H.r.n. Jones-Application of Combustion Principles to Domestic Gas Burner Design (1990)

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THE APPLICATION OF COMBUSTION PRINCIPLES TO

DOMESTIC GAS BURNER DESIGN

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THE APPLICATION OF COMBUSTION PRINCIPLES TO

DOMESTIC GAS BURNER DESIGN

H.R.N. Jones , MA, PhD, CEng, MIGasE

British Gas Teaching Fellow, University of Cambridge

E. & F.N.Spon, London and New York

in association with British Gas plc

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First published 1989 by E. & F.N. Spon Ltd

11 New Fetter Lane, London EC4P 4EE and 29 West 35th Street, New York, NY 10001

in association with British Gas plc Research and Technology Division, 152 Grosvenor Road,

London SW1V 3JL This edition published in the Taylor & Francis e-Library, 2005.

“To purchase your own copy of this or any of Taylor & Francis or Routledge's collection of thousands of eBooks please go to

www.eBookstore.tandf.co.uk.”

© 1989 British Gas plc

ISBN 0-203-47313-2 Master e-book ISBN

ISBN 0-203-78137-6 (Adobe e-Reader Format) ISBN 0 419 14800 0 (Print Edition)

All rights reserved. No part of this book may be reprinted, or repro duced or utilized in any form or by any electronic, mechanical or other

means, now known or hereinafter invented, including photocopying and recording, or in any information storage and retrieval system,

without permission in writing from the publisher.

British Library Cataloguing in Publication Data Jones, H.R.N. (Howard Richard Neil), 1957

The application of combustion principles to domestic gas burner design

I. Household gas appliances. Combustion equipment 1. Title 683’.88

ISBN 0-419-14800-0 zvv

Foreword

It is a great pleasure to write a foreword to this excellent book. The gas industry is almosttwo hundred years old and for the first hundred years was dependent on a veryelementary understanding of the combustion process. The advent of the premixed flameby Bunsen, the concept of burning velocity and the development of the chemistry of thecombustion mechanism by Hinshelwood and by Semenov in the 1920s and 1930s formedthe basis of our present day understanding of the fundamentals of combustion. Bone andTownend’s Flame and Combustion in Gases in 1927 was the first major text oncombustion and flame structure and heralded an enormous expansion of combustionresearch. However, our understanding of flame propagation, with its complex interactionof chemistry and aerodynamics, still remains incomplete at the present time.

The ever increasing demand for higher efficiency and precise controllability coupled with more stringent pollutant emission requirements imposes considerable pressures oncombustion research and development. The gas industry enjoys a strong position here,but nevertheless recognises the need to continue to develop its technology. The advancesover the past two decades have been considerable as is made clear by this textbook, butfurther developments in our understanding of combustion aerodynamics will make asignificant impact in future years. This book gives a comprehensive account of onespecialist area of combustion, that of domestic gas burner design, which is only brieflydiscussed in existing combustion and flame textbooks. It covers not only existing burnersbut also future domestic burners, and is therefore greatly welcomed.

Professor A.Williams Department of Fuel and Energy The Houldsworth School of Applied Sciences University of Leeds

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Preface

The use of gas as a fuel has increased dramatically throughout the world during the lastthirty years. In many countries this has been primarily due to the development of naturalgas resources as a replacement for the older and, in many cases, declining manufacturedgas industries. Great Britain is no exception, especially in the domestic sector, where,since conversion to natural gas, there has been a marked rise in both the number ofcustomers and annual gas consumption. Gas combustion research has played a key role in this rapidly expanding market, particularly with respect to advances in appliance design.These advances are manifest in two areas: firstly through the increase in applianceefficiency which has been demanded in an ever more energy-conscious age; secondly, and arguably more important, through developments in burner design, without whichneither the remarkable success of conversion nor the introduction of today’s compact and highly efficient appliances could have been achieved.

Much fundamental work in combustion and burner design for domestic natural gasappliances has been published in the form of scientific papers. This volume bringstogether much of that work, and is aimed at both the student and the qualified engineer.After an opening chapter, which provides relevant definitions and some basic combustionchemistry, four chapters discuss, in some detail, design aspects of the three mostcommonly encountered types of burner (non-aerated, partially aerated and fully aerated)together with some more unusual burner systems. Chapter 6 considers the application of these design aspects to various domestic appliances, in particular the constraints whichmust be placed on any practical design of burner, while in Chapter 7, burner control systems are discussed with special emphasis given to safety, ease of appliance operationand methods of maintaining high efficiency.

A book such as this would be impossible to produce without the assistance of a numberof people at the British Gas Watson House Research Station who have worked on manyof the aspects described herein. I wish to thank, in particular, Mr J.J.F.Flood, Mr K.J.A.Hargreaves, Mr J.A.Harris and members past and present of the Gas CharacteristicsGroup and the Advanced Combustion Systems and Controls Group for many usefuldiscussions on individual topics. Additionally, thanks are due to Dr N.C.Ross, DrR.South, Dr A.H. Curran, Dr M.B.Green, Dr M.C.Patterson (all of Watson House), Mr N.Flicker (Midlands Research Station), Prof. A.Williams (Leeds University), MrR.Wakefield (British Gas HQ) and Dr R.Pritchard (Salford University) for theircomments on the completed manuscript.

Special thanks are due also to Miss N.O’Brien for her care and patience whilst typing the original draft manuscript and for coping so admirably with subsequent textualrearrangements and revisions.

H.R.N.Jones January 1989

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Contents

1 Combustion fundamentals and gas properties 12 Non-aerated burners 183 Partially aerated burners 314 Fully aerated burners 795 Future domestic burners 1096 Burners in appliances 1267 Burner controls 151

References 178

Subject index 182

Author index 188

Index of symbols 194

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CHAPTER 1 Combustion fundamentals and gas properties

1.1 General

The term combustion is applicable to many phenomena: afterglows, cool flames, flash photolysis, fireballs, fuel cells, as well as the more traditional concept of the stationaryflame. All are characterized by the emission of energy and (usually) light resulting fromchemical reaction between a fuel and an oxidant*. The physical and chemical propertiesof any flame are greatly dependent on the exact conditions of combustion: gas flow maybe turbulent or laminar; fuel and oxidant may or may not be premixed to some extentprior to burning; variations in temperature and pressure can have a large effect.

Engineers involved with domestic gas appliances can normally restrict their area of interest to gaseous fuels burning in air at approximately atmospheric pressure. The prime

1.1 General 1.2 Flame structure 1.2.1 Aerated flames 1.2.2 Non-aerated flames 1.3 Combustion properties 1.3.2 Minimum ignition energy 1.3.3 Burning velocity 1.3.4 Flame stability 1.4 Gas characteristics 1.4.1 Basic flame chemistry 1.4.2 Minor combustion products 1.4.3 Properties of gas mixtures 1.4.4 Classification of gases and appliances 1.4.5 Interchangeability of gases 1.4.6 Prediction of interchangeability 1.4.7 Test gases

objectives of the burner designer are to ensure that (a) the correct mixture of gas and airis supplied, (b) ignition is controlled and reliable, (c) the resultant flame is of the requiredshape and structure and is stable, and (d) the appliance is inherently safe. In order toachieve these, an understanding of the general structure and chemistry of a flame is mostimportant. This chapter is intended purely to introduce a number of concepts anddefinitions in the areas of flame structure and gas characteristics, many of which aredeveloped further in succeeding chapters. Readers requiring a more comprehensivetreatment should consult one or more of the many excellent fundamental text-books available (e.g. refs. 1–5).

1.2 Flame structure

Stationary flames are customarily divided into three classes depending on how the fueland oxidant come into contact. In a diffusion or non-aerated flame, combustion proceeds by diffusion of oxygen into the unburnt fuel stream as the fuel is injected into thecombustion chamber. Alternatively, the unburnt fuel stream may be supplied with air(known as primary air) before combustion occurs. If all the air required for completecombustion is provided as primary air, then the flame is said to be fully aerated or fullypremixed. If only part of the total air required is supplied in the primary air, then theflame is said to be partially aerated, and the remaining air (known as secondary air)diffuses into the hot combustion gases downstream of the flame front. The three types offlame differ both in appearance (Figure 1.1) and with respect to the dominant physical and chemical processes present. All have been extensively used by combustion engineers,the properties of each being particularly suited to certain applications.

1.2.1 Aerated flames

The simplest aerated flame is that of the Bunsen burner with the air hole open, whereprimary air is entrained into the gas stream prior to combustion at the top of the burnertube. All aerated flames are generally considered to consist of three distinct regions orzones:

1. Preheat zone. In the preheat zone, unburnt gases are taken from ambi

* Oxidant here includes not just oxygen and air but such special cases as the reactions of luorine and other halogens with hydrogenous fuels.

ent temperature to a temperature (about 700 to 1000 K) where combustion is initiated.Heat for this is supplied by radiation and conduction against the direction of flow fromfurther downstream. Carbonaceous fuels may pyrolyse in the preheat zone, therebyaffecting the nature of the fuel. The preheat zone thickness is about 1 mm in atmosphericpressure flames.

2. The reaction zone. Combustion actually takes place in the reaction zone or flame

The application of combustion principles to 2

front, which is a region of intense chemical activity just 0.1 mm thick at atmosphericpressure. It is clearly visible in hydrocarbon flames as a turquoise-blue inner cone. In this region, temperature increases to over 2000 K and the gas expands rapidly such that gasvelocity increases to about seven times that of the unburnt mixture.

3. The burnt gases. The reactive chemical fragments formed in the reaction zone give rise to stable combustion products such as water vapour, carbon monoxide and carbondioxide; additionally, depending on whether the flame is partially or fully aerated, excessfuel or oxygen will remain. In a partially aerated hydrocarbon flame, the unused fuelcontinues to react with secondary oxygen diffusing into the hot gases, thereby forming awell-defined violet outer mantle which extends downstream from the reaction zone. In a fully aerated flame, no further oxygen is required, so no outer mantle exists, althoughthere is usually a visible afterglow extending downstream from the reaction zone.

1.2.2 Non-aerated flames

If a Bunsen burner is operated with the air hole closed, then no primary air can beentrained and neat gas issues from the burner tube (Figure 1.1). Combustion takes place

Figure 1.1 Schematic showing structure and air supply for non-aerated, partially aerated and fully aerated flames.

as oxygen from the surrounding atmosphere diffuses into the emerging gas stream. Theresultant flame is not as well structured spatially as an aerated flame, there being noeasily defined regions. Unlike aerated flames, chemical reaction is not confined to a well-defined reaction zone. However, mathematical treatments usually find it convenient tolocate a reaction zone as an infinitesimally thin region where the fuel/oxygen ratio is

Combustion fundamentals and gas properties 3

stoichiometric*, such that inside the enclosed volume the mixture is fuel-rich, while outside that volume the mixture is fuel-lean.

1.3 Combustion properties

In addition to the appearance of a flame on a burner, there are a number of importantconcepts which must be understood before any detailed analysis of the physicalproperties of a flame can be undertaken. Firstly, for ignition to occur, the gasconcentration in a gas/air mixture must be within a certain range, i.e. between theflammability limits. Secondly, a finite amount of energy must be supplied to the mixturebefore a nominally flammable mixture will ignite. Thirdly, once ignited, the flame frontwill propagate through the unburnt mixture at the burning velocity of that mixture.Burning velocity in particular is a vital parameter in discussion of the stability of flameson burners.

1.3.1 Flammability limits

For all fuel/air mixtures there are composition limits outside which ignition will notoccur, that is to say, if there is either too little fuel or too little oxygen, chemicallyreactive collisions between fuel and oxygen molecules are too infrequent for combustion

*A fuel/oxidant mixture is said to be stoichiometric when there is, theoretically, just sufficient oxidant present for complete combustion of the fuel.

Table 1.1 Flammability limits of selected fuels in air at 25°C and atmospheric pressure, (Source: Zabetakis 6 )

Gas Flammability limit (% gas by volume)

Lower Upper

Hydrogen 4.0 75

Methane 5.0 15.0

Ethane 3.0 12.4

Propane 2.1 9.5

Butane 1.8 8.4

Ethylene 2.7 36

Propylene 2.4 11

But-l-ene 1.7 9.7


to be supported. The upper and lower limits are dependent on a number of factors, not allof which are properties of the mixture. These include gas composition, temperature,pressure, dimensions of the vessel in which combustion occurs and the point and natureof ignition in the vessel. Table 1.1 lists flammability limits for a selection of gaseousfuels, and shows the rather narrow range of flammability for hydrocarbon fuels comparedwith hydrogen. Zabetakis 6 provides a very thorough review of the subject.

1.3.2 Minimum ignition energy

Figure 1.2 Minimum Spark ignition energy for selected fuel gases in air at 25 °C, 1atm pressure.(After Lewis and von Elbe 4 .)

Even if a gas/air mixture is within its flammability limits, a finite amount of energy isstill needed for ignition to occur, that is the initiating chemical reactions require anactivation energy. If the mixture is raised to a high enough temperature (about 600 °C for


methane), ignition can be spontaneous (termed ‘autoignition’). For most practical purposes, the activation energy must be supplied by an external source (e.g. pilot flame,electric spark). The minimum ignition energy is dependent on the mixture composition,pressure, temperature and the ignition method used. Figure 1.2 shows the variation of minimum spark ignition energy with composition for a number of fuel/air mixtures atroom temperature and atmospheric pressure. The relatively high ignition energies forhydrocarbon fuels were of particular concern during conversion to natural gas.Theoretical treatments are given by Lewis and von Elbe 4 and Williams 5 , while the more practical design aspects are discussed in Chapter 7.

1.3.3 Burning velocity

When a fuel/oxidant mixture has ignited, the flame front produced propagates through theremaining unburnt gases at a rate dependent on the mixture composition, pressure andtemperature. The burning velocity is a fundamental property of the mixture and is linkedto the overall chemical reaction rate in the flame. Burning velocity is defined as thevelocity normal to the flame front, relative to the unburnt gas, at which an infinite one-dimensional flame propagates through the unburnt gas mixture.

Figure 1.3 Burning velocity for selected fuel gases in air at 25 °C, 1 atm pressure. (After Rose and Cooper 7 .)


Burning velocity is greatly dependent on gas composition, rising from almost zero at eachflammability limit to a maximum near the point of stoichiometry (see Figure 1.3). Because burning velocity is related to chemical kinetics and molecular transport properties, it is possible, in principle, to derive values of burning velocity as a function oftemperature and pressure by mathematical modelling 5 . In practice, because of the great complexity of both physical and chemical processes, the best predictive models are onlyapproximate, such that experimental determination provides the best source of data.

The preferred definition of burning velocity refers to an infinite one-dimensional flame front. In reality, this cannot be achieved; the burner itself acts as a heat sink and the flamefront is very often curved (cf. that on a Bunsen burner). Both of these factors can affectpropagation through the fuel/air mixture. Consequently, experimentally determinedburning velocities are dependent on the method used. Andrews and Bradley 8 have reviewed the variation in measured burning velocities for methane/air mixtures and havederived correction factors for each experimental method. They currently recommend amaximum burning velocity for methane/air of 0.45 m s−1 at 298 K and 1 atm pressure.

1.3.4 Flame stability

From the foregoing sections, it will be apparent that the processes within a stationaryflame are in dynamic equilibrium. Propagation of the flame front upstream through theunburnt mixture is balanced by the downstream flow of unburnt gases, such that theflame appears stationary. More precisely (see Figure 1.4), at any point in the flame front, the local burning velocity equals the component of the stream velocity normal to theflame front. If either the flow rate or the burning velocity changes slightly, the shape ofthe flame front varies in order to restore the balance. This leads to the concept of flamestability, which is discussed in much greater detail in later chapters, and is of vitalimportance to burner designers. If the flow rate is too high or the burning velocity toolow, blow-off (flame lift) occurs whereby the flame comes away from the burner andextinguishes. Alternatively, if the flow rate is too low, or the burning velocity too high,lightback may occur, whereby the flame propagates upstream into the burner tube. Eithersituation can result in unsatisfactory appliance performance and may be hazardous. It isthe task of the burner designer to ensure that the gas/air flow velocity and burningvelocity are correctly matched, in order that safe operation of a gas appliance is achievedboth under normal running conditions and when variations in gas and air supply ortolerances in manufacture cause a departure from the design point.


Figure 1.4 Relationship between unburnt mixture flow and burning velocity.

1.4 Gas characteristics

Much of the experimental data on burning velocity, flammability, etc, presented above isderived from work involving a single fuel (e.g. methane) and an oxidant (air). Thedomestic gas appliance engineer requires information on mixtures of gaseous fuels, sincenatural gas, its manufactured substitutes and town gases are all gas mixtures. This sectionconsiders some of the properties of gaseous fuel mixtures and introduces the concept ofgas interchangeability.

1.4.1 Basic flame chemistry

In purely chemical terms, any understanding of flames must account for the transitionfrom fuel molecules to combustion products. A full chemical reaction scheme is beyondthe scope of this book, but it suffices to say that a very large number of molecularfragments are generated in the reaction zone. The complexity of the combustion ofmethane, arguably the simplest hydrocarbon, is shown by the work of Westbrook 9 , who used 75 separate chemical reactions in his computer model. However, the overall trend,regardless of fuel composition, is a gradual oxidation and dehydrogenation as thereactions proceed, yielding finally H2O, CO2 and a little CO and H2.

For stoichiometric combustion of methane in air, we have the following overall chemical equation (assuming dry air containing 20.95% O2 7 ):


where “N2” represents oxygen-free air (inerts). A general equation for hydrocarboncombustion can be similarly derived:

The theoretical air requirement (TAR) is defined as the number of volumes of airrequired for stoichiometric combustion per volume of fuel, i.e. in this case 4.77 (x+ y/

4). Primary aeration is the fraction of TAR supplied as primary air, expressed as apercentage.

In order to ensure completeness of combustion, practical systems employ air flow rates greater than would be required for stoichiometric combustion. A frequently usedparameter is excess air, which may be defined as the amount by which the total (i.e.primary and secondary) air flow exceeds the TAR. It is normally expressed as apercentage, i.e. 100% excess air indicates an air flow rate of double the TAR.

1.4.2 Minor combustion products

The above stoichiometric combustion equations represent an ideal case. In practice, thereare other important by-products which are formed in small quantities. Three suchproducts of particular interest to gas engineers are carbon monoxide (CO), nitrogenoxides (NO x ) and sulphur oxides (SO x ).

Sulphur emissions caused concern in Great Britain when sulphur-containing town gases were distributed. Current British natural gases are virtually sulphur-free, so SO xemissions are now negligible. Carbon monoxide formation is essentially a sign ofincomplete combustion due to lack of oxygen or low flame temperature. However, tracesof CO will be produced even at high excess air levels. Maximum permissible COemissions are set by appliance safety standards. Nitrogen oxides (chiefly NO and NO2) are formed in trace amounts by the partial oxidation of atmospheric nitrogen entrained inboth primary and secondary combustion air. NO x formation is generally favoured by hightemperatures 3 and low primary aeration (see Chapter 6).

All three oxides are undesirable in that they are pollutants and may be hazardous tohealth when present in sufficient quantities. It must be stressed, however, that such ahazard does not arise with gas appliances that are designed, installed and operatedcorrectly.

1.4.3 Properties of gas mixtures

Domestic gas supplies, whether manufactured or natural gas, are multi-component

(1.1)

(1.2)


mixtures liable to variation in composition, rather than the single components hithertodescribed. Clearly an appliance designer needs to know the properties of the particularmixture(s) that he wishes or is required to use.

The density, calorific value and theoretical air requirement of a mixture can be derived from the equation:

where P is the parameter for the mixture, x i the mole fraction of component i, and P i the value of the parameter for component i.

Of particular interest is the Wobbe number (W) defined as:

It should be noted that, because of the square root term, the Wobbe number cannot becalculated directly, i.e.

The significance of Wobbe number is made clearer by considering the fluid mechanics ofa gas stream emerging from a tube or nozzle. Volume flow ( ) is proportional to gas pressure (p), relative density (σ) and orifice area (A) according to the equation:

Thus, for a particular orifice at constant supply pressure, heat input is directlyproportional to Wobbe number. Consequently, Wobbe number is a better measure thancalorific value of heat input to the appliance since it indicates the effect of gascomposition changes on appliance heat input with a constant pressure gas supply and isespecially useful in comparing gaseous fuel mixtures.

In Great Britain, it became customary to use an empirical flame speed factor for gas mixtures rather than using burning velocity. The Weaver flame speed factor relates the

(1.3)

(1.4)

(1.5)

(1.6)

(1.7)


maximum burning velocity of the fuel to its TAR, and is defined as the burning velocityof a stoichiometric fuel/air mixture expressed as a percentage of the burning velocity ofthe same mixture of hydrogen/air. Weaver 10 derived the mathematical expression:

where F i is the Weaver coefficient for component i, while x O , x N , x i are the mole fractions of O2, inerts and component i respectively.

Table 1.2 gives values of density, calorific value, Wobbe number and Weavercoefficient for a number of gases which may be encountered in gaseous mixtures.

1.4.4 Classification of gases and appliances

In recognition of the wide variation in distributed gases, the International Gas Union

(1.8)

Table 1.2 Physical properties of selected gases (assumed ideal) at 15 °C, 1 atm pressure, (Sources: Rose and Cooper 7 , Weaver 10 , Gilbert and Prigg 11 (subject to rounding errors).)

Gas Density /kg m−3

Relative density (air=1)

Gross calorific

value /MJ m−3

Wobbe number /MJ

m−3

Weaver coefficient

TAR

Hydrogen 0.085 0.070 12.10 45.86 339 2.38

Methane 0.679 0.554 37.71 50.68 148 9.52

Ethane 1.272 1.038 66.07 64.86 301 16.67

Propane 1.865 1.522 93.94 76.15 398 23.81

Butane 2.458 2.006 121.80 86.00 513 30.95

Ethylene 1.186 0.968 59.72 60.69 454 14.29

Propylene 1.780 1.452 87.09 72.27 674 21.43

Butylene 2.373 1.936 114.62 82.37 890 28.57

Carbon monoxide

1.185 0.967 11.97 12.17 61 2.38

Carbon dioxide

1.861 1.519 0 0 0 0

Nitrogen 1.185 0.967 0 0 0 0


(IGU) has proposed 12 a classification scheme based on Wobbe number, whereby gasesare divided into three families with one or more groups per family:

1. First family gases. Gases, often town gas or manufactured gas, having a Wobbenumber in the range 22.6 to 29.8 MJ m−3 * .

