessel, below the level of the silicone oil. The tensile test samples were placed on top of the compression set fixtures. Some were above and some were below the level of the silicone oil. A set of tests was conducted in which two thermocouples were placed inside a small pressure

vessel filled with air at ambient pressure (high pressure agent could not be used because of leaks due

to the presence of the thermocouples). One was placed near the bottom and the other near the top of

the vessel. A third thermocouple was immersed in the silicone oil bath which was heated to 150 °C.
750 SECTION 8. ELASTOMER SEAL COMPATIBILITY

These tests show that temperature within the pressure vessel may have varied from 140 to 149 °C

when the temperature of the silicone oil was 150 °C. During the actual experiments, this variation was probably less due to the larger heat conduction (relative to air) of the supercritical fluids at 5.86

MPa.

This setup was used for high temperature exposures of FC-116 (hexafluoroethane), FC-218

(octafluoropropane), FC-31-10 (decafluorobutane), and FC-318 (octafluorocyclobutane). As will be

shown in the following sections, high temperature exposures to these agents were severe for most of

the elastomers and lubricants. Consequently, the results show that exposures throughout the entire

temperature range of 140-150 °C are severe.

8.3.2 Elastomers

8.3.2.1 Experimental Materials. The o-rings used in this study were Parker No. 2-214 with a nominal cross section diameter of 3.2 mm (1/8 in.), actual cross section diameter of 3.53 ± 0.10

mm, and a mean outside diameter of 3.2 cm. The Parker compounds considered are listed in Table 7.

Table 7. Elastomers used in durability experiments

Elastomer Vendor Designation

Silicone Parker S604-70^

Nitrile (standard industrial) Parker N674-70

Fluorosilicone Parker LI 120-70

Fluorocarbon Parker VI 164-75

Neoprene Parker CI 185-70

Nitrile (low temperature industrial) Parker N103-70

^The number following the dash in the designation (compound number) represents the Shore

hardness of the elastomer.

8.3.2.2 Compression Set Measurements. Compression set tests were conducted per standard

test methods (ASTM, 1989 and 1990b). Measurements were taken after the one, 2, 4, and 6 week

exposure times. For this purpose, eight compression set fixtures were designed and built. Each

fixture consisted of two 18 cm diameter and 12.7 mm thick plates, seven spacers 19 mm in diameter

and 2.65 ± 0.01 mm thick, and seven sets of 10 mm bolts, nuts, and washers. The plates were made

of stainless steel 304, ground to a maximum roughness of 250 nm, chrome plated and polished. The

spacers were also made of stainless steel 304 and their thickness was specified to be 75 % of the actual

mean value of the o-rings (0.75 x 3.53 mm).

Specimens were prepared by cutting 52 mm sections from sample o-rings. The original thickness

of each specimen was measured using a hand micrometer with a 6 mm diameter hemispherical tip.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 751

Measurements were taken at four marked locations equally distributed around the circumference in both the radial and axial direction. The average reading was used for subsequent calculations. Two

specimens of each compound were placed into a fixture and closed by tightening the bolts in a criss- cross manner so that the plates were drawn together uniformly. The compression set fixtures were

placed in a vessel at the beginning of each high temperature exposure. After the pressure vessel was removed from the oven, the compression set fixture was removed from the vessel and the specimens were placed on a poor conducting surface (wood table) and allowed to cool at room temperature for 30 minutes. After the rest period, the thickness measurements were made as described above at the same marked locations. The same specimens were used for all exposures (one, 2, 4. and 6 weeks) to a given agent.

The compression set Cg was calculated using the average thickness readings before and after the

exposure and is expressed as a percentage of the original deflection as follows:

C« = ^^-^ X 100 , (15)

where and are the original and final thickness, respectively, = 2.65 mm is the thickness of the spacers used in the compression set fixtures. This calculation was made for both specimens of each

compound after each exposure. The average of these two values was reported as the compression set

Table 8 shows the experimental values of compression set for each of the compounds after exposures of one, 2, 4, and 6 weeks at 5.86 MPa and 150 °C in each fluorocarbon agent. In some

instances, the specimen was split or broken after a given exposure time. For these cases, the

compression set measurements were still taken but the value is superscripted in Table 8. In Table 9, compatibility ratings of the agents based on compression set measurements of each

elastomer are shown. Bad compatibility was defined as a condition when the compression set exceeded 90% after a 2 week exposure. Elastomers have good compatibility if the compression set was less that 90% after 4 weeks. If the compression set was less than 90% after 2 weeks but

exceeded 90% after a 4 week exposure, the agent was considered to have fair compatibility with the

elastomer and represents a marginally acceptable system. If a specimen was split or broken, the agent was considered to have bad compatibility with the elastomer.

As specified in Option 1 of the standards for compression set (ASTM, 1989 and 1990b), the

value of compression set for each elastomer/agent system reported in Table 8 is the average of the

compression set calculated from measurements of two samples. The experimental standard deviation

s fCg) and the standard uncertainty of the mean u (Cq) - s (Cg) iVn, where « = 2 is the number of

samples, were also calculated. The coefficient of variation is CV = u (Cg) / Cg x 100, the expanded

uncertainty is U = k u (Cq), where k = 2, and the percentage of relative expanded uncertainty is k CV. The maximum value of CV was 10%. Hence, for an estimated compression set of 90%, the

measurand (true value) is 90 ± 18%. This level of confidence suggests that an elastomer with an

estimated compression set value of approximately 90% after 4 weeks may fall into the good or fair category. However, in order for an elastomer assigned a good rating to actually have bad compatibility, the estimate for compression set based on a separate set of data after 2 weeks must also exceed

90% which is unlikely because compression set decreases monotonically with decreasing exposure

time. Similar arguments can be made for elastomers with estimated compression set values of

approximately 90% after 2 weeks.

8.3.2.3 Tensile Testing. Tensile tests were conducted per standard test methods (ASTM, 1990a

and 1990b). For each agent, twenty sample o-rings of each compound were tied together with wire
752 SECTION 8. ELASTOMER SEAL COMPATIBILITY

Table 8. Compression set measurements after 1, 2, 4, and 6 week exposures at 5.86 MPa and

150 °C in various fluorocarbon agents

U1 1 1 Ui.J li 1 ILl (weeks)236fa 32/ 125227ea 22 134a 116 1 94 1 ?s 9z 11 8

0

1 1 ft

S604-70 (1) 27 34 25 16 35 30 11 33 41 41 31

S604-70 (2) 38 36 34 28 37 41 26 36 48 50 35

S604-70 (4) 52 48 48 66 47 53 35 50 50 61 45

S604-70 (6) 56 69 57 109' 59 61 59' 58 52 62 51

N674-70 (1) 81 79 77 76 84 59 98 70 77 78 61

N674-70 (2) 92 92 84 96' 87 78 103 84 88 89 82

N674-70 (4) 96 93 93 110' 98 90 104 91 92 98 94

N674-70 (6) 98 99a 97 113' 102 96 104 96 95 98 97

L1120-70 (1) 53 66 51 48 60 64 30 65 68 82 60

LI 120-70 (2) 65 78 56 62 54 73 51 65 74 87 65

L1120-70 (4) 88 87 72 86 72 83 69 80 76 91 11

LI 120-70 (6) 94 104' 83 99 78 88 89 90' 79 91 87

VI 164-75 (1) 31 41 21 46 36 19 14 35 25 25 62

VI 164-75 (2) 47a 63 33' 100' 40' 31 30 46' 39 39 80

VI 164-75 (4) 67' 82' 60' 100' 55' 55 59' 72a 56 64 92

VI 164-75 (6) 94a 99' 77' 100' 94' 70' 91' 90' 58' 66' 95

C1185-70 (1) 69 67 66 68 72 54 96 68 67 71 49

CI 185-70 (2) 76 76 75 77 79 68 100 70 79 78 70

CI 185-70 (4) 84 78 81 95 86 79 105 80 83 89 79

CI 185-70 (6) 86 81 84 100 94 84 107 98 82 90 84

N103-70 (1) 79 81 79 79 82 70 88 73 78 82 62

N103-70 (2) 87 91 87 89 85 82 91 83 87 91 80 N 103-70 (4) 93 95 91 110 96 92 94 92 90 97 89 N 103-70 (6) 95 98 94 113 99 95 98 83 95 97 93

'specimen was split or broken
SECTION 8. ELASTOMER SEAL COMPATIBILITY 753

Table 9. Compatibility of elastomers based on compression set measurements after exposures at 5.86 MPa and 150 °C in various fluorocarbon agents

Agent S604-70 N674-70 L 11 20-70 VI 164-75 Cl 185-70 N103-70

HFC-236fa good^ bad^ good good bad bad

HFC-32/125 good bad fair'^ bad fair bad

HFC-227ea good bad good good fair bad

HCFC-22 good fair good good fair fair HFC- 134a good bad good good fair bad

FC-116 good fair good fair good bad

HCFC-124 good bad good good bad bad

HFC-125 good bad good bad good bad

FC-218 good fair good fair good bad

FC-31-10 good bad good fair fair bad

FC-318 good fair good good good fair

^compression set is less than 90% after four weeks

^compression set is less than 90% after two weeks but exceeds 90% after four weeks

^compression set exceeds 90% after two weeks

and stored in a pressure vessel. Ultimate elongation, tensile stress, and modulus measurements were

taken after one, 2,4, and 6 week exposure times. A type TT-B Instron tensile testing instrument was

used for this purpose. Grips for testing o-rings were designed and built from low carbon steel. The

grips consisted of ball-bearing spools 8.9 mm in diameter and were capable of being brought within

19 mm center-to-center distance at closest approach. Stresses were minimized by lubricating the contact surfaces of the spools with Super-Lube o-ring lubricant.

After the exposure was completed, the pressure vessel was removed as described in the previous

section and five samples of each compound were removed, placed in sealed plastic bags and labeled

with the agent, compound, exposure time and date. All tensile tests were conducted within a 2 week

period near the end of the short term exposures in August 1993. Each specimen's thickness W was measured at four points equally distributed around the circumference in the radial and axial direction

using a dial indicator with a contact force of 0.2 N. The average reading was used for calculation.

The internal diameter of each specimen was measured using a stepped cone with diametric intervals of

0.50 mm.

The grips were brought together and the o-ring specimen was installed with minimal stretching.

The crosshead and recorder were set to a speed of 500 ± 50 mm/min. A load cell with full scale

load ranges of 2, 5, 10, 20, and 50 kg was used. Chart paper with a grid of 2.54 x 2.54 mm was

used to plot load vs. jaw travel. The charts were labeled with the initial jaw position starting point,

load range, agent, compound, sample number, temperature, date, average thickness and inside
754 SECTION 8. ELASTOMER SEAL COMPATIBILITY

diameter. The force F was recorded at rupture and at an elongation of 100% (if the ultimate

elongation exceeded 100%).

The tensile strength T.S. is the tensile stress at rupuire and is calculated as

T.S. = - , (16)

where F is the breaking force and A is twice the cross-sectional area calculated from the thickness W

as A-^. (17)

The ultimate elongation U.E. is expressed as a percentage of the original inside circumference as follows:

U.E. = X 100 , (18)

where D is the distance between centers of the spool grips at the time of rupture. G is the circumfer- ence of one spool (tp x spool diameter), and C is the inside circumference (tt x inside diameter) of the

specimen. Assuming neo-Hookean behavior (Mooney, 1940: Treloar, 1948: Atkin and Fox, 1980),

the modulus Cj is expressed as

=

, (19)

2AiX^ - 1/A)

where F is the tensile force and X is the stretch ratio in the direction of extension. The center-to- center distance of the spool grips D for a prescribed stretch ratio X is calculated as D ^ -[XC - G] . (20)

The median of three specimens was taken as the characteristic value. Table 10 shows the percent

decrease in ultimate elongation for each of the compounds after one, 2, 4, and 6 week exposures at 5.86 MPa and 150 °C in each fluorocarbon agent. In some instances, the o-ring could not be

installed without breaking the specimen because of embrittlement due to the high temperature

exposure. For these cases, the specimen was considered to have a 100% decrease in ultimate

elongation.

In Table 1 1 , ratings for the compatibility of elastomers are shown based on tensile testing after exposures at 5.86 MPa and 150 °C in various fluorocarbon agents. Bad compatibility was defined as a condition when the decrease in ultimate elongation exceeded 65 % after a 2 week exposure. Elasto- mers have good compatibility if the decrease was less that 65% after 4 weeks. If the decrease in ultimate elongate was less that 65 % after 2 weeks but exceeded 65 % after a 4 week exposure, the

agent was considered to have fair compatibility with the elastomer and represents a marginally

acceptable system.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 755

Table 10. Percent decrease in ultimate elongation measurements after 1, 2, 4, and 6 week exposures

at 5.86 MPa and 150 °C in various fluorocarbon agents

Compound

(weeks)

236fe 32 /125 227ea 22 134a 116 124 125 218 31-10 318

S604-70 (1) -7.1 -8.4 7.52 5.75 1,77 23 12,8 8,41 -1,3 7.08 3,1

S604-70 (2) -1.3 8.85 10.6 8.85 22.1 8.85 11.5 7,08 6,19 3,1 12,8

S604-70 (4) 18.1 13.7 -1.8 8.85 -13 10,6 -9,7 5,31 9,29 12,8 3.1

S604-70 (6) -3.1 14.6 -6,2 100 -2.7 18,6 0 10,6 4,42 15 13,3

N674-70 (1) 50.3 57.1 32.9 40.3 43,4 36,6 92 48,9 37.7 46,9 23.4

N674-70 (2) 83.9 87.3 80.7 52.9 90.7 64,9 98,2 74,6 59.1 76 52.6

N674-70 (4) 100 98 100 100 98 84,5 100 95,1 73,7 87.9 73.2

N674-70 (6) 100 98.7 99.8 100 100 93 100 98.7 81.6 95.8 83.5

LI 120-70 (1) -0.6 29.3 17.8 48.9 42.5 17,8 -13 8.62 47.9 25.3 2,87

LI 120-70 (2) 5.17 37.4 6.9 29.9 25.3 49.7 -4.6 22,4 31 45.2 29.3

LI 120-70 (4) 42.5 80.1 41.4 41.4 60.1 64,2 39.7 56,3 43,2 53.4 52.6

LI 120-70 (6) 78.8 100 41.4 62.3 71.3 46,8 46 75,1 43,2 61,3 60.1

VI 164-75 (1) 38.1 54.8 53.9 44.2 50 42,2 34,5 59.8 50.9 47.9 58.6

VI 164-75 (2) 55.6 68.3 41.2 55.1 43,1 60.6 28,6 69.6 52.8 51.7 41.1

VI 164-75 (4) 64.5 76.1 62.5 52.8 61.7 73.2 48.9 70.3 72,2 68.3 47

VI 164-75 (6) 76.8 92.5 68.3 65,7 72.9 77.4 46 87.7 76,2 80 66

C1185-70 (1) 36.7 37.9 36.7 26,2 27.7 31,3 78.4 23,8 25 27,7 12.9

CI 185-70 (2) 65.8 62.6 63.9 38.7 63,9 27,3 88.6 41 41.4 46,9 36.3

CI 185-70 (4) 85.6 78.9 85.2 100 95,5 55,1 97.4 63.9 48.8 69,6 48.8

CI 185-70 (6) 97.4 86.6 95.9 100 100 69.5 98.5 88 58.6 77,7 57.4

N103-70 (1) 45.2 48 29.4 13.1 36,1 41,3 95.2 60.8 50,8 39,3 9.13

N103-70 (2) 83.6 88.5 79.9 34.5 94.4 73,7 100 79 72.9 78,5 45.2

N103-70 (4) 100 99.1 99.6 100 100 90 100 94.4 83.5 90,4 72.5

N103-70 (6) 100 100 100 100 100 95 100 100 89 96,4 82.7

Since the standards for tensile testing (ASTM, 1990a and 1990b) specify the characteristic value

as the median of three specimens, an expression of uncertainty is not easily determined without

additional analysis and/or testing. However, arguments similar to those in Section 8.3.2.2 on

compression set can be made. That is, an elastomer with an estimated percent decrease in ultimate

elongation of approximately 65 % after 4 weeks may fall into the good or fair category but bad
756 SECTION 8. ELASTOMER SEAL COMPATIBILITY

Table 11. Compatibility of elastomers based on tensile testing after exposures at 5.86 MPa and

150 °C in various fluorocarbon agents

Agent S604-70 N674-70 LI 120-70 VI 164-75 CI 185-70 N103-70

HFC-236fe good^ bad' good bad good fair''

HFC-32/125 good bad bad bad good bad

HFC-227ea good fair good bad good feir HCFC-22 bad bad good bad fair feir

HFC- 134a good fair good bad good fair

FC-116 good fair good bad good fair HCFC-124 bad bad good bad bad bad

HFC- 125 good fair bad bad good fair

FC-218 good fair good bad good fair

FC-31-10 good fair fair bad good bad

FC-318 good fair good good good good

^decrease in ultimate elongation is less than 65% after 4 weeics

''decrease in ultimate elongation is less than 65% after 2 weeks but exceeds 65 % after 4 weeks

"^decrease in ultimate elongation exceeds 65% after 2 weeks

compatibility is unlikely. Similar arguments can be made for elastomers with an estimated percent

decrease in ultimate elongation of approximately 65 % after 2 weeks.

8.3.3 Lubricants. The same types of lubricants used in the swelling experiments (listed in Table 2),

i.e., Krytox 240AC, Braycote 600 and 807, were subjected to high temperature exposures and then

tested for changes in viscosity.

8.3.3.1 Viscosity Measurements. The original rheological properties of each grease were

characterized using the Rheometrics Mechanical Spectrometer Model 800. Approximately 10 mg of

grease was evenly placed in a cone and plate fixture with a diameter of 25 mm, cone angle of 0.02

radians, and gap of 0.023 mm. The instrument was used to perform steady shear testing of the

sample. During the test, the instrument subjects the sample to a rotational shear at a given shear rate 7 and measures the resulting torque. The instrument analyzes the sample's rotational response and

calculates the shear stress r and steady shear viscosity rj. An open 2 ml vial was filled with approximately 100 mg of grease and placed in the pressure

vessel. After each exposure approximately 25 mg of grease was removed and stored in a closed 2 ml

vial and labeled with the agent, lubricant name, exposure time, and date. All rheology tests were

conducted near the end of the short term exposures. Each grease was subjected to rate sweep

measurements. The rate sweep test was used to determine the response of the sample to a range of

steady shear rates at room temperature (24 + 1 °C). The test was performed with strain rates of 7 =

0.010, 0.016, 0.025, 0.040, 0.063, and 0.100 s"^ which correspond to six points in a logarithmic
SECTION 8. ELASTOMER SEAL COMPATIBILITY 757

decade. Figure 9 shows typical results for Krytox 240AC after high temperature exposures to HFC- 125 (pentafluoroethane). Note the absence of systematic time dependence of viscosity at 150 °C. The viscosity of the grease increases after 1 week exposure to HFC- 125, but then decreases

after the remaining exposures. This suggests that the diff'erences are due to experimental uncertainty

and no significant chemical degradation occurs. However, in some instances, viscosity measurements

were not possible because mobile substances were extracted and the greases became powder-like.

Table 12 lists the number of weeks after which the grease's viscosity was still measurable, i.e., had not become powder-like. Table 13 shows the ratings for the compatibility of lubricants based on

the viscosity measurements after exposures at 5.86 MPa and 150 °C in various fluorocarbon agents.

Bad compatibility was defined as a condition when the grease became powder-like after a 4 week

exposure. Lubricants have good compatibility if the grease did not become powder-like after 6

weeks. If the grease became powder-like after 6 weeks, the lubricant was considered to have fair compatibility with the agent.

Due to the nature of the criteria regarding the powder-like nature of the lubricants, an expression

of uncertainty is not readily determined without additional testing. However, arguments similar to those in Section 8.3.2.2 on compression set can be made. That is, a lubricant whose viscosity was

still measurable after 6 weeks may fall into the fair category but bad compatibility is unlikely.

Similar arguments can be made for lubricants which were powder-like after 4 weeks.

8.4 Summary and Conclusion

In order to investigate the potential for deterioration of the o-ring and its lubricant in the fire suppressant storage system, short term exposure experiments have been conducted and data have been

generated on the changes in the properties of various elastomers and greases after exposure to the

eleven fluid agents. Isopiestic measurements and durability testing proved to be sensitive methods for characterizing the compatibility of elastomers and lubricants with fluorocarbon agents. However,

exposures between 140 and 150 °C were found to be severe for most of the elastomers and lubricants. Additional testing at lower temperatures and for longer times is required to better simulate the

in-service conditions. In Table 14, the compatibility of elastomers and lubricants are shown based on the durability

ratings in Tables 9, 11, and 13 (G = good, F = fair, B = bad). If the ratings for compression set and tensile testing were different, the worse compatibility is listed. In parentheses, the ratings based on the swelling compatibility only (Table 6) are listed if diff'erent than the durability ratings. The

swelling ratings for crosslinked 55% butadiene-45 % acrylonitrile (N206) and 85% butadiene-15%

acrylonitrile (N926) are listed with the standard industrial nitrile (N674-70) and low temperature

industrial (N 103-70) compounds, respectively These ratings are summarized below for the 11 agents. If the ratings for durability and swelling compatibility were different, the worse compatibility

is reported.

HFC-236

Lubricants Krytox 240AC, Braycote 600 and 807 have good compatibility and silicone is a fairly

compatible elastomer. The nitrile, fluorosilicone, fluorocarbon, and neoprene elastomers are incompatible.
758 SECTION 8. ELASTOMER SEAL COMPATmiLITY

10

# unexposed

V 1 week

2 weeks

4 weeks

6 weeks

0

10' 0'

strain Rate (s

Figure 9. Viscosity measurements of the lubricant Krytox 240AC after various exposure times

in HFC-125 (pentafluoroethane) at 150 °C.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 759

Table 12. Number of weeks of exposure at 5.86 MPa and 150 °C in various fluorocarbon agents

after which viscosity measurements were still possible, i.e., the grease was not powderlike

Grease 236fe 32

/125

227ea 134a 116 124 125 218 31-10 318

Krytox 6 6

-) T 0 4 6 6 0 0 0 240AC wks wks wks wks wks wks wks wks wks wks wks

Braycote 6 6 6 2 1 6 6 6 2 0 0 600 wks wks wks wks wk wks wks wks wks wks wks

Braycote 6 6 6 4 0 6 6 6 1 0 0 807 wks wks wks wks wks wks wks wks wk wks wks

HFC-32/125

Lubricants Krytox 240AC, Braycote 600 and 807 have good compatibility and silicone and

neoprene are fairly compatible elastomers. The nitrile, fluorosilicone, and fluorocarbon elastomers

are incompatible.

HFC-227

Braycote 600 and 807 are fairly compatible lubricants and silicone, fluorosilicone, and neoprene

are fairly compatible elastomers. Krytox 240AC lubricant and nitrile, fluorosilicone, and fluorocarbon elastomers are incompatible.

HCFC-22

Braycote 807 is a fairly compatible lubricant and fluorosilicone is a fairly compatible elastomer.

Krytox 240AC and Braycote 600 lubricants and silicone, nitrile, neoprene, and fluorocarbon

elastomers are incompatible.

HFC- 134a

Silicone, fluorosilicone, and neoprene are fairly compatible elastomers. Krytox 240AC, Braycote

600 and 807 lubricants and nitrile and fluorocarbon elastomers are incompatible.

FC-116

Krytox 240AC, Braycote 600 and 807 are fairly compatible lubricants. Elastomers silicone,

fluorosilicone, and neoprene have good compatibility and nitrile (N674-70) is fairly compatible.

Fluorocarbon and nitrile (N 103-70) elastomers are incompatible.

HCFC-124

Lubricants Krytox 240AC, Braycote 600 and 807 have good compatibility and fluorosilicone is a

fairly compatible elastomer. Silicone, nitrile, and fluorocarbon elastomers are incompatible.

HFC- 125

Lubricants Braycote 600 and 807 have good compatibility and Krytox 240AC is a fairly good

lubricant. Elastomers silicone and neoprene have good compatibility. Fluorosilicone, nitrile, and

fluorocarbon elastomers are incompatible.
760 SECTION 8. ELASTOMER SEAL COMPATIBILITY

Table 13. Compatibility of lubricants based on viscosity measurements after exposures at 5.86 MPa

and 150°C in various fluorocarbon agents

Agent Krytox 240AC Braycote 600 Braycote 807

HFC-236fa good good good

HFC-32/125 good good good

HFC-227ea bad'^ good good

HCFC-22 bad bad fair^ HFC- 134a bad bad bad

FC-116 fair good good

HCFC-124 good good good

HFC- 125 good good good

FC-218 bad bad bad

FC-31-10 bad bad bad

FC-318 bad bad bad

^grease was not powder-like after 6 weeks

^grease was not powder-like after 4 weeks but was powder-like after 4 weeks

•^grease became powder-like after 4 weeks

FC-218

Elastomers silicone, fluorosilicone, and neoprene have good compatibility and nitrile (N674-70)

is fairly compatible. Krytox 240AC, Braycote 600 and 807 lubricants and fluorocarbon and nitrile (N 103-70) elastomers are incompatible.

FC-31-10

Silicone, fluorosilicone, and neoprene are fairly compatible elastomers. Krytox 240AC, Braycote

600 and 807 lubricants and nitrile and fluorocarbon elastomers are incompatible.

FC-318

Elastomers silicone, fluorosilicone, fluorocarbon, and neoprene have good compatibility and the

nitrile elastomers are fairly compatible. Krytox 240AC, Braycote 600 and 807 lubricants are incompatible.
SECTION 8. ELASTOMER SEAL COMPATIBILITY 751

Table 14. Compatibility of elastomers and lubricants based on the durability ratings in Tables 10,

12, and 14 (G = good, F = fair, B = bad)

Agent 240AC 600 807 S604- N674-70 L1120- V1164- C1185- N103-70 70 (N206) 70 75 70 (N926)

HFC-236fe G G G (F)'' B (F) G (B) B B B

HFC-32/125 G G G G (F) B (G) B (G) B (G) F (G) B (G)

HFC-227ea B (F) G (F) G (F) G (F) B (G) G (F) B (F) F (G) B HCFC-22 B (F) B (G) F B (F) B G (F) B (F) B (G) F fB) HFC- 134a B (G) B (F) B (G) G (F) B (G) G (F) B (G) F(G) B (G)

FC-116 F G (F) G(F) G F (G) G B (G) G B (G)

HCFC-124 G G G B B (F) G (F) B (F) B (F) B (F)

HFC- 125 G(F) G G G B (G) B (G) B (F) G B (G)

FC-218 B (G) B (G) B (G) G F (G) G B (G) G B (G)

FC-31-10 B(F) B B G (F) B (F) F (G) B (F) F (G) B (F)

FC-318 B (F) B (F) B (F) G F (G) G G G F

^If the ratings for compression set and tensile testing were different, the worse compatibility is listed

''If different, the rating for swelling compatibility (Table 7) is listed in parentheses

8.5 References

Allied Signal, Inc., Thermodynamic Properties of Refrigerants (Genie), SML Services, Inc. (1989).

ASTM, Designation D 395-89, "Standard Test Methods for Rubber Property - Compression

Set," Annual Book of ASTM Standards, 09.01, 24 (1989).

ASTM, Designation D 412-87, "Standard Test Methods for Rubber Properties in Tensions,"

Annual Book of ASTM Standards, 09.02, 100 (1990a).

ASTM, Designation D 1414-90, "Standard Test Methods for Rubber 0-Rings," Annual Book of

ASTM Standards, 09.02, 181 (1990b).

Atkin, R.J. and Fox, N., An Introduction to the Theory of Elasticity , Longman, New York

(1980).

Boyer, R.F., "Deswelling of Gels by High Polymer Solutions," J.Chem.Phys., 13, 363 (1945).

Brotzman, R.W. and Eichinger, B.E., "Volume Dependence of the Elastic Equation of State. 3. Bulk-Cured Poly(dimethylsiloxane)," Macromolecules 15, 531 (1982).

Brotzman, R.W. and Eichinger, B.E., "Swelling of Model Poly(dimethylsiloxane) Networks,"

Macromolecules, 16, 1131 (1983).

E.I. du Pont de Nemours and Co., Inc., Thermodynamic Properties of "Freon" 116 Refrigerant

(1968).
762 SECTION 8. ELASTOMER SEAL COMPATIBILITY

E.I. du Pont de Nemours and Co., Inc., Thermodynamic Properties of HFC-236fa Refrigerant,

(1993).

Flory, P.J., "Thermodynamics of High Polymer Solutions," J. Chem. Phys., 10, 51 (1942).

Flory, P.J., Principles of Polymer Chemistry, Cornell University Press, Ithaca, New York

(1953).

Flory, P.J., "Statistical Thermodynamics of Liquid Mixtures," J. Am. Chem. Soc, 87, 1833

(1965).

Flory, P.J., "Thermodynamics of Polymer Solutions," Disc. Faraday. Soc, 49, 7 (1970).

Flory, P.J. and Erman, B., "Theory of Elasticity of Polymer Networks 3 . ," Macromolecules 15,

800 (1982).

Flory, P.J., OrwoU, R.A., and Vrij, A., "Statistical Thermodynamics of Chain Molecule

Liquids. I. An Equation of State for Normal Paraffin Hydrocarbons," J. Am. Chem. Soc, 86, 3507

(1964a).

Flory, P.J., OrwoU, R.A., and Vrij, A., "Statistical Thermodynamics of Chain Molecule

Liquids. II. Liquid Mixtures of Normal Paraffin Hydrocarbons," J. Am. Chem. Soc, 86, 3515

(1964b).

Flory, P.J. and Rehner, J., Jr., "Statistical Mechanics of Cross-Linked Polymer Networks,"

J.Chem.Phys., 11, 521 (1943).

Gallagher J., McLinden, M., Huber, M., and Ely, J., NIST Standard Reference Database 23:

Thermodynamic Properties of Refrigerants and Refrigerant Mixtures Database (REFPROP), Version

3.0, U.S. Department of Commerce, Washington, DC (1991).

Gee, G., Herbert, J.B.M. and Roberts, R.C., "The Vapor Pressure of a Swollen Cross-Linked

Elastomer," Polymer 6, 51 (1965).

Horkay, F. and Zrinyi, M., "Studies on Mechanical and Swelling Behavior of Polymer Networks on the Basis of the Scaling Concept. 4. Extension of the Scaling Approach to Gels Swollen to Equilibrium in a Diluent of Arbitrary Activity," Macromolecules 15, 1306 (1982).

Horkay, F, Hecht, A.-M. and Geissler, E., "The Effects of Cross-Linking on the Equation of

State of a Polymer Solution," J. Chem. Phys., 91, 2706 (1989).

Huggins, M.L., "Some Properties of Solutions of Long-Chain Compounds," J. Phys. Chem., 46,

151 (1942).

James, H.M. and Guth, E., "Theory of Elastic Properties of Rubber," J. Chem. Phys., 11,455

(1943).

Jandel Scientific, SigmaPlot Scientific Graphing Software, Version 5.0, (1992).

McKenna, G.B. and Crissmann, J.M., "Thermodynamics of Crosslinked Polymer Networks: The

Anomalous Peak in Swelling Activity Measurements," Macromolecules 1992, ed by J. Kahovec, VSP,

Utrecht, The Netherlands, pp. 67-81 (1993).

McKenna, G.B., Flynn, K.M. and Chen, Y., "Experiments on the Elasticity of Dry and Swollen

Networks: Implications for the Frenkel-Flory-Rehner Hypothesis," Macromolecules 22, 4507 (1989).

McKenna, G.B., Flynn, K.M. and Chen, Y, "Swelling in Crosslinked Natural Rubber:

Experimental Evidence of the Crosslink Density Dependence of x," Polymer 31,1937 (1990).

Mooney, M., "A Theory of Large Elastic Deformation," J. Appl. Phys., 11, 582 (1940).

Pennings, A.J. and Prins, W., J. "A Versatile Osmometer for Polymer Gels and Solutions with

Applications to Cellulosic Gels," Polym. Sci., 49, 507 (1961).

Press, W.H., Flannery, B.P, Teukolsky, S.A., and Vetterling, W.T., "Numerical Recipes,"

Cambridge University Press (1986).

Robin, M.L., Thermodynamic and Transport Properties of FM-200, Great Lakes Corporation

(1992)
SECTION 8. ELASTOMER SEAL COMPATffilLITY

Shih, H., Flory, P.J., "Equation-of-State Parameters for Poly(dimethylsiloxane)," Macromolecules, 5, 758 (1972).

Treloar, L.R.G., "Stresses and Birefringence in Rubber Subjected to General Homogeneous

Strain," Proc. Phys. Soc, 60, 135 (1948).

Wilson, L.C., Wilding, W.V., and Wilson, G.M., Thermo-Physical Properties of Perfluorobutane, Wiltech Research Company (1992).

765

9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Emil Braun, Richard D. Peacock, Glenn P. Forney, George W. MulhoUand, and Barbara C. Levin

Building and Fire Research Laboratory

9.1 Potential Human Exposure

Although these agents are typically employed in unoccupied sections of an aircraft, the possibility of

human exposure still exists during handling, storage, and transport. Thus, it is important to know if the accidental release of the 12 agents in areas of typical occupancy would result in differing threats

to life safety. At least two topics are important in assessing the impact of a potential release of an

agent: 1) how does the agent distribute in an occupied space upon an accidental release, and 2) how

does this release affect personnel who may be exposed? For the former, a series of tests was

conducted to study the release of four of the twelve agents in a sealed compartment to measure the

airborne concentration of agent that results from complete venting of containers of typical size into spaces of typical volume. These tests were augmented with field modeling to extend the range of the

test results to other compartment geometries. For the latter, published toxicological results for chronic or acute exposure are summarized. It is important to note that in these tests, no humans were

exposed.

9.1.1 Full-Scale Tests. In order to evaluate the effects of an agent release in a closed compartment

on potential human occupants, a series of experiments were conducted under full-scale geometries to determine

• agent distribution within the compartment, and

• temperature distribution within the compartment

for near (within the jet plume) and far field conditions over a short period of time, i.e., acute exposure conditions. The test duration was 30 minutes. Temperature and gas concentrations at various levels within the compartment were measured during this "exposure" period. All agent

release tests were conducted with the cylinder oriented such that the agent liquid was at the bottom of

the cylinder and release occurred from the top. Cylinders were filled with a measured amount of

agent and pressurized to a nominal 4,200 kPa with nitrogen gas. Since resources (i.e., time, funding,

and agent availability) were limited for this phase of the project, tests were conducted on a subset of

agents that spanned a broad range (greater than a factor of 3) of molecular weights and vapor

pressures. While furnishing detailed data on the release process for a specific subset of agents, it was

the also the intent of these experiments to provide data that could be used to verify the computational

modeling of the release process. Therefore, in addition to halon 1301, only four agents were selected

for evaluation. These agents are listed in Table 1 along with the amounts and initial pressurization

states. Halon 1301 was used as a reference agent to determine if the observed conditions were

significantly different than those currently experienced through accidental releases of halon 1301.