2. Second family gases. Gases, often natural gas, having a Wobbe number in the range39.1 to 55.0 MJ m−3, and further divided into two groups, viz. Group H (45.7 to 55.0 MJ m−3) and Group L (39.1 to 45.0 MJ m−3).

3. Third family gases. Gases, usually liquefied petroleum gas (LPG), having a Wobbe number in the range 73.4 to 87.6 MJ m−3, and divided, in practice, into butane-based mixtures and propane-based mixtures.

This classification of gases led to the IGU proposal (now adopted in European Standards) to classify appliances according to the gas families which they are designed touse. Appliances are labelled Category I, II or III, depending on whether they are intendedto burn gases from a single family, two families or all three families. Category I and IIappliances are further defined, with a subscript, according to which particular gas familyor group

*First family gases are further subdivided into three groups; these are not discussed here, but are described in references 12 and 13.

may be used. The IGU 12 provides fuller details of appliance adaptations allowed or required in order to change from one gas group to another. This classification is nowincluded in harmonized standards produced by CEN (European Standards Organization).

Table 1.3 Composition manufactured gases. (Sources: Gilbert and Prigg 11 , Prigg and Rooke 14 .)

Source H2 CO CH4 C2H4 CO2 N2 Others

Debenzolized coal (thermal decomposition) 53 8 30 3 2 4 trace

Onia-Gegi (fuel oil+steam) 48 23 17 – 6 2 4

Lurgi (coal+O2) 61 23 13 – 2 1 trace

Carburetted water gas (CO reduced) 46 4 23 5 16 5 1

Natural gas (continuous reforming) 42 8 33 – 6 11 trace

G4 Reference gas 51 15 19 2 4 6 3


1.4.5 Interchangeability of gases

The composition of both manufactured gases and world resources of natural gas variesgreatly. Table 1.3 shows a selection of manufactured gas compositions from different production processes that were distributed in Great Britain prior to conversion. Table 1.4illustrates the widely differing composition of natural gas reserves. Any gas not suitablefor distribution needs to be treated in such a way that it is brought within specification forsatisfactory combustion; otherwise appliance performance may deteriorate in one of thefollowing ways:

1. The heat input may be too high or too low. 2. Aerated burners may light back. 3. Flames may lift from non-aerated and aerated burners. 4. Appliances may give incomplete combustion (i.e. emission of carbon monoxide and/or

formation of soot).

Each of these conditions can result in impaired appliance performance and dissatisfiedusers.

In practice, it is not economically viable to treat all gases in such a way that a single composition gas is distributed. Provided that the appliance performance remainssatisfactory with regard to the four considerations above, gases of differing compositionsmay be distributed without cause for concern. Such alternative gases are said to beinterchangeable.

1.4.6 Prediction of interchangeability

The only entirely satisfactory method of testing the suitability of a gas mixture fordistribution is to burn the gas on a range of appliances and compare with applianceperformance using a reference gas. This is, however, a costly and lengthy procedure, somethods of prediction have been devised for both first and second family gases. Only avery brief review is given here.

Table 1.4 Composition (vol %) of selected natural gas deposits

Source CH4 C2H6 C3H8 N2 Others

Southern Basin 94 3 1 1 1

Brent 77 10 6 ½ 6½

Forties 35 39 25 trace 1

Morecambe 85 4 1 8 2

Groningen (Holland) 81 3 trace 14 2


Figure 1.5 Limits for satisfactory operation for G4-type manufactured gas. (After Gilbert and Prigg 11 .)

Gilbert and Prigg 11 proposed a plot of Wobbe number against Weaver flame speed factor for first family gases (Figure 1.5). On the diagram were plotted numerous data points (corresponding to each gas composition) and experimentally determined limits ofacceptability with regard to flame stability and incomplete combustion. This results in a region on the prediction diagram within which all gas compositions give satisfactoryappliance performance. Consequently, it is possible to replace the normal gas with analternative supply of somewhat different composition but similar burning properties.

Prediction of interchangeability was of great importance when using manufactured first family gases, because compositions could vary greatly (Table 1.3). The first supplies of natural gas from the Southern Basin of the North Sea were all of very similarcomposition (approximately 94% CH4, 3% C2H6, 1% N2 +other traces). Thus, appliances were designed to burn that particular composition, although it was recognized that futuresupplies might be different and that substitute natural gas (SNG) of various compositionsmight also be required as natural gas reserves were depleted. Subsequent discoveries ofhydrocarbon-rich gases (e.g. Brent) as well as lean gases (e.g. Morecambe) have shownthat prediction of interchangeability is a necessary tool for present-day gas treatment and distribution. In the future, prediction of gas properties will still be required, sincemanufactured SNGs will almost certainly consist of a complexhydrocarbon/hydrogen/inert mixture with burning properties similar to methane.


Figure 1.6 Limits of operation for family 2H natural gas. (After lessen et al 16 .)

Figure 1.7 The three-dimensional prediction diagram for interchangeability of family 2H gases.(After Dutton 19 .)


Initial work 15 , 16 in Great Britain on the prediction of the combustion characteristicsof second family gases used the same plot as for first family gases, that is Wobbe numberagainst flame speed factor (Figure 1.6). However, because burning velocity is mostaffected by hydrogen content and natural gases contain no hydrogen, the abscissabecomes very insensitive to changes in composition. The approach now followed 17 – 19

is based on gas composition. It is possible to represent any gas mixture by a four-component equivalent based on methane, propane, nitrogen and hydrogen. The three-dimensional diagram currently used (Figure 1.7) plots Wobbe number against the sum ofpropane+nitrogen against hydrogen concentration. A volume of acceptability can bedefined within which all gases are interchangeable and suitable for distribution. With thismodel, it is now possible to predict, with greater accuracy than ever before, the likelycombustion characteristics of any hydrocarbon-based mixture, and what treatments, if any, are needed in order to render the mixture suitable for distribution.

1.4.7 Test gases

In order to assess the effect of changes in gas composition, appliance gases performanceis checked during British Standard Institution certification against a number of gaseswhich represent limiting compositions of gas which could be distributed for short periodsunder emergency conditions.

The International Gas Union 12 recommended a series of test gases of fixed compositionfor each of the three gas families. As well as a reference gas, each family has associatedlimit gases, which are intended to test the susceptibility of appliance burners to sooting,incomplete combustion, lightback and flame lift. Table 1.5 lists the compositions and

Table 1.5 Current family 2H test gases used in Great Britain.

Designation Composition (vol %) Calorific value (dry) Relative

Wobbe number

Test gas GB IGU CH4 C3H8 H2 N2 /MJ m−3 density /MJ m−3

Reference NGA G20 100 – – – 37.71 0.554 50.68

Incomplete combustion

NGB G21 87 13 – – 45.02 0.680 54.61

Lightback NGC G22 65 – 35 – 28.75 0.384 46.37

Lift NGD G23 92.5 – – 7.5 34.88 0.585 45.62

Burn-back sooting

NGC2 G24 68 12 20 – 39.34 0.573 51.96

Note: Values of calorific value and relative density calculated using Table 1.2 and equation (1.1) (subject to rounding errors); gases assumed ideal.


properties of the family 2H test gases currently used in Great Britain.


CHAPTER 2 Non-aerated burners

2.1 Historical background

Until Bunsen’s invention of the aerated burner in 1855, gas was used only for lighting,since the luminous, non-aerated flame provided illumination that was far superior to candlelight. Further development of Bunsen’s principle, together with introduction of theincandescent gas mantle with its more compact, soot-free flames, helped establish gas as a major fuel for both lighting and heating in Victorian Britain.

The aerated burner had, however, several disadvantages. Lightback to the injector was encountered when gas suppliers did not maintain the correct gas composition. Burnerswere also noisy and susceptible to blockage (linting) by airborne dust. As a result of thesedifficulties, non-aerated burners were designed with well-defined, virtually soot-free flames, and from 1945 the majority of British domestic gas appliances (except cookers)were fitted with non-aerated burners. The development of quiet, clean, trouble-free appliances (particularly gas fires) led to a large increase in the domestic use of gas.

The advent of natural gas necessitated a reversion to the aerated burner, because of the inherent problems of flame instability and sooting with non-aerated hydrocarbon burners. The only non-aerated flames currently used in British domestic appliances are somepilots and the decorative flames in some fuel-effect gas fires. However, this chapter isincluded in recognition of the great historical importance of non-aerated burners.

2.1 Historical background 2.2 Fundamental aspects 2.2.1 General structure 2.2.2 Mathematical theories 2.2.3 Diffusion flame stability 2.3 Non-aerated burners for natural gas 2.3.1 Wedge-cavity burners 2.3.2 Pinhole burner 2.3.3 Matrix burners 2.3.4 Other burners 2.3.3 Matrix burner 2.3.4 Other burners

2.2 Fundamental aspects

Despite the historical and practical importance of non-aerated burners, systematic work on the fundamental physical and chemical processes occurring in non-aerated flames has emerged only relatively recently. Early chemical predictive models were often inaccurateowing to the need to consider both kinetics and diffusion. Mathematical treatments of thephysical properties proved difficult because, unlike many aerated flames, it was notusually possible to reduce a diffusion flame to a one-dimensional problem. Recent developments in mathematical and computational techniques have now permittedsolution of more complex problems (see, for instance, the review by Williams 5 ).

For gas appliance engineers, prediction of flame height is of particular importance, since it is that parameter that ultimately dictates the selection of combustion chambersize, and needs to be taken into account in assessing the interchangeability of gases.

2.2.1 General structure

Qualitatively, the physical structure of a diffusion flame is fairly well understood. Thefuel issues from the burner port and combustion occurs as oxygen diffuses towards theflame axis from the surroundings. It is usually assumed that gas concentration is zerodownstream from the flame front, while oxygen concentration is zero upstream from theflame front. Figure 2.1 shows schematically concentration profiles of fuel, oxidant andcombustion products for a radial cross-section of a typical diffusion flame.

The most obvious visual property of many hydrocarbon/air diffusion flames is their propensity to soot. Soot is observed as a yellow emission from the centre of the flame andis formed by local pyrolysis of the fuel in the oxygen-free flame core. If sufficient oxygen can diffuse to the flame core, the soot is subsequently burned in the tail of the flame,whereas if the oxygen supply is too low, soot will not be oxidized, in which case a sootystream will be evident. In a gas appliance, such soot release cannot be tolerated, sincedeposition could lead to gradual blockage of the heat exchanger and flue. The designermust therefore ensure a sufficient supply of air in order to maintain satisfactorycombustion.

Non-aerated burners 19

Figure 2.1 Schematic concentration profiles of fuel, oxygen and combustion products for a radial cross-section of an axisymmetric diffusion flame.

2.2.2 Mathematical theories

The first theoretical treatment for the prediction of diffusion flame height and shape wasgiven in 1928 by Burke and Schumann 20 . They considered the case of one tube setinside another, such that fuel flowing in the inner tube comes into contact with airflowing concurrently in the outer tube. A number of assumptions were made, in particularthat the air and gas velocities are constant and parallel to the flame axis, that there is asingle very fast chemical reaction, that temperature and diffusivity are constant, that axialdiffusion may be neglected, and that the flame ends at the same point as the visible flame.It can be shown that, for a circular flame port, the problem reduces to a differentialequation of the form:

where r and y are the radial and axial co-ordinates, c ry is the concentration of the flame gases at (r,y), and D is an overall diffusion coefficient. Subsequent workers (e.g. references 21–23, reviewed by Gaydon and Wolfhard 3 ) have extended the original work to laminar and turbulent flames burning in free air by employing empirical or semi-empirical relationships which attempt to compensate for the variation in flametemperature and buoyancy.

(2.1)


The Burke—Schumann approach was developed for circular port burners and tested extensively only for that geometry. Roper and co-workers 24 − 26 found large errors when predicting flame heights for other geometries, such as were used in gas appliances.Consequently, a more rigorous mathematical model was developed, which redefined theend of the flame as that point on the flame axis where the fuel and oxidant are instoichiometric proportions. The model also allowed for the natural buoyancy of the hotflame gases. If the effect of buoyancy is to increase mass flow rate, then continuityrequires that the flow streamlines converge towards the axis. The radial concentrationgradient therefore increases, thereby increasing the rate of diffusion. Roper derivedequations 24 − 26 which were subsequently experimentally verified 25 , 26 for square and circular port, straight slot and curved slot burners. The equations are valid not only fornon-aerated flames, but also for the outer diffusion-controlled mantle of partially aerated flames. Only a brief summary is given here; readers requiring further information shouldconsult the original papers.

Roper 24 showed that circular and square port burners are special cases, in that flame height is independent of axial flow velocity and therefore independent of buoyancyforces. This would explain the apparent success of the Burke—Schumann theory for this geometry. It seems that for circular and square ports, the interaction outlined abovebetween mass flow rate and diffusion rate is such that the effect of any change in axialvelocity is counterbalanced by a change in diffusion time. Flame height thereby remainsunaltered. For slot burners, equations for flame height do show a dependence on axialvelocity. Roper 24 , 26 derived a number of equations depending on whether axial velocityis momentum-controlled or buoyancy-controlled or in a transition region where neithereffect can be ignored.

2.2.3 Diffusion flame stability

It is generally assumed that combustion is diffusion-controlled throughout a diffusion flame. There is, however, considerable evidence to suggest that some premixing ofoxygen and fuel occurs at the base of the flame near the burner port This can be causedby turbulence at the gas/air boundary where the gas stream emerges from the burner 27 . Indeed, a thin reaction zone is very often visible at the base of a diffusion flame, and ithas been shown 28 that if nitrogen is passed around the burner port, then it becomes almost impossible to stabilize the flame. Thus, some premixing appears essential for astable diffusion flame. Perhaps as a consequence, many authors ascribe the problems ofnon-aerated methane flame stability to the low burning velocity of hydrocarbonscompared with hydrogen, although a pure fuel is outside its flammability limits, andtherefore possesses no burning velocity. Edmondson 29 has suggested that the criteria for diffusion flame stability are essentially similar to aerated flames. The effect of the widelydifferent rates of diffusion of hydrogen and hydrocarbons only become important furtherdownstream.


2.3 Non-aerated burners for natural gas

Because reliable predictive mathematical models are a relatively recent development,most non-aerated burners, whether burning town gas or natural gas, were designed using educated guesswork rather than any firm theoretical basis. For example, it had beenrealized for several years that a circular port produces a cylindrical flame with a relativelylow surface area to volume ratio, whereas use of a thin slot burner produces a very thin,essentially flat flame with a high surface area to volume ratio. Use of a slot burner,therefore, increases the rate of molecular diffusion into the flame and the rate ofcombustion, such that the flame is relatively compact and almost non-luminous (soot-free). Burners using such fan-shaped flames became extremely widespread with town gas appliances.

Attempts to produce a non-aerated natural gas burner comparable to existing town gasdesigns have generally failed because of the difficulty of accommodating the combustioncharacteristics of methane and, more especially, SNG mixtures. This section reviews anumber of designs which were developed and tested. The main reasons for lack ofsuccess, viz., flame instability and sooting, are also highlighted.

Figure 2.2 The Bray wedge-cavity burner, showing gas flowlines and the position of the flame. 7(After Edmondson 29 .)


2.3.1 Wedge-cavity burners

The easiest method of producing a flat flame is by use of a thin slit rather than a roundhole. In practice, it is difficult to maintain the very tight production tolerances requiredfor slits much less than 1 mm in width. The most successful flat flame burner wasdeveloped by Bray; gas passes through a suitably shaped cavity on the underside of anorifice plate which contains a single hole. The principle of operation has been describedby Minchin 30 . The cavity (see Figure 2.2) has two mutually perpendicular planes of symmetry and is shaped such that gas can approach the orifice in one plane from anydirection, but is constrained between parallel walls in the other plane. The resultantdynamic forces produce a fan-shaped flame, the plane of which is perpendicular to theaxis of the cavity. Flame shape is largely dependent on the shape of the cavity rather thanthat of the orifice.

Figure 2.3 Stabilization of a town gas flame on a Bray wedge-cavity burner. (After Edmondson 29 .)


Prigg and Shah Jahan 31 considered fan shape and flame lift from Bray jets at a detailed, but empirical level. Edmondson 29 subsequently gave a simple qualitativeexplanation of Bray jet flame stability based on aerodynamic theory 32 , which predicts that the greater the flow rate from a given orifice, the thinner the resultant gas jet. Figure 2.3 shows schematically the effect of increasing the gas flow on a town gas Bray jet, viewing the fan-shaped flame edge-on. At very low flow rates (Stage 1), the flame is wider than the orifice length, and is stabilized by interdiffusion of gas and air, whichcreates a flammable mixture at the base of the flame. As gas flow rate increases andflame thickness decreases (Stages 2 and 3), the flame forms at a more central and, due tothe orifice shape, a lower position. Mixing occurs here by entrainment at the flame baseas well as diffusion, so that stability is achieved by a balance between gas flow rate andmixture burning velocity. The entrained air flows parallel to the jet structure, such that atStage 3, a counter-current gas/air flow is set up which ensures good mixing and excellentflame stability. Only at very large gas flow rates is instability observed. Experiments withmethane as a fuel gas fail to show any reverse air flow. Because of the different transportproperties involved, a methane flame is thicker and is positioned somewhat further fromthe orifice, such that the necessary stabilizing reverse air flow is never established. Figure 2.4 shows this effect quite clearly with a plot of flame base position against flame thickness for methane and a town gas at a number of volume flow rates.

Figure 2.4 Position of methane and a town gas flame on a Bray wedge-cavity burner.Volume flow rate is given in cm3 s−1 next to each data point. (After Edmondson 29 .)


Van der Linden 33 achieved some success in stabilizing a Bray jet burning methane byadding six small retention flames to the top of the original jet (Figure 2.5). The retention flames burn about 20% of the total gas flow and are stable because of the low flow ratecompared with the main port. They then act as a permanent ignition source for gasemerging from the main port. The resultant flame is thicker and less well defined than atown gas flame. While satisfactory operation was obtained when burning methane in stillair, the burner was found to be intolerant of cross-draughts and changes in gas composition 34 . Subsequent work therefore concentrated on new designs.

Figure 2.5 A stabilized Bray wedge-cavity burner for natural gases. (After Culshaw and Prigg 39 .)

2.3.2 Pinhole burners

Because of the difficulties in producing a stabilized Bray jet for natural gas, workersconcentrated on developing new designs which would be more suited to lowburningvelocity mixtures. The first successful new method for stabilizing a methane flame wasdeveloped in Holland at the VEG Gas Instituut 35 . The pinhole burner, an example of which is shown in Figure 2.6, consists of a row of pinholes providing the main flames,surrounded by larger holes for auxiliary retention flames. Baffles within the burnerensure that the retention ports are supplied with gas at a lower pressure and hence lowergas velocity than the main ports.


Figure 2.6 A pinhole burner with retention ports for natural gases. (After Harris and Wilson 36 .)

Pinhole burners were found to operate satisfactorily using distributed British natural gas 34 , although there were engineering problems with regard to pinhole tolerances andthe distribution of retention gas within the burner. Major problems were encounteredwhen hydrogen-containing SNGs were burned. Although a non-aerated burner will not light back, the flames can burn very close to the surface, resulting in excessive burnertemperatures when the fuel gas contains a significant proportion of hydrogen. Such high temperatures lead both to burner overheat and to pyrolysis of the emerging fuel, resultingin deposition of soot (known as burn-back sooting) round the burner ports. Figure 2.7shows a severe case of near total blockage. It was therefore recommended 36 that pinhole burners should not be used to burn gases with more than 10% hydrogen content.


Figure 2.7 An example of the build-up of soot on a pinhole burner.

2.3.3 Matrix burner

The matrix burner 37 was developed by British Petroleum for the combustion of wastehydrocarbon gases, but was found to be capable of burning virtually any gas, includingboth town and natural gases, as a diffusion flame. After that initial work many prototypes were tested, but all operate on the same principle (see Figure 2.8). Fuel gas discharges from narrow slots or ports adjacent to much larger ports which supply all the air requiredfor complete combustion. The flames produced are generally compact, stable and semi-lifted, forming a matrix of blue cones about 1 mm downstream from the plane of theburner. Consequently, the matrix burner combines the advantages of a fully aeratedburner (short flames, compact combustion chamber) with the advantages of non-aerated operation (no lightback, linting, or noise during use). It is probable that interactionsbetween neighbouring cones help stabilize the system. In order to maintain the correct airflow and adequate mixing of air and gas, air is usually supplied by use of a small fan.Westwood and South 38 have considered the effect of variation in air flow rate, heat inputand gas composition on combustion using a matrix burner. Three regimes were identifiedwhere, in terms of CO emission, combustion became unsatisfactory (CO/CO2 > 0.02): at low air levels owing to lack of air; at high air levels owing to quenching of thecombustion reactions; at low gas input because of partial flame extinction. Figure 2.9summarizes their results and also includes a shaded area within which soot depositionand burner overheat occur with an SNG containing 20% hydrogen. With these highburning velocity mixtures, the flames stabilize closer to the burner. Two effects result:firstly, the burner can overheat causing permanent damage; secondly, there is a reductionin intermixing time for gas and air resulting in fuel-rich combustion and soot deposition in the heat exchanger. The area of satisfactory operation of a matrix burner is therefore


Figure 2.8 The principle of operation of a matrix burner. (After Harris and Wilson 36 .)

Figure 2.9 Combustion diagram for a matrix burner operating with natural gas. (After Harris and Wilson 36 and Westwood and South38.)


considerably reduced. Another major drawback concerns the very tight tolerances inmanufacture and assembly of the burner. Unless the gas ports are drilled with greatprecision and a very uniform air distribution is maintained, it is difficult to ensure acompact flame in the correct position. These two problems of combustion andmanufacture have diverted attention away from further developments in matrix burnerdesign.