9.1.2 Instrumentation. A compartment, measuring 2.45 m by 3.66 m by 2.45 m high, was

constructed and instrumented as shown in Figure 1 . Because of an extended doorway entrance the
766 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Table 1 . Agents tested in the full-scale geometry

Compound Formula Molecular

Weight

Vapor Press. @ 25 °C

(MPa)

Test Mass

(g)

FC-116 C2F6 138 _a 465

Halon 1301 CFjBr 149 1.61 420

391 HCFC-22 CHF2CI 87 1.05 189

297 HFC-227 C3HF7 170 0.47 108 HFC-32/HFC-125 CH2F2/C2HF5 67 1.67 227

"above critical pressure

total volume of the compartment was 22.7 m-'. A 1 liter stainless steel flask identical to those used in the stability tests was filled with test agent at approximately 4,200 kPa. The flask was attached to a support stand located in the center of the compartment. The support stand consisted of a swagelock

mating nut, pneumatic actuator and valve, and a 10 mm ID tube that extended to a release point 0.9

m from the floor (Figure 2). Each experiment involved installing the charged agent cylinder into the

support frame, sealing the compartment, and actuating the pneumatic valve to release agent into the

compartment. The agent cylinder was left open through the duration of the test, 30 minutes.

Thermocouple data were collected every ten seconds and evacuated flasks were filled in a specified

pattern (Table 2) to ensure that if stratification or settling occurred it could be detected.

Three types of measurements were made during the release of each agent. Thermocouple

measurements were made to determine the thermal efl'ects of a rapid release of agent. Gas concentrations were measured using evacuated flasks as well as FTIR measurements to determine agent

distribution in the compartment.

9.1.2.1 Temperature. Thermocouples were located on two vertical strings. One string of

thermocouples was located in a corner of the compartment, 0.6 m from each adjoining wall. A

second string of thermocouples was located 0.5 m north of the release point. Each thermocouple

string consisted of eight thermocouples spaced 0.3 m apart starting at the ceiling. Thermocouples

were also located 0.3, 0.6, and 0.9 m above the release point as well as one thermocouple 0.15 m

above and 0.08 m north and another thermocouple 0.3 m above and 0.15 m north of the release

point. Thermocouple data was collected every 10 seconds.

9.1.2.2 Gas Concentration. Using evacuated flasks, gas were sampled at three locations in the

compartment. One sampling probe was placed 1.23 m above the release point (0.31 m from the

ceiling). The other two sampling probes were located 0.46 m from the south wall. For tests involving agents FC-116, halon 1301, and HCFC-22 the two probes were located 0.92 m and 1.98 m,

respectively, from the ceiling. For agents halon 1301, HCFC-22, HFC-227, and HFC-32/HFC-125

the lower probe was placed 44.5 mm from the floor (approximately 2.41 m from the ceiling). This was done to determine if agent settling was occurring. Each sampling probe was connected to a four
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

Sample Flask and Support Frame

Q Thermocouple Trees - 8 TCs

0.3 m Intervals - ceiling=0.0

• Thermocouples - 3 TCs centered

2 TCs offset

2.45 m

3.66 m

Figure 1. Floor plan and instrument layout for halon 1301 replacement release tests.
768 SECTION 9 HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Pneumatic Valve Exit Tube

Sample Flask

Figure 2. Schematic of agent sample flask, support frame, manual valve, pneumatic valve, and

exit tube in relationship to the floor of the compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 769

flask manifold. Room atmosphere was drawn through the manifold by a vacuum pump operating at 2.5 1/m. In order to reduce atmospheric mixing in the compartment, gas samples drawn through each

manifold were returned to the compartment at the same level from which they were taken at approximately the same location. Prior to a test, 12 flasks, 125 ml each, were evacuated. Four

flasks were connected to each manifold. The vacuum pump for each manifold operated continuously,

except during the time it took to fill a flask. The content of each flask was analyzed by a Hewlett

Packard HP5730A gas chromatograph equipped with a flame ionization detector (FID). A column

packed with 60/80 Carbopack B/ 5 % Fluorcol was used to separate the agent from the background

atmosphere. For the two component agent (HFC-32/HFC-125), peaks were detected representing

each component. The carrier gas was nitrogen flowing at 30 ml/min. Each injection represented 500

/xl of gas drawn from a flask. An approximate calibration was determined for each agent by injecting

500 ^1 samples drawn from a flask containing a known amount of agent dispersed in a know volume

of air. Measurement errors associated with this calibration procedure consist of the error in measuring the volume of the calibration flask, the weight of the calibration flask before and after the

addition of the agent, and the volume of gas drawn from the flask for injection into the GC carrier

gas. A balance with a reading resolution of 0.1 mg was used to determine the mass of agent added to the flask. The balance was also used to determine the volume of the flask by measuring the mass of

water needed to completely fill the flask. An error of 0.1 mg in the mass of agent in 125 ml flask represents a variation of approximately 0.8 mg/1 in agent concentration or, for agent concentration

range of 5 to 21 mg/1, represents a variation of about 4 to 15%. Errors associated with the GC

system are a combination of syringe volume variation and GC system variation. These errors can be

approximated by computing the variation in integrator area calculations. This error was found to vary

from 1 to 4%. Combining these errors results in a final error approximation ranging from 5 to 19%.

Relative gas concentrations were also monitored in-situ with a Fourier Transform InfraRed

(FTIR) spectrometer. The FTIR was relocated for each test depending on the intent of the measurement. For tests of FC-116, halon 1301 (test 1301-1), and HCFC-22 (test 22-1), the FTIR was

located such that the sampling beam was approximately 1.2 m from the ceiling aimed diagonally

across the compartment. For tests of HFC-32/HFC-125 and HCFC-22 (test 22-2), the FTIR was

located 0.45 m from the ceiling with the beam parallel to and 0.52 m from the north wall. For the

retest of halon 1301 (test 1301-2), the FTIR was placed along the north wall 0.05 m from the floor. While the FTIR provided better time resolution of agent concentration than the evacuated flasks,

spatial resolution was provided by the data obtained from the use of evacuated flasks to sample the

enclosure atmosphere as specified above. It was hoped that comparable trends would be observed by

both measurement techniques.

9.1.3 Experimental Results

9.1.3.1 Temperature. The data from the thermocouple string located in the corner of the

compartment displayed only small deviations from ambient conditions, less than 1 °C. These far field temperature measurements indicate that the rapid release of any of the agents at the loading concentrations used in these experiments produced little deviation from ambient conditions. Near field measurements below the point of release also showed deviations of less than 1 °C from ambient

conditions for all tested agents. Data from these locations will not be shown or discussed further. Figures 3 to 9 show the temperature data for each test at the following three general sets of locations:
770

Table 2

SECTION 9. HLTvUN EXPOSURE AN'D ES\'lROSMEST.\L IMPACT

Description of evacuated flask location, sampling sequence and timing

Sequence Number Location

Distance from Ceilmg

I'm)

Desimation Sample Time

12

o.3r

0.31'

0.91^

1.98^-'

0.31'

0.91^

1.98^^

0.91^

1.98^'

0.31'

1.98^'

0.91^

2.41

2.41

2.41

2.41

b.d

b.d

b.d

b.d

I.l

I. 2

II. 1

III. l

I. 3

II. 2

III. 2

II. 3

III. 3

I, 4

111,4

II. 4

30

90

120

180

300

420

480

600

720

1320

1500

1680

' Centered over exit tube

^ 0.46 m from South Wall and 1.43 m from East Wall

Height for agent release test: halon 1301. FC-116, HCFC-22

Height for agent release test: halon 1301, HCFC-22, HFC-227, HFC-32 'HFC-125

1 . directly above the exit tube,

2. for a thermocouple 0.15 m above and 0.08 m nonh and another thermocouple 0.3 m above

and 0.15 m north of the release point, and

3. for the upper four thermocouples 0.5 m north of the release point.

To provide the most visually clear set of data for each test, we have used different temperature

ranges in the graphs. Care must therefore be exercised when comparing results between agents.

Figure 10 schematically shows the relationship between the thermocouples in the vicinity of the

exit tube, including all of the thermocouples displayed in Figures 3 to 9. For all tests, the largest

deviations from ambient occurred directly above the exit tube with the minimum temperature recorded

by the thermocouple closest to the exit tube. Based on the thermocouple data in Figures 3 to 9. the release pattern can clearly be discerned. In every "A" graph of Figures 3 to 9, the change in minimum temperature decreases as the distance from the exit tube increases. In the "C" graphs, the deviation increases as the height above the exit tube level increases, Funhermore, the data from the two off-axis thermocouples displayed in the graphs labeled "B" show no significant response to the
SECTION 9 HL"SL\N EXPOSITIE .AVD ENATROS"StENT\L IMPACT 771

(A) Time (,s) IB)

E -

Ceiling thermocouple

1000 Time (s) (C)1500

Figure 3. Temperature data FC-1I6: A-at 0.3. 0.6. and 0,9 m above tube; B-one 0.15 iii above. 0.08 ni offset and one 0.3 m above. 0.15 m offset; C-tbur 0.5 m offset. 0.0. 0.31. 0.62. and 0.93 m from ceiling.
772 SECTION 9. HUMAN EXPOSURE AND ENXTRONMENTAL 1MR\CT

1500

22 I 1 ^ 1—

0 500 1000 1500 Time (s) (C)

Figure 4. Temperature data Halon 1301, test 1: A-at 0.3, 0.6, and 0.9 m above

tube; B-one 0.15 m above, 0.08 m offset and one 0.3 m above, 0.15 m

offset; C-four 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 773

Figure 5. Temperature data Halon 1301, test 2: A-at 0.3, 0.6, and 0.9 m above

tube; B-one 0.15 m above, 0.08 m oflFset and one 0.3 m above, 0.15 m

offset; C-four 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
774 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

22 I 1 1 1—

0 500 1000 1500 Time (s) (C)

Figure 6. Temperature data HCFC-22, test 1: A-at 0.3, 0.6, and 0.9 m above

tube; B-one 0.15 m above, 0.08 m offset and one 0.3 m above, 0.15 m

offset; C-four 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 775

Figure 7. Temperature data HCFC-22, test 2: A-at 0.3, 0.6, and 0.9 m above

tube; B-one 0.15 m above, 0.08 m offset and one 0.3 m above, 0.15 m

offset; C-four 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
776 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

Figure 8. Temperature data HFC-227: A-at 0.3, 0.6, and 0.9 m above tube; B-one

0.15 m above, 0.08 m offset and one 0.3 m above, 0.15 m offset; Cfour 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

30

27

24

X 21

15

0 31 m above release point

J L. 500 1000 1500

Time (s) (A)

26

o

25

f- 23

22

5=F

Ceiling thermocouple

_i_ 500 1000 Time (s) (C)

1500

Figure 9. Temperature data HFC-32/125: A-at 0.3, 0.6, and 0.9 m above tube; B- one 0.15 m above, 0.08 m offset and one 0.3 m above, 0.15 m offset; C-four 0.5 m offset, 0.0, 0.31, 0.62, and 0.93 m from ceiling.
778 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

0.5

I

Thermocouple

Exit Tube

Figure 10. Schematic drawing showing the relationship of the near field thermocouples to the

exit tube in the north-south vertical plane of the agent sample flask.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 779

Figure 1 1 . Theoretical representation of agent plume in relationship to the thermocouples

located in the north-south vertical plane of the agent sample flask.
780 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Table 3. Summary of the extent of temperature depression of two thermocouple

locations

Compound Thermocouple @ 0.31 m above exit tube Thermocouple @ ceiling 0.5 m north

of exit tube

Maximum

Depression

rc)

Time to Recover to 90% of Ambient

(s)

Maximum

Depression

(°C)

Time to Recover to 90% of Ambient

(s) FC-116 15.8 170 3.3 130

Halon 1301 1 6.8 140 1.1 390

2 6.6 180 1.1 130

HCFC-22 1 4.6 190 1.0 170

2 4.4 135 1.0 185

HFC-227 4.0 180 0.8 80

HFC-32/HFC-125 7.4 185 1.4 90

sudden release of agent. This suggests that the release plume is very narrow and quickly reached the

compartment ceiling. The plume spreads rapidly along the ceiling and expands down towards the

floor of the compartment. This is shown idealized in Figure 11. Based on the thermocouple data,

this process occurs rapidly.

Table 3 summarizes the magnitude of the maximum depression from ambient temperature for the

thermocouple located 0.3 m above the exit tube and the time to return to within 90% of ambient

conditions. The same data are also tabulated for the ceiling thermocouple of the 0.5 m north thermocouple string. Maximum depression directly above the exit tube ranged from 15.8 °C for agent FC116 to 4.0 °C for agent HFC-227. The maximum depression of the ceiling thermocouple was 3.3 °C

for FC-116 to 0.8 °C for HFC-227. The 90% recovery time to near ambient conditions, with one

exception, ranges from 80 s to 190 s. The recovery time data displays a classically exponential

increase in temperature consistent with the natural convective heating of an object.

9.1.3.2 Concentration. In-situ FTIR and atmospheric samples captured in evacuated flasks represented the two methods employed in this study to determine agent concentration in the compartment. The FTIR apparatus is a line-of-sight measurement of the gas concentration of the released

agents. The evacuated flasks are point measurements made at specific time intervals and, because

three distinct sampling probes were employed, specific locations as previously described.

Figures 12 to 18 show the measured gas concentrations as a function of time for the three probe

locations. Also noted on each figure (dashed line) is the average agent concentration in the compartment assuming (a) complete release of the agent from the cylinder, (b) even distribution within the

compartment, and (c) no losses from the compartment.

While an efl'ort was made to ensure that the compartment was completely sealed, the rapid decay

in the early data (first half of the test period) indicates that significant leakage may have occurred.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 781

25

20 -

o

15

O

10

o

V

0.31 m

0.91 m

1.98 m

Maxiinum Average Concentration

300 600 900

Time (s)

1200 1500 1800

Figure 12. Agent concentration for agent FC-1 16 as a function of position and time with the maximum average concentration of agent shown (—) assuming uniform distribui; c within the compartment.
782 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

20

00

d

o•-I

cd -pd

a>od

o

o

1

15

O

10 - o

V

o

V

o

o

V

0.31 m

0.91 m

1.98 m

Maximum Average Concentration

300 600 900

Time (s)

1200 1500 1800

Figure 13. Agent concentration for halon 1301 test 1 as function of position and time with the

maximum average concentration of agent shown ( —) assuming uniform distribution

within the compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

20

15

10

1 1 1 1 1

'

-

V V

o V

o

° V

o 0.31 m

V 0.91 m

2.41 m

Maximum Average Concentration

1 1 1 1 300 600 900

Time (s)

1200 1500 1800

Figure 14. Agent concentration for halon 1301 test 2 as a function of position and time v/itb. ihe

maximum average concentration of agent shown (-—) assuming uniform distribution

within the compartment.
784 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

10

8

O

6

-

O V

V

o

O 0.31 m

V 0.91 m

1.98 m

Maximum Average Concentration

0

0 300 600 900

Time (s)

1200 1500 1800

Figure 15. Agent concentration for HCFC-22 test 1 as a function of position and time with the

maximum average concentration of agent shown ( —) assuming uniform distribution

within the compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 785

20

15 -

O 0.31 m

V 0.91 m

2.41 m

MeLximum Avereige Concentration

O

10

O V

o

07

300 600 900

Time (s)

1200 1500 1800

Figure 16. Agent concentration for HCFC-22 test 2 as a function of position and time with the

maximum average concentration of agent shown (—) assuming uniform distribution

within the compartment.
786 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

10

8

O 0.31 m

V 0.91 m

2.41 m

Maximmn Average Concentration

4 - 07 O V

o V

_L 300 600 900

Time (s)

1200 1500 1800

Figure 17. Agent concentration for agent HFC-227 as a function of position and time with the

maximum average concentration of agent shown ( —) assuming uniform distribution

within the compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT 787

10

2

-

O 0.31 m

V 0.91 m

2.41 m

Maxiinuin Average Concentration

a

o

V

_L _L 300 600 900

Time (s)

1200 1500 1800

Figure 18. Agent concentration for agent HFC-32/125 as a function of position and time with

the maximum average concentration of agent shown ( —) assuming uniform

distribution within the compartment.
788 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

However, once the turbulent action of the initial agent release had dissipated, further leakage would

only affect the final compartment concentration and not the ultimate distribution of agent within the

compartment.

Table 4 summarizes the average agent concentration for each release test, based on the average

of the last sampling time period for the three locations (this represents a time period of 6 minutes) as well as the theoretical average concentration. The coefficient of variation, computed as the ratio of

the standard deviation to the sample mean, for all tests was 3 to 19%. This was within the measure- ment variation of the GC analysis procedure. The data also shows that the final measured average

agent concentration was always less than the theoretical average value. This could be due to (a) not

all of the agent being released into the compartment (i.e., residual agent remaining in cylinder after the discharge) or (b) leakage of agent from the compartment.

In general, the data show that the initial concentration in the compartment was below the

expected average agent concentration assuming complete release of agent and that the concentration at a given level within the compartment decreased as a function of time. All the data indicate that an

average steady-state concentration was reached within the 30 minute test period. Table 5

compares the final average agent concentration within the compartment with the 300 s agent

concentration 0.31 m from the ceiling. The data show that the 300 s data was never more than

20.0% from the final compartment average agent concentration. In 4 out of 7 cases the 300 s data was less than 15% above the final average concentration. Within measurement errors, the data

appear to indicate that a uniform agent distribution was developed within a short period of time

(approximately 300 s). In the 30 min duration of the experiments, it is not likely that significant amounts of agent would

be left in the cylinders. Leakage from the compartment is likely, especially during the initial, high

velocity release of the agent. The tests, however, indicate that computing the theoretical average

concentration would represent a conservative estimate of occupant exposure conditions shortly after an

accidental release of agent.