2.3.4 Other burners

A number of other designs have also been developed 39 , all of which suffered from problems of manufacture, intolerance of hydrogen-containing gases or poor flame stability. Two of particular interest are considered here.

In the Uniplane burner (Figure 2.10), gas is discharged through a slit at the base of awall, at the top of which the flame is stabilized and shielded. It was found that withhydrogen-containing gases, the flame stabilized at the slit rather than at the top of the wall, such that air entrainment was reduced and the flame became sooty. A retention gridwas therefore fitted which successfully anchored the flame above the slit. Despite thesuccess in stabilizing a variety of gases, the design ultimately proved unsatisfactorybecause of an inability to guarantee performance on production burners.

A ‘rod-stabilized’ burner was developed (Figure 2.11) whereby a metal rod was placed just above the burner. Downstream from the rod, frictional drag induces the formation ofeddies in which the flame can be partially stabilized. This enables the rod to heat upsufficiently to act as a hot bluff body to which the flame can anchor itself. The majordisadvantage with the burner was that stability was achieved only when the rod was hot;partial flame lift was evident for a considerable period after ignition. Soot deposition wasalso observed on the rod when hydrocarbon-rich gases were burned.

Figure 2.10 The Uniplane non-aerated natural gas burner with and without shield. (After Culshaw and Prigg 39 .)


Figure 2.11 The rod-stabilized non-aerated burner. (After Culshaw and Prigg 39 .)


CHAPTER 3 Partially aerated burners

3.1 Introduction 3.2 Injectors 3.2.1 Discharge from an orifice 3.2.2 Coefficient of discharge 3.2.3 Injector ignition noise 3.3 Air entrainment 3.3.1 Physical mechanism 3.3.2 Mathematical theory—background 3.3.3 The basic air entrainment equations 3.3.4 Mixing tube design 3.3.5 Final entrainment equations 3.4 Flame port design/flame stability 3.4.1 The combustion diagram 3.4.2 Flame lift 3.4.3 Lightback 3.4.4 Yellow tipping and incomplete combustion 3.5 External design variables 3.5.1 Effect of gas composition 3.5.2 Manufacturing tolerances 3.5.3 Vitiation 3.5.4 Draughts 3.5.5 Linting 3.6 Summary procedure for burner design

3.1 Introduction

The previous chapter has described the problems associated with non-aerated burners for natural gas in terms of poor flame stability, intolerance to SNG with a high hydrogencontent, and difficulties with maintaining the very tight manufacturing tolerances in massproduction. Conversion to natural gas therefore resulted in a large-scale reintroduction of the partially aerated burner for all domestic gas appliances, and included the replacementof existing town gas non-aerated burners with aerated methane burners. Such changes were, however, anticipated; fundamental work had been undertaken by the gas industryfrom the mid-1950s onwards on both the theoretical and the practical aspects of aeratedburner design.

Figure 3.1 The general layout of an atmospheric injection aerated burner.

Figure 3.1 shows schematically the basic design of a typical aerated burner. Gas emerges from an injector nozzle consisting of one or more small holes. On leaving the injector, thegas entrains primary air by a momentum-sharing process between the emerging gas and ambient air. The gas/air mixture enters a mixing tube, which may be shaped in the formof a tapered venturi or may have parallel sides. As its name suggests, the mixing tube isdesigned to ensure thorough mixing of gas and air, such that a constant air/gas ratio ismaintained throughout the burner head; the mixture must then be distributed uniformly tothe burner ports. Optional features include an aeration shutter or mixing tube restrictorwhich controls primary air entrainment, and baffles or gauzes within the burner bodywhich aid good mixing and prevent lint from clogging the burner ports. Examples ofburners (described further in Chapter 6) include drilled bar burners (such as those used incooker ovens and grills), box burners (boilers and space heaters), circular cooker hotplateburners, jetted burners consisting of an array of small individual aerated burner jetsscrewed into a manifold, and a number of other variants.


This chapter considers the fundamental aspects of the four main components of the aerated burner, that is the injector, air entrainment, mixing tube design and burner portgeometry. The final section examines the effect of such external parameters as linting,draughts and vitiation on burner performance. Much of the theory is equally applicable toboth partially and fully aerated burners. Practical examples and problems of the latter arediscussed in Chapter 4.

3.2 Injectors

An injector needs to be carefully designed and positioned in order to ensure the correctsupply of gas and air to the burner. The size and shape of the injector orifice control thegas flow rate and hence heat input for a given gas composition and supply pressure. Thesiting of the injector with respect to the mixing tube affects air entrainment, so theinjector must be positioned with a high degree of precision during manufacture of aburner or assembly of an appliance. This section covers various fundamental andpractical aspects of injector design.

3.2.1 Discharge an orifice

In physical terms, an injector serves to convert the potential energy of a from highpressure gas supply into the kinetic energy of an emerging gas jet. Mathematically,therefore, by conservation of energy and assuming no losses at the nozzle, we have (perunit mass):

where is the volume flow rate from the orifice (in m3 s−1), A j is the orifice area (m2), gis the acceleration due to gravity, and h is the height (m) of a column of gas required toexert the gas pressure at the orifice. The gauge pressure, p (Pa), is given by hρ g g, where ρ g is the gas density (kg m−3). Equation (3.1) becomes:

where ρa is the density of air (1.225 kg m−3 at standard conditions7), and σ is the gas relative density. With more conventional units ( in m3 s−1, A j in mm2, p in mbar), the expression reduces to:

(3.1)

(3.2)

Partially aerated burners 33

The heat flow q=CV , where CV is the gas calorific value, so the theoretical heat flowfrom the orifice (in watts) is given by:

where W is the Wobbe number in MJ m−3. This forms the basic discharge flow equationfrom any orifice, and may be used both for the injector and the flame ports.

3.2.2 Coefficient of discharge

In practice, the flow of gas from an orifice is less than predicted because of frictionallosses and the vena-contracta effect, although the latter should be negligible in a well-designed, tapered nozzle. It is usual to represent these two terms as a coefficient ofdischarge, c d such that

Figure 3.2 Discharge coefficients for a number of injector geometries. (After Prigg 40 .)

The discharge coefficient is dependent on the orifice shape and on the Reynolds numberfor flow through the orifice.

(3.3)

(3.4)

(3.5)


1. Effect of orifice shape. If the orifice tube is too long compared with its diameter, energy loss due to friction is high; if the tube is too short, energy is lost by the suddencontraction and expansion of the gas passing through. Losses can, however, be reduced ifthe entrance to the orifice is tapered. Figure 3.2 gives values of c d for a range of orifice dimensions. Highest values

Figure 3.3 The effect of Reynolds number on the discharge coefficient from a typical injector orifice and flame ports. (After Prigg 40 .)

(> 0.9) are achieved 40 when the orifice length is about the same as the diameter andwhen the angle of approach is about 35°.

2. Effect of Reynolds number. The discharge coefficient is dependent on flow conditions in the orifice as expressed in terms of the Reynolds number (Re)*. Figure 3.3shows that for a typical orifice with laminar flow (Re < 2000), c d is strongly dependent on Re, whereas for turbulent flow, c d may be considered to be virtually constant. Using gas pressures of several millibars, single-hole injectors will normally operate in the turbulent flow regime, such that any variation in discharge coefficient can usually beignored 41 . If a multihole injector with the same total cross-sectional area is used, then the Reynolds number may be reduced sufficiently, owing to the decrease in diameter, forflow to become laminar within the injector 42 − 43 . In the latter case, special consideration needs to be given to the discharge coefficient of the injector.


3.2.3 Injector ignition noise

High frequency noise (>3 kHz) in aerated burners originates almost entirely in theinjector and is due to shear forces between the high velocity gas jet and the surroundingair. Noise originating at the injector may be

*Reynolds number is a measure of the relative magnitude of the inertial and viscous forces in a fluid flow. For flow in a cylindrical pipe, it is defined as where ρ is the density, v the flow velocity, d the pipe diameter and µ is the dynamic viscosity.

amplified by the flame (pyro-acoustic amplification) during operation of the appliance. Dance and Sutherland 42 found that injector noise levels are greater in turbulent flow thanlaminar flow, and concluded that multihole injectors with a sufficient number of holes toensure laminar flow conditions should be used. The leading edges of orifices should alsobe chamfered to prevent eddy formation within the injector. Such requirements for smalland smooth orifices are most easily met using ceramic materials.

3.3 Air entrainment

The mechanism of air entrainment has been studied experimentally and theoretically formany years, and is of vital interest to burner designers because the quantity of primary airtaken up has a considerable effect on burner port design requirements, flame stability,shape and temperature, and, ultimately, the design of the combustion chamber itself. Thissection gives a summary of various mathematical treatments applicable to the domesticaerated burner. For complete theoretical detail, the reader is advised to consult theoriginal references.

3.3.1 Physical mechanism

The basic air entrainment mechanism may be described by reference to Figure 3.1. The gas stream emerges as a free jet from the orifice at a rate dependent on gas pressure,orifice dimensions and gas composition (equation 3.5). Momentum transfer occursbetween the jet and the surrounding air, resulting in entrainment and expansion into theentrance of the mixing tube. It is normally assumed that, owing to turbulence, mixing isvirtually complete at the throat, at which point, depending on the precise geometry, thestatic pressure of the gas/air stream maybe reduced to almost that of the surroundingatmosphere. Further downstream, the static pressure must increase sufficiently toovercome resistances to flow within the burner and at the flame ports. A pressureincrease can only be induced by a decrease in momentum and velocity, and this can onlyarise from a gradual expansion within the mixing tube. It is normally achieved by asuitably shaped diffuser section in the burner. The gas/air mixture then flows into the


burner head prior to discharge from the flame ports.

3.3.2 Mathematical theory—background

Two basic techniques have been used in various attempts to treat air entrainmenttheoretically. Silver 44 , using an energy balance, obtained good agreement with experiment except at low flow rates, where not even qualitatively correct predictionscould be made. Waight 45 added to Silver’s equations a term incorporating energy lossduring expansion in the mixing tube, and was able to model successfully turbulent tunnelburners for furnaces. Von Elbe and Grumer 46 used a force-momentum equation which gave good agreement for all flow rates, but required the determination of empiricalcoefficients for each burner. The first comprehensive theory was presented by Simmonds47 , who combined elements of both approaches to obtain good correlation withexperiment at all flow rates without the need for empirical constants. Prigg 40 and Francis48 subsequently developed a simplified design procedure specifically for optimum performance. As such conditions are of most interest to gas appliance designers, a similarapproach has been adopted by Pritchard et al. 49 and is also used here. Readers requiringa more general approach should follow the work of Simmonds 47 .

3.3.3 The basic air entrainment equations

In deriving equations for air entrainment, a number of simplifying assumptions are made: 1. Flow is turbulent in the mixing tube such that there are no radial components of

velocity, temperature or composition. 2. Air is entrained perpendicular to the axis of the mixing tube, and therefore

contributes no axial momentum. This is not entirely true; however, Pritchard et al. 49

consider that the approximation is valid because any increase in momentum due to axialentrainment of air is effectively balanced by an increase in head loss and friction losses inthe throat.

3. Flow is incompressible, i.e. density is not dependent on pressure. This is a fairapproximation for the pressures used in domestic burners. The effect of temperature isalso ignored in the initial analysis, but is discussed in Section 3.3.3.6.

4. The pressure distribution at the burner head is uniform. In practice, it is not,especially in bar burners where there is a variation in flow and pressure along the axis ofthe burner as gas is discharged. Goodwin et al. 50 have developed a procedure for bar burners and some of their equations are presented in Section 3.3.3.7.

5. The combustion chamber is at atmospheric pressure, i.e. the effects of flue pull orfan pressure are ignored.

6. Buoyancy of the unburnt gas/air mixture is neglected. This is a fair assumptionexcept at low flow rates in vertical mixing tubes 47 .

3.3.3.1 Momentum and energy balances

The principle behind the derivation is the application of a momentum balance to the


mixing process and energy balances to the pressure increase behind the burner ports andto the discharges from the injector and from the burner ports. For the physical propertiesof the burner assembly, p, A and c d representing pressure, area and discharge coefficient,are used with subscripts o, j, t and p for ambient, injector outlet, throat and diffuser exitrespectively. For gas and air, ρ and V represent density and volume flow rate, whilesubscripts g, a and m, are used for gas, air and the mixture.

Neglecting stream contraction, a force-momentum balance for the mixing process between the injector outlet and the throat gives:

An energy balance between the throat and the diffuser outlet (i.e. just upstream of theburner ports) gives:

where C L is a friction loss coefficient for the throat and diffuser, expressed as a fraction of the kinetic head in the throat of the mixing tube.

Energy balances for discharge from the injector and the burner ports yield, respectively,

Combining (3.6) and (3.7) to eliminate p t

3.3.3.2 Injector/throat area ratio

Francis 48 introduces a dimensionless pressure efficiency η, defined as the ratio of the static pressure behind the burner ports to the dynamic pressure of the gas jet issuing fromthe injector, i.e.

(3.6)

(3.7)

(3.8)

(3.9)

(3.10)


which, on combining with equation (3.8), gives:

In order to simplify the algebra, two other parameters will be defined:

Therefore, equation (3.10) can be written in the form:

Optimum, performance for aerated burners is given by the maximum value of η, i.e. the greatest static pressure behind the burner ports for a particular gas rate from a giveninjector. Using elementary calculus,

So, η is a maximum when λθ=1, i.e. when

If R is the entrained air to gas volume ratio, then a=R g and m= g(1+R). Substitution into (3.13) gives:

(3.11)

(3.12)

(3.13)

(3.14)


From a mass balance, we obtain for the mixture density

where σ is the relative density of the gas. Substituting this expression into (3.14) gives:

This equation relates the optimum injector/throat area ratio to properties of the gas andgas/air mixture.

3.3.3.3 Throat/burner port area ratio

Taking equation (3.10) and substituting (3.15) and m= g (1+R), we have, after rearranging,

For an optimized system, we can substitute equation (3.16):

Also, taking (3.9) and performing the same substitution for ρm and m,

Equating (3.18) and (3.19) yields the expression:

This equation shows that the optimum throat to burner port area ratio is independent of

(3.15)

(3.16)

(3.17)

(3.18)

(3.19)

(3.20)


flow rate except for the dependence of both c dp and C L on Reynolds number (see Section 3.3.4). Note that if there were no losses (c dp =1; C L =0), then the optimum conditions would correspond to A t=A p.

3.3.3.4 Injector/burner port area ratio

By combining equations (3.16) and (3.20), it is clear that the optimum injector/burnerport area ratio is given by

Substitution of this into equation (3.5) gives a relation between burner port loading,air/gas ratio and injector pressure:

3.3.3.5 Air/gas ratio

Although, in practice, a designer would construct a burner to give a required primaryaeration (i.e. R is known), there is some merit in expressing R in terms of area ratios. Equations (3.16) and (3.21) both provide quadratic equations in R which can be solved. From (3.16):

Since (1–σ)2 is small compared with

(3.21)

(3.22)


we can approximate the solution to

or, substituting (3.20),

The equations predict that the primary aeration is independent of flow rate (i.e. gassupply pressure), except for any variation in c dp and C L . Figure 3.4 shows the variation of R with gas pressure for a burner designed for operation at 60% primary aeration. Other than at low flow rates, where c d p and C L are affected by the low Reynolds numbers, theexpected behaviour is observed. Equations (3.23) and (3.24) can be compared with theexpression derived by Prigg 40 for an ideal burner:

It can be seen that both (3.23) and (3.24) reduce to Prigg’s equation if c d p=1, C L =0 (i.e. no losses), and if the approximation

is used. This latter is a fair assumption for natural gas where σ ~ 0.6.

(3.23)

(3.24)

(3.25)


Figure 3.4 The effect of gas supply pressure at the injector on primary aeration for a burner with a design aeration of 60%. (After Goodwin et al. 50 .)

3.3.3.6 The effect of temperature

The analysis hitherto has assumed that the gas, air and gas/air mixture are all at standardconditions (15 °C, 1 atm). As a gas/air mixture passes through a hot burner assembly, the mixture temperature will rise, leading to an increase in volume, a decrease in density andan increased flow resistance through the flame ports. Prigg 40 derived a modified form of equation (3.25):

where T 0 and T 1 are the absolute temperatures of standard conditions and the air/gas

(3.26)


mixture respectively. The equation predicts a decrease in R as the burner heats up after ignition. This has been confirmed experimentally by Harris and Prigg 41 (see Figure 3.5), where a burner temperature of 350 °C to 400 °C was measured, corresponding to an air/gas temperature of 200 °C.

Figure 3.5 The effect of temperature on the entrainment of primary air for a typical partially aerated burner. (After Harris and Prigg 41 .)

3.3.3.7 Parallel-sided mixing tubes and bar burners

The mathematical analysis presented above assumed that a diffuser is used to aidexpansion and that the mixture pressure is identical at each burner port. Similarderivations are possible for parallel mixing tubes, where eddy formation increases frictionlosses, and for bar burners, where the ports are spaced along the supply tube and mixtureflow and pressure vary along the tube. Fortunately, a suitable correction factor to the losscoefficient is all that is required for these arrangements. Only the results are quoted here.

In a parallel mixing tube with no diffuser, Francis 48 suggests that the only adaptation required is replacement of the term (1+C L ) by (2+C L ) in all equations, and that the


same values of C L as before may be used.

Goodwin et al. 50 have produced a detailed treatment of bar burners. In their analysis,equation (3.21) becomes

where F is a correction factor dependent on the burner bar and port areas and thelength/diameter ratio. Values of F are plotted in Figure 3.6. An expression was also derived for A t /A p (cf. 3.20):

Figure 3.6 Correction factor for bar burners as a function of the bar and port areas for a range of length/diameter ratios. (After Goodwin et al. 50 .)

where K is the average pressure rise along the burner bar in velocity heads. It is dependent on the burner bar dimensions and needs to be determined empirically. For bars

(3.27)

(3.28)


of length/diameter ratios of 10 and 20, Goodwin et al. 50 quote values of K as 0.90 and 0.66.

Without a diffuser, (1+C L ) should be replaced by (2+C L ) in equation (3.27), while equation (3.28) becomes 50 :

3.3.4 Mixing tube design

Section 3.3.3 has derived a number of equations which relate optimum area ratios toprimary aeration, gas properties and the loss coefficients c d p and C L . In this section, the loss coefficients are discussed with particular reference to the design of the throat andmixing tube, and recommendations are made for minimizing friction losses within theburner.

3.3.4.1 Port discharge coefficient

Discharge coefficients were introduced in Section 3.2.2 for injectors, where, for a single port, values for c d of at least 0.85 can be expected. For flame ports, where the pressures and velocities are much lower, flow is usually laminar, and the discharge coefficient isrelatively low. For a typical circular port burner, Prigg 40 quotes a range of values for c dof 0.6 to 0.7 (Figure 3.3).

3.3.4.2 Friction loss coefficient

The loss coefficient C L was defined as the sum of the individual losses in the throat andthe diffuser sections of the mixing tube, and will be dependent on tube length, throatdiameter and the angle of taper in the diffuser. Friction losses in the throat can beapproximated to the loss for steady flow in a pipe, that is 4fL/D, where/is the friction factor (dependent on Reynolds number), and L/D is the ratio of throat length to diameter. For the range of mixing tubes in common use, Francis 48 gives an average value for the throat loss coefficient of 0.1 to 0.2. For a gradual diffuser, values of diffuser loss havebeen tabulated 48 as a function of angle of taper (Table 3.1). The optimum included angle is in the range 5° to 10°, giving a diffuser loss coefficient of 0.15. Hence, a typical burner will have a value of C L in the range 0.25 to 0.35.

(3.29)


Figure 3.7 The variation in primary aeration with length of the mixing tube. (From Prigg 40 after Berry et al. 51 .)

Although authors have identified two separate loss coefficients, it should be rememberedthat the two are interdependent. The longer the parallel throat, the flatter the velocityprofile at the entry to the diffuser so that losses in the diffuser will be low. If a shortthroat is used, pressure losses can be very high in the diffuser unless a long shallow taperis used. Consequently, in practice, there is an optimum combination of throat and taper.Berry et al. 51 considered the mixing tube and diffuser as one unit and obtained byexperiment the effect on air entrainment of mixing tube length both downstream andupstream of the throat (see Figure 3.7). They found that the distance from the throatentrance to the injector should be about 2 to 2½ times the throat diameter, and that

Table 3.1 Diffuser frictional loss coefficient as a function of angle of taper. (Source: Francis 48 .)

Included-angle of taper 3° 5° to 8° 10° 14° 20°

Diffuser loss coefficient 0.18 0.14 0.16 0.25 0.45


mixing tube length should be about 10 to 12 throat diameters. In practice, appliancegeometry often dictates the size of the burner. Decreasing the mixing tube length from 12to 6 throat diameters gives a 10% reduction in air entrainment.

3.3.5 Final entrainment equations

If values of c d and C L are substituted into the equations derived earlier, a number of relatively simple relationships are produced for optimized burner performance.

The simplest equation relates throat area to burner port area; using values from above,

which corresponds closely to the experimentally determined (and frequently used) valueof 0.7 quoted by Prigg 40 . Equations (3.16), (3.21) and (3.24) can be rearranged to:

For a typical partially aerated natural gas burner operating at 40% to 60% primaryaeration (R=4 to 6) and taking σ=0.6, A t needs to be 50 to 100 times A j, while A p needs to be 60 to 150 times A j, depending on the primary aeration required. These values applyto a parallel throat and gradual diffuser with no change in temperature. Similarexpressions may be derived for burners without a diffuser, for bar burners and fortemperature dependence (cf. Sections 3.3.3.6 and 3.3.3.7).

3.4 Flame port design/flame stability

Equations derived above enable the burner designer to calculate the burner port areaneeded for a given size of injector and vice versa. The equations do not, however, offerany assistance on how the port area should be distributed at the burner head. In otherwords, there is no information on port size (depth and width), interport spacing, oroptimum burner port loading (i.e. the heat input to port area ratio). These three veryimportant factors are governed by combustion considerations, in that, as well as matching

(3.30)

(3.31)

(3.32)

(3.33)


injector, throat and port area to the required volume of entrained air, the flame portdistribution must be matched with a flame that is resistant to lift, lightback andincomplete combustion.

This section covers fundamental aspects of flame stability as applied to typical partially aerated burners, and gives practical design guidance for ensuring satisfactorycombustion performance. Although many attempts have been made to analyse flamestability mathematically, only a qualitative approach is adopted here. Relevant theoreticaltreatments are referenced at appropriate points in the text. Much fuller reviews are given by Griffiths and Weber 52 and Harris and South 53 .

3.4.1 The combustion diagram

Discussion of flame stability is aided considerably by use of a combustion diagram.Figure 3.8 shows the general form of the combustion diagram in which areas ofsatisfactory operation of an aerated burner are depicted as a function of primary aerationand burner port loading. The diagram shows qualitatively those three areas whereunsatisfactory combustion results:

1. Flame lift at high primary aeration, due to the increased flow rate not being balancedby a similar increase in burning velocity.

2. Lightback at low heat input, due to the opposite of (1).

Figure 3.8 Schematic combustion diagram for a typical aerated burner. (After Harris and South 53 .)


3. Yellow tipping and incomplete combustion at low primary aeration, due to oxygenstarvation in the flame.

The exact size and location of each shaded area is dependent on the burner configuration. Careful consideration of the design parameters can minimize the extent towhich any of the above three conditions may occur. In other words, the good burnerdesigner will ensure that there is a large area of satisfactory operation and that theoperating point is somewhere near the middle of that area.

3.4.2 Flame lift

The concept of flame lift was introduced in Section 1.3.4, where it was attributed to the stream velocity through the burner not being balanced by the burning velocity of thefuel/air mixture. With a partially aerated burner, burning velocity will increase as primaryaeration rises (Figure 1.3), but this is more than outweighed by the increase in volume flow through the burner, so flame lift may occur. This section briefly considers themechanism of flame lift, and then discusses practical methods by which the effects of liftmay be minimized.

3.4.2.1 Theories of flame lift

Although single-port burners are not in use in domestic appliances (except in some types of pilot), they provide the basis for theoretical mechanisms of flame stability. Thequalitative conclusions do, however, hold for multiport burners, although precisecorrelation can be complicated by the interactions of adjacent flames.

The earliest generally accepted model of flame lift, due to Lewis and von Elbe 4 , invoked a stabilizing region within the preheat zone of the flame in terms of the velocitygradient in the boundary layer. Figure 3.9 shows the effect of increasing the gas velocity.The flame will be stable at a distance y from the stream boundary where the stream velocity u matches the burning velocity S. As u is increased, the flame stabilizes further from the burner (positions 1 to 3), until at position 4, S and u cannot be balanced, resulting in lift. The main drawbacks of this approach were that it assumed deep burnerports to establish a parabolic flow profile, and that no account was taken of the outerdiffusion flame in partially aerated flames. Later attempts attributed lift to excessiveflame curvature or stretch at the base of the flame due to shear forces and aerodynamicquenching. This was advocated by Karlovitz et al. 54 for turbulent flames and was subsequently adapted by Reed 55 , 56 for laminar aerated flames.

Although Edmondson and Heap initially supported flame stretch 57 , they subsequently held certain reservations 58 , while Melvin and Moss 59 decisively rejected flame stretch and also questioned the validity of the original boundary velocity gradient theory.Günther and Janisch 60 also discussed flame stretch and showed that the stabilizing region exists not in the preheat zone, but in the reaction zone. They suggested that as gasvelocity increases to the point of lift, the stabilizing zone moves from a position insidethe reaction zone to a position near the edge of the base of the flame. More recenttheoretical treatments 61 , 62 are based on a balance between chemical and convective


effects. A local Damköhler number, representing a ratio of chemical reaction rate to bulk mass flow rate, and local flame temperature are used as the criteria for lift. As the gasvelocity increases, the Damköhler number decreases such that below a critical value andassociated characteristic temperature, chemical activity effectively vanishes and the flamelifts off.

Figure 3.9 An illustration of the Lewis and von Elbe theory for flame lift. (After Harris and South 53 .)

Although the above theoretical concepts may provide qualitative guidance on how toimprove the stability of a flame, they are of limited quantitative application to the burnerdesigner, since, in general, they are based on a single-port burner, and assume a much more detailed knowledge of the flow characteristics of the flame port than is generallyavailable.

3.4.2.2 Application to multiport burners

Nearly all domestic gas appliances are of multiport design, albeit with many differentport configurations. Their susceptibility to flame lift is dependent on port size andinterport spacing, which in turn influence the interaction of flames from adjacent burnerports. The effect of the latter is clear from the qualitative observation that as flow rateincreases, lift is first noticeable at the edges of the flamestrip, where interaction fromneighbouring flames will be least. The following sections consider the effect on flame liftof variation in port size, shape, depth, the operating heat input, primary aeration, theposition of retention ports (if any), and the flamestrip temperature.


3.4.2.3 Effect of flow rate

On the combustion diagram for a single-port burner (Figure 3.8), the lift limit is represented as a gradual decrease in the critical primary aeration as heat input isincreased. However, a multiport burner behaves somewhat differently. At low aerations(up to about 65% primary aeration), flame interaction induces a single inner cone overthe array of flame ports. As the primary aeration is increased to over 75%, separate innercones become visible above each port. Figure 3.10 shows photographs of this effect, while the lift limit on the combustion diagram for a multiport burner is shown schematically in Figure 3.11. Practical experience 41 , 49 , 52 has indicated that, as a rule of thumb, burners operating at 50% primary aeration should have a design port loading inthe range 9 to 14 W mm−2. Higher loadings are possible only if some form ofstabilization is used, e.g. retention flames (see 3.4.2.5).

Figure 3.10 Photographs showing (left) a single induced inner cone, and (right) multiple inner cones.


Figure 3.11 Schematic illustration of the flame lift limit on a combustion diagram for a multiport burner. (After Harris and South 53 .)

3.4.2.4 Effect of port geometry

The influence of port design on flame lift has been considered in great detail by Griffithsand Weber 52 , who produced a number of empirical correlations for lift susceptibility asa function of burner port dimensions. Other workers (e.g. references 41 and 53) haveused a more qualitative approach and established empirical guidelines for lift-resistant burners.

Figure 3.12 summarizes graphically much of the work presented by Harris and South53 . The three graphs show limiting primary aeration as a function of port diameter,interport spacing and port loading. Naturally, the greater the primary aeration at lift, theless susceptible is the burner to lift at its nominal operating point. The three graphstogether identify a number of important trends which amplify the work of Griffiths andWeber 52 :

1. An increase in port loading increases the tendency to lift. This has been mentionedalready when considering the fundamental balance between burning velocity and streamvelocity.


Figure 3.12 The effect of port spacing, port size and port loading on the flame lift limit (After Harris and South 53 .)


2. For circular ports, an increase in port diameter reduces the tendency to lift, i.e. a fewlarge ports are better than many small ports. The discharge coefficient is lower for smallports than large ports. Consequently, even if the measured geometric port area remains constant, then for a constant heat input, the effective port loading increases as port sizedecreases, such that flame lift is more likely with small ports. With rectangular ports, fora given slot width, an increased slot length reduces the tendency to lift.

3. An increase in interport spacing increases the tendency to lift. This can be attributedto a reduction in interaction from neighbouring flames as the ports become more widelyspaced. So, for spacings of greater than 6 mm, little effect of interaction is observed;reduction from 6 mm to about 1.5 mm can double the primary aeration at which liftoccurs. For burner ports arranged in rows, where the row spacing differs from the portspacing in any row, the smallest spacing controls lift susceptibility 52 .

The detailed data presented by Harris and South 53 were all derived using circular ports of 1.2 mm depth. The shape and depth of the gas ports can also have an effect on flamestability. In general, ports with sharp corners are more likely to lift than ports withrounded corners. Circular ports are better in this respect than square ports and rectangularslots with a low aspect ratio (i.e. length/width ratio). However, long rectangular slots(aspect ratio >4) can be more stable than circular ports of the same port loading. Only asmall fraction of the total flow will be affected by the sharp corners at each end of theslot. The remainder has a very flat flow velocity profile so that there are no points of highvelocity that could destabilize the flame. Recirculation currents at the base of the flamesin a series of long slits also constructively interfere to stabilize the flame further. Forcircular ports, such currents are oriented in different directions and cannot stabilize theflame to the same extent.

The effect of variation in port depth is shown schematically in Figure 3.13. Unless the entrance to a gas port is suitably rounded, the flow may detach itself from the port wall(Figure 3.13a), thereby adversely affecting stability. Reattachment can take place if theport is deep enough (Figure 3.13b). Burner ports should, therefore, have a depth/diameter ratio of more than two if detachment is to be avoided 53 . In practice, thick burner materials add to production costs, so satisfactory flame stability often depends on thedesign and control of the port piercing process.


Figure 3.13 Schematic representation of flow streamlines through a burner port, showing wall attachments in deep ports. (After Harris and South 53 .)

3.4.2.5 Provision of retention flames

The concept of subsidiary retention flames was introduced in Chapter 2, where problems with lift from non-aerated hydrocarbon flames were described. The same principle can beapplied to partially aerated burners, whereby small retention flames are inherently morestable than the main flame because of their low efflux velocity.

Retention systems are most frequently used where a high port loading is required, orwhere there is to be a large turndown ratio (e.g. the domestic cooker hotplate). Because ofthe somewhat narrow range of port loading for a given aeration over which performanceis satisfactory (cf. Figure 3.8), then without some method of stabilization, it would be impossible to provide a burner with the variation in heat input that is required in mostkitchens. A wide variety of designs exists and, when properly optimized, retention flamescan increase the port loading at flame lift by up to a factor of five 53 . In practice, a supply pressure of 20 mbar limits the maximum port loading to about 25 to 35 W mm−2 for 40 to 50% primary aeration (equation 3.25).

Figure 3.14 reproduces results of Harris and South 53 , who reported the effect on stability of the separation between retention and main ports, the angle of the flames, andthe retention rate expressed as a percentage of total heat input. Figure 3.14a shows the effect of retention rate and port spacing on flame stability for a vertically firing system.There is clearly an optimum separation depending on the retention rate employed, and liftstability is very sensitive to small changes in separation. Figure 3.14b compares vertical flames with flames inclined at 45° with 10% retention. Successful retention is evident over a much wider range of port separation when the flames are inclined. Figure 3.14cshows the change in lift point as the retention flames are lowered below the plane of themain burner. Lift stability is fairly constant with inclined flames but falls off quickly atlarge separations. Although these results were presented for one particular burner, they do


Figure 3.14 The effect of various parameters on the performance of retention flames. (After Harris and South 53 .)

demonstrate clearly that retention flames offer great improvements in flame stability,although such improvements can be very sensitive to small changes in port position. Aretention flame misplaced by only 1 mm can reduce the primary aeration at which liftoccurs by up to 15 percentage points.


3.4.2.6 Effect of burner temperature

During operation of an appliance, the burner will become heated, which in turn willpreheat the gas/air mixture as it passes through the burner assembly. There are twoeffects, both of which are slight but beneficial:

1. The burning velocity of the gas/air mixture increases with temperature 3 . This leads to better flame stability at high aeration and high port loading.

2. As the gas/air mixture rises in temperature, it expands and flow resistance in theburner increases, thereby decreasing air entrainment by as much as 10% (Section 3.3.3.6). Consequently, the flames will become more stable as the appliance warms up. This effect will be absent at ignition and will decrease if the injector also is subject to arise in temperature. Clearly, satisfactory (lift-free) ignition must be provided, so, althoughincreased burner temperature is beneficial, it can be ignored by the designer with respectto flame lift.

3.4.3 Lightback

The aerated burner possesses an inherent disadvantage compared with non-aerated burners, in that lightback can occur through the mixing tube, resulting in unsatisfactoryappliance performance and possibly damage to the burner due to overheating. Lightbackoccurs when the burning velocity exceeds the flow velocity through the burner, that is theopposite condition to flame lift. Consequently, it is more frequently a problem with lowflow rates through the burner, and needs special consideration with burners operatingwith a large degree of turndown, for instance cookers. This section considers lightbackand the associated area of flame quenching, whereby lightback can in principle beavoided completely by suitable burner design.

3.4.3.1 Theories of flame quenching

Since burner designers aim to produce a burner where flames do not light back undernormal conditions, it seems appropriate to consider flame quenching, a phenomenonknown since the early nineteenth century when Davy showed that use of a fine wiregauze in a miner’s lamp prevented ignition of any ambient gas/air mixture. Davybasically demonstrated that if the burner ports are small enough, flames will notpropagate through, owing to quenching of the combustion processes.

Just as a flame is stable in a kinetic sense when there is a balance between burning velocity and gas flow velocity, there is also a thermal balance whereby heat generated inthe reaction zone is dissipated to the surrounding air, upstream into the preheat zone andto the burner assembly. If the balance is upset such that heat losses exceed heatgeneration then the flame will be quenched, i.e. it will extinguish. Such a situation occurswhen the port diameter is very small. If the burning velocity is gradually increased, or theflow velocity decreased, the reaction zone stabilizes progressively closer to the burner,and heat loss to the burner will begin to increase. If the port size is large, lightback can


occur before the flame is quenched; if the port size is small, heat loss to the burnerbecomes too great, thermal equilibrium cannot be achieved, and the flame extinguishesbefore lightback occurs.

For a particular gas/air mixture, there is a critical port diameter known as the quenching diameter, below which the flame will not light back. Many workers (seereview by Potter 63 ) have attempted to model flame quenching because of its importantinfluence on gas safety. Friedmann 64 identified a relationship between quenching diameter, burning velocity and thermal diffusivity. This was subsequently simplified byPotter 63 to give the simple relation for circular ports:

Where S u is the burning velocity, d o the quenching diameter and α the thermal diffusivity of the unburnt gas/air mixture. From this expression, it is apparent that thelimiting burner port diameter is dependent on the nature of the fuel and the primary aeration. Figure 3.15 shows quenching diameter as a function of primary aeration for methane at 20 °C. The plot is almost parabolic with a minimum near stoichiometric,reflecting the change in burning velocity. For primary aerations of less than about 65%,the gas/air mixture is outside the upper flammability limit at ambient temperatures, solightback should not occur. Table 3.2 lists minimum quenching diameters for a number ofgases, and again shows the clear connection with burning velocity (cf. Figure 1.3).

Figure 3.15 The effect of primary aeration on quenching diameter for natural gas/air mixtures in cylindrical tubes at ambient temperature and pressure, (After Lewis and von Elbe 4 .)

(3.34)



The degree of quenching depends on the shape of the burner port, since heat loss to theburner must be related to the flame shape and the distance between the reaction zone andburner port. For non-circular ports, Berlad and Potter 65 derived a number of mathematical expressions which related port dimensions to the quenching diameter of anequivalent circular port:

(a) Rectangular slot, length L, width W

(b) Infinite slot, width W

(c) Annular port, radii d 2, d l where d 2>d l

(d) Equilaterally triangular port, side length L

(e) Elliptical port, axes d 2, d l where d 2>d l

Table 3.2 Minimum quenching diameters for various gases at 1 atm and 20 °C. (Source; Harris and South 53 .)

Gas H2 CH4 C2H4 C2H6 C3H8 NGC (G22) Limiting port diameter, mm 0.8 3.5 1.8 2.5 2.9 2.7

(3.35)

(3.36)

(3.37)

(3.38)


Perhaps the most useful is equation (3.35) for rectangular slots, noting that it reduces to(3.36) as L becomes infinite. In practice, (3.36) can be used provided that L>6W, in which case errors become negligible. Thus, we have the very simple relation forrectangular ports:

whereby a quenching slot width may be easily calculated.

3.4.3.3 Effect of burner temperature

Any rise in burner temperature preheats the unburnt gas/air mixture, thereby increasingits burning velocity. While this has a slight beneficial effect with regard to flame lift (cf.Section 3.4.2.6), it also reduces the quenching diameter (cf. equation 3.34) and maypromote problems with lightback. Harris and South 53 presented experimental data which showed that for methane, increasing the burner temperature from ambient to 400 °C decreases the minimum quenching diameter from 3.5 mm to 2.3 mm. Additionally,preheating will widen the flammability limits 6 , such that, assuming a gas/air mixturetemperature of up to 200 °C, primary aerations less than 55% would be needed to prevent lightback on grounds of non-flammability.

3.4.3.4 Theories of lightback

The discussion above has shown that if the burner ports are small enough, lightback willnot occur. However, in Section 3.4.2.4, it was concluded that small ports are more proneto flame lift than large ports. In practice, because of the need to guard against flame lift, itis not feasible to use burner ports that are sufficiently small to eliminate lightback.Lightback, therefore does need to be understood with a view to minimizing theprobability of occurrence.

The so-called classical theory of lightback, attributed to Lewis and von Elbe 4 , is similar to the same authors’ theory of flame lift, in that it correlates lightback with the boundary velocity gradient. The argument presented in Section 3.4.2.1 and Figure 3.9maybe used in reverse. Thus, for a particular fuel gas, as the stream velocity u is decreased, the flame stabilizes closer to the burner, until, at some critical point, S and ucannot be balanced, and the flame propagates upstream. The drawbacks with thisapproach are the same as mentioned earlier, viz., it is assumed that burner ports aresufficiently large and deep for fully developed parabolic flow. Lightback does not appear

(3.39)

(3.40)


Figure 3.16 The transition from boundary to axial lightback as port size decreases.

to have received as much critical attention as flame lift, as a consequence of whichfewimprovements have been made to the Lewis and von Elbe theory. France 66 has studied lightback experimentally in shallow small diameter ports. He concluded that theboundary velocity gradient theory broke down because, as port size decreases, the pointof lightback remains at approximately the same distance from the port wall, that is to say,it moves towards the axis of the flame where the velocity gradient is no longer constant(Figure 3.16). Additionally, a port depth of at least twenty port diameters is required to ensure fully developed parabolic flow 67 . With practical burners, where port diameterand depth are usually roughly equal, parabolic flow is only partially developed, resultingin a much flatter velocity profile. Consequently, small shallow ports are more susceptibleto axial lightback than small deep ports (see Figure 3.17).

3.4.3.5 Practical design considerations

The lack of a suitable theoretical treatment for lightback using the form of burnercommonly employed in domestic appliances does not preclude drawing up guidelines toaid the burner designer to guard against lightback occurring. From the considerationsabove, it is clear that the risk of lightback can be eliminated by using very small ports andlow primary aeration. In practice, port sizes larger than the quenching diameter need to beused. Lightback can be avoided provided that the flow rate is high enough and the burnerports are deep enough. Attention must be paid to the effect of elevated burnertemperatures and the possible need for turndown.


Figure 3.17 The effect of non-parabolic flow on lightback in small ports. (After Harris and South 53 .)

3.4.3.6 Lightback on ignition and extinction

Even though a burner may have been correctly designed with respect to the lightbackcriteria discussed hitherto, and may operate satisfactorily at equilibrium, some systemsremain prone to the phenomenon of lightback on ignition and extinction. Because of theirtransient nature, neither has been investigated experimentally in great detail, althoughqualitative explanations have been given by Culshaw and Prigg 39 and Harris and South53 . The criterion for lightback on ignition or extinction remains the same as conventionallightback, viz., it occurs when the burning velocity exceeds the flow velocity through theburner port. Both phenomena are due to transient changes in these two parameters duringthe ignition or extinction sequence. Harris and South 53 , using high speed Schlieren photography, established experimentally that lightback on ignition with box burnersusually occurs at those ports near the end of the mixing tube, where there are rapidfluctuations in stream velocity, due probably to turbulence within the burner. Figure 3.18shows the variation of stream velocity and burning velocity with time after ignition. Thestream velocity varies rapidly and randomly, while the burning velocity rises from zero(residual air in burner), through a maximum as the mixture ratio reaches stoichiometric,and then approaches its equilibrium value. Experiment showed that burning velocity canmomentarily exceed the gas velocity during the ignition sequence, but because of therandom variation in stream velocity, it is impossible to predict accurately whether or notlightback on ignition will occur with any particular burner. Despite such uncertainty, thegeneral guidelines set out for conventional lightback apply equally here, viz., lightback ismore probable with large, shallow ports or with high burner temperatures. The latterwarrants special consideration if usage patterns include reignition of a burner which may


still be quite hot from a previous ignition.

Figure 3.18 Schematic representation of a mechanism for lightback on ignition. (After Harris and South 53 .)

Lightback on extinction occurs when the gas is turned off, if the flow velocity decreases rapidly enough to fall below the burning velocity before all the gas has beenburned. Factors to consider in addition to the usual criteria are the size of the burner bodyand the mode of valve closure. Physically large burners will take longer to burn off anyresidual gas, while a slow-acting valve or a valve with a large turndown ratio may operateat low volume flow, thereby increasing the probability of lightback on extinction. It is of interest to note that lightback on extinction has been proposed as a method of overcominglinting by using the flame to burn off lint that collects within the burner (see Section 3.5.5).

3.4.4 Yellow tipping and incomplete combustion

The maximum satisfactory primary aeration for a particular burner is generallydetermined by its susceptibility to flame lift. The minimum primary aeration isdetermined by the onset of incomplete combustion and sooting, which are accompaniedby emission of carbon monoxide and yellow-tipped flames. Incomplete combustion is an indication either that the primary or secondary air supply (or both) is insufficient, or thatthe flame is being quenched by impingement on a cool surface, or that there isaerodynamic quenching by the surrounding air (e.g. with a badly designed fanned draught


appliance). This can be caused by bad design or such factors as vitiation of the air supply,linting and changes in gas composition. This section considers burner design criteria forthe prevention of incomplete combustion. External factors (vitiation, etc.) are discussedin later sections.

3.4.4.1 Effect of flow rate

The general combustion diagram (Figure 3.8) shows that the primary aeration at the onsetof yellow tipping is virtually independent of port loading. Since the cross-section area of a flame hardly varies as gas flow is increased, the radial diffusion length for secondary airto the core of the flame remains roughly constant. Thus the primary aeration required justto prevent yellow tipping also remains approximately constant 52 . At very low flow rates, where flame length and width are almost equal, axial diffusion upstream from the flametip also becomes important, so lower primary aerations are attainable without the risk ofsooting.