A comparison of the data for agent HCFC-22, Figures 15 and 16, shows how quickly agent

concentrations can change at various levels in the test compartment. While the data for the two

release tests show a dramatic diff'erence for the first 600 s after agent release, the final measured

average agent concentration did not significantly differ. This was probably due to a combination of

convective mixing, diffusion, and compartment leakage. Tests on halon 1301 (Figures 13 and 14) were also repeated to determine the effects of probe

location on the measured agent concentrations. For the data in Figure 13, the lower probe was

1.98 m from the ceiling. For Figure 14, the lower probe was located 2.41 m from the ceiling (nearly

at the floor). Unlike HCFC-22 agent release tests, which showed a short term stratification of agent,

the halon 1301 tests showed no such stratification. There was no significant difference between the

data at all levels between these two tests. To further determine the time response of agent mixing and, therefore, concentration within the

compartment, an open path FTIR was installed to measure the compartment atmospheric concentration

of each agent at a specified level within the compartment. The results from the FTIR represent line- of-sight average values and are reported in Figures 19 to 22 as the integrated area of absorption peaks

in the spectral data for each agent. These data present a finer time resolution on agent concentration

than the evacuated flask data previously presented. Data for agent release tests involving halon 1301

and the second HCFC-22 test were lost because of instrumentation failures associated with the FTIR.

However, the data presented does indicate that a significant amount of mixing occurs very rapidly

after release of the agent. In all cases, an initial peak was detected, the magnitude of the peak
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 7g

Table 4. Final agent concentration for 30 minute release test

Compound

Theoretical

Average

Measured

Average^

Final Instantaneous Value

Specific Locations

(mg/1) (mg/1)

Sample 1,4 (mg/1)

Sample 11,4 (mg/1)

Sample 111,4 (mg/1)

FC-116 21.0 11.8 + 1.2' 11.9 10.5 12.9 Halon 1301 1 19.0 8.0 + 0.8 8.3 7.1 8.6

2 17.0 7.6 + 0.5 7.6 7.1 8.0 HCFC-22 1 8.0 4.0 ± 0.4 4.3 3.5 4.2

2 13.0 5.1 + 1.0 5.9 5.4 4.0

HFC-227 5.0 3.5 ± 0.1 3.6 3.5 3.5 HFC-32/HFC-125'' 5.0 3.2 ± 0.1 3.3 3.1 3.3

^ Average of last sampling flask at each location.

^ Weighted average values listed for mixture.

Standard deviation of three samples

Table 5. Comparison of final average agent concentration within the compartment to agent

concentration at sample probe located 0.31 m from ceiling at 300 s

Compound

Final Average

Concentration

(mg/1)

Probe 0.31 m

at 300 s (mg/1)Difference

(%)

FC-116 11.8 13.7 16.1

Halon 1301 1 8.0 7.9 -1.3

2 7.6 8.5 11.8

HCFC-22 1 4.0 4.8 20.0

2 5.1 5.2 2.0

HFC-227 3.5 4.0 14.3

HFC-32/HFC-125 3.2 3.8 18.7
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300

elapsed time (sec.)

Figure 19. Time dependent agent concentration for agent FC-116 as measured by

an open path FTIR located 0.3 m above and slightly off center of the

exit tube with measurement beam diagonally across compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT 791

elapsed time (sec.)

Figure 20. Time dependent agent concentration for agent HCFC-22 as measured by

an open path FTIR located 0.3 m above and slightly off center of the

exit tube with measurement beam diagonally across compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

I j

I

I

i 0 100 200 300 400 500 600 700 800 900 1000 1100 1200

elapsed time (sec)

Figure 2 1

.

Time dependent agent concentration for agent HCFC-227 as measured

by an open path FTIR located 0.3 m above and slightly oflF center of the

exit tube with measurement beam diagonally across compartment.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT 793

Figure 22. Time dependent agent concentration for agent FC-32/FC-125 as measured by an open path FUR located 0.3 m above and slightly off center of the exit tube with measurement beam diagonally across compartment.
794 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

depending upon the actual position of the FTIR sampling beam relative to the release point, that quickly dissipated, within 60 s. The data for agents HCFC-22, Figure 20, and HCFC-227, Figure

21, show a relatively steady agent concentration after the dissipation of the initial peak. Figure 19,

agent FC-116, shows a small initial peak, which may be due to a combination of scan average and

short duration of the initial release for this agent, followed by a steady but small decline in agent

concentration. These measurements were made with the FTIR beam located approximately 0.3 m

above the release point traversing the diagonal of the compartment slightly off center of the exit tube. For agent HFC-32/HFC-125, Figure 22, the FTIR was placed such that the sampling beam traveled

along the north wall approximately 0.35 m from the ceiling/wall junction. The data also show a steady but small decline in agent concentration after an initially small peak that lasted less than 60

seconds. Qualitatively, the FTIR data is in agreement with the evacuated flask data previously

presented. At the location of the FTIR line-of-sight measurement, immediately after agent release, the

agent concentration began to decay to a relatively steady state concentration. Since the FTIR

instrument was not calibrated for agent concentration, only relative concentration numbers are reported. Therefore, the FTIR data is not directly comparable to the evacuated flask data. Nevertheless, the relative decay rates calculated from the data in figures 19 and 22, 8.5 x 10 '^ and 1.0 x 10'-^, respectively, are within an order of magnitude of those determined by the evacuated flask measurements.

9.1.4 Modeling Results. In an efi'ort to generalize the results of the full scale agent release tests discussed in the previous section, various modeling techniques ranging from an analysis of agent

diffusion characteristics to performing 3-D fluid flow calculations using Harwell-Flow3d (CFD 1990) were employed to model the dispersal behavior of halon alternatives escaping from a pressurized

storage container into a closed compartment. Techniques for determining diffusion coefficients are

available in the literature (Bennett and Myers, 1974). These techniques were used to describe agent

diffusion characteristics as a function of temperature and molecular weight. The modeling technique

used to simulate agent dispersion in a compartment is to divide the enclosure into a collection of small

rectangular boxes or control volumes. Harwell-Flow3d is a member of a class of computer models

known as field models that perform this operation. The conditions in each control volume are

initially ambient. Agent is then released in several control volumes over time. The resulting flow or exchange of mass, momentum and energy between control volumes is determined so that these three

quantities are conserved. The momentum conservation equations are equivalent to Newton's second

law of motion and are referred to as the Navier-Stokes equations. The energy conservation equation

is equivalent to the first law of thermodynamics. These fluid flow equations are expressed mathematically as a set of simultaneous, non-linear partial differential equations. After being discretized, the

resulting finite volume equations are solved iteratively using a variant of Newton's method for computing coupled non-linear algebraic equations. Details of the fluid flow are realized by performing these calculations for each control volume throughout the compartment.

Two main factors afi'ect the dispersion of an agent release within a compartment, convection

caused by high gas velocities at the exit hole of the pressurized bottle and molecular diffusion.

These velocities were predicted by Harwell-Flow3d to decrease to a negligible size soon after the

bottle had finished evacuating. Figures 23 and 24 show shaded velocity contours in a vertical plane

of the bottle evacuation at the end of the bottle evacuation and 65 seconds later. Note how the

velocity contours in Figure 24 are small compared to the ones in Figure 23. This was also observed

experimentally. Therefore, the significant dispersion occurs during the initial agent release, with

diffusion controlling further dispersion.

Bennett and Myers (1974) outline a procedure for determining the difliision coeflficient for a binary mixture (air and the halon alternative in this case) given the temperature, and molecular
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 795

Speed, Plane IJ, Slice 2, File ha?, Time 25.0 s

2.50

2,25

2 00

1.75

1.50

1-25

1.00

75

.50

,25

Figure 23. Velocity magnitude shaded contours 25 seconds after discharge has begun in an ASTM standard room.
Figure 24. Velocity magnitude shaded contours 90 seconds after discharge has begun in an ASTM standard room.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 797

weights of the two gasses. Incorpera and De Witt (1990) list diflFusion coefficients for various gases

in air ranging from 0.62 x 10'^ m^/s for naphthalene, CioH^q, (molecular weight of 132) to 0.16 x

10"^ m-/s for CO2 (molecular weight of 46). Since the molecular weights of halon alternatives range

from 50 to 240 it is expected that the diffusivity of a halon alternative would be close to this range of

0.62 X 10"^ m-/s to 0.16 x 10""^ m-/s. A value of 7.2 x 10"^ was used in the field modeling studies

in this report. The distance x that a gas diffuses (assuming the gas is free to diffiise in three

dimensions) in r=600 seconds for difiiision coefficient within this range can be estimated using the

formula Ax = \/6DAt to be from 0.15 m to 0.24 m. This is a fraction of the room dimensions.

Thus, for a quiescent compartment, steady conditions should be reached quickly, with slow changes

thereafter. This is consistent with experimentally observed results. Two scenarios were modeled using Harwell-Flow3d. Both scenarios involved a storage bottle vented upwards in the center of a compartment with dimensions 2.4 m x 3.4 m x 2.4 m. The

compartment in both scenarios contained a small leak. In the first scenario, the room was free of

obstructions. In the second scenario, the compartment contained two solid baffles in order to partially

block the initial flow from the pressurized bottle. Figures 25 and 26 illustrate the results of the

simulation after the initial release from the bottle. For the compartment without blockages, the

dispersion is relatively uniform throughout. This is consistent with the results from the experimental

observations. For the room with blockages, the initial flow is restricted resulting in a more

concentrated region contained within the space between the storage bottle and the first baffle. This

shows that, for storage areas with exposed structural members along the ceiling or for high bay rack

storage areas, vertical obstructions can result in much higher agent concentrations than would

normally be anticipated based the total compartment volume. Care must therefore be exercised in the

placement and storage of large quantities of any agent.

9.1.5 Exposure Limits. The preceding sections indicate that dispersion of these chemicals

throughout an unobstructed volume is relatively rapid. Should an accidental discharge occur, the

threat to any trapped people would be determined by the toxicity of the fluid. Table 6 presents

recommended chronic and acute limits, as available, for the 12 agents. These provide working

guidelines (rather than definitive exposure limits) for safe exposures. For many of the agents, little data is available. For each agent, the CAS number provides unique identification of the chemical and

much of the material safety data sheet (MSDS) information is summarized in Table 6. In addition,

information on human exposure to the agents is included, to the extent such information is available. This information falls in basically four categories: acute exposure (eff'ects of short-term, usually

high-level exposure to the chemical), chronic exposure (long-term low-level exposure to the

chemical), contact exposure (eff'ects of exposure to the skin or eyes), and carcinogenicity.

For exposure during an accidental release, acute exposure efi'ects are most important. These

efl'ects are most often reported by an LC5Q value. This is the level which caused 50% of the test animals in a given experiment to die during and/or after the stated exposure. In all cases where the

toxicity information is known, the levels which cause lethal eff'ects are high (greater than 24 %).

However, information for most of the agents is not known. Thus, more information will be required

for those agents chosen for further smdy to place the results in appropriate context. Chronic exposure data are typically used to determine maximum allowable exposure for workers

exposed to chemicals over long periods of time. These limits are established by the Occupational

Safety and Health Administration (OSHA), American Conference of Governmental Industrial

Hygienists (ACGIH), or National Institute of Occupational Safety and Health (NIOSH). Chronic

exposure limits are typically reported by time-weighted averages of maximum allowable concentration

over an 8 hour workday or 40 hour work week.
Figure 25. Mass fraction shaded contours 600 seconds after discharge has begun (570 seconds

after pressurized bottle has finished evacuating) in an ASTM standard room.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 799

Figure 26. Mass fraction shaded contours 600 seconds after discharge has begun (570 seconds

after pressurized bottle has finished evacuating) in an ASTM standard room containing baffles that restrict flow.
800 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Table 6. Available Toxicity Information for Candidate Replacement Agents

Agent HFC-32 CAS # MSDS Exposure Toxicity Information

75-10-5 yes Acute No limits established by OSHA^ ACGIH^ or NIOSFf . Freons in high concentration 9/91 may cause pulmonary irritation, narcosis, dizziness, incoordination, confusion, nausea.

vomiting, tremors, and rarely, coma. When oxygen deficiency has been corrected, these

effects appear to reverse.

Chronic No limits established by OSHA\ ACGIH^ or NIOSH'.

Contact Skin No adverse effects reported. Due to rapid evaporation, frostbite with redness.

tingling, pain and numbness may occur. Skin may become hard and white and

develop blisters.

Eyes No adverse effects reported. Due to rapid evaporation, frostbite with redness, pain and blurred vision may occur.

Carcinogenicity No data available

a OSHA = Occupational Safety and Health Administration b ACGIH = American Conference of Governmental Industrial Hygienists.

c NIOSH = National Institute of Occupational Safety and Health

Agent HFC-32/HPC-125

CAS tt MSDS Exposure Toxicity Information

75-10-5 354-33-6

yes 9/91 4/89

See above for HFC-32

See below for HFC- 125

Agent HFC-227 CAS tt MSDS Exposure Toxicity Information

431-89-0 yes 9/90

Acute

Chronic

Contact

Carcinogenicity

No limits established by OSHA, ACGIH, or NIOSH. The toxicological properties of

this material have not been determined. Preliminary testing indicates the acute inhalation LC50 in rats is greater than 241,000 ppm, 4 hour LCjq is greater than 800,000.

Stimulants such as epinephrine may induce ventricular fibrillation. No limits established by OSHA, ACGIH, or NIOSH.

Skin No data available

Eyes No data available No data available
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

Table 6. (continued) Available Toxicity Information for Candidate Replacement Agents

Agent HCFC-22

CAS # MSDS Exposure Toxicity Information

75-45-6 ves Acute Relatively non-toxic by inhalation.. Rat LCjg: 350,000 ppm (15 min). Mouse LC50:

10/91 280,000 ppm (30 min). Dog LCjq: 700,000 ppm. Central nervous system depressant, simple asphyxiant. Stimulants such as epinephrine may induce ventricular fibrillation.

Chronic 1000 ppm (3540 mglm") = OSHA TWA'', ACGIH TWA, and NIOSH recommended TWA. 1250 ppm (4375 mg/m^) NIOSH recommended STEL^ 500 ppm (1770 mg/m^)

DFG MAK' TWA. 1000 ppm (3540 mg/m^) DFG MAK 60 min peak, momentary

value, 3 times/shift.

Contact Skin No adverse effects reported due to the gas. The liquid may cause frostbite; skin could turn red, tingle, and may become hard and white and develop

blisters. Pain and numbness may develop.

Eyes No adverse effects have been reported from the gas. Due to rapid evaporation,

the liquid may cause frostbite with redness, pain and blurred vision.

Carcinogenicity No information for humans. Male rats had increased incidences of fibrosarcomas and

zymbal-gland tumors. Not found in females.

TWA Time Weighted Average: Time-weighted average concentration for a normal 8 hour workday or 40 hour work week.

^ STEL Short Term Exposure Limit: 15 min time weighted average exposure which should not be exceeded at any time during a work day even if the 8 hour time weighted average is within established threshold limit value. Exposures at the STEL should

not be longer than 15 min and should not be repeated more than 4 times per day with at least 60 min between successive exposures at the STEL.

'^DPG MAK Maximum concentration values in the work place from the Federal Republic of Germany.
802 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

Table 6. (continued) Available Toxicity Information for Candidate Replacement Agents

Agent HFC- 134a

CAS # MSDS Exposure Toxicity Information

811-97-2 yes 11/91

Acute

Chronic

Contact

Carcinogenicity

Presently available data are indicative of a low order of toxicity. Nonflammable gas with low water solubility and low chemical reactivity.. LC5o(4 hr), rat, > 500,000 ppm.

Acute symptoms are indicative of central nervous system depression with anesthetic

effects induced at sub-lethal levels. The cardiac sensitizirig potential is very low (thresh- old - 75,000 ppm in dogs).

rats, 6 h/day, 5 days/week, 90 days, NOEL = 50,000 ppm.

Skin No data available

Eyes No data available

rats, one year study, oral doses of 300 mg/kg body weight (in corn oil). After 125

weeks, no carcinogenic potential found.

a NOEL No Observable Effect Level.

Agent FC-116

CAS ft MSDS Exposure Toxicity Information

76-16-4 yes 5/92

Acute

Chronic

Contact

Carcinogenicity

No limiting data available. Simple asphyxiant. Inhalation of high concentrations may

cause disorientation and narcosis.

No limits established by OSHA, ACGIH, or NIOSH.

Skin No adverse effects reported, due to rapid evaporation, liquid may cause

frostbite with redness, tingling and pain or numbness. Severe exposures: skin can become hard and white and develop blisters.

Eyes No adverse effects reported Due to rapid evaporation, liquid may cause

frostbite with redness, pain and blurred vision.

No data available
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 803

Table 6. (continued) Available Toxicity Information for Candidate Replacement Agents

Agent HCFC-i:24

CAS # USDS Exposure Toxicity Information

2837-89-0 yes 7/91

Acute

Chronic

Contact

Carcinogenicity

Very low acute toxicity. Moderate cardiac sensitizer for dogs exposed to 25,000 ppm

plus given epinephrine. In a two weeic subchronic inhalation study. 100,000 ppm

caused mild anesthesia in rats, but recovery occurred in 15 min. Guinea pigs exposed

for 30 min, one hour, and two hours at concentration ranging from 9000 to 207,000 ppm had no deaths. Anesthetic effects occurred at concentrations greater than 47,000 ppm, but the guinea pigs recovered following the exposure. High concentrations may

cause heart irregularities, unconsciousness or death. Vapors decrease oxygen availability. Frostbite could occur from liquid exposure. The effects in animals from a single exposure by inhalation include central nervous system effects, anesthesia and decreased blood pressure. Repeated exposures increased liver weights, and caused anesthetic

effects, irregular respiration, poor coordination, and nonspecific effects such as de- creased body weight. However, histopathological evaluation showed no irnsversible

effects. No limits established by OSHA, ACGIH, or NIOSH.

Skin No data available

Eyes No data available An impurity in FC-124, FC-133a, has been shown to be a potential carcinogen in one

published study.

Agent HFC- 125

CAS # MSDS Exposure Toxicity Information

354-33-6 yes Acute Not determined. Exposure may be harmful and could cause frostbite. Vapors can cause 4/89 headache, nausea, giddiness, unconsciousness.