The effect of port geometry on yellow tipping has been studied by Griffiths and Weber 52

for single-port and multiport burners. Since yellow tipping is dependent to a great extenton diffusion of secondary air, any design feature which increases diffusion times will alsoincrease the probability of yellow tipping.

As port size is increased, flame thickness will also increase, such that diffusion timesbecome longer and more primary air is needed to prevent yellow tipping (see Figure 3.19). Griffiths and Weber 52 found empirically that the limiting primary aeration (A) is directly proportional to the cube root of port diameter for a single circular port:

where K is a constant for the gas being used and is equal to 18 for methane, if A is expressed in percentage units and d in mm. This relation holds for port sizes up to a critical diameter (21.5 mm for methane) above which there is no dependence on portdiameter. In such cases, the diffusion path length has become so great that no secondaryair reaches the flame core and the appearance of yellow tipping is entirely dependent onprimary aeration. No dependence on port depth was found.

Figure 3.19 The effect of circular port diameter on the yellow tipping limit for selected fuel gases. (After Griffiths and Weber 52 .)

(3.41)


For a rectangular port of length L and width W,

where n is the ratio of port length to width, L/W. As with circular ports, this relation holdsfor values of L and W up to 21.5 mm, and there is no dependence on port depth.

Griffiths and Weber 52 acknowledged that multiport burners were very much more complex because of flame interaction and the effect of port size and interport spacing (inboth directions). Coalescence of adjacent flames on multiport burners will reduce accessfor secondary air. Thus, increasing port spacing will increase the tendency to soot, until aspacing is reached at which coalescence is less marked, secondary air is able to diffuse in and there is a gradual reduction in sooting tendency (see Figure 3.20).


Figure 3.20 The effect of the spacing between adjacent rows on the yellow tipping limit for a multiport burner. (After Griffiths and Weber 52 .)

3.5 External design variables

Preceding sections have dealt with parameters such as burner dimensions on which thedesigner has a direct influence. There are other ‘external’ variables which the designer must be aware of, even if they are beyond his control. Five such variables are discussedhere: gas composition changes, manufacturing tolerances, vitiation, the effects ofdraughts, and linting.

3.5.1 Effect of gas composition

The discussion hitherto has been based largely on combustion of pure, or nearly pure,methane. The effect of transient or permanent variation in gas composition must beconsidered by the designer in order that the burner can tolerate such changes withoutunduly affecting appliance performance.

The qualitative effects of changes in gas supply will be clear from the theories of lift, lightback and incomplete combustion already presented:


1. The presence of higher hydrocarbons will increase the theoretical air requirementand the risk of incomplete combustion.

2. The presence of hydrogen will increase the burning velocity of the gas and theremay be danger from lightback.

3. The presence of inerts (nitrogen, carbon dioxide) decreases the burning velocity andincreases the tendency to lift.

Assessment of the characteristics of multi-component mixtures is now aided bycomputer models 17 – 19 and empirical equations are now available which predict thelikelihood of flame lift 68 and sooting 69 . In practice, the gas supplier must distribute gasof a suitable composition, but the designer should be aware of variations that might occurand must produce appliances which will satisfactorily burn the reference test gas, as wellas the various limit test gases which assess the overall burner characteristics.

3.5.2 Manufacturing tolerances

Ideally all burners manufactured to a particular design should operate with the same heatinput and primary aeration, and, consequently, all examples of a given appliance shouldperform identically. In practice, there are inevitable production tolerances which willaffect appliance performance. For instance, gas supply from the injector is dependent onorifice diameter and the roughness of the internal surfaces. Any misalignment ordisplacement of the injector can reduce the quantity of air entrained while any roughnesson the internal surface of the burner will increase flow resistance and reduce airentrainment.

Variation in gas supply rate may be calculated using the basic injector equation (3.5),but the effect of injector displacement is much more difficult to quantify. Harris andSouth 53 showed experimentally that quite small changes in a number of these parameterscan have a large effect on primary aeration (see Table 3.3). The magnitude of each effect will depend largely on individual burner design, data presented here being for aproduction cooker hotplate burner.

The designer should be aware of the possible magnitude of production tolerances. A burner that is too sensitive to dimensional variation may provide unsatisfactory operationfor the customer and could ultimately be hazardous if lightback or incompletecombustion became prevalent.


3.5.3 Vitiation

The air supplied to a burner should ideally be pure, containing 20.9% oxygen. If, as aresult of recirculation of combustion products, the combustion air becomes deficient inoxygen, that air is said to be vitiated. Vitiation of the primary air reduces the burningvelocity of the gas/air mixture such that flame lift is encouraged. Vitiation of thesecondary air reduces the rate of diffusion of oxygen into the outer diffusion flame,thereby increasing flame length and the risk of incomplete combustion.

Reed and Wakefield 70 and Harris and South 53 have reviewed the effects of vitiation in great detail. Two causes of vitiation were identified: firstly, operation of applianceswhereby combustion products are discharged into the room air, thereby vitiating primaryand secondary air to an extent dependent on heat input, room size and ventilation rate;secondly, local recirculation of combustion products within the appliance duringoperation, whereby either the primary or secondary air (but not usually both) can becomehighly vitiated. Figure 3.21 shows the reduction in maximum burning velocity as afunction of vitiation, expressed in terms of oxygen concentration.* It can be clearly seen that the burning velocity drops quite sharply as oxygen concentration decreases. Sincesmall changes in burning velocity can have an appreciable effect on flame stability,methane burners can be particuarly sensitive to the effects of vitiation.

Table 3.3 The effects of dimensional tolerances on primary aeration for a cooker hotplate burner. (Source: Harris and South 53 .)

Tolerance factor Nominal dimension or position

Change in primary aeration per unit dimensional change

Injector radial displacement

0 mm 3.8%/mm

Injector yaw 0° 2.1%/1°

Injector diameter 1.22 mm 6.3%/0.1 mm

Throat diameter 14.7 mm 1.4%/mm

Injector axial displacement

14.7 mm 0.4%/mm

Burner resistance 30.2 mm2 0.6%/10 mm2


Figure 3.21 The effect of vitiation on the maximum burning velocity of methane. (After Harris and South 53 .)

Figure 3.22 The effect of primary vitiation on the flame lift limit for a single-port burner. (After Harris and South 53 .)


Most experimental work on vitiation and flame stability has been carried out usingdeep, single-port burners 71 , 72 , where the separate and combined effects of primary and secondary vitiation have been observed. Results for a methane burner are shown inFigures 3.22 to 3.24, which show a clear reduction in the flame lift limit as oxygenconcentration decreases. There is a near linear relationship between the lift limit andoxygen concentration when the primary air is vitiated (Figure 3.22). For secondary

Figure 3.23 The effect of secondary vitiation on the flame lift limit for a single-port burner. (After Harris and South 53 .)

vitiation, the reduction in flame stability is much less marked (Figure 3.23), and is observed only when the flame is partially aerated. That there should be any effect eventhen on lift can be attributed to the known increase in stability due to secondary diffusionnear the base of the flame 28 , 71 . If that diffusion is restricted by vitiation, then stability will be affected. Simultaneous vitiation of primary and secondary air (Figure 3.24) has a more than additive influence on stability, particularly at low port loadings.

*Vitiation level is often referred to using CO2 concentration. This is rather misleading and not used here since oxygen deficiency is the most important factor and ‘1% CO2’ represents different oxygen concentrations for different gases, because of the varying compositions of combustion products.


With multiport burners 53 , 73 , similar general effects are observed, but are complicatedby flame interaction, which itself will be affected by changes in flame shape associatedwith vitiation. Figure 3.25 shows schematically the effect of simultaneous primary andsecondary vitiation on lift stability for a typical multiport burner, using as a base thecombustion diagram described in Figure 3.11 (Section 3.4.2.3).

Figure 3.24 The effect of simultaneous primary and secondary vitiation on the flame lift limit for a single-port burner. (After Harris and South 53 .)

As well as increased susceptibility to flame lift, vitiation of the air supply alsopromotes incomplete combustion. Primary vitiation reduces primary aeration and burningvelocity and will lengthen the inner cone height of a partially aerated flame. Thesecondary diffusion flame will also be lengthened by secondary vitiation and the excessresidual fuel which is present as a result of vitiation of the primary air. Consequently, theoverall flame height can increase considerably. Harris and South 53 suggest that a reduction in oxygen concentration from 20.9% to 16.5% will roughly double the heightof a methane flame (if it has not already lifted). If the combustion chamber is not largeenough to accommodate such increases in flame size, quenching of the flame on coolersurfaces can lead to incomplete combustion.


Figure 3.25 Schematic illustration of the effect of vitiation on the flame lift limit for a multiport burner. (After Harris and South 53 .)

3.5.4 Draughts

The susceptibility of burners to flame instability caused by draughts is an importantconsideration in appliances such as cookers where burners are exposed to ambientconditions, rather than being enclosed in a combustion chamber. Draughts within abuilding are primarily dependent on external wind speed, but will also be influenced byuse of doors and windows and by attenuation inside the dwelling. Each of these factors isdiscussed by South 74 , who concluded that hotplate burners will frequently encounterwind speeds of 2 m s−1 and that burners should be designed to withstand such draughts.

Whether or not a cross-draught causes flame extinction will depend on burner primaryaeration, port loading, interport spacing and the physical shape of burner and appliance.The basic effect of a cross-draught on a hotplate burner is shown schematically in Figure 3.26. The circular symmetry ensures that part of the burner is shielded whatever thedirection of approach. Detachment occurs initially on the windward side of the burner,and moves progressively round the burner as wind speed increases. Figure 3.27 gives the wind speed for complete extinction as a function of port loading for a number of differentburners operating at 58% primary aeration. There is a clear increase in susceptibility toextinction at burner turndown. In practice, this imposes an additional limit to the amountof turndown that may be incorporated into a burner. At turndown the efflux velocity ofthe gas/air mixture from ports on the windward side of the burner head may be reduced tothe point where lightback or flame quenching occurs. With the latter, the draught maycause flow of air into the burner head. This dilutes the gas/air mixture and increases theflow rate through the leeward ports of the burner, resulting in flame lift.


Figure 3.26 An illustrative example of the effect of a cross-draught on a cooker hotplate burner.

Figure 3.27 Draught resistance as a function of port loading for a selection of production hotplate burners, (After Harris and South 53 .)


Draught resistance can be improved by careful consideration of those parameters alreadydiscussed in references to flame lift, e.g. port size, port spacing, retention flames. Figure 3.28 plots wind speed at extinction against the parameter (S—D) where S is the interport separation and D is the diameter, for a burner at full rate and on turndown. An optimum value of (S—D) exists at which very good draught resistance is possible even onturndown.

It should be pointed out that this optimum port configuration for draught stability is not necessarily the same as for normal flame lift or vitiation resistance, and that practicaldesigns will need to reach a compromise solution or be biased towards that problemwhich is most likely to be encountered. The special requirements of particular applianceswill be dealt with in a later chapter.

3.5.5 Linting

Linting of aerated burners by airborne dust and fibres constitutes a major servicingproblem as well as causing a deterioration in appliance performance. Lint enters theburner with the primary air stream and tends to accumulate at the injector, in the burnerbody and at the burner ports (Figure 3.29). This restricts both primary and to some extentsecondary air entrainment, resulting in flame lengthening and eventually incomplete

Figure 3.29 Linted burner

combustion and sooting. In extreme cases, soot deposition can block a heat exchanger orflue pipe and combustion products may be emitted into the surrounding air.

A linted burner is more prone to incomplete combustion, but less susceptible to flamelift. Although regular servicing should identify and rectify linted appliances, it isimportant that burners should be resistant to blockage for reasonable periods of operation.Effort has therefore been expended on development of lint-resistant burners 36 , 75 . Only


the main conclusions are discussed here; a selection of lint-resistant burners is shown in Figure 3.30. Burners A, B and C are modified examples of standard box burners. In burner A, a cylindrical gauze traps any lint contained in the primary air. Baffles withinburner B reduce the momentum of the lint so that it drops out into a removable tray. Inburner C, the primary air inlets are sufficiently close to the flame for lint to bepreferentially incinerated rather than enter the burner body. All three modificationsincrease flow resistance in the burner, and this will affect primary air entrainment.

Burners D, E and F are of the jetted type and have been designed so that internalobstructions and sudden changes in direction of flow are avoided.

Any airborne lint is carried through the burner and incinerated. Exposure to anartificially linted atmosphere shows great improvements over similar burners ofconventional design (Figure 3.31). It should be pointed out, however, that such burners still require servicing and cleaning; only the time which elapses before combustionbecomes affected has been increased.

Although burners with more resistance to lint provide a possible solution, there areother approaches which ought to be considered. Appliances which draw combustion airfrom outside the room (e.g. balanced flue boilers) will obviously not attract householddebris. If such room-sealing is not feasible, air intakes should be at high level rather than floor level, where much of the heavier dust and fibres tend to collect. With fannedappliances, an air filter may be fitted, but will require regular cleaning or replacement.Lightback on extinction has been advocated whereby lint within the burner is incineratedas the flame front propagates upstream within the burner body.

None of these methods is a guarantee against the linted burner, but it should bepossible, using careful design, to reduce the problem to a more acceptable level.

3.6 Summary procedure for burner design

This chapter has covered many aspects of the design of partially aerated burners. It is notintended to be comprehensive; many sections could form substantial publications on theirown if all published data were presented. It is perhaps worth summarizing in the form ofa generalized procedure for design of a partially aerated burner:

1. Decide the maximum heat input for the appliance and an approximate primaryaeration (~ 50%).

2. Calculate injector size (equation 3.5) from heat input and gas pressure. The fullsupply pressure of 20 mbar need not be used; a low pressure gas regulator upstream ofthe injector enables pressure to be reduced. If the appliance is to be operated at more thanone heat input, the designer must decide whether to use one injector at full and partialrates, or a number of injectors, each of which may be on or off. The benefits of eitherapproach are discussed in Chapter 6.

3. Calculate the mixing tube throat area for each injector using equations from Sections3.3.3 or 3.3.5, depending on the accuracy required. Qualitative guidelines on tube length and diffuser angle were given earlier. In practice, mixing tube length is often ultimatelydetermined by the physical size of the appliance.


Figure 3.30 A selection of lint-resistant burners. (After Harris and Wilson 36 .)


Figure 3.31 A comparison of lint resistance of conventional and experimental burners. (After Harris and Wilson 36 .)

4. Calculate the total burner port area and the burner port loading. For burner portloadings of greater than about 12 W mm−2, retention flames may be needed to preventflame lift. The exact dimensions of individual ports are largely a matter of personalchoice and constraints imposed by the size and shape of the appliance. Due regard mustbe given to the risk of lift, lightback and incomplete combustion during normal operation,and the possible effects of variation in gas composition, vitiation, draughts and linting.

From these guidelines, it is possible to construct a burner which should give stable flames and is sufficiently resistant to adverse conditions. It is very often not possible tomeet the optimum conditions owing to physical constraints imposed by the nature of theappliance. The special needs of individual appliances and the relation with burner designare discussed in Chapter 6.


CHAPTER 4 Fully aerated burners

4.1 Introduction

At present, nearly all burners used in domestic gas appliances are partially aerated, that isonly part of the air needed for complete combustion is entrained as primary air. The needfor an adequate supply of secondary air for partially aerated burners raises a number ofdesign considerations:

1. Secondary air openings in the appliance must be sized and positioned so that thecorrect air flow distribution is maintained.

2. Burner ports must be located with due regard to access for secondary air. 3. In order to ensure sufficient secondary air, it is necessary to introduce a large

quantity of excess air (usually at least 40%), which reduces appliance efficiency. 4. The combustion chamber size is limited by the need to provide enough space for the

flames not to impinge on relatively cool surfaces. With fully aerated (premixed) burners, which require no secondary air, none of the

4.1 Introduction 4.2 Air supply 4.2.1 Atmospheric injection 4.2.2 Mechanical aeration 4.3 Burner design 4.3.1 Ribbon burners 4.3.2 Punched metal burners 4.3.3 Radiant burners 4.3.4 Burner noise 4.4 External design variables 4.4.1 Vitiation 4.4.2 Draughts 4.4.3 Linting 4.5 Summary

above restrictions applies. The more compact fully aerated flame allows combustionchambers to be smaller, while the lower excess air levels that are possible permit muchhigher appliance thermal efficiencies. However, the design of fully aerated burnerspresents a number of problems regarding air supply and flame stability. Although it ispossible to entrain enough primary air using an injector/mixing tube system, there arepractical difficulties (discussed below) with such an arrangement. It is therefore moreusual to use an external energy source, such as a fan. The burner itself also needs to bedesigned with great care, because of the risks of flame lift, lightback and burner overheat,which are associated with the high primary aeration and the consequent increased burningvelocity of the gas/air mixtures used. This chapter considers these special requirementsfor fully aerated burner systems. Thorough, though somewhat dated, reviews arecontained in American Gas Association Research Bulletins 76 , 77 .

4.2 Air supply

Partially aerated burners take up primary air by using the pressure of the gas supplybehind an injector to entrain from the surroundings a fraction of the air required forcomplete combustion. Because of the relatively low-energy gas supply, primary aerations of greater than 100% are difficult to achieve. Consequently, many fully aerated systemsemploy a fan, which can deliver air at a predetermined flow rate. This section extends theair entrainment theories of Chapter 3, as far as they are applicable to fully aeratedsystems, identifies the limitations of air injection and considers some of the basicrequirements for a fan-powered burner system.

4.2.1 Atmospheric injection

The equations derived in Sections 3.2 and 3.3 are applicable to both partially and fully aerated burners. Therefore, by substituting suitable values of R, the air/gas ratio, optimum design area ratios may be obtained for a fully aerated burner. In order to ensure complete combustion and allow for tolerances in manufacture and gas composition, fully aeratedburners usually have a design primary aeration in the range 110% to 120%, i.e. R is roughly 11 to 12 for natural gas. The friction loss and discharge coefficients should beapproximately the same as before, so equations (3.31) and (3.32) maybe used to calculatethe optimized throat/injector and port/injector area ratios. Taking σ=0.6 and R=11 to 12, it is found that the throat area should be 300 to 400 times the injector area, while theburner port area needs to be 400 to 550 times the injector area. Compared with the arearatios derived for partially aerated burners, it is clear that physically very large burnersare required for fully aerated operation.

An associated problem concerns the maximum port loading achievable usingatmospheric injection. Equation (3.22) relates burner port loading to gas supply pressureat the injector:


Clearly, as R increases, the burner port loading decreases for a given supply pressure.This is shown in Figure 4.1 as a plot of port loading against primary aeration for three gas supply pressures. While port loadings of up to 35 W mm−2 are possible with a typical partially aerated flame, full aeration imposes a limit of about 5 W mm−2. At such low port loadings, lightback and burner overheat become evident, especially if turndown isalso required.

Figure 4.1 Maximum burner port loading by atmospheric injection as a function of primary aeration and gas supply pressure.

The effect of preheating the gas/air mixture also becomes critical. If a burner is designed to operate at, say, 130% primary aeration at equilibrium, then under coldconditions (e.g. on ignition) it will require up to 175% primary air 77 . Aerations of such magnitude can cause severe problems with flame lift as well as reducing applianceefficiency.

As a consequence of these inherent design problems, atmospheric injection is littleused for fully aerated burners in domestic gas appliances. The major exception is thosegas fires which use surface combustion radiant plaques, where a low port loading gives

Fully aerated burners 81

an attractively large radiating surface area and a flame that burns very close to thesurface, thereby heating the surface to incandescence. Radiant plaques using atmosphericinjection may also be used in cooking appliances, where flameless radiant heating isparticularly adaptable for ceramic hotplates and grills.

4.2.2 Mechanical aeration

A more satisfactory method of obtaining the required gas/air mixture is by use of anexternal energy source, such as a fan. Although fans have been used in industrial aeratedburners for many years, their incorporation into domestic gas appliances is comparativelyrecent. Most of those fitted to date are used with partially aerated burners, where the fanprovides the supply of air and ensures removal of combustion products. However,atmospheric injection using the fan-induced air stream remains the method by whichprimary air is entrained. Fans may be positioned either upstream of the burner (forceddraught), thereby pressurizing the appliance with respect to atmosphere, or downstreamof the burner and heat exchanger (induced draught), thereby placing the appliance undernegative pressure. In the latter case, the fan will need to withstand the elevatedtemperatures and the mildly corrosive environment of the combustion products;consequently, this arrangement is much less commonly used at present, despite the virtualelimination of any risk of combustion products leaking into the room in which theappliance is installed.

The correct air/gas ratio is obtained by matching the respective air and gas flow rates. The gas flow rate will generally be regulated by a valve or governor, while the air flowrate will be determined by matching the flow/ pressure characteristic of the fan with thatof the appliance and its flue system. The designer must, therefore, choose a fan which ispowerful enough both to overcome all the flow restrictions in the appliance and flue, andto provide and maintain the correct flow of air. Figure 4.2 shows a generalized layout for a fully aerated fanned draught combustion system. Air and gas are introduced into amixing tube, the sides of which may be parallel or slightly tapered. The mixture is thenled directly to the burner, where the combustion process occurs. Downstream of the heatexchanger, the combustion products are finally directed to the flue for disposal. Flues forfanned draught appliances do not depend on the buoyancy of hot combustion products fortheir operation, and can, therefore, be directed in an upward, horizontal or downwarddirection.


Figure 4.2 Schematic representation of a fanned draught combustion system.

As mentioned above, a satisfactory fanned draught design matches the air flow fromthe fan with the gas supply. Because the pressure losses in the appliance are greatlydependent on the precise design of the pipework and the combustion chamber, it is notpossible to present here a detailed quantitative assessment.

Working with industrial fully aerated burners, Francis and Jackson 78 derived a series of design equations, the general principles of which are equally applicable to domesticappliances. The approach is similar to Francis’ earlier work 48 on atmospheric injection (see Chapter 3), in that static pressure changes are calculated throughout the appliance in terms of loss coefficients for the mixing tube, combustion chamber and flue system.Because the coefficients are critically dependent on the exact design, it is not possible togive exact values here, other than to say that the pressure drop across a typical appliance(excluding the flue) is about 1 mbar.

The friction losses along the flue are dependent on flue dimensions, the number of bends and the temperature of the flue gases. For turbulent flow in smooth pipes, thepressure drop (∆p in Pa) is given by:

where L and d are the pipe equivalent length* and diameter, and µ, ρ and v are the flue gas dynamic viscosity, density, and velocity (all in SI units). As an example, consider a10 kW boiler operating at 130% primary aeration (30% excess air) with a flue gastemperature of 100 °C. Substituting appropriate values for µ and ρ, a pressure drop of about 0.1 mbar m−1 for a 75 mm diameter flue, and 0.5 mbar m−1 for a 25 mm diameter flue is obtained. Thus, for large-diameter flues (greater than 75 mm), the heat exchanger

(4.1)


offers the greatest resistance to flow, whereas for miniature flues, the flue can be the mostresistive component.

With a knowledge of the pressure drop in the system and the volume flow of airrequired, a suitable fan can be chosen. Commercially available centrifugal fans of theflow rate required develop a static pressure of up to 2.5 mbar, which limits their use withsmall-diameter flues to either short lengths (e.g. through-the-wall balanced flues) or low heat input appliances (e.g. 2 kW space heaters). Fans that generate a higher staticpressure (>5 mbar) are under development 79 which will enable several metres of small-diameter flue to be used with higher rated domestic boilers.

4.3 Burner design

The design of a satisfactory fully aerated burner head presents a number of problems tothe designer:

1. Such burners are physically much larger than partially aerated burners because ofthe much greater volume flow rates involved.

2. The design and arrangement of burner ports must prevent flame lift, lightback andburner overheat. Incomplete combustion should not be a problem provided that there isno flame quenching and good gas/air mixing is maintained.

3. The high primary aeration and associated flow rates can create noise problems in theburner and the combustion chamber.

One advantage of the fully aerated burner is that because of the small compact flames,which are momentum-controlled rather than buoyancy-controlled, it is possible to mount the burner so that it fires upwards, sideways or downwards. This provides much greaterflexibility for siting of the burner within an appliance. For condensing boilers, adownward-firing burner placed at the top of the appliance eliminates the problem of condensate falling back on to the burner; in cooker grills, downward-firing radiant burners ensure an even heat distribution above the grill pan.

This section considers a number of burner types with respect to their susceptibility to flame instability and burner overheat. Because most fully aerated burners are still at thedevelopment stage, only preliminary results, which in places are necessarily incomplete,can be given here.

*The equivalent length of any combination of pipe, bends and fittings is the length of straight pipe that would produce the same pressure drop as that combination.

4.3.1 Ribbon burners

Ribbon burners consist of a series of corrugated metal strips between which are locatedstraight metal separators (Figure 4.3). Ribbon burners provide a large port area within a given space, thereby reducing the physical size of the burner and the combustionchamber. Parameters which need to be considered include port depth, ribbon dimensions(gauge, pitch and amplitude), the separator width, packing pressure and the evenness of


the flamestrip. Pearson et al. 80 have examined the variation of these parameters and theireffect on flame stability and burner temperature. A summary of their results together withsome previously unpublished data from the British Gas Watson House Research Stationis presented here.

Figure 4.3 Exploded and assembled view of a ribbon burner. (After Pearson et al. 80 )

4.3.1.1 Flame lift

Flame lift has been found to be a major limitation of all fully aerated burners; systematicstudies 80 have attempted to identify the effect of the various design parameters on thetendency to lift. Lift is here regarded as unacceptable when the flames have lifted fromthe outer ribbons and the inner flames are just starting to lift. Results presented at this stage refer to a simple ribbon burner operating in the open. The effect of enclosing theburner and using flame retention is discussed in a later section.


1. Number of ribbons. Of all the factors discussed here, the number of ribbon elementsappears to be one of the most critical. Figure 4.4 shows the lift limit as a function of port loading and the number of ribbons. It is clear that the primary aeration at lift rises as thenumber of ribbons increases from one to four. For six or more ribbons, the lift pointremains virtually constant and is dependent only on port loading. This is consistent withthe observations for partially aerated burners (Section 3.4.2.4) where interaction with adjacent flames greatly enhances stability. Optimum performance is therefore obtainedwith a burner containing more than a certain critical number (n c ) of ribbon elements, usually seven or more 80 .

Figure 4.4 The variation in flame lift limit with number of elements for a ribbon burner.

2. Separator width. The only other parameter that greatly influences stability is theseparator width. Separator width is somewhat akin to interport spacing in a traditionalmultiport burner, so variation in separator width can affect the degree of mutualstabilization from neighbouring ports. Figure 4.5 shows the effect of separator width on flame stability for a burner already optimized with regard to the number of ribbons. Forall port loadings, there is a range of separator widths (1.25 to 2.25 mm) for whichresistance to lift is a maximum 80 . Outside these limits, a decrease in flame stability is observed. For large separator widths, this may be attributed to the reduced stabilizinginteraction; at low separator widths,it is probable that stabilization due to localrecirculation of combustion products near the base of the flame is restricted by the veryclose presence of adjacent flames. Figure 4.6 shows a burner where one half uses separators of 1.25 mm thickness giving stable flames, while the other half uses 0.5 mm


separators, which give a very unstable flame.

Figure 4.5 The variation in flame lift limit with separator width for a ribbon burner.

3. Ribbon pitch and gauge. Varying the gauge or pitch of the ribbons has little effecton stability, providing that the port loading remains constant and that the burner containsat least n c ribbons. However, n c does vary slightly with ribbon pitch. A practical maximum value of ribbon pitch is 4 mm, for which n c is 5 ribbons. A typical ribbon pitch of 2.5 mm gives 7 ribbons as the critical number.

4. Port depth. The ports generally used with ribbon burners are of such a depth thatproblems of flow detachment in shallow ports discussed in Section 3.4.2.4 are not expected. Experimental evidence suggests that port depth has little effect on flame lift;only at high port loadings with port depths of less than 5 mm has any reduction in flamestability been noted.

Even if the preceding guidelines are followed in the construction of a simple ribbon burner, it should be noted that the maximum port loading attainable is about 4 W mm−2

for a primary aeration of 130% (Figure 4.5). Increased stability can be achieved by use of a suitably designed flame retention system (see below).


Figure 4.6 The effect of separator width on flame lift from a ribbon burner. (After Pearson et al. 80 )

4.3.1.2 Lightback

Little systematic work has been carried out on lightback for ribbon burners. However, thegeneral theoretical and practical guidelines presented in Section 3.4.3 for partially aerated burners should remain valid. The deep ports generally used for ribbon burners shouldhelp guard against lightback, although the larger port size and higher burning velocity ofthe gas/air mixture may increase the risk. The probability of lightback is an importantfactor to consider if any degree of turndown is required.

4.3.1.3 Burner temperatures

Flame temperatures are higher for a near-stoichiometric flame than for a typical partially


aerated flame. This, together with the simple observation that the combustion zone isvery near the burner surface, can lead to problems with burner overheat in fully aeratedsystems. Excessive burner temperatures (greater than about 500 °C for conventional metallic materials) may cause material deterioration, burner distortion and an increasedprobability of lightback. Figure 4.7 shows isotherms for a typical ribbon burner as a function of primary aeration and port loading. Maximum temperatures are found at about1 W mm−2 and for aerations near stoichiometry. Although the shape of the isotherms remains the same for all ribbon burners examined, the burner temperature is greatlydependent on burner geometry and the orientation of firing. For example, a downward-firing burner may operate up to 120 °C hotter than the same burner firing upwards, because of the buoyancy of the hot combustion products. Temperatures quoted in thissection were measured for an upward-firing burner with thermocouples placed along the axis of the central separator.

Figure 4.7 The effect of primary aeration and port loading on the temperature of a typical ribbon burner. (After Pearson et al. 80 )

1. Number of ribbons. Figure 4.8 shows the burner temperature plotted as a function ofport loading and the number of ribbon elements for a primary aeration of 110%. Clearly,the greater the number of ribbon elements, the higher the burner temperature. This wouldbe expected, since the centre of a wide burner will be further away from its relativelycooler surroundings than a narrow burner, so conduction and radiation of heat away fromthe burner is more difficult. Thus, the flamestrip should consist of as few elements as


Figure 4.8 Ribbon burner temperature as a function of port loading and number of ribbon elements.

Figure 4.9 Ribbon burner temperature as a function of port loading and separator width.


Figure 4.10 Ribbon burner temperature as a function of port loading and ribbon pitch.

possible to minimize burner overheat. This is in contrast to the recommendation regarding flame lift; in practice, a compromise between the two effects is needed.

2. Separator width. The effect of separator width on burner temperature is shown in Figure 4.9. Temperature is seen to increase significantly with separator width, indicatingthat overheat may be minimized by employing thin separators. This may be attributed tothe same conduction and radiation mechanism mentioned above, and may be exacerbatedby recirculation of hot combustion products between the ports at high separator widths.

3. Ribbon pitch. Whereas variation in ribbon pitch has little effect on tendency to lift, itcan greatly influence the flamestrip temperature. Figure 4.10 shows that, for a given ribbon, as port loading is increased, burner temperature rises, reaches a maximum andthen steadily decreases. As pitch increases, the maximum burner temperature increaseswhile the port loading at which that maximum temperature occurs decreases.Consequently, at port loadings of greater than about 2.5 W mm−2, a high ribbon pitch will minimize burner temperature; at low port loadings, a small ribbon pitch should be used. Ifa burner is to be used at a number of heat inputs, a small pitch is preferable, since it is the


Figure 4.11 Ribbon burner temperature as a function of port loading and port depth.

temperatures generated at low port loadings that are most likely to cause burnerdeterioration.

4. Port depth. Figure 4.11 shows that as port depth is increased from 5 mm to 15 mm,burner temperature reduces significantly. Heat flow by conduction is dependent ontemperature gradient and cross-section area. Consequently, because deeper ribbons andseparators provide a greater cross-section for heat flow to the burner manifold, heat lossesare higher and the burner is cooler.

5. Flamestrip packing pressure. Because a ribbon burner is composed of a number ofindividual elements, the compression (packing pressure) applied across the flamestripbecomes an important design parameter Loose-fitting ribbons can create additional gaps between the elements, such that port area is larger and the flames burn closer to theflamestrip. Conduction to the edge of the burner will also be reduced if adjacent strips arenot quite in contact. As a consequence, the burner temperature is higher with loose-fitting ribbons, and the probability of lightback is increased. The extent of compression isvirtually impossible to quantify. It is to be recommended, however, that ribbons arepacked as tightly as possible, and that close tolerances are maintained.


6. Flamestrip evenness. The multi-component nature of a flamestrip can

Figure 4.12 Photograph showing burner overheat due to an uneven flamestrip.

lead to unevenness due to vertical displacement of individual elements with respect toeach other. A downward displacement of a ribbon or separator has little effect on burnertemperature; an upward displacement can lead to severe overheating of the misalignedcomponent. This is shown photographically in Figure 4.12, where displaced elements raise parts of the burner to red heat. As a design guide, experimental results suggest thatif the flamestrip is compressed tightly, then evenness can vary within limits of ± 0.5 mm without adversely affecting flamestrip temperature.

4.3.1.4 Combustion diagram

The preceding sections on lift, lightback and burner temperature can be summarized as acombustion diagram (Figure 4.13) in which the area of satisfactory operation is shown as a plot of primary aeration against port loading for a typical ribbon burner. The diagramshows how the three fault conditions can combine to restrict the range of satisfactoryperformance, especially with regard to turndown. Consider point X as a nominaloperating point (110% primary aeration, 3 W mm−2). The line AA represents the performance characteristic on turndown with the aeration held constant (i.e. gas and airsupplies are modulated). Heat input may be varied without adverse effect within therange 2 to 4 W mm−2, which, bearing in mind manufacturing tolerances, would give arealistic turndown of no more than 1.5:1. Line AA also indicates that if the flowdistribution is not uniform through the burner, then local variations in port loading canlead to local areas of overheat or lift on the burner head. If the air supply is not modulated


with the gas rate, then primary aeration varies considerably, such that, as well asefficiency being reduced by excess dilution air, turndown is not practicable (line BB).

Figure 4.13 illustrates two important design considerations: the need for some form offlame retention system to increase the operating area (Section 4.3.1.5), and the need for some form of air/gas ratio control to optimize efficiency and the degree of turndown(Chapter 7).

Figure 4.13 Combustion diagram for a typical non-optimized ribbon burner.

4.3.1.5 Flame retention

The principles of flame retention, already discussed for non-aerated and partially aerated burners, may also be applied to fully aerated burners as a method of improving flamestability at high primary aeration and high port loading. Two retention methods have beenfound to be suitable 80 : using auxiliary flames and wall attachment. Auxiliary flames use low-velocity flames to anchor the higher rated main flames, whereas wall attachmentrelies on the well known observation that flames can be stabilized by attachment toadjacent surfaces (such as the wall of a combustion chamber). Both principles were usedin Chapter 2 for the retained Bray jet and the rod-stabilized burner, which were developed as non-aerated methane burners.

1. Auxiliary flames. Figure 4.14 shows that, regardless of main port loading, maximumstabilization is obtained by use of an auxiliary port loading in the range 1 to 2 W mm−2, although all auxiliary loadings above 0.5 W mm−2 give some degree of stabilization. The number of ribbons has also been found to be important; satisfactory performance wasachieved over the whole operating range when three or four auxiliary ribbons were


employed in addition to the six or seven ribbons already recommended for the mainburner.

2. Wall attachment. The separation between the retaining wall and the outermost flame ports appears to be the critical parameter for satisfactory wall attachment in that amaximum separation exists above which no

Figure 4.14 The effect of main and auxiliary flame port loading on main flame lift

stabilization is evident. This maximum separation is dependent on port loading, andvaries from 5 mm at 6 W mm−2 loading down to 2 mm at low port loadings. For allseparations less than the critical distance, stability appears to remain roughly constant.With respect to design guidelines, a separation of 2 mm or less is recommended, in orderthat all port loadings used will give stable flames.

Figure 4.15 shows the improvement in performance which can be obtained with anoptimized fully aerated ribbon burner (cf. Figure 4.13). A much larger area of satisfactory operation is now apparent. This provides for a greater turndown ratio (up to 6:1 at 130%primary aeration) without the risk of a deterioration in performance. The choice betweenwall attachment and auxiliary flames is left to the designer. Both methods are roughlycomparable in effectiveness, so, ultimately, factors such as cost, component reliabilityand the physical construction of the appliance determine which form of retention is mostsuitable.


Figure 4.15 Combustion diagram for a typical optimized ribbon burner with flame retention.

4.3.1.6 Effect of gas composition

Because of the relatively small area of satisfactory operation for fully aerated burners, theeffects of variations in gas composition are especially important for this type of burner.Flame lift can be eliminated by suitable choice of operating point and incorporation of aflame retention system. There remains the problem of burner temperature which has beenstudied extensively for an optimized ribbon burner.

Changes in gas composition can be considered in two ways: 1. Gas supplied at a fixed supply pressure; air supplied at constant rate (either

atmospheric injection or by fixed-speed fan). Figure 4.16 shows the effect of a simple exchange of one gas for another, by plotting burner temperature against nominal portloading (i.e. the port loading obtained with G20 at the supply pressure) at 120% nominalprimary aeration. Any change in TAR due to variation in gas composition will also affectthe true primary aeration, since the air flow rate is fixed. The true primary aeration isgiven by:

(4.2)


Figure 4.16 shows that G21, the sooting and incomplete combustion test gas, producesthe highest burner temperature. This is because of its higher heat input, TAR and burningvelocity compared with G20; the true primary aeration is reduced, the flame burns closerto the surface and the burner temperature increases. G23, the lift test gas, gives lowestburner temperatures for the opposite reasons, viz., lower heat input, TAR and burningvelocity. For G22, the lightback test gas, the higher burning velocity is balanced by thelower TAR (hence higher true primary aeration) such that burner temperature is virtuallyunchanged compared with G20. G24 generally gives higher temperatures than G20

Figure 4.16 Ribbon burner temperature as a function of port loading for various test gases at 120% nominal primary aeration.

because of its higher burning velocity. Different orientations of firing give the sametrends, although absolute values may vary because of buoyancy effects.

2. Gas supplied at a fixed supply pressure; air may be varied by air/gas ratio control(see Chapter 7 ) to maintain constant true primary aeration. Burner temperature for various gases under these conditions is shown in Figure 4.17. Comparing with Figure 4.16, the trends observed are quite different. G21 gives the lowest temperatures because,in view of its high TAR, a much increased flow rate through the burner is required tomaintain primary aeration. The flames therefore stabilize further from the burner, therebylowering the surface temperature. G22 now produces the highest burner temperatures.The flames burn very close to the surface because, in addition to the higher burningvelocity, the flow rate through the burner is considerably reduced owing to the lowerTAR With G23, the lower TAR and volume flow rate balances the lower burning


velocity, while for G24, as in Figure 4.16, slightly higher temperatures are recordedbecause of its greater burning velocity than G20. Once again, orientation of firing has noeffect on the general trends observed for the various gases, but may influence the actualtemperature values.

Figure 4.17 Ribbon burner temperature as a function of port loading for various test gases at 120% true primary aeration.

4.3.2 Punched metal burners

Punched metal burners are employed frequently as partially aerated in boilers, gas firesand water heaters. This does not preclude their burners use as fully aerated burners,provided that due regard is given to the problems of flame stability. Minchin 81

demonstrated that full aeration was feasible for town gas, although burner overheat andlightback were noticed at low port loadings. Little systematic work has been undertakenfor fully aerated punched metal burners because of the advantages offered by otherburner configurations. This section therefore highlights the main problems that need to beconsidered in punched metal designs.

The blow-off limits for multiport burners were discussed in Section 3.4.2 and presented graphically in Figures 3.12 to 3.14. The conclusions given there are equally applicable to fully aerated burners, that is to say, stability is best achieved by low portloading, large ports and small interport spacing. Figure 3.12 shows that the area of satisfactory operation (i.e. an aeration at lift of at least, say, 130%) is confined to portsizes of more than 2 mm and interport spacing of less than 1 mm. Such port sizes can


increase the risk of lightback, while interport spacings of less than 1 mm may seriously weaken the burner head. The low port loadings required to eliminate the risk of flame liftfrom a fully aerated burner cause the flames to burn very close to the burner. This canlead to high burner temperatures and the possibility of lightback and material damage tothe burner head. The very shallow ports usually associated with punched metal burnersare particularly vulnerable to high burner temperatures (over 600°C has been recorded), which in turn promotes lightback.

In summary, although punched metal burners offer advantages over ribbon burners in terms of costs, they must be carefully designed in order to minimize problems with flamelift and burner overheat.

4.3.3 Radiant burners

Radiant burners have been used for overhead heating for many years. The principle isrelatively simple: a fully aerated flame burns very close to a ceramic surface, therebyheating the surface to incandescence. Thus, whereas the burners described hitherto haverequired careful design to minimize the effects of high burner temperatures, radiantburner design encourages a high surface temperature, subject to the limitations of theceramic. Two categories of radiant burner are considered here: firstly, radiant plaques,where discrete holes through the ceramic form the burner ports; secondly, burnersconsisting of layers of porous ceramic fibres or foam through which the gas/air mixturemay pass. Other types have been developed 82 , particularly for industrial applications, but are not discussed here.

Satisfactory radiant output requires the flame to burn very close to the surface of the ceramic. Therefore, for a particular volume flow, optimum performance is to be expectednear stoichiometry (100% primary aeration) where the burning velocity and flametemperature are at a maximum. In practice, a primary aeration of 110% to 115% iscustomary, in order that the flame remains fully aerated in the event of any variation inair or gas supply (see Section 4.3.3.3). Because radiant output is crucially dependent onmaintaining the correct primary aeration, turndown can be achieved only by interlinkedcontrol of gas and air flow rates. With atmospheric injection this is providedautomatically, because primary aeration is independent of injector gas pressure (Section 3.3.3.5). For fanned draught operation, electrical control of air/gas ratio is required.

Theoretical aspects (not included here) of radiant heat transfer and radiant burners are given by Kilham and Lanigan 82 , De Werth 83 and Weil 84 .

4.3.3.1 Radiant plaques

The radiant plaque in its simplest form is shown in Figure 4:18; the gas/air mixture passes through a flat ceramic plate containing an array of small (~ 1.2 to 1.6 mm)diameter flame ports. Radiant efficiency (i.e. the radiant output as a fraction of the totalheat input) has been measured as about 18% 82 . This can be increased in three ways.

Firstly, a wire mesh gauze may be placed a few millimetres above the surface of the plaque. The gauze reflects heat back to the plaque, thereby increasing the surface


temperature. This in turn produces greater preheating of the gas/air mixture, increasedmixture burning velocity and a flame burning closer to the surface. The gauze itself isalso heated to red heat. Kilham and Lanigan 82 recorded 36% radiant efficiency when agauze was fitted. One drawback with a gauze is that if, for any reason, the flamesimpinge on the gauze (e.g. linting, vitiation), quenching of the flames together withcarbon monoxide emission may occur. Deterioration in the gauze material with use isalso known.

Figure 4.18 A Schwank flat radiant plaque.

A second method for increasing efficiency is to use a ‘profiled’ plaque surface. Two commercially available plaques are shown in Figure 4.19. The example by Tennant features a pyramidal surface with the gas ports situated in the ‘valleys’; that by Schwank consists of a number of indentations in an originally flat surface. The principle ofoperation is, however, similar. A degree of mutual heating is possible between adjacentraised areas, resulting in the same effect as is seen with the radiant gauze 85 . Radiant efficiencies of nearly 40% have been measured for profiled plaques burning natural gas.