Chronic No limits established by OSHA, ACGIH, or NIOSH.

Contact Skin No data available

Eyes No data available

Carcinogenicity No data available
804 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Table 6. (continued) Available Toxicity Information for Candidate Replacement Agents

Agent FC-218 CAS tt MSDS Exposure Toxicity Information

76-19-7 yes 10/91

Acute

Chronic

Contact

Carcinogenicity

No limits established by OSHA, ACGIH, or NIOSH. Asphyxia may occur with symptoms such as headache, dizziness, incoordination, dyspnea on mild exertion, sweating, malaise, tremors, convulsive movements, irregular breathing and death. No limits established by OSHA, ACGIH, or NIOSH.

Skin No data available

Eyes No data available No data available

Agent FC-31-10 CAS # MSDS Exposure Toxicity Information

355-25-9 yes 7/92

Acute

Chronic

Contact

Carcinogenicity

The toxicity of this material has not been determined. May be harmful if inhaled,

ingested or by skin absorption No limits established by OSHA, ACGIH, or NIOSH.

Skin No data available

Eyes No data available No data available

Agent FC-3I8 CAS # MSDS Exposure Toxicity Information

115-25-3 yes 6/92

Acute

Chronic

Contact

Carcinogenicity

No limits established by OSHA, ACGIH, or NIOSH. Simple asphyxiant. Stimulants such as epinephrine may induce ventricular fibrillation. No limits established by OSHA, ACGIH, or NIOSH.

Skin No adverse effects reported. The liquid could cause frostbite accompanied by

redness, tingling, pain or numbness. The skin may harden and turn white and

develop blisters.

Eyes No adverse effects from gas. The liquid could cause frostbite accompanied by

redness, pain and blurred vision. No data available
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT 805

Table 6. (continued) Available Toxicity Information for Candidate Replacement Agents

Agent CF3I

CAS # MSDS Exposure Toxicity Information

2314-97-8 yes Acute The toxicity of this material has not been determined. May have adverse effects if in- 8/90 haled, ingested or absorbed by the skin.

Chronic

-

No limits established by OSHA, ACGIH, or NIOSH.

Contact Skin Exposed skin may show signs of frostbite.

Eyes No data available

Carcinogenicit> No data available

Table 6 will be updated over the next two years as new information becomes available. A final version will appear in the September, 1995 final report.

9.2 Environmental Requirements for Candidate Replacements for Halon 1301

In addition to being eff'ective in extinguishing fires, a replacement for halon 1301 must not cause

greater damage to human health and the environment. Environmental issues include ozone depletion,

global warming, water and air pollution. Health and safety efi'ects include acute toxicity from an

accidental release (some presented in Section 9.1, above), long term carcinogenic efi'ects from

repeated exposures, and fires resulting from the flammability of the chemical.

The U.S. Environmental Protection Agency has established the Significant New Alternative

Policy Program (SNAP) under Section 612 of the Clean Air Act to evaluate alternative substances to ozone-depleting chemicals. The program includes a wide range of applications including fire and

explosion protection agents, which is the application of interest for this report. The SNAP Program

requires a manufacturer to provide specific information in order for EPA to determine the suitability

of the agent as a replacement for halon 1301. The information includes not only the environmental/safety parameters but also information pertaining to the amount of substance that would be

released and the practicality of using the replacement chemical in terms of availability, cost, and

amount required relative to halon 1301. In addition the relevant physical properties of the replacement compound must be included.

Here we consider each of the candidate replacements chemicals relative to the SNAP protocol.

We also consider halon 1301 as a reference point and N2 as a neutral chemical. This review focuses on major issues that might affect the application of a specific chemical rather than on a detailed

analysis for each chemical. Information on each of 14 chemicals is contained in a database entitled

"Environmental/Health Information on Candidate Replacements for halon 1301" and is also included

in hard copy form as Appendix A. The data base is available by request from the editors of this report. The key findings are included in Section 9.2.2 below. Some of the candidate chemicals have

already gone through a preliminary screening by EPA, and this information is identified as well. In the paragraphs below, each of the pertinent elements of the SNAP protocol is described. TIi;- remainder of the section discusses the major implications of the tabulated information.
806 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

9.2.1 SNAP Protocol. The information required under the SNAP Protocol (Federal Register, 1993)

regarding the environmental and health impact of each chemical is described below. The protocol

also requires detailed confidential business information and process-specific information, which is not discussed here. Information about each entry is provided below.

1. Name and Description of Substitute. The chemical formula and the standard nomenclature for refrigerants are given.

2. Physical and Chemical Information. This includes molecular weight, normal boiling

point, vapor pressure, and water solubility This information is pertinent both to the

performance of the replacement chemical and to its dispersion in the environment.

3. Application. How will the substance be used? In this case the application is for engine

nacelle and the dry bay protection. The engine nacelle is the region enclosing the engine.

The fire suppression agent would be directed into the nacelle and not into the engine

itself. Dry bays are compartments adjacent to fuel cells and can be located in either wing

or fuselage areas. They may contain electronic, hydraulic or mechanical equipment. In both of these applications it is important that the agent extinguish the fire rapidly.

4. Ozone Depletion Potential (ODP). Stratospheric ozone is critical to blocking ultraviolet

(UV) radiation from the sun. An increase in the UV solar radiation will have deleterious

eff'ects including increased incidence of skin cancer. The ODP value is a dimensionless

number with a value of 1 corresponding to the ODP of CFC-11, CHF2CI. The ideal

value is zero. The values listed in Appendix A are obtained from The Scientific Assess- ment of Stratospheric Ozone (Pyle et al., 1991).

5. Atmospheric Lifetime. The global warming potential (GWP) of a chemical is determined by its absorption of thermal radiation from the earth and by its tropospheric

lifetime relative to the value for carbon dioxide (IPCC, 1990). For the compounds

selected here, the atmospheric lifetime provides a good estimate of the GWP. An ideal

candidate would have a lifetime of a year or less; whereas, a molecule lasting for 1,000

years would be considered a problem.

6. Flammability. The substance should not be flammable at ambient oxygen concentration

at atmospheric pressure and temperature. This would create special needs in terms of the

storage and use of the chemical. Of course, one would not expect a fire suppressant

agent to be flammable.

7. Toxicity data. The toxicity data include acute exposure leading to death, the chronic

effects of a long term, low level exposure, cardiac sensitization such as an increase in the

pulse or irregular heartbeats, and the tendency to cause cancer or to have a mutagenic effect. Instead of determining the absolute health impact of a chemical, it is often easier to compare its impact with that of halon 1301.

8. Release. The expected amount of chemical released per event. The amount released for an engine or dry bay fire is on the order of a fraction of a kg for halon 1301.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 807

9. Replacement Ratio. This ratio is based on the volume of an agent (as a saturated liquid

@ 25 °C) relative to the volume of halon 1301 necessary to suppress various flame

arrangements. Our value corresponds to the average of spray burner and detonation tube

tests discussed in Section 4. This is thought to be the most appropriate number for dry

bays. A diff'erent estimate for the replacement ratio is obtained for engine nacelles and

the resulting values (see Section 4) are as much as 50% greater than the dry bay values.

10. Availability. The issue is whether the chemical is currently (as of September 30,

1993) available in the amount required for replacing halon 1301.

11. Cost. The cost is a variable, which fluctuates with the market. The tabulated values are our best estimates as of September 30, 1993 based on the purchase of 909 kg (2000 lbs)

of the agent.

12. Required Technological Changes. This refers to changes in the method for dispensing

the agent relative to halon 1301.

9.2.2 Discussion of Key Results in Database. A summary of the key issues for each agent is included in Table 7, with details in Appendix A. Important aspects of the information in the database

are discussed in the sections below.

9.2.2.1 Physical/Chemical Information. The database includes information for 16 difl'erent compounds. It is helpful to divide the compounds into various classes. Two chemicals stand out from the others: N2 and NaHC03. Nitrogen was chosen to provide test data for a gas with no

expected chemical efi'ect on a fire and no environmental impact since the earth's atmosphere is 80%

nitrogen. The NaHC03 dry powder releases CO2 when heated.

The other compounds involve carbon bonded to halogens (fluorine, chlorine, bromine, and

iodine), to hydrogen, and/or to other carbon atoms. The bond dissociation energy (BDE) decreases

monotonically with increasing atomic weight of the halogen for the carbon halogen bonds in CF3X

from a value of 546 kJ per mole for C-F (McMillan et al., 1982), to 331 for C-Cl, to 301 for C-Br,

to 223 for C-I. The last three BDE's were computed by the method described in Section 6.1.3.1.5

(Nyden, 1993). The C-H BDE in CF3H is 447 kJ/mole (McMillan et al., mi) and the C-Cl BDE

in CHF2CI is about 330 kJ/mole. It is convenient to classify fire extinguishing agents with regard to the bond most easily broken. We have the following classes in order from weakest to strongest:

C-I bond (CF3I)

C-Br bond (CF3Br)

C-Cl bond (HCFC-22, HCFC-124)

C-H bond (HFC-32, HFC- 125, HFC-32/HFC-125, HFC-227ea, HFC-236fa, HFC- 134a)

C-F bond (FC-116, FC-218, FC-3110, FC-318)

As discussed below, the smaller BDE correlates with a shorter atmospheric lifetime; for example, the

estimated lifetime of CF3I is about 2 weeks compared to an estimated 10,000 years for FC-116

(Wuebbles, 1993).

Another qualitative trend is the decreasing vapor pressure and increasing boiling point with

increasing molecular weight. For example the normal boiling point is -78 °C for FC-116 (C2F6)

compared to -2 °C for FC-3110 (C4F10). The vaporization characteristics are important for the two

phase dispersal as the fluid leaves the nozzle.
808

Table 7.

SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Summary of Issues Regarding Candidate Chemicals

I^UIIIIIICIILS

halon 1301 Not acceptable because of high ozone depletion potential

CF3I ODP not Icnown; toxicity and corrosivity not known

sodium bicarbonate Corrosive to aluminum; technology for rapidly dispersing the powder

must be developed

perfluorinated compounds, (FC116, FC-218, FC-318, FC3110)long atmospheric lifetime; high price for FC-318

HCFC-22, HCFC-124 Both on SNAP short list; however, finite ozone depletion potential and

cardiac sensitivity

HFC-32 Not acceptable because of flammability

HFC-125, HFC-227 Both on SNAP short list; slight cardiac sensitivity

HFC-32/HFC-125 High HF production from flame; cardiac sensitivity

HFC- 134a On SNAP short list; high HF production from flame, cardiac sensitivity

HFC-236fe Possible cardiac sensitivity; more difficult to disperse because of high

boiling point

Because of their high vapor pressures, these chemicals are not expected to persist as a liquid pool

or dissolved in water. The water solubility for most of the compounds is small. The most soluble is HCFC-22 (CHF22CI) with a water solubility of 3.3 kg/m^ followed by CF3Br with a value of 0.3

kg/m?. The available data indicate that the other compounds are less soluble. Even the moderate

solubility of HCFC-22 is not a water pollution issue for a release in an engine nacelle or dry bay.

9.2.2.2 Ozone Depletion Potential. As can be seen in the results tabulated in Appendix A, the ozone depletion potential of most of the compounds is zero. The exceptions are CF3Br, which is being proscribed because of its high ozone depletion potential of 16; CF3I, which is currently being

evaluated, and the two chlorine containing molecules, HCFC-22 and HCFC-124, with ozone depletion

potentials of 0.055 and 0.022, respectively. These values are considered small enough relative to what they replace that both of these chemicals are on the short list of proposed acceptable alternatives

under SNAP Program.

9.2.2.3 Atmospheric Lifetime. As discussed above, the tropospheric lifetime is used as a

surrogate for the global warming potential. Based on the heuristic bond energy analysis given above, one would expect CF3I to have the shortest atmospheric lifetime since it has the smallest bond energy. This expectation is in qualitative agreement with the statement in Section 5 that the absorption spectra

of CF3I is shifted to the red relative to the other CF3X's. Work in progress also indicates a very

short lifetime of about 2 weeks (Wuebbles, 1993). Since NaHC03 releases CO2 when heated, its
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 809

impact on the environment is determined by the CO2 life-cycle and is not considered to be an issue. Chlorine containing molecules are next in increasing atmospheric lifetimes followed by the hydrogen

containing molecules. The lifetime of the perfluorinated compounds is the longest, with Wuebbles

(1993) estimating 50,000 years for CF4. Because the five candidate perfluorinated compounds

considered here have C-C bonds, which are weaker than the C-F bond, the lifetime is expected to be

less than 50,000 years but may still be on the order of tens of thousands of years. This long lifetime

is a serious issue with regard to their suitability as candidate replacement chemicals.

9.2.2.4 Flammability. The chemical HFC-32 (CH2F2) is flammable at an ambient oxygen

concentration. In a mixture with HFC- 125, HFC-32 is not flammable. While the compounds are not flammable, one should note that the increased number of C-C and C-H bonds will increase the heat

release of these molecules in a flame environment.

9.2.2.5 Replacement Ratio. In all cases except NaHC03 more material will be required than

for the currently used suppression agent CF3Br; however, in the worst case only about twice as much

is required. While this is a narrow range, it could still be important because of the weight and

volume limitations on an aircraft. The replacement ratio computed here is on a volumetric basis. The

corresponding mass ratio, that is the mass of an agent relative to the mass of halon 1301 necessary to suppress various flame arrangements, can be determined from the tabulated results and the densities

given in Sections 2 and 4 of this report.

9.2.2.6 SNAP Alternative. There are five compounds from the above list that are on EPA's

short list as proposed acceptable alternatives under the SNAP Program for total flooding of unoccupied areas. These are HFC-125 (C2HF5), HFC-227ea (C3HF7), HCFC-22 (CHF2CI), HFC-134a

(C2H2F4), and HCFC-124 (C7HF4CI). Many of the other chemicals in the database we are considering are similar to these. The exceptions are the perfluorinated compounds which have a significantly

longer atmospheric lifetime than any of the other chemicals.

9.2.2.7 Required Technological Changes. For each agent there will likely be some changes

required relative to the current halon 1301 technology in terms of size of container, material for container, valve type, lubricant, etc. However, for one of these, NaHC03, a completely difl'erent dispersal system must be developed, since the NaHC03 is in the form of a powder. To be efl'ective,

the powder must be distributed essentially as fast as a gas.

9.2.2.8 Availability. The issue of availability is a critical one in the case of iodotrifluoromethane, CF3I, which, as of this writing, is made in amounts of a few kg per batch. One company

(DuBoisson, 1993) uses the so-called Hunsdiecker reaction (Banks, 1971). This involves the careful

pyrolysis of an anhydrous mixture of silver trifluoroacetate and iodine:

CF^CO^g + /2 - CF3/(>90%) + CO2 + ^gl (1)

The reaction is initiated by gently heating the reagents to about 100 °C. Once initiated the reaction is exothermic and the dissipation of the heat is one of the difficulties in scaling up this process above a few kilogram batch. Another difficulty is the recovery of the Agl from the vessel for reprocessing.

The current technology involves collecting the CF3I in cooled tubes connected to the reactor vessel

and breaking the reactor vessel each time to recover the Agl.
810 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

Another company (Newhouse, 1993) has just begun manufacturing CF3I using a 200 / reactor with a batch product of about 25 kg. They claim that there is not a major impediment to scaling up

the operation. However, until this is demonstrated the availability of CF3I will remain an issue.

9.2.2.9 Combustion Products. One of the major unwanted products of combustion is HE

Based on cup burner experiments. Appendix A shows that there is more than a factor of 10 variation

in the amount of HF released per gram fuel consumed with the two poorest performers,

HFC-32/HCF-125 and HFC134a. The HF production of the other compounds is a factor of two or more less than these compounds. For the CF3I there will also be the issue of the production of HI

and I2.

9.2.2.10 Price. There is a tremendous range in the price from as low as $3.50 per kg for HCFC-22 to $210 per kg for FC-318 and $330 per kg for CF3I. A major factor in the high price is the limited production of some of the compounds. The cost is likely to drop significantly with greater

production.

9.2.2.11 Toxicity. These chemicals are intended for use in areas external to the inhabited areas of the airplane. Therefore, people will not be exposed to the suppressant during fire suppression.

However, there is a chance of accidental exposure during the installation and maintenance of the

system. In such a case, safety considerations following rapid release of the two-phase flow include:

• low temperatures arising from the flow expansion, and - • asphyxiation from release into a confined space.

Chronic effects from repeated low level exposures should also be considered.

Our toxicity data are abstracted from the Air Force literature survey (Jepson, 1993). The

perfluorinated chemicals, which are apparently unreacted in the body, appear to be the least toxic of

the candidate chemicals and compare favorably with halon 1301 in acute testing. On the other hand,

the most toxic appear to be the chlorine-containing chemicals (HCFC-22 and HCFC-124) because of

their elevated levels of cardiac sensitization relative to the other chemicals. There is also a growing

popular concern about the toxicity of any chlorine-containing chemical. It is noteworthy that, even

so, these chemicals are on the short list for proposed alternatives under the SNAP Program. The

chemical HFC- 134a, C2H2F4, has about the same cardiac sensitization as halon 1301 and has been

shown to undergo a metabolism similar to anesthetic agent halothane. This is of some concern since

repeated low level exposures of personnel to the anesthetic agent halothane can produce viral-like

hepatitis. The perfluorinated halocarbons do not produce the viral-like behavior because they are biologically inert. Toxicity data for CF3I are just beginning to be generated at this time.

Little is known about the mutagenic/carcinogenic potential of the agents for humans. FC-318,

FC-218, and HCFC-22 have all demonstrated some mutagenic activity in an in vitro system. Only

HCFC-22 has been tested in a mammalian system; based on the results it is classified as Group 3 (the

agent is not classifiable as to its carcinogenicity in humans).
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT y

9.3 Conclusions

• Experimental measurement, hand calculations, and 3-D fluid flow modeling of an accidental

release of a halon alternative show that a reasonably uniform concentration of agent is obtained in a compartment shortly after release of the agent. Thus, simple ideal gas calculations of agent concentrations should suffice for estimating agent concentration following an

accidental release.