Surface coatings (e.g. metal oxides or salts) have also been studied as a method for improving radiant output from flat plaques 83 . Any increases that are observed generallycompare unfavourably with fitting a gauze or modifying the surface geometry; somemetal coatings can decrease radiant output. At the time of writing, the effect of coatingson a profiled plaque has not been thoroughly investigated. As well as being critically dependent on primary aeration, the radiant efficiency also depends on the flame portloading. Figure 4.20 plots radiant efficiency as a function of port loading for three radiant plaques, and shows that the highest radiant efficiency is obtained with port loadings ofabout 0.4 W mm−2. This must, however, be treated with some caution if optimum radiant appearance is required. As the port loading is increased up to 4 W mm−2, the radiant efficiency decreases steadily, but the radiant output (in terms of watts per unit plaque


Figure 4.19 Profiled radiant plaques by Schwank (top) and Tennant

area) continues to rise. Figure 4.21 illustrates this by plotting radiant output against portloading for the same three plaques. As a consequence, plaques may appear moreattractive a high port loadings. This should be borne in mind for gas fire applicationwhere a customer is likely to be influenced more by appearance than by the more abstractradiant efficiency.


Figure 4.20 Radiant efficiency as a function of port loading for three radiant plaque burners using methane.

Figure 4.21 Radiant output as a function of port loading for three radiant plaque burners using methane.


Flame stability considerations for radiant plaques centre on lightback. For port loadings commonly used, lift is evident only at primary aerations in excess of 160%. Thiscompares with an optimum operating point of 110% primary aeration and severedeterioration in radiant appearance above 130%. Consequently, flame lift should not beencountered if plaques are used as radiant heaters.

The very high surface temperatures may lead one to expect that lightback could be a problem. Kilham and Lanigan 82 measured the temperature profile within a 12 mm thick plaque, and found that, because of the very low thermal conductivity of the ceramic, thetemperature at the back of the plaque is approximately 40 °C (cf. front surface temperature of 850 °C). Consequently, lightback should not occur for burner ports currently used in commercially available plaques, since the port diameter (–2 mm) is less than the quenching diameter.

One factor to consider with radiant plaques is the fragility of the ceramic materialsused. Poor assembly, mishandling or excessive burner temperature during operation caninduce cracking of the material which reduces service life. If severe cracking is evident,lightback can become a serious problem. One known cause of cracking is the internalstress due to differential expansion of the ceramic plaque and the burner mounting.Designers must allow for some degree of expansion during appliance operation.

4.3.3.2 Porous medium burners

Burners of this type consist of a porous ceramic block through which the premixedgas/air stream passes. The block may be in the form of either a hard ceramic foam, orsofter ceramic fibres held together by a binder on a support frame 86 , 87 (see Figure 4.22). Burners can be moulded into a variety of shapes; such versatility makes the burnersuitable for a number of uses, especially in the industrial sector 86 − 90 .

The operating characteristics of the porous medium burner are broadly similar to those of the radiant plaque, the optimum aeration being 110% primary air. There are, however,two noteworthy differences.

Firstly, burner port loading cannot be used as a design parameter because there are no discrete ports. Heat input is therefore generally quoted with respect to the total burnersurface area. Commercially available fibre burners 86 , 87 operate at a surface loading of 0.3 to 0.4 W mm−2, which, as for radiant plaques, optimizes radiant efficiency but notnecessarily radiant appearance.

Secondly, the mechanism of lightback differs from burners with discrete flame ports. Owing to the porous nature of the burner material, three combustion regimes can beidentified 88 :

1. ‘Free flame’, where flames burn above the surface of the burner without significantemission of radiant heat.


Figure 4.22 Schematic illustration of the ceramic fibre burner (after Schreiber et al. 86 ) and photograph of burner and .

Figure 4.23 Typical performance of a porous medium radiant burner using methane and propane. (After Coles and Bagge 88 .)

2. Surface combustion, where combustion occurs at or just under the burner surface,raising the latter to incandescence.

3. Unstable interstitial combustion, where the burning velocity is greater than the flowvelocity and the flame propagates, often very slowly, upstream through the porousmedium, resulting ultimately in lightback behind the burner block.

The latter phenomenon has been observed particularly when burning propane 88 ,


which has a higher burning velocity and lower quenching diameter than methane andinduces a higher burner surface temperature. Foam burners, with their larger pore sizesare prone to this form of lightback. Figure 4.23 plots surface temperature against surface loading for stoichiometric methane/air and propane/air mixtures. Coles and Bagge 88

found that lightback is less likely with small pore sizes, i.e. closely packed fibres. It needsto be borne in mind, however, that close packing will increase the pressure needed to passthe required flow through the medium. This is particularly important in domesticapplications, where only low pressures are available.

Porous medium burners, like radiant plaques, are susceptible to damage from mechanical abrasion and excessive burner temperatures. If individual fibres aredislodged, lightback is unlikely unless damage is so severe that a visible crack appears.Even with slight damage, radiant appearance may become patchy.

4.3.3.3 Effect of gas composition

Assuming that plaques and porous medium burners are being used at or near theiroptimum operating point for high radiant output, there should be no occurrence of flamelift when burning any of the limit gases. Lightback should also be absent provided thatthe ports are small enough to prevent any such effect when using G22. Variations in gascomposition will, however, affect burner temperature and therefore radiant output andappearance.

The effect of gas composition on ribbon burners was discussed in Section 4.3.1.6, and it is to be expected that the same trends will be evident for radiant burners. Currentapplications for radiant plaques (mainly space heating) generally use atmosphericinjection, so case (1) in Section 4.3.1.6 will apply (cf. Figure 4.16). A change from G20 to G21 will increase radiant output, which should cause no ill effect unless the maximumworking temperature of the ceramic is exceeded. Conversely, a change from G20 to G23will reduce radiant output. For commercially available plaques, a reduction in radiantoutput of up to 15% may be expected for G23. Radiant appearance is somewhatsubjective, but customer complaints could arise if a satisfactory appearance were notmaintained on low Wobbe number gases.

A similar argument may be presented for burners using air/gas ratio control (cf. Figure 4.17). Under such conditions, G22 will give an increase in radiant output, while G21 willgive a reduced radiant output.

4.3.4 Burner noise

Noise generated by combustion systems has been recognized as a potential nuisance formany years 42 , 43 , 91 , 92 . In general, noise emission from a flow of gas is due toturbulence, which causes local fluctuation in velocity and pressure. Thus, laminar flamesare practically silent, whereas large turbulent industrial burners can be very noisy. Highfrequency noise generated by injectors was discussed in Section 3.2.3; the present section considers the resonant noise problems associated with fully aerated burners. It is intendedas an introduction to a field which is still not well understood either qualitatively or


quantitatively at a theoretical level. Depending on the detail required, further reading isprovided by the references quoted above and by acoustics textbooks. Pritchard et al. 49

and Putnam and Brown 93 also provide a useful introduction to noise generation and suppression.

Two types of noise can be generated within a combustion system. Broad-band noise in the range 125 Hz to 1 kHz can originate in the fan, pipework and in theburner/combustion chamber; noise originating in the latter is often termed ‘combustion roar’. Sound intensity is dependent on the flow velocity and the level of turbulence. Consequently, the precise design of the burner can influence the nature and level oftransmitted sound. In general terms, broad-band noise is minimized by using a largeburner with low flow velocity.

Resonance in a combustion system is heard as high intensity noise at a discretefrequency (usually below 500 Hz). It can be caused by a number of mechanisms which are amplified by the flame itself. Resonance occurs when a pressure fluctuation due toturbulence in the flame or elsewhere coincides with a natural frequency of the system.The sound waves set up propagate in such a manner that the flame itself is excited andliberates more acoustic energy at that frequency. Such a cycle is similar to the positivefeedback principle of an amplifier.

The general characteristics of fully aerated burners often create conditions favourable for resonance. Compact, intense fully aerated flames burning at or just above 100%primary aeration promote pulsations at a much higher frequency than the less intensepartially aerated flames 94 , 95 , and it is perhaps unfortunate that the frequency of such pulsations is often close to that of the combustion chamber or burner manifold. Oneimportant source of turbulence which may exacerbate the problem is vortex-shedding due to shear forces between the moving fluid and a stationary object. This includes not onlyobstructions to flow downstream from the burner, but also formation of vortices as thegas/air mixture emerges from the burner ports into a relatively slow-moving medium 27 .

Two general approaches may be adopted if resonance is encountered. The traditional method involves some form of silencer or filter which can be added to a resonatingsystem 49 . These are often expensive and therefore not generally suitable for low-cost domestic applications. Some examples are discussed in Section 5.3 with particular reference to pulsed combustors. The alternative method, which aims to eliminateresonance at source, involves a theoretical assessment in terms of flow characteristics andthe geometry of combustion chamber and burner ports. If a successful model is produced,then it should be possible to offer design guidelines for burners and combustion chambersin as simple a form as can currently be provided for good flame stability.

4.4 External design variables

Five so-called ‘external’ design variables were introduced in Section 3.5 with regard to partially aerated burners. The effects of manufacturing tolerances and variation in gascomposition on the performance of fully aerated burners have been discussed in Section 4.3 with reference to individual burner types. The other three, vitiation, draughts and


linting, are considered here. The general principles of Section 3.5 remain equally valid for fully aerated burners; consequently, this section consists purely of an extension ofthose principles to fully aerated operation. Readers may wish to refer to Section 3.5before proceeding.

4.4.1 Vitiation

Section 3.5.3 introduced vitiation and its effect on the flame stability of aerated burners. The data presented in Figures 3.22 to 3.24 are applicable to fully aerated burners, and show a clear reduction in the flame lift limit as oxygen concentration decreases. Becausethe burning velocity of the air/ gas mixture is reduced, flames will burn further from thesurface, so burner temperature will decrease and (for radiant burners) radiant output maybe reduced to an unsatisfactory level. In cases of severe vitiation, primary aeration may drop to below 100%, such that flame lengthening will occur and, in a sealed combustionchamber, incomplete combustion is inevitable. Flame quenching by impingement will benoticeable in compact combustion chambers, where the base of the heat exchanger maybe close to the burner, and with radiant plaque burners that use a gauze to enhance radiantoutput.

4.4.2 Draughts

At the time of writing, no work has been published on draught stability of fully aeratedburners. However, the general principles given in Section 3.5.4 can be extended with some confidence, although the detailed guidelines concerning port separations may not bevalid.

There are two general points concerning fully aerated burners which deserve mention: firstly, the port loadings used with fully aerated burners are relatively low and might beexpected to decrease stability in draughts; secondly, fully aerated burners that are proneto lift will also be susceptible to draughts. Thus, a ribbon burner or punched metal burnercan be expected to show poor draught resistance unless it incorporates some form offlame retention. The flames on radiant burners, which operate well within the lift limit,are generally under momentum-control and can be expected to be quite tolerant ofdraughts, although radiant output may be affected.

4.4.3 Linting

Linting of fully aerated burners could be a major problem because of the much greater airflow rates through the burner. However, there are three design features which should helpguard against serious deterioration in appliance performance:

1. Fanned draught appliances will not be fitted with injectors, thus eliminatingcomponents especially susceptible to lint.

2. Fans will usually be fitted with a filter, so burners in forced draught appliancesshould be lint-free, although the filter will require regular cleaning or replacement.

3. Current trends towards room-sealed appliances, especially boilers, should minimize


ingress of foreign particles and fibres. Any linting which does occur will have the same effect as with partially aerated

burners, i.e a reduction in primary aeration resulting eventually in incompletecombustion. With radiant burners, linting can reduce radiant output and, hence, radiantappearance.

4.5 Summary

This chapter has covered various design considerations for fully aerated burners. Becausesuch burners are still under development for use in domestic appliances, it has beenimpossible to cover every facet of fully aerated burner operation. However, from thediscussion above, it should be possible to construct a burner which gives satisfactoryperformance over a wide range of conditions. The special requirements of individualappliances are covered in Chapter 6.


CHAPTER 5 Future domestic burners

5.1 Introduction

The foregoing three chapters have dealt with what may be termed ‘conventional’ burners, that is burners or adaptations thereof that have been used or may soon become a featureof British domestic gas appliances. While most burners generally perform very well,there is a continuing need to examine new methods of gas utilization in order to improvecombustion performance, thermal efficiency, reliability and installation flexibility.

Although concerned with ‘future’ burners, this chapter involves concepts that in some cases have been known for many years. Their application to domestic natural gas systemsis, however, only at a developmental stage in Great Britain. Three types of burner areconsidered here: catalytic combustion, pulsed combustion, and fluidized combustion. Thechapter is intended to be only a brief outline of burner systems. Full theoretical andpractical design details are provided in the references quoted throughout the text.

5.1 Introduction 5.2 Catalytic combustion 5.2.1 Principles of operation 5.2.2 Performance 5.2.3 Future developments 5.3 Pulsed combustion 5.3.1 Principle of operation 5.3.2 Pulsed combustor design 5.3.3 Noise suppression 5.3.4 Future development 5.4 Fluidized combustion 5.4.1 Principle of operation 5.4.2 Design considerations 5.4.3 Future development 5.5 Other unusual burners

5.2 Catalytic combustion

All burners considered so far have been characterized by a flame and a release of energywhich raises the gas temperature to about 1800 °C. It is also possible to burn gas flamelessly at much lower temperatures (500 °C) with the same heat release on a catalytically active surface. Space heaters working on this principle, using liquefiedpetroleum gas, have been commercially available for many years and are in widespreaduse as mobile heaters in Europe and North America. Combustion of natural gas has, untilrecently, proved difficult because of the nature of the fuel and problems with the catalyst.

5.2.1 Principles of operation

Like conventional burners, catalytic heating elements (known as ‘pads’) may operate with a supply of either neat gas or a gas/air mixture. Outside the USA, developments fordomestic appliances 85 , 96 , 97 have concentrated on the first of these, that is diffusive catalytic combustion; only this form of catalytic combustion is considered here.

Figure 5.1 shows a schematic cross-section of a typical diffusive catalytic heating element. Gas is fed to the back of the housing and diffuses through a fibrous pad onwhich the catalyst particles have been evenly distributed. Combustion air is supplied tothe gas by diffusion from the surrounding atmosphere into the front layer of the pad.Combustion is not spontaneous but needs to be initiated by preheating the pad to about250 °C. An electrical heating element (embedded in the pad and requiring an external power supply) or a traditional pilot flame may be used for this purpose. Above 250 °C, chemical reaction is self-sustaining; typical surface temperatures during operation are in the range 400 °C to 500 °C. Theoretical modelling of the chemical and diffusionprocesses is beyond the scope of this publication, but has been discussed by Dongworthand Melvin 98 .

Catalytic pads are generally made from a fibrous material of high porosity, such asasbestos wool, ceramic wool, glass fibre and, more recently, alumina fibre 96 , 97 . Catalyst compositions and details of the deposition processes are in many casescommercial secrets, but catalysts usually include 85 precious metals (such as platinum, palladium and rhodium) or transition metal oxides (such as those of cobalt andmanganese). Of these, platinum-containing catalysts have found greatest favour.


Figure 5.1 Schematic illustration of the construction of a diffusive catalytic heater. (From Radcliffe and Hickman 96 .)

5.2.2 Performance

There are a number of advantages offered by catalytic combustion. High radiantefficiencies (over 50%) are achievable 96 (cf. up to 35% for gas fire box-radiants and surface combustion radiant plaques), and emissions of gaseous pollutants (carbonmonoxide, nitrogen oxides) are very low, even under highly vitiated conditions. Becausecombustion is flameless, flame

Future domestic burners 111

Figure 5.2 Methane combustion efficiency as a function of surface loading for a number of catalytic pads when new. (From Radcliffe and Hickman 96 .)

stability considerations are not relevant and the pads are extremely resistant to draughtsand changes in gas composition 99 . A possible aesthetic drawback, however, is thatduring operation, the catalytic pads remain somewhat drab in appearance; Britishconsumers generally seem to prefer a bright radiant glow for focal point space heaters.

The performance of catalytic heaters is usually measured in terms of the fraction of fuel gas that is consumed during the combustion process. It is acknowledged that fuel gascan ‘slip’ through the pad without being burned. Combustion efficiency (ε) for fuel gas X may be defined as

(5.1)


where brackets indicate volume concentration in the combustion products. The degree ofslippage (100−ε) depends on the fuel, pad material, gas flow rate and the composition andage of the catalyst. Figure 5.2 presents data of Radcliffe and Hickman 96 who measured methane combustion efficiency of eight pads with platinum catalyst as a function of gasflow (heat input). For each pad, there is an optimum heat input at which combustionefficiency is at its greatest. The best performer was 95% efficient, but peak efficiencies ofjust 80% were common. Even though such fuel slippage would be unlikely to provehazardous, the waste of fuel is unacceptable. Methane is the most difficult gaseoushydrocarbon fuel to burn catalytically; combustion efficiencies of over 95% can beachieved using LPG 96 .

Radcliffe and Hickman 96 also reported a marked deterioration in methane combustionefficiency with prolonged operation. Figure 5.3 shows the decrease in methane combustion efficiency with use. Dongworth and Melvin 98 have attributed this to poisoning of the catalyst by deposition of carbon particles resulting from the thermalcracking of the fuel. Such carbon deposition can screen the catalyst sites from the flow ofgas, thereby deactivating the catalyst.

Figure 5.3 The variation in methane combustion efficiency with age of a pad using a platinum-based catalyst. (From Pearson 99 .)


More recent catalyst development in Japan 97 suggests that it is possible to produce a catalyst and associated pad with a methane combustion efficiency of 98% and with very little deterioration with use. The new catalyst gives higher rate constants for methanecombustion at a given temperature, better dispersion of the catalyst particles on the pad,and carbon deposition was not observed even after 5000 hours in use (Figure 5.4).

Figure 5.4 The variation in methane combustion efficiency with surface loading and age of pad with the improved Osaka Gas catalyst. (From Sadamori et al. 97 )

5.2.3 Future developments

Development of British catalytic heaters for natural gas has received low priority duringrecent years, despite the advantages outlined above. Of the drawbacks mentioned, theneed for an external electrical supply should not be seen as a problem; the current trend isfor more electrical hardware in gas appliances (e.g. fanned draught space, heaters).Assessment of appearance during operation is highly subjective and should not precludefurther development. On a technical level, the future for catalytic combustion of methaneseems to rest largely on the development of catalysts that give very low slippage rateseven after many years use. The latest Japanese work suggests that this is possible,


although at this stage an assessment of the economic viability of producing high costcatalysts is not available.

5.3 Pulsed combustion

Pulsed (or pulsating) combustion has interested engineers for several decades. The firstsystems were developed and patented during the 1900s, but were not generallyconsidered to be practicable. However, in the late 1920s, aeronautical engineers realizedthat pulsed combustion could be used as a method of aircraft propulsion. Work inGermany by Schmidt led directly to the V-1 flying bomb, arguably the best knownapplication of pulsed combustion. Pulse-jet engines were also developed until about 1950in the United States, where much valuable work was published by Reynst 100 . As superior jet engines became available, pulsed combustion disappeared from aeronautics.However, industrial engineers took up the principle, and a number of pulsed combustorswere designed, mainly for burning pulverized coal. The gas industry has been activelyinvolved in pulsed combustion since about 1960, when prototype high efficiency boilersand water heaters were developed in the United States 85 , 101 .

The major drawback of pulsed combustion in domestic applications is that the process is excessively noisy. However, recent interest in energy conservation has prompted areappraisal both of the process itself and of methods of noise suppression, particularly inthe United States (e.g. refs. 93, 102–105).

5.3.1 Principle of operation

A typical pulsed combustor consists of a combustion chamber with an open tailpipe atone end and a valve assembly at the other which controls entry of fuel and air into thechamber. Valves may be in the form of either mechanical flaps or an aerodynamic(fluidic) device, which has no moving parts and behaves like a one-way valve. The latter type of valve permits larger flow rates of gas and air, but flapper valves have generallybeen preferred because of the better control of gas and air flow 101 , and it is this type of pulsed combustor that is discussed here. The tailpipe is usually formed into or contains asuitably shaped heat exchanger.

Two simple geometries of pulsed combustor are commonly referred to (Figure 5.5): firstly, the Schmidt burner, where the combustion chamber and tailpipe are of the same diameter such that the whole system resonates like an organ pipe with the source at oneend; secondly, the Helmholtz burner, where the combustion chamber is of much largerdiameter than the resonance tube. Although the acoustic and heat transfer properties ofthe two burner types are different the principle of operation is the same.

Figure 5.5 Schematic illustrations of (A) Schmidt and (B) Helmholtz pulsed combustors. (Redrawn from Griffiths et al. 85 )


Figure 5.6 shows a schematic illustration of the pulsed combustion operating cycle. During start-up, air (from a fan) and gas are admitted into the combustion chamber(Figure 5.6A). Air and gas streams may be separate, or they may be mixed before entry into the chamber. The gas/air mixture is ignited, usually by a spark plug. Ignition createsa positive pressure wave which closes the air and gas flaps and forces the combustionproducts along the tailpipe (Figure 5.6B). A negative pressure (relative to the incoming mixture) is then induced in the chamber, the flaps reopen and a new charge of air and fuelis admitted (Figure 5.6C). A small proportion of the hot combustion products will also re-enter the chamber from the tailpipe, if the unit is designed correctly. It is thought thatthese hot gases initiate combustion of the new gas/air mixture without the need for aspark, so that the cycle is repeated. The frequency of the cycle (typically in the range 50Hz to 100 Hz) is dependent on the resonant frequency of the system, which in turndepends on the shape of the combustion chamber, dimensions of the tailpipe, the velocityof sound in the exhaust gases and the burning velocity of the gas/air mixture. The last twothemselves vary with the pressure and temperature within the combustor. Whenresonance is established, the combustor is self-powered, that is the starting fan and sparkgenerator may be removed, and combustion continues provided that the valve designadmits the correct air/gas ratio.

Figure 5.6 Schematic illustration of the operating cycle of a pulsed combustor. (From Griffiths et al. 85 )


The pulsed combustion boiler possesses many distinct advantages when compared with a conventional system. The burner is relatively simple in construction and it is self-powered, requiring no external power (except for initial ignition) for either combustion orflueing. The whole system is compact, gives a high rate of heat release per unit volume,and very high convective heat transfer in the tailpipe because of the scrubbing action ofthe oscillatory flow. The major drawbacks are noise emission, the limited turndowncharacteristic, and poor performance using high burning velocity gases.

Subsequent sections here consider various design parameters and methods of suppressing noise.


5.3.2 Pulsed combustor design

A complete design guide for pulsed combustion is beyond the scope of this publication.This section is intended to be an indication of the parameters which need to be consideredby the designer. Far more detailed quantitative guidelines are given elsewhere forcombustion chamber design 101 , 104 , 106 , aspects of heat transfer and efficiency 104 , 106− 108 , and the problems of noise generation and suppression 103 − 105 , 109 , 110 . The last is also considered below in Section 5.3.3.

The gas input rate will be determined by the intended use of the appliance. The dimensions of the combustion chamber and the gas valve must then be designedaccordingly. The air flow rate is determined by the air valve design and the pressureregime in the combustor. Francis et al. 106 provide guidelines for gas and air inletdiameters for both Schmidt and Helmholtz combustors. The designer must ensure thatsufficient air is admitted for combustion to be complete. A design point of 20% to 30%excess air is recommended; although complete combustion will occur with less than 10%excess air 85 , an allowance must be made for variations in gas supply and manufacturingtolerances.

The pressures generated within a combustor for a given heat input are dependent on chamber size and tailpipe diameter, whereas these two parameters and tailpipe lengthdetermine the operating frequency 101 . In general, the smaller the chamber and tailpipediameter, the greater the pressure generated. In most instances, tailpipe length (and henceoperating frequency) has little effect on pressure for a given gas input. However,combustors with short tailpipes (frequency in the range 150 to 200 Hz) are usually moredifficult to start than those with longer, lower frequency tailpipes. For a given chamberand gas rate, there is a minimum tailpipe length below which the combustor may notoperate satisfactorily.

The major advantage of pulsed combustion is that a high thermal efficiency is achievable with a compact appliance. The pressures generated can be used to force hotgases through more restrictive heat exchangers than are used with conventional burners.The oscillatory flow also gives heat transfer coefficients up to three times greater than forsteady flow 111 . Consequently, pulsed combustors can extract more heat from the combustion products without the need for large heat exchangers. It must be remembered,however, that for efficiencies of greater than 85%, some provision must be made fordisposal of condensate formed in the tailpipe. This will also influence the range ofmaterials that may be used.

A significant disadvantage with pulsed combustors is the small turndown ratio that is attainable. For both types of pulsed combustor, an increase in gas flow rate automaticallyincreases the air flow rate because of the greater negative pressure peak duringcombustion 106 . The system is, therefore, to some extent self-proportioning. However, the air rate increase usually fails to match the increase in gas rate, such that excess airdecreases and combustion may become incomplete. Conversely, if the gas rate decreases,the air rate will also decrease, although not enough to prevent the excess air level fromrising. Eventually, the gas/air mixture becomes too fuel-lean to sustain combustion and


the burner shuts down. Because of problems with turndown, current commerciallyavailable pulsed combustors generally operate at a fixed gas rate. Some authors have,however, claimed turndown ratios of between two and five to one with laboratorycombustors 101 , 106 .

5.3.3 Noise suppression

Because successful operation of a pulsed combustor requires resonance to be established,noise emission is an inherent major problem, which has precluded the use of pulsedcombustion in British domestic appliances. Current work on noise emission covers twoareas: theoretical aspects of noise generation and oscillatory combustion, and practicaldesign guidelines for the reduction of noise emission. The simplest method of noisesuppression is to surround the combustor with an acoustically insulating material. Thereare two other methods of noise reduction which attempt to control noise emission atsource: acoustic decouplers and reactive silencers.

5.3.3.1 Acoustic decouplers

An acoustic decoupler is essentially a filter in the form of a chamber placed in the outletor inlet pipe (see Figure 5.7) which attenuates sound above a critical frequency and transmits sound below that frequency. The critical frequency is determined by thedimensions of the decoupling chamber and the mean flow rate 102 , 103 . Decouplers should be designed so that the decoupler frequency is about half of the fundamentalfrequency of the unsilenced combustor 105 .

Figure 5.7 Schematic illustration of a Helmholtz combustor with an acoustic decoupler in the tailpipe.


5.3.3.2 Reactive silencers

In a reactively silenced system, provision is made for the generated wave pattern to beacoustically cancelled by a similar wave pattern which is 180° out of phase with the original. Chaplin 112 gives an introduction to the principles and applications of noisecancellation (anti-noise generation). Three methods are outlined here (see Figure 5.8):

1. A tuned sidebranch is designed to reflect a sound wave so that it cancels asubsequent incident wave. This is achieved with a sidebranch length of a quarter thefundamental wavelength.

2. The quinc tube splits the flow in the tailpipe, thereby increasing the path length forpart of the flow. If the extra path length is equal to half the fundamental wavelength,antiphase cancellation will occur at the downstream junction.

Figure 5.8 Schematic illustrations of three methods for reactive silencing of pulsed combustors: (A) quarterwave side branch; (B) quinc tube; (C) paired pulsed combustors.

3. Two identical pulsed combustors operating at 180° out of phase and connected to a common tailpipe should destructively interfere and reduce noise emission. Briffa et al.111 found that paired Schmidt combustors give better noise reduction than paired Helmholtz combustors. Sran and Kentfield 104 have reported a 20 dB reduction in sound


levels from paired Helmholtz combustors. All three methods have drawbacks. The quarter wave sidebranch and quinc tube are

both tuned to a precise frequency. If the fundamental frequency changes because of variation in combustion conditions, noise suppression may become ineffective. The extrapipework can also occupy considerable space within an appliance, especially for low-frequency (long-wavelength) combustors. Paired combustors are difficult to synchronizebecause of slightly different operational characteristics of nominally identicalcombustors.

5.3.4 Future development

Pulsed combustion offers compact high-efficiency gas appliances which can produce significant fuel savings. The major drawback is that noise emission is considered toogreat for domestic gas appliances. However, current work is providing both theoreticaland practical guidance for future pulsed combustor design. With the current interest inenergy conservation, there is considerable scope for further developments in quieterpulsed combustion systems for domestic use.

5.4 Fluidized combustion

Fluidized combustion is a low-temperature (800 °C to 1000 °C) flameless combustion process which is capable of high thermal efficiency within a fairly compact appliance.Until recently, the process has been used chiefly in the industrial sector (cf. pulsed combustors), but it is now being considered for the domestic market 99 . Advantages of fluidized combustors are low pollutant emissions, tolerance of a wide range of gascompositions, and very good stability at high excess air levels. Disadvantages includenoise emissions, methods of ignition, flame sensing (particularly on startup), erosion ofthe heat exchanger, and manufacturing costs.

This section describes the principles of operation and design considerations for satisfactory performance of a fluidized combustor. Davidson and Harrison 113 , 114

provide comprehensive treatments of fluidized processes, while a basic theoreticalintroduction is given by Kay and Nedderman 67 .

5.4.1 Principle of operation

Figure 5.9 shows a schematic cross-section of a fluidized bed, which comprises a bed ofsand particles or similar refractory material resting on a porous distribution plate. Thegas/air mixture is introduced below the distributor and passes through the bed. At lowflow rates, the bed is not disturbed by passage of gas. As flow rate increases, the pressuredrop across the bed increases until it becomes equal to the weight of particles per unitcross-section area. The particles can then be supported by the flow and the bed is said to be fluidized. The velocity at which this occurs is termed the minimum fluidizing velocity


(u mf). Further increases in flow rate have little effect on pressure drop across the bed, but above a critical velocity (the terminal velocity, u t), particles can be ejected (elutriated) from the top of the bed. When the fluidizing condition is attained, the bed expands and

Figure 5.9 Schematic illustration of a fluidized bed.

assumes the appearance of a boiling liquid, hence the term fluidization. If the gas/air mixture is ignited above the bed, the particles are gradually heated. At

about 500 °C, combustion occurs within the bed itself. Combustion at this stage is somewhat sluggish and is accompanied by explosions from bubbles within the bed. Thestarting sequence is excessively noisy for a domestic burner and not easily controlled. Atthe optimum operating temperature (about 900 °C), combustion noise is much reduced and the bed glows uniformly at red heat. Satisfactory performance requires the bed to bemaintained in the range 800 °C to 1000 °C. Below 800 °C, combustion can become noisy and efficiency may deteriorate. At temperatures above 1000 °C, particles may fuse together and bed performance will be seriously impaired.

5.4.2 Design considerations

The task of the burner designer is to achieve and maintain fluidizing conditions and toensure that the bed temperature is within the permitted range. The gas/air velocity isusually chosen to be about twice the minimum fluidizing velocity for the cold bed. This isusually well below the terminal velocity, so elutriation should not be a problem unlessparticle attrition and formation of fines occur, in which case dust emission will be


evident. Although u mf decreases sharply as the bed temperature increases, it is notnecessary to modulate the flow rate when steady combustion is achieved.

The bed depth determines the pressure drop for a given flow rate, and must therefore be matched with the power source for the supply of air (usually a small fan). A bed depthof a few centimetres is usually satisfactory. The bed also needs to be cooled in order toensure that the temperature does not exceed 1000 °C. In domestic boiler applications, a cooling water jacket or a heat exchanger immersed within the bed would extract surplusheat for use in the heating system. Material selection is critical, because the agitatingparticles can cause severe erosion of the heat exchanger surface. About half of the totalheat generated during combustion maybe removed in this way. Additional removal ofheat from the bed may cause the bed temperature to drop below 800 °C, thereby leading to incomplete combustion and possibly extinction. Further heat may be extracted fromthe hot combustion products by placing a second heat exchanger above the bed (Figure 5.10). This may take the form of a non-combusting fluidized bed surrounding water-cooled tubes. Heat transfer coefficients are very high in fluidized beds, thereby permittingcompact, highly efficient appliances.

Figure 5.10 Schematic two-stage fluidized combustor.


There remains the problem of noise emission. Because there is no equivalent of the fundamental frequency emitted by a pulsed combustor, noise is far more difficult tosuppress systematically in a fluidized combustor. At present, little attention appears to bedevoted to this aspect of operation.

5.4.3 Future development

Like pulsed combustion, fluidized combustion can offer high thermal efficiency andcompact appliances. The problems of noise and the sluggish starting cycle requiresolution before domestic fluidized combustors could be considered acceptable. Inaddition, production costs are likely to be higher than for pulsed combustors, with whichfluidization will inevitably be compared. The system will almost certainly be developedfurther, particularly industrially, but at the time of writing commercial viability in thedomestic sector remains questionable.

5.5 Other unusual burners

Many experimental designs of gas burner have been proposed and developed byengineers in attempts to improve appliance design and performance. Although patents areissued for a great number of designs, they have met with varying degrees of success, andmost are not feasible for domestic gas appliances.

Figure 5.11 A submerged combustor.


Some differ only in geometry from designs already discussed. Of particular note areburners which use impinging jets of gas and air. Such forced mixing greatly reducesdiffusion time and increases the rate of combustion, thereby producing much shorterflames. Partially aerated burners, with the secondary air supplied by enforced mixingrather than diffusion, can give flames that are almost as compact as those on a fullyaerated burner, and give lower emissions of carbon monoxide than a conventionalpartially aerated system. The spiral burner 85 is a development of this principle, andmaintains complete combustion down to 5% primary aeration and 10% total excess air.

In an attempt to increase efficiency, methods of direct contact between combustion products and the heat transfer fluid have been assessed. Submerged combustion is anexample of this concept (Figure 5.11). A fully aerated burner contained within acylindrical tube fires downwards so that the combustion gases are forced through areservoir containing the heat transfer fluid (e.g. water). As well as increased efficiency ofheat transfer, the burner is relatively simple and cheap to manufacture. Disadvantages atpresent are the complex and expensive control system required, the gradually increasingacidity of the circulating water, and the high pressures required to force the combustiongases through the fluid. At present, submerged combustion is most suited to industrialand commercial applications 49 .

Two other novel systems currently under consideration for domestic applications are the gas engine and the fuel cell. Gas engines have been used industrially for many yearsas part of total energy schemes 49 . Engines are now being assessed in Great Britain andabroad for domestic gas-fired heat pumps 115 based on either the internal combustion orStirling cycles. Disadvantages at present include noise, exhaust gas emissions (CO, NOx,etc.) and the likely cost of the system. However, there is considerable scope for furtherimprovement.

The fuel cell is essentially a device which produces electricity from a fuel without needing a thermal intermediate such as a boiler or other generator. It operates like aconventional electrochemical cell, where reduction and oxidation reactions occur at therespective electrodes. For hydrocarbon/air fuel cells, methane is convertedelectrochemically to carbon dioxide at the anode, and oxygen reacts to form water at thecathode. Fuel cells are characterized by the electrolyte and electrodes used. Electrolytesare typically strongly acidic (e.g. phosphoric acid) or alkaline (e.g. molten carbonate orhydroxide). Because CO2 will react with an alkaline system, hydrocarbon fuel cellsusually employ concentrated phosphoric acid with platinum electrodes. The chiefdisadvantage of fuel cells at present is the cost of manufacture, and, for hydrocarboncells, the need for a commercially satisfactory catalyst.

Gas engineers will undoubtedly continue to explore new combustion methods or adaptprevious designs for new applications. Despite the best of intentions, the success of anyconcept is ultimately determined by the extent to which the design is used incommercially marketable appliances.


CHAPTER 6 Burners in appliances

6.1 General

Each of the different types of burner described in the last four chapters is used in variousforms and configurations in order to suit particular applications. For example, althoughboth gas fires and cookers are generally fitted with partially aerated burners, thedimensions, shape and general design of each burner reflects the usage requirements ofeach appliance.

Previous chapters have considered how flames can be stabilized on a burner in terms of burner port dimensions, port loadings, aeration, etc. The final design is also dependent onlimitations imposed by the appliance itself (e.g. geometry, controls, means of ignition andturndown) and by appliance performance and safety requirements (as laid down bynational Standards authorities). This chapter discusses British requirements that arerelevant to the burner designer and outlines the design features and accepted practice forburners in appliances. Where relevant, a brief comparison is made with accepted practicein other countries.

6.1 General 6.2 British Standards requirements 6.3 General appliance characteristics 6.3.1 Carbon monoxide emission 6.3.2 Nitrogen oxide emission 6.4 Individual appliance requirements 6.4.1 Boilers and water heaters 6.4.2 Individual space heaters 6.4.3 Cookers 6.4.4 Other appliances

6.2 British Standards requirements

Although not legally required to do so, most manufacturers of gas appliances in GreatBritain submit their appliances for certification according to the relevant BritishStandards 116 – 118 . Appliances are tested under a number of operating conditions forsafety of operation (including details of construction, electrical safety, combustionperformance) and thermal efficiency. For full details of test procedures and requirements,the reader should consult the appropriate British Standard. It is, however, worthsummarizing here those tests which are of particular concern to the burner designer. Theprecise requirements depend on the appliance and the fuel; any figures quoted below arefor family 2H gases and should be used only as a guide.

1. Gas rate. Using reference gas G20 at normal operating conditions, the gas consumption of a burner shall correspond with the manufacturer’s declared (rated) heat input within certain limits of tolerance. For most appliances, the limits of tolerance are±5%, but can be as wide as ±10% for boilers and refrigerators that are fitted with gasgovernors.

2. Combustion. For most appliances, the combustion ratio (CO/CO2) shall not exceed 0.02 using G20 for heat inputs between the minimum operational rate (the definition ofwhich depends on the appliance) and an overload condition (usually 115% or 120% ofthe rated heat input). For European and American Standards, combustion performancerequirements are linked to the concentration of carbon monoxide in the combustionproducts. For all appliances, combustion must remain satisfactory when (a) G21 is usedin place of G20 (with no change in operating burner pressure), (b) G20 is used at ratedheat input under prescribed adverse flueing conditions (e.g. wind at terminals of balancedand closed flue systems, down-draughts in open flues).

3. Flame stability. No lightback or flame lift shall be permitted when using G22 or G23 respectively. The range of gas rates for which this requirement must be met depends on the appliance. Flames must also remain stable using G20 when subjected to cross-draughts and, for gas fires, excessive up-draught or down-draught in the flue.

4. Thermal efficiency. Appliance thermal efficiency is largely determined by the sizeand construction of the heat exchanger, the design of which is beyond the scope of thispublication. However, minimizing the excess dilution air in the combustion chamber willmaximize the amount of heat that can be extracted from the combustion process. Mostappliances are required to operate at or above a prescribed minimum thermal efficiency.

6.3 General appliance characteristics

Within the framework already discussed for flame stability and BS requirements, it ispossible to list a number of criteria which need to be met by all gas appliances:

1. Uniform flame distribution and height, giving an even heat distribution to that whichis being heated.

Burners in appliances 127

2. Complete combustion, subject to the carbon monoxide and soot emission clauses inthe appropriate Standard.

3. No flame lift or lightback within the full range of operating conditions. 4. No excessive noise, particularly during the ignition or extinction sequence. 5. Materials used and the standard of construction should be such that burner lifetime is

acceptable (at least twelve to fifteen years). Additionally, there is currently much concern regarding emission of nitrogen oxides

(NOx) into the indoor and outdoor environment. Although the major global sources of outdoor NOx are power stations and automobile exhausts, gas appliances and cigarette smoke can make a significant contribution in the domestic environment 119 , 120 . With air quality and emission Standards currently or soon to be in force in many countries, manydesigners are now developing ‘low NOx’ burners.

The following sections consider appliance CO and NOx emission characteristics in general terms; they provide design points that give satisfactory combustion under normaloperating conditions and a safety margin which allows for adverse operating conditions.The individual requirements of domestic appliance burners are discussed in Section 6.4.

6.3.1 Carbon monoxide emission

One method of depicting the carbon monoxide emission of an appliance graphically is asa plot of CO/CO2 emission ratio against excess air. Figure 6.1 shows such a plot for typical partially and fully aerated burners. The curves presented should be taken only as aguide; the exact characteristic for any burner is greatly dependent on the design of thewhole appliance. It is, however, worth noting the observed trends. As excess air isdecreased, a point must be reached where combustion deteriorates owing to oxygenstarvation. This can be caused either by a general lack of oxygen or by poor mixing in theprimary gas/air stream leading to a localized oxygen deficiency. Because partially aeratedburners rely on diffusion of secondary air into the flame for completion of combustion,the critical excess air level is somewhat higher (~ 15%) than for fully aerated burners.With the latter, the more effective forced mixing can maintain satisfactory combustiondown to less than 10% excess air.


Figure 6.1 Combustion ratio as a function of burner excess air for fully and partially aerated burners.

At high excess air levels, combustion performance is critically dependent on the natureof the burner. Partially aerated burners using atmospheric injection show little or nodependence on excess air above about 40%. Primary aeration and the flow rate throughthe burner will remain constant, such that flame stability should not be affected. Only at afew hundred per cent excess air does incomplete combustion due to aerodynamic flamequenching occur.

For fully aerated burners without injection, all the air supplied is transported as primary air through the flamestrip. Consequently, any increase in excess air will increaseprimary aeration and, above a certain level, will result in a decrease in flame stability.Flame lift will be accompanied by flame quenching on the relatively cooler heatexchanger and emission of carbon monoxide. The point at which such an effect isobserved will depend on the design parameters discussed in Chapter 4. For a typical optimized burner with flame retention, combustion performance may be expected todeteriorate above about 60% excess air (cf. Figure 4.15). This upper limit of satisfactory operation will depend on port loading, direction of firing, burner type, etc. (see Chapter 4).

Fully aerated burners that use atmospheric injection can be considered to perform somewhere between the two extremes. Because primary aeration remains constant, thereis little change in flow rate through the burner as excess air increases, and consequentlyno serious occurrence of flame lift. Because combustion is complete without the need for


secondary air, any secondary air that is introduced merely acts to quench the flame. As aresult, the onset of poor combustion may be at a somewhat lower excess air level than fora partially aerated injected burner.

Figure 6.2 The effect of wind pressure, gas rate and fan voltage on the operating excess air for a centrifugal fanned draught partially aerated closed-flue boiler.

A combustion diagram in the form of Figure 6.1 defines an optimum operating point in terms of excess air. In order to maximize thermal efficiency, the design excess air shouldbe as low as can be achieved without exceeding the critical emission ratio, both atnominal operating conditions and after taking into account the various limits of tolerance.For a partially aerated burner, the minimum is about 45% excess air. This allows for thevariation of excess air that can be expected under such conditions as gas overload, changein gas composition and, with fanned draught appliances, variation in fan voltage and theeffect of wind pressure at the flue terminal.

Typical data for a closed-flue boiler fitted with a centrifugal fan are shown in Figure 6.2. It can be seen that if a lower excess air level than about 45% was chosen for nominal operating conditions, then there would be a risk of incomplete combustion at gasoverload or high terminal wind pressures. Balanced-flue forced draught appliances are specifically designed to minimize the effect of wind on combustion; gas rate variationthen becomes the determining factor in setting the operating point. Natural draughtappliances (both balanced and open flue) are also designed to withstand wind effects andare inherently self-proportioning such that the rate of air entrainment varies in order to


compensate for any change in gas rate. They do not, however, compensate for variationin gas composition.

Figure 6.3 The effect of wind pressure, gas rate and fan voltage on the operating excess air for a centrifugal fanned draught fully aerated closed flue boiler.

Figure 6.3 shows a similar plot for a fanned draught fully aerated burner. Because afully aerated burner is more tolerant of low excess air levels (cf. Figure 6.1), the design point can be decreased to about 35% excess air without the risk of incompletecombustion at high wind pressures or gas overload. The susceptibility to lift at increasedexcess air levels suggests that fanned draught fully aerated burners may not be able tosatisfy the present underload requirement.

6.3.2 Nitrogen oxide emission

The mechanism of NOx formation remains far from clear, despite receiving a great deal of attention in recent years (see review by Gaydon and Wolfhard 3 ). In general terms, NO may be formed by two mechanisms. Firstly, by the Zeldovich mechanism, which isthe slower of the two at the flame temperatures that are of interest here:


Secondly, in fuel-rich hydrocarbon flames, by the fast ‘prompt