• For compartments which include baffles to the gas flow, the diffusion controlled flow after agent release may confine the released agent within the baffles near the release point of the

agent.

• As of this report, only five candidate agents have been designated as acceptable alternative

chemicals to halon 1301. These are: HCFC-124, HCFC-22, HFC-134A, HFC-227, and

HFC-125.

9.4 References

Banks, R.E., Fluorocarbons and their Derivatives, Elsevier Publishing Co., Inc., New York,

1971.

Barrow, G.M., Physical Chemistry, McGraw-Hill, Inc., New York, 1961.

Bennett, CO. and Myers, J.E., Momentum, Heat, and Mass Transfer, McGraw-Hill, Inc., New

York, second edition, 1974.

CFD Department, AEA Industrial Technology, Harwell Laboratory, Oxfordshire, United

Kingdom. HARWELL-FL0W3D Release 2.3: User Manual, 1990.

DuBoisson, R., PGR, Gainesville, Florida, personal communication, 1993.

Federal Register, Volume 58, 90, pp. 28102-28104, May 12, 1993.

Incorpera, FP. and De Witt, D. P., Fundamentals of Heat and Mass Transfer, John Wiley and

Sons, New York, third edition, 1990.

Intergovernmental Panel on Climate Change (IPCC): Climate Change; The IPCC Scientific

Assessment, Cambridge University Press, Cambridge, UK, 1990.

Jepson, G.W., Memorandum for Capt. Jepson, WL/FIVS (Mr. Bennett), May 21, 1993.

McMillan, D.F., and Golden, D.M., "Hydrocarbon Bond Dissociation Energies," Annual Review

of Physical Chemistry 33 ,493 (1982)

Newhouse, S., Pacific Scientific, Duarte, CA, personal communication, 1993.

Nyden, M., National Institute of Standards and Technology, personal communications, 1993.

Pyle, J. A., Solomon, S., Wuebbles, D., and Zvenigorodsky, S., "Ozone Depletion and Chlorine

Loading Potentials," in World Meteorological Organization, Global Ozone Research Monitoring

Project - Report No. 25, Scientific Assessment of Ozone Depletion: 1991, 1991.

Wuebbles, D.J., Lawrence Livermore National Laboratory, personal communication, 1993.

SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

Appendix A. Halon Replacement Properties

HALON Replacement Chemical Properties Tabl

Date Recorded: 09/29/93

Formula: CF3Br Designation: Halon 13 01

Molecular Weight: 148.9 kg/kmole

Normal Boiling: 215 K

Vapor Pressure: 1610 kPa @ 2 98 K

Water Solubility: 0.3000 kg/m3 @ 101 kPa

Ozone Depletion

Potential : 16

Atmospheric Life

Time (yrs) : 110

Flammability : NO

Replacement Ratio: 1.0 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability: UNLIKELY AFTER 1995

Price: $61.60 per kg @ 13 kg

Required Technological

Changes : NONE

Environmental- -Release : ABOUT 4 KG PER INCIDENT

Combustion Products: 0.08 G HF/G C3H8 BURNED

Toxicity Acute : Cardiac

:

Carcinogenic

Mutagenic

Pyrolysis
814 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: CF3I Designation: lodotrifluoromethane

Molecular Weight: 196.0 kg/kmole

Normal Boiling: 251 K

Vapor Pressure:

Water Solubility:

Ozone Depletion

Potential

:

• Atmospheric Life

Time (yrs) : 0 . 04

Flammability

:

Replacement Ratio: 1.2 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability: CURRENT CAPACITY ABOUT 2 5 KG PER BATCH

Price: $3 3 0.00 per kg @ 2 5 kg

Required Technological

Changes

:

Environmental --Release: ABOUT 4 KG PER INCIDENT

Combustion Products

:

Toxicity Acute:

Cardiac

Carcinogenic

Mutagenic

Pyrolysis:
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: NaHC03 Designation: Sodium Bicarbonate

Molecular Weight : 84.0 kg/kmole

Normal Boiling:

Vapor Pressure:

Water Solubility: 69.000 kg/m3 @ 101 kPa

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs)

:

Flammability : NO

Replacement Ratio: 0.8 based on HALON 13 01

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability: YES

Price: $2.00 per kg @, 1000 kg

Required Technological

Changes: MAJOR CHANGES REQUIRED TO RAPIDLY DISPERSE A POWDER

Environmental- -Release : ABOUT 4 KG PER INCIDENT

Combustion Products : NaOH

Toxicity Acute: NOT A CONCERN

Cardiac : NOT A CONCERN

Carcinogenic: NOT A CONCERN ; Mutagenic : NOT A CONCERN

Pyrolysis: NOT A CONCERN
816 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Designation: FC-116

138.0 kg/kmole

195 K

3070 kPa @ 293 K

Formula: C2F6

Molecular Weight

Normal Boiling

Vapor Pressure

Water Solubility

Ozone Depletion

Potential

Atmospheric Life

Time (yrs)

Flammability

Replacement Ratio

SNAP Alternative

Application

Availability

Price

Required Technological

Changes

:

0 . 0 10000

NO

3.2 based on HALON 1301

ENGINE NACELLE /DRY BAY

$37.30 per kg @ 450 kg

Environmental - -Release : ABOUT 4 KG PER INCIDENT

Combustion Products

Toxicity Acute

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

0.7 G HF/G C3H8 BURNED

RAT, 1 hr @ 800,000 ppm - initial hyperactivity

followed by hypoactivity , hyperemia, and closed

eyes

.

dog, up to 600,000 ppm - no sensitization

not mutagenic in exposed E. coli.

rat, exposure of pyrolysis products from acetone

fire resulted in dyspnea, salivation, lacrimation,

and gasping. Substantial weight loss.
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C3F8 Designation: FC-218

Molecular Weight: 188.0 kg/kmole

Normal Boiling: 236 K

Vapor Pressure: 880 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs) : 10000

Flammability : NO

Replacement Ratio: 1.6 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability

Price: $39.60 per kg @ 654 kg

Required Technological

Changes

Environmental - -Release

Combustion Products

Toxicity Acute :

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

ABOUT 4 KG PER INCIDENT

1 . 1 G HF/G C3H8 BURNED

rat, 1 hr @ 800,000 ppm - initial hyperactivity

followed by hypoactivity , hyperemia, and closed

eyes

.

Biochemical mutants were produced in E. coli
818 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C4F10 Designation: PC- 31- 10

Molecular Weight: 238.0 kg/kmole

Normal Boiling: 271 K

Vapor Pressure: 265 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs) : 10000

Flammability : NO

Replacement Ratio: 1.7 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability

:

Price: $33.00 per kg @ 600 kg

Required Technological

Changes

:

Environmental- -Release : ABOUT 4 KG PER INCIDENT

Combustion Products

Toxicity Acute

Cardiac

Carcinogenic

:

Mutagenic

Pyrolysis

1.2 KG HF/G C3H8 BURNED

rat, 16 hrs @ 800,000 ppm no effect

dog, no cardiac sensitization noted at up to

4 0 0,000 ppm
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 819

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Designation

:

200.0 kg/kmole

266 K

310 kPa @ 298

Formula: CYCLO C4F8

Molecular Weight

Normal Boiling

Vapor Pressure

Water Solubility

Ozone Depletion

Potential

Atmospheric Life

Time (yrs)

Flammability

Replacement Ratio

SNAP Alternative

Application

Availability

Price

Required Technological

Changes

FC-318

K

0 . 0 10000

NO

1 . 7 based on HALON 13 01

ENGINE NACELLE/DRY BAY

LIMITED PRODUCTION

$209.00 per kg @ 45 LBS

Environmental - -Release

Combustion Products

Toxicity Acute

Cardiac

:

Carcinogenic

Mutagenic

Pyrolysis

ABOUT 4 KG PER INCIDENT

rat, 4 hr @ 800,000 ppm, no deaths, signs of

irritation

mouse, up to 400,000 ppm alone - no arrhythmias;

200,000 ppm + epinephrine - arrhythmias

5/1300 visible mutation (0.38%) compared to 0.008%

spontaneous control rate; sex- linked recessive

lethal mutation was not significant.
820 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: CH2F2 Designation: HFC-32

Molecular Weight: 52.0 kg/kmole

Normal Boiling: 221 K

Vapor Pressure

Water Solubility

Ozone Depletion

Potential

Atmospheric Life

Time (yrs)

Flammability

Replacement Ratio

SNAP Alternative

Application

Availability

Price

Required Technological

Changes

0 . 0

7YES

ENGINE NACELLE/DRY BAY

Environmental - -Release

Combustion Products

Toxicity Acute

Cardiac

:

Carcinogenic

Mutagenic

Pyrolysis

ABOUT 4 KG PER INCIDENT

rat, 4 hr @ 111,000 to 760,000 ppm - signs of

lethargy, loss of mobility, spasms and gnawing of

cage during exposure. No abnormal signs 30 min to

14 days after exposure

.

1/12 dogs sensitized @ 250,000 ppm; same dog was normal after 200,000 ppm.

Ames assay - negative

no information
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMR\CT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: CH2F2/C2HF5 Designation: HFC-32/HFC-125

Molecular Weight: kg/kmole

Normal Boiling: 220 K

Vapor Pressure: 1670 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs)

Flammability : NO

Replacement Ratio: 2.5 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability

Price: $33.00 per kg @ 680 kg

Required Technological

Changes

Environmental- -Release

Combustion Products

Toxicity Acute

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

ABOUT 4 KG PER INCIDENT

6 . 2 G HF/G C3H8 BURNED
822 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C2HF5 Designation: HFC-125

Molecular Weight: 120.0 kg/kmole

Normal Boiling: 224 K

Vapor Pressure: 1380 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs) : 40.5

Flammability : NO

Replacement Ratio: 1.8 based on HALON 13 01

SNAP Alternative: + Application: ENGINE NACELLE/DRY BAY

Availability

:

Price: $31.46 per kg @ 490 kg

Required Technological

Changes

:

Environmental - -Release : ABOUT 4 KG PER INCIDENT

Combustion Products: 2 . 1 KG HF/G C3H8 BURNED

Toxicity Acute : Cardiac

:

Carcinogenic

:

Mutagenic: Ames assay - negative, human lymphocyte - negative,

CHO cell assay - negative

Pyrolysis

:
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 823

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C3HF7 Designation: HFC-227

Molecular Weight: 170.0 kg/kmole

Normal Boiling: 257 K

Vapor Pressure:

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs) : 13

Flammability : NO

Replacement Ratio: 1.6 based on HALON 1301

SNAP Alternative: + Application: ENGINE NACELLE/DRY BAY

Availability

:

Price: $49.50 per kg @ 654 kg

Required Technological

Changes

Environmental - -Release

Combustion Products

Toxicity Acute

Cardiac

;

Carcinogenic

Mutagenic

Pyrolysis

:

ABOUT 4 KG PER INCIDENT

1 . 7 G HF/G C3H8 BURNED

rats, 4 hrs @25,000 or 53,000 ppm - no deaths,

irreg. breathing, slight lacrimation, red ears, no post exposure effects.
824 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C3H2F6 Designation: HFC-236FA

Molecular Weight: 152.0 kg/kmole

Normal Boiling: 272 K

Vapor Pressure: 270 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs)

:

Flammability : NO

Replacement Ratio: 1.6 based on HALON 1301

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability

:

Price

Required Technological

Changes

Environmental- -Release : ABOUT 4 KG PER INCIDENT

Combustion Products: 1.9 G HF/G C3H8 BURNED

Toxicity Acute

:

Cardiac

Carcinogenic

Mutagenic

:

Pyrolysis
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 825

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C2H2F4 Designation: HFC-134A

Molecular Weight: 102.0 kg/kmole

Normal Boiling: 247 K

Vapor Pressure: 670 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.0

Atmospheric Life

Time (yrs) : 15.6

Flammability : NO

Replacement Ratio: 1.9 based on HALON 13 01

SNAP Alternative : + Application: ENGINE NACELLE/DRY BAY

Availability

Price: $11.55 per kg @ 680 kg

Required Technological

Changes

Environmental- -Release

Combustion Products

Toxicity Acute

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

ABOUT 4 KG PER INCIDENT

3 . 9 G HF/G C3H8 BURNED

rat, 4 hr LC50 > 500,000

Threshold level - 75,000

ppm

ppm

Ames assay - negative, clastogenicity - negative;

rat cytogenetics up to 50,000 ppm negative, mouse

dominant lethal up to 50,000 ppm - negative.
826 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Designation: HCFC-22

87.0 kg/kmole

232 K

1050 kPa @ 298 K

3.3000 kg/m3 @ 101 kPa

Formula: CHF2CL

Molecular Weight

Normal Boiling

Vapor Pressure

Water Solubility

Ozone Depletion

Potential

Atmospheric Life

Time (yrs)

Flammability

Replacement Ratio

SNAP Alternative

Application

Availability

Price

Required Technological

Changes

:

Environmental - -Release

Combustion Products:

Toxicity Acute :

Cardiac

:

Carcinogenic

:

Mutagenic

Pyrolysis

0 . 055

15 . 8 NO

2.2 based on HALON 1301

+ ENGINE NACELLE /DRY BAY

$3.50 per kg @ 680 kg

ABOUT 4 KG PER INCIDENT

1.3 G HF/G C3H8 BURNED

rats, guinea pigs; 2 hr exposure - 75,00 - 100,000

ppm -excitation and/or ataxia; 200,000 ppm - CNS

depression; 300,000 - 400,000 ppm - death,

dogs and monkeys; 50,000 - 100,000 ppm caused early

respiratory depression, bronchoconstriction, tachycardia, myocardial depression and hypotension.

Inadequate evidence of carcinogenicity in humans;

some evidence of tumors in male mice exposed for a long period of time.

Ames assay, BHK21 cell transformation - positive

results; no induction of unscheduled DNA synthesis

in human heteroploid EUE cells treated with a 2 0 mM

soln. generated at 500 ml/min (1:1 air) in presence

or absence of SIO

.

Combustion products highly toxic
SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT 827

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: C2HF4C1 Designation: HCFC124

Molecular Weight: 137.0 kg/kmole

Normal Boiling: 26 0 K

Vapor Pressure: 380 kPa @ 298 K

Water Solubility:

Ozone Depletion

Potential: 0.022

Atmospheric Life

Time (yrs) : 6.9

Flammability : NO

Replacement Ratio: 1.6 based on HALON 1301

SNAP Alternative: + Application: ENGINE NACELLE/DRY BAY

Availability

:

Price: $17.05 per kg @ 790 kg

Required Technological

Changes

:

Environmental- -Release

Combustion Products

Toxicity Acute

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

:

ABOUT 4 KG PER INCIDENT

0 . 6 G HF/G C3H8 BURNED

Approx. lethal cone. - 230,000- 255,000 ppm.

dog; threshold - 26,155 ppm.

Ames assay - negative; human lymphocyte assay - negative; CHO cell assay - negative; Micronucleus

assay - negative.
828 SECTION 9. HUMAN EXPOSURE AND ENVIRONMENTAL IMPACT

HALON Replacement Chemical Properties Table

Date Recorded: 10/01/93

Formula: N2 Designation: NITROGEN

Molecular Weight: 28.0 kg/kmole

Normal Boiling: 77 K

Vapor Pressure:

Water Solubility:

Ozone Depletion

Potential

:

Atmospheric Life

Time (yrs)

Flammability

Replacement Ratio:

SNAP Alternative:

Application: ENGINE NACELLE/DRY BAY

Availability

:

Price: $2.00 per kg @ 450 kg

Required Technological

Changes

:

Environmental - -Release

:

Combustion Products;

Toxicity Acute

;

Cardiac

Carcinogenic

Mutagenic

Pyrolysis

:

ABOUT 4 KG PER INCIDENT
829

10. SUMMARY AND RECOMMENDATIONS

William L. Grosshandler, Richard G. Gann and William M. Pitts Building and Fire Research Laboratory

The main objective of this research program was to provide guidance to the sponsors on which

materials to evaluate in the full-scale fire suppression test plan at Wright-I^tterson AFB. Specifically,

the recommendations were to include the following:

1. the best two agents from the core list of twelve FCs, HFCs and HCFCs provided by the

Air Force when NIST began the project in 1992,

2. the best practical alternative from among materials not on the list of twelve, and

3. whether or not sodium bicarbonate powder was a viable option.

New metrics have been developed along the way which can be used to rate other fire fighting compounds for a variety of applications.

The methodology for combining the results of literally hundreds of tests for the purpose of

defining an overall performance rating of the chemicals is described in this section. A justification

for each agent recommended using the methodology and the ranking of all of the agents evaluated by

these methods are included. The report ends with a summary of the continuing investigation into the

formation of by-products during the suppression process and long-term materials compatibility for the

agents selected by the sponsors for full-scale testing. A list of recommendations for further research

is provided.

10.1 Rationale for Selection

The agents have been rated using the criteria given below. The desirable magnitude of each variable

is given in parentheses.

• saturation liquid density at 25 °C (high)

• temperature at which storage vessel reaches liquid-filled condition (high)

• boiling point at 101 kPa (low)

• discharge rate from vessel (high)

• dispersion and mixing with air (high)

• distance required for agent vaporization during sudden discharge (short)

• concentration required to extinguish laminar flame in opposed-flow diffiision (OFDF) burner

for two air temperatures, multiple fuels, and variable strain rates (low)

• concentration required to suppress cup burner flame for multiple fuels (low)

• concentration required to extinguish turbulent spray flame for two temperatures and fuels, and

variable injection rates (low)

• concentration required to suppress a high-speed turbulent flame in detonation tube for two

equivalence ratios (low)
830 SECTION 10. SUMMARY AND RECOMMENDATIONS

agent residue (low)

corrosion of metals by neat agent for multiple materials (low)

swelling of polymers by neat agent for multiple materials (low)

durability of elastomers and greases by neat agent for multiple materials (high)

decomposition during storage (low)

concentration of corrosive byproducts during suppression (low) ozone depletion potential (low)

atmospheric lifetime (short)

relative potential human exposure (low)

availability (high)

price of chemical (low)

Many of the above criteria did not lead to differentiation among the alternative agents to a

statistically significant degree. The discriminating factors could be lumped into four categories:

agent dispersion and evaporation characteristics, required storage volume, environmental factors, and

operational issues.

10.1.1 Dispersion and Evaporation. The rate of agent discharge from the storage vessel was

discussed in Section 3.4 for different initial conditions. The average volume flow rate, V^, of the

liquid agent from a 2/3-filled vessel initially at room temperature and 4.1 MPa is summarized here in

Table 1. All of the alternative agents leave the vessel at about the same volume flow rate, 10.5 ±

0.6 l/s, which is significantly faster than the discharge rate of halon 1301. Once leaving the bottle, more of an agent is required aboard an aircraft if it disperses and evaporates slowly to ensure that a

fire anywhere within the protected space can be extinguished. Measurements taken during a discharge

under conditions similar to a large open bay were discussed in Section 3.5. From those measurements, the dispersion, mixing and evaporation behaviors of the alternative agents were evaluated.

The rankings from that study are also listed in Table 1. Unlike the dry bay discharge which is limited to the heat contained in the stored liquid for vaporization of the agent, more opportunity is available for evaporating the liquid droplets in engine

nacelle applications. Heat transferred from the piping system and air can provide the additional

energy for the vaporization process as long as the surrounding temperature is sufficiently above the

agent boiling point. The minimum temperature required to ensure that enough agent is in a fully

gaseous state, T^-^, can be estimated from an energy balance on the protected volume assuming the mass fraction of agent, Yj, is the amount required to suppress a JP-8 cup burner flame. It can be

shown that

^ _ l^A^VW_ (1)

n-Y)C +YC . I ^ p,air I p,i

where the specific heats are for the liquid agent, and air, and ^y^p enthalpy of

vaporization at the boiling point, 7),. Values of 7"^,^ are listed in Table 1 for each of the agents.

HFC-134a, FC-318, FC-31-10, HCFC-124 and HFC-236fa require temperatures greater than 25 °C

to vaporize fully, making them suspect for low temperature operation. The FC-116 is stored above

its critical point, but if one were interested primarily in low temperature operation, its 7"^,^ value

equal to -43 °C would make FC-116 a strong candidate. FC-218, HFC-125 and CF3I are the only

other compounds with T^^^ less than 0 °C.
SECTION 10. SUMMARY AND RECOMMENDATIONS

Table 1 . Agent discharge, dispersion and evaporation parameters

Agent Sect. 3.5 Rank Tmm' °C^ V',, 1/s CFjBr n.r. -42 7.6 FC-116 n.r. -43 a FC-218 1 -7 10.8 HFC- 125 1 -6 9.8 HFC-32/HFC-125 3 +20 10.0 HCFC-22 4 + 21 10.2 HFC-134a 5 + 29 10.7 HFC-227ea 6 + 19 11.1 HCFC-124 7 + 28 10.9

FC-318 8 + 28 10.8

FC-31-10 9 + 29 10.3

HFC-236fe 10 +37 10.8

CF3I n.r. -1 a

a data not taken n.r. not ranked

The Jakob number, Ja, was defined in Section 3 as the amount of heat available from cooling the

liquid agent from ambient temperature to its normal boiling point, divided by the agent's heat of

vaporization. A high value for Ja and a low value for r^,„ are both desirable attributes which often,

but not always, go hand-in-hand. For example, the Jakob numbers for FC-218 and CF3Br are 0.59

and 0.51, respectively, while T^j^ for CF3Br is 35 °C less than the minimum temperature for FC218. The implication is that the perfluorocarbon will flash more completely from a storage vessel at room temperature than halon 1301 in a dry bay application, but FC-218 will vaporize less efficiently

in the engine nacelle at lower operating temperatures.

10.1.2 Storage Volume. The required storage volume is estimated from the measured suppression

concentration and the density of the agent within the storage vessel at ambient temperature. The fuel type and temperature of the air were found to be of secondary importance in the measurements of the

relative mass of the agents. While the absolute amount of agent necessary to quench the flames

varied with the apparatus, the strain rate, the rate of injection, and the equivalence ratio, a single

parameter called the volume factor, VF, is useful in comparing the relative suppression performance

of the various agents. VF is a measure of the liquid agent volume in the storage vessel normalized by

the liquid volume of halon 1301 required to suppress a flame under identical conditions; i.e..
832 SECTION 10. SUMMARY AND RECOMMENDATIONS

Y 1-Y,1301 VF = ( )( ) (2)

1301

The saturation liquid densities at 25 °C, pj, were taken from Section 2 of this report. Table 2 lists the values of VF for each apparatus. The OFDF values were found from the ambient temperature

experiments with JP-8, taken at a strain rate of 100 s ^ The cup burner values were measured with

heptane as the fuel. For the spray burner, the VF was calculated from the mass of agent injected in the ambient temperature JP-8 experiments. The performance in the detonation tube was based upon

the average of the mass of agent required to reduce the pressure ratio by one half and to 10% of the

maximum increase, with data restricted to the lean experiments. The uncertainty in the values of VF

is estimated to be + 10%.

Independent of the apparatus, the volume factors for FC-116 and the HFC-32/125 mixture are the highest {i.e., the poorest performers). The best performer in the table is sodium bicarbonate.

(Note that the flame suppression number, FSN, is used to rate the powder rather than VF. The direct

comparison is valid as long as the powder density is close to the density of liquid halon.) lodotrifluoromethane is the most eff'ective of the gaseous agents in suppressing the cup and spray burner

flames, and HCFC-124 is the best of the core alternatives. FC-218 stands out as the best agent in suppressing the high-speed deflagration. Except for the high value of VF for HCFC-22 in the

deflagration tube, there is little basis upon which to say that any one of the remaining agents is better than another in suppressing the flames using this metric.

10.1.3 Environmental Factors. Primary environmental concerns include acute toxicity, ozone

depletion potential, and atmospheric lifetime of the alternative agents. None of the core chemicals

have been identified as presenting an acute toxicity hazard, so that consideration of application to normally unoccupied spaces like the engine nacelle and dry bay does not lead to any discrimination

among them. Toxicity data on CF3I is incomplete, but preliminary studies (as of this writing) have

not uncovered any major health hazards which would prevent its application to unoccupied spaces. All but two of the core agents have an ODP of zero. The two chlorinated compounds, HCFC-22

and HCFC-124, have an ODP less than one tenth of the present EPA limit. Iodides are not currently

regulated, but preliminary estimates indicate that the ODP for CF3I released in the lower troposphere

is small because it easily photodissociates. Any of the chemical released directly into the strato- sphere, however, would react eflSciently with the local concentration of ozone. The general ranking

(from best to worst) based upon ODP is as follows:

Short atmospheric lifetime is considered a positive attribute in an alternative agent, but as of this writing it is not mandated by governmental regulations nor international agreements. The estimated

tropospheric lifetime varies by over four orders-of-magnitude, with CF3I having the shortest (weeks)

and the perfluorocarbons having the longest (1000s of years). Among the remaining compounds, the

HCFCs have a shorter lifetime than the HFCs, which are less long-lasting than CF3Br (estimated

lifetime is 66 years). The following ranking applies, from best to worst:

FCs = HFCs = 0 < CF3I < HCFCs < < CFjBr (ODP)

CF3I < < HCFCs < HFCs < < FCs (atmospheric lifetime)

Because sodium bicarbonate is a powder, atmospheric chemistry is a non-issue. The powder is also considered to be non-toxic.
SECTION 10. SUMMARY AND RECOMMENDATIONS

Table 2. Volume factors for flame suppression in diff'erent apparatus

Agent DFDP Cup Burner Spray Burner Deflag. Tube

U.zj 0.45*' 0.85'' a

a 0.9 0.8 1.4

Z.J 2.2 1.8 1.5

J.U 2.8 1.8 2.5

rirL.-ZjOia, "3 1 J. 1 2.3 1.8 1.4

T TT/^ ITT HFC-227 3.2 2,4 1.8 1.4 FC-218 3.6 2.8 2.1 1.2

FC-31-10 3.9 2.6 2.1 1.4 HFC-134a 4.0 2.8 2.0 1.8 HFC- 125 4.1 2.8 2.0 1.7

FC-318 4.1 2.6 2.0 1.4

HFC-32/125 5.3 3.5 2.5 2.5

FC-116 7.1 6.0 4.0 2.5

^ data not taken FSN for powder

10.1.4 Operational Issues. Factors which involve maintenance or logistics can be evaluated

separately from fire suppression performance and environmental impact. These include agent residue,

stability under storage, accidental discharge under storage, corrosion from combustion byproducts,

compatibility of the agent with metals and seals, and availability of the chemical. Table 3 compares

the agents on operational issues. A rating of "B" in the table implies the compound has no significant

shortcomings nor advantages over the other agents in that particular performance criteria. A "C"

indicates a weakness or below average performance, but not necessarily a fatal flaw. An "A" rating means above average performance or a particular operational advantage demonstrated by the compound.

None of the core agents left an undesirable residue when discharged, none segregated in space to a significant degree during a simulated leak, and little decomposition was observed during one-month,

high temperature storage. All produced a considerable amount of undesirable HF during suppression.

There was some variation in the interaction of the core compounds with metals and polymers, but

storage material combinations do exist which are compatible with most of the agents. Industrial sources exist for the core agents, so that neither lack of availability nor price of the agent were issues

of major concern. lodotrifluoromethane from the initial batch received was observed to be slightly more aggressive

than the core agents in attacking some metals and polymers, but the limited amount of short-term

testing which was completed did not uncover insurmountable materials compatibility issues. Iodine
834 SECTION 10. SUMMARY AND RECOMMENDATIONS

Table 3. Rating of agents on various operational issues: A, above average performance or

particular advantage; B, average performance or neutral; C, below average, or particular

disadvantage.

Agent

AvaildUlllLy Pnct

Stability

in

Metals Compatibility

Polymers

Compatibility

Combustion

Byproducts Residue

rL-1 10 D

DD

AA

AA A B A

rjD D

AAD B B A

D a

AA L A B A

r>D

DDAA L

DD B A

HFC-125 B B A A B A

HFC-32/125 A B B B A C A

HFC- 134a A B A A B C A

HFC-227ea B B A A A B A

HFC-236fe B B

* B A B A

HCFC-22 A A A C B B A

HCFC-124 B B A B A A A

CF3I C C B/C + + + #

NaHCOj A A B C

* C C

* insufficient data for rating # highly acidic impurity present in some samples

was observed as a decomposition product when the CF3I was subjected to the high temperature

storage test. No other significant decomposition products were identified. (However, noticeable

acidic residue was present in batches of CF3I received later from a second supplier.) U.S. production

of CF3I, as of September 30, 1993, was of the order of 50 kg/yr, and bulk quantities sufficient for

full-scale testing were unavailable. The renewed interest generated in iodofluorocarbons could cause

industrial production to increase substantially. However, as of this report, bulk production of suffi- ciently pure chemical has not been demonstrated.

All of the NaHC03 which is not carried away from a fire remains as an undesirable residue.

Instability of the carbonate under high temperature storage was not fully explored, but some metal

corrosion was evident. A major product of combustion from the powder is NaOH, which interacts aggressively with aluminum.
SECTION 10. SUMMARY AND RECOMMENDATIONS

10.2 Ranking of Agents

835

10.2.1 Agents Recommended Against Full-Scale Engine Nacelle and Dry Bay Testing. Cup

burner tests with HFC-32 confirmed previous reports that the compound is flammable. The decisior was made, in consultation with the sponsors, to replace HFC-32 with HFC-236fa in the test matrix.

During the course of the evaluation, seven chemicals were identified which either performed

unimpressively or were deficient in other ways. Considering the available alternatives, we do not recommend these for full-scale testing. These chemicals are listed below with the major reasons for our non-recommendation:

Agent Major Deficiencies

FC-116 highest volume factor in all flame configurations; critical temperature below

20 °C leading to uncertainty in design of storage vessel and speed of dis- charge

HFC-32/HFC-125 high volume factor in all flame configurations; highest over-pressures in detonation tube; high acid gas levels in combustion products

HFC-236fa highest vaporization temperature; lowest Jakob number

FC-318 high vaporization temperature; low Jakob number; below average metals

compatibility; long atmospheric lifetime

FC-31-10 high vaporization temperature; low Jakob number; below average metals

compatibility; long atmospheric lifetime

HCFC-22 highest ODP; poorest compatibility with metals

HFC- 134a overall mediocre performance; higher levels of HF in combustion products

10.2.1 Ranking of Remaining Agents. The performance of all the candidates regarding environmental and operational issues was judged to be sufficient enough that those criteria, alone, were

inadequate to disqualify any of the remaining agents. The relative weighting of the suppression and

dispersion criteria depended upon the application; i.e., full-scale engine nacelle or dry bay testing.

10.2.1.1 For Full-scale Simulated Engine Nacelle Testing. The release of the agent into the

engine nacelle is crew-initiated, and the agent is normally distributed to diff'erent engines or regions of

a single nacelle via piping manifolds. The fire is relatively slowly growing, and conditions which

favor the transition to a detonation are unlikely to be present. The highest premium is placed upon

fire suppression efficiency (low VF) for the three atmospheric pressure test flames. The time for dispersion is relatively long, so that heat is transferred from the air to the spray to assist vaporization.

Thus, T^jn is more appropriate than Ja as a measure of dispersion efl'ectiveness. Table 4 summarizes

these parameters and the recommended action for each agent regarding full-scale engine nacelle
836

Table 4.

SECTION 10. SUMMARY AND RECOMMENDATIONS

Ranking of agents for full-scale engine nacelle test matrix

i^VWld&t V 1 T' min

Environmental Operational

Tcci IPC Recommendation

HCFC-124 2.2 + 28 °C -1- test HFC- 125 3.0 - 6 °C + + test FC-218 2.8 - 7 °C -1- 1st back-up

HFC-227ea 2.5 + 19 °C + + 2nd back-up

NaHCOg 0.5 powder + partial test matrix

CF3I 0.8 - 1 °C 0 partial test matrix

testing. A " + " indicates overall positive attributes in that category and a

"-" indicates that there is at least one serious concern in applying the agent.

The volume factor for the HCFC-124 is the lowest of all the core agents evaluated. Full-scale

testing at low temperatures is required to ascertain if a Tj^j^^ of -I- 28 °C might hamper the agent's

effectiveness. The negative environmental factor is the non-zero ODP of the chlorine-containing

compound. Nonetheless, HCFC-124 is recommended for full-scale testing, especially to see if the

low volume factor is validated.

HFC-125 and FC-218 have similar values for Tj^j^ and only slightly different values for VF.

The HFC-125 is a more conservative choice for full-scale testing because the FC-218 has a much

longer atmospheric lifetime. HFC-227ea is a logical back-up because it has the second best volume

factor.Sodium bicarbonate has a volume factor half the size of halon 1301, and less than 1/4 the best of

the core agents. There are three concerns with using sodium bicarbonate:

1. powders do not disperse as effectively as gases; hence, the performance of NaHC03 is likely

to be lower in an engine nacelle than is indicated by the bench-scale tests;

2. NaOH, a by-product of the suppression process, reacts strongly with aluminum, which leads

to the corrosion of exposed aluminum surfaces; and

3 . high temperature storage leads to a deterioration of the flow properties of the powder.

Full-scale engine nacelle tests could determine if the first concern is warranted. In connection

with the second concern, it should be ascertained whether the nacelle exhaust would put decomposition products in contact with aluminum. If so, aluminum test coupons strategically placed in the

engine nacelle test fixture could be used to assess the significance of corrosion. The third problem
SECTION 10. SUMMARY AND RECOMMENDATIONS 837

requires additional research to determine the best solution, but it is unlikely, by itself, to be a cause

for rejecting sodium bicarbonate.

lodotrifluoromethane also has a volume factor less than halon 1301. It leaves little residue (if it is pure) and appears to have a short tropospheric lifetime. Only a limited full-scale test matrix is recommended because:

1 . greatly increased production levels are required to meet the full-scale test program and fleet needs;

2. toxicity, stability and corrosion test results are incomplete, and

3. detailed atmospheric chemistry and the implications of stratospheric release need to be further

investigated.

10.2.1.2 For Full-scale Simulated Dry Bay Testing. Different criteria apply to the perfor- mance of an agent in dry bay fire suppression. Here, time is of the essence. Agent release is initiat- ed automatically within milliseconds of a sensed threat. Following ignition, the heat release can

create increases in pressure if expansion is restricted, leading to flame acceleration, the generation of

turbulence due to high speed flow over obstructions, and the possibility of a strong shock wave.

Effectiveness in damping the pressure build-up in the detonation/deflagration tube, averaged with the

results in the spray burner, was used to determine an appropriate volume factor for the dry bay.

Because the discharge occurs so quickly, there is no time for heat transfer from the surrounding air to vaporize the liquid agent; hence, the rankings on dispersion from Section 3.5 apply. The environmental factors and operational issues are identical to the engine nacelle application, but differences in the amount of agent released in a year and the cleanup procedure following a false discharge affect the way this information is factored into the recommendations. Table 5 gives the ranking of the

remaining agents for full-scale simulated dry bay fire testing.

FC-218 and HFC- 125 are recommended for testing because they have the highest Jakob

numbers. None of the core agents has a volume factor below FC-218. The negative associated with

long atmospheric life is of less concern for dry bay applications because the total annual release is much lower than in engine nacelle applications. HFC-125 had a slightly higher VF than FC-218, and

produced high over-pressures in the detonation/deflagration tube at concentrations below the suppression level. Testing is recommended, but should similar over-pressures be observed in the simulated

dry bay, then HFC-227 is the recommended back-up. The VF for HCFC-124 is the same as for FC218, but the HCFC is difficult to vaporize. Considering that it has a non-zero ODP, FC-218 remains

the better choice.

The CF3I has a VF only 10% greater than halon 1301, but a very low Jakob number. A unique

pressure/concentration behavior was observed with this agent in the detonation/deflagration tube

which needs further investigation. If sufficient supplies of CF3I can be arranged, full-scale testing in the simulated dry bay is recommended, especially at low temperatures. The ongoing toxicity,

stability, corrosion and atmospheric chemistry studies should be closely monitored. When one compares VF in the engine nacelle and dry bay applications, the diff'erence in performance between

CF3I and the best core agent is much greater in the engine nacelle application. Adding in the low Ja number makes CF3I somewhat less attractive for the dry bay application.

The recommendation against sodium bicarbonate in the dry bay program is based upon the much

higher premium placed on dispersion to counter the fire threat. A destructive fire can develop in tens of milliseconds, requiring that the fire suppressing agent be distributed throughout the volume in an

equivalent time. The gasified agents can easily penetrate a reasonable amount of clutter in the dry
838 SECTION 10. SUMMARY AND RECOMMENDATIONS

Table 5. Ranking of agents for full-scale dry bay test matrix

Agent Average VF

Sect 3.5

Rank

Environmental

Factors

Operational

Issues Recommendation

FC-218 1.6 1 + test HFC- 125 1.8 1 + + test HFC-227ea 1.6 6 + + 1st back-up

HCFC-124 1.6 7 do not test

CF3I 1.1 not ranked 0 partial test matrix

NaHC03 incomplete

data

powder + do not test

bay, but powders can impact and stick on surfaces and do not follow the carrier gas streamlines,

greatly reducing the mass of material which actually is delivered to the fire zone. Clean-up could be more difficult in the dry bay, and since exposed aluminum surfaces are likely to be plentiful, they

would be susceptible to NaOH corrosion.

A summary of the recommendations for all of the core agents, CF3I and NaHC03 is given in Table 6. The inter-governmental technology transition team will weight the NIST recommendations

with other practical concerns to arrive at the final set of chemicals for full-scale testing.

10.3 Ongoing Studies

The original research program begun in 1992 contained several projects which were to continue

beyond September 30, 1993. These tasks were designed to enhance the selection and use of the

optimal suppressant(s). Intermediate data will be available late in 1994 to assist the sponsors in evaluating the three candidate suppressants for each application. Knowledge developed during the

continuation period will guide optimal use of agents selected for deployment.

10.3.1 By-Product Formation.

10.3.1.1 Instability Under Storage. In the one-month exposures, HFC- 125, FC-218 and

HCFC-227 showed no decomposition at 150 °C in the presence of any of the eight metals. By

contrast, CF3I did show some iodine formation and metals discoloration at 150 °C. During the next two years, the four selected chemicals will be exposed for about 18 months at three temperatures (22,

125, and 150 °C) to three metals: Nitronic-40 (21-6-9 stainless steel), 304 austenitic stainless steel,

and titanium 15-3-3-3.
SECTION 10. SUMMARY AND RECOMMENDATIONS

Table 6. Summary of recommended actions by agent

839

Agent Recommended Actions Cautions HFC- 125 Perform full-scale test matrix in nacelle. Screen for high pressures in dry bay, and complete

matrix if pressure is not a problem.

Over-pressures may occur in dry bay tests.

HCFC-124 Perform full test matrix in engine nacelle simulator. Note high T^^, and non-zero ODP. Care need- ed when choosing elastomers to reduce chance

of swelling. FC-218 Perform full test matrix in dry bay simulator. Use as a back-up for engine nacelle.

L-uiig diiiiuapiici n, iiiciiiiic, propenies or greases

affected at high temps. HFC-227ea Consider it as a back-up to HFC-125 in engine nacelle and dry bay test program.

Relatively high T^^ may limit dispersion

NaHCOj Test in engine nacelle simulator. Truncate test matrix if mass concentrations consistently exceed HFC-125.

Complications associated with powders (particle dispersion, corrosive by-pnaducts of combustion,

Lciiipcidiuic aiuidgc^ neeu to De orr-sei oy

superior full-scale fire suppression effectiveness

to warrant fleet adoption.

CF3I Perform reduced test matrix in engine nacelle simulator

to verify superior performance. Perform tests in dry

bay consistent with availability of chemical.

Supply for testing limited: pilot scale production

capability for 1995 not proven; toxicity, atmospheric chemistry, stability, and materials compatibility not fully documented HFC-134a Not recommended for full-scale testing. Mediocre performance compared to alternative HFCs, higher HF levels

FC-116 Not recommended for full-scale testing. Critical temperature below ambient limits stor- age

HFC-236fe Nor rpcnmmpndpH for fiill-^ralp fp*;tin(T T-tlon Olgli I limitc M 1 cr^f^rc ir\n 1

fy^jj^ IIIIIIIA UlspCIMUlI HCFC-22 Not recommended for full-scale testing. Incompatible with many materials, non-zero ODP

FC-31-10 Not recommended for full-scale testing. High T^jj,, some materials incompatibility

FC-318 Not recommended for full-scale testing. High T^jj,, materials issues

HFC-32/125 Not recommended for full-scale testing. High over-pressure, high VF

HFC-32 Not recommended for full-scale testing. Flammable

These were selected in conjunction with the sponsors and the manufacturers of suppression

storage containers and systems. Samples will be analyzed for decomposition every 1 to 2 months

using FTIR, with stretched intervals if no decomposition is observed.

CF3I will be exposed to the same three metals, with and without the presence of metallic copper,

a suggested inhibitor of the agent's decomposition, and perhaps at a fourth temperature. We expect

that FTIR analyses will need to be supplemented by gas chromatography, mass spectroscopy or a combination of the two. Should there be any decomposition of any of the chemicals in the presence of any of the metals,

approximate first-order decomposition rate constants will be determined. These results will enable
840 SECTION 10. SUMMARY AND RECOMMENDATIONS

evaluation of the viability of storing this chemical for extended periods. Alternatively, the results can be used to determine how often cylinders will need to be refilled. The work will be completed by

September, 1995.

10.3.1.2 Combustion By-Products. The objective of this project is to develop an understanding of acid gas formation in hydrocarbon flames suppressed by three halon 1301 replacements, HFC125, HFC-227 and HFC-218. This will provide the basis for a predictive capability for the amount

of acid gas formed during the suppression of dry bay and engine nacelle fires, allowing the rank

ordering of the three agents with respect to by-product production for these situations. CF3I will be

tested to verify that the amounts of HF formed are comparable to that formed with CFjBv.

The general approach is to obtain global information on the final products in simple burners

which permit operation over a range of conditions imitating those of dry bay and engine nacelle fires. This will allow examination of the efi'ects of fuel and inhibitor structure and concentration. Fluid

dynamic effects will be examined through a comparison of difiiision and premixed flames under

laminar and turbulent conditions. Since the characteristic time for chemistry is on the order of one

millisecond, and fluid dynamic times are usually longer, the chemistry important in HF formation can be studied under steady-state configurations more amenable to species sampling and variation of the

individual operating parameters over a range of conditions which simulate those estimated for the dry

bay and engine nacelle fires. The results from the steady-state measurements will then be used to predict the HF formed in flames suppressed under transient conditions. To test these predictions,

experiments will be conducted where the agent is added rapidly to a flame so that the flame extinguishes, and the HF formed will be measured. These results will be compared to the predictions

based on the steady-state measurements.

A second component will develop correlations, based on experimental results and global properties, of acid gas formation in inhibited flames based on inhibitor chemical formula and concentration,

fuel type and consumption rate and flame type. The extent of the correlations will depend upon the

importance of the chemical kinetic rates of inhibitor and inhibitor fragment decomposition. Measuring the concentration of acid gases formed in the flames described above and comparing these with

the estimated concentrations from a simple global model will help elucidate the degree of kinetic

control in the acid gas formation process. Third, we will develop understanding of the chemical mechanism of acid gas formation in hydrocarbon flames to provide a fundamental basis for a true predictive capability. This entails (a) detailed numerical calculations of flame structure using a chemical kinetic code and (b) validation of

the mechanism through comparison of model predictions and experimental measurements of

temperature and species concentrations in the flame. The latter will be obtained in a laboratory flame

using micro-probe gas sampling with mass spectrometry.

10.3.2 Metals Corrosion. The recent data show that a selection of metals are compatible with the

down-selected chemicals over a one-month period. Additional testing will be conducted to enable

judgments concerning relative performance of the chemicals over their storage lives in the suppression

system.

Samples of the three metals will be exposed to the four chemicals plus halon 1301 at both

ambient temperature and 150 °C at 5.4 MPa. Periodic (initially monthly) examination will provide a time history of the advancing of any corrosive efl'ects. Slow strain rate tensile tests will be performed

at ambient conditions and 150 °C and 5.4 MPa to provide information relating to liquid phase interactions.Electrochemical experiments of the three metals and four chemicals will be conducted to:
SECTION 10. SUMMARY AND RECOMMENDATIONS 841

• improve the quality of corrosion ratings by helping to resolve the scientific uncertainties

inherent in short-term exposure tests;

• enable assessment of the effects of variations in the exposure conditions that could cause rapid

runaway corrosion that would otherwise be missed by the exposure test matrix due to the time

limitations; and

• enable assessment of the potential for galvanic attack and local breakdown of protective films

to a level not possible with exposure tests.

It is also important to assess the potential for damage from the combustion by-products of the

suppressants. Thirty-day surface exposures of key metals (identified by the sponsors) will be

conducted in aqueous solutions of HE Up to three concentrations will be used, depending upon the

magnitude of the predicted difference in acid production for the four chemicals. The studies will be

conducted at ambient temperature and pressure.

10.3.3 Compatibility with Seals and Elastomers. The recent research showed some differentiation

among the chemicals in their interactions with the selected elastomers and greases, and these were

factored into the down-selection process. It was also observed that the 150 °C exposures resulted in degradation of many of the seal materials by some of the chemicals. Subsequent information from the

Air Force indicated that all-metal storage is used for exposures above about 70 °C.

In the continuing research, mechanical property measurements and additional short-term

exposures will be performed at 50, 75, and 100 °C for the four agents and seven elastomers, the

original six plus an ethylene-propylene-diene terpolymer (EPDM). These will enable a valid

activation energy analysis over the full temperature range of interest. Compression set measurements

and tensile testing (ultimate elongation, tensile strength, and modulus at 100% elongation) will be

conducted after exposure times of 1, 2, and 4 weeks. The analysis of these measurements using an

activated process model will provide reliable estimates of the elastomer resistance to the fluid exposure in service conditions.

Since CF3I was not included in the initial experimental plan, swelling measurements will be con- ducted at 35, 70, 105, and 150 °C for all elastomers (crosslinked/uncrosslinked sets of fluorocarbon,

standard and low temperature nitrile, silicone, fluorosilicone, and neoprene) and greases (Braycote

600, Krytox 240AC, Braycote 807) at four different vapor pressures. Swelling measurements on a crosslinked/uncrosslinked set of EPDM rubber will also be performed.

Following completion of the above studies, samples of seven elastomers will be exposed for about 18 months to the four fluids at 75 °C. Tensile testing and compression set measurements will

be made after 8, 16, 32, 48 and 78 weeks.

10.4 Complementary Research Underway

There are additional projects now beginning at NIST that will provide knowledge important to the

implementation of an agent selection decision. The objectives are summarized in this section.

10.4.1 Discharge Optimization. The research documented above shows that the effectiveness of a

fire-fighting agent can depend as much on how the material is delivered to the fire zone as it does on

its molecular composition. A particular chemical which leads to flame extinction in a laboratory
842 SECTION 10. SUMMARY AND RECOMMENDATIONS

burner at very low concentrations may, in fact, be a poorer choice than one which exhibits no

chemical activity if the inert material can be transported to the critical portions of an actual fire in a much more efficient manner. This project will develop the understanding needed to ensure that the

agents chosen by the inter-governmental technology transition team for full-scale testing are delivered

to the fire in an efficient manner so that an unbiased evaluation of system effectiveness can be made.

The generic scientific data will also aid fire suppression system suppliers and airframe manufacturers

in their eff'orts to design protective systems for existing and future aircraft. To accomplish this, NIST

will:

• determine the practical limits to the design of system hardware which constrain what can be

done;

• expand temperature-pressure data on the recommended fluids, pure and in mixtures with

nitrogen, to enable the development of correlations for total pressure as a function of fill density and temperature for pressures up to 40 MPa;

• for dry bay applications, determine how the initial conditions in the storage bottle and the

design of the release mechanism and exit orifice can be modified to enhance the discharge

rate of the recommended fluid agents directly into an unconfined space; and

• for engine nacelle applications, determine the performance of the recommended fluid agents

flowing through various piping geometries and components in transient operation.

10.4.2 Alternative Agent Concentration Monitors. It is essential to know that the fire suppression

system will deliver an appropriate concentration of suppressant to the regions where a fire might

occur. This requires a probe capable of extremely fast (milliseconds) measurements at many points.

Our research has indicated that the initial flows may be either liquid or two-phase and that the mixing

patterns are extremely complex.

The objective of this efl'ort is to evaluate possible methods for real-time measurements of

concentrations of alternative fire fighting agents for dry bay and nacelle fire applications. No analytical measurement techniques have yet been applied to such a complex problem. It will therefore be

necessary to extend current measurement techniques or to develop completely new approaches. If one

or more feasible approaches are identified early in the investigation, a demonstration system will be

developed for characterization under actual test situations. This will be accomplished as follows:

• search the concentration measurement literature to identify measurement techniques for the

required concentration measurements and two-phase flow characterization and recommend

approaches for further research;

• test and evaluate an infrared sensing instrument recently developed by a company under

contract to the Air Force for measuring alternative agent concentrations; and

• conduct theoretical and experimental feasibility evaluation of aspirated hot-film probes,

recommending whether to pursue further development of this measurement technique.

10.4.3 Engine Nacelle Simulations. There are a number of diff^erent aircraft and operating conditions for which a new suppressant for nacelle fires is needed. Because testing cannot be performed
SECTION 10. SUMMARY AND RECOMMENDATIONS 843

for all possible aircraft and conditions, knowledge is needed which will provide guidance in the

extension of the Wright Patterson AFB full-scale data to untested systems and conditions.

Extinguishment occurs when a critical amount of agent is transported to the flame, where it is entrained into the primary reaction zone. The extinction process is afl'ected by a number of parameters, including the velocity of the air flow, the type and quantity of fuel, the system temperature,

agent properties and concentration, and the flow field geometry {e.g., the location of obstacles in the

flow field). Re-ignition is also dependent on these system parameters and should be considered

independently from the extinction phenomena.

The objective of this research is to produce organized guidance for adjusting the needed

concentration of candidate fire suppressants over a range of engine nacelle fire conditions. This will

include

• review of the range of parameters which characterize the temperature and flow field in various engine nacelle types in an effort to understand better the possible fire scenarios and to develop an appropriate test program;

• use of the coaxial spray burner described earlier to measure (in a well-controlled and reducedscale environment) the eflTects of air flow, air temperature, fuel flow, pre-burn time, system

pressure, agent type (CF3I, HFC-125, and HFC-227ea), rate of application, and the flow field geometry (e.g., effect of an obstacle in the flow field) on the amount of chemical needed to suppress a fuel-spray fire;

• measurement of the impact of increased agent temperature on flame stability;

• computational modeling and optimizing the agent concentration as a function of location, rate, and orientation of agent injection, including possible multiple-location injection and differ- ences in nacelle geometry, with wind tunnel validation; and

• compilation of the knowledge and data gained through the study into tools to assist sup- pression system designers, including information regarding concentration/duration requirements as a function of fire zone conditions.

The engine nacelle simulations will complement and provide guidance to the ongoing full-scale

effort at Wright-Patterson.

10.5 Recommendations for Further Research

A significant amount of resources, eflFort, and science have gone into selecting the best candidates for

full-scale evaluation, and it is expected that the additional testing at Wright-Patterson AFB will be

successful in identifying one or two chemicals which will be suitable for replacing halon 1301 for inflight suppression of engine nacelle and dry bay fires. It is equally likely that some inconsistencies

will be revealed in scaling up one hundred-fold from the laboratory apparatus to the actual test

articles. More understanding is required of the fundamental principles controlling the suppression of a turbulent flame, the dispersion of a vaporizing spray in an enclosure, the chemistry of halogenated

species in the flame and in the atmosphere, and modeling the interaction of all these phenomena to
844 SECTION 10. SUMMARY AND RECOMMENDATIONS

increase the generality of the laboratory results and suggest improvements in the full-scale test protocol. Specifically, pursuit of the following topics is recommended:

• Inhibition mechanisms by both halocarbons and non-halogenated entities, including chemical

kinetics measurements, studies of flame chemistry and dynamics, and computer modeling of

chemically perturbed combustion.

• Innovative suppression, such as by non-intrusive flame perturbations or species chemically

generated in situ.

• Atmospheric chemistry of flame extinguishants other than chlorinated and brominated species,

including laboratory measurements of the rate processes and computer modeling of the

atmospheric impacts.

• Turbulent two-phase flow, both experimental characterization and numerical modeling of the

dynamics.

• Standardized suppression test methods at both small and medium scale, including the

appropriate instrumentation.

True understanding of a phenomenon requires a sustained eff^ort, and the flexibility to allow

research to explore multiple paths. In the environment of continuously changing regulations and

evolving technologies, the understanding which would be generated by the above research is essential so that optimal agents and suppression strategies can be developed in a rational and efficient manner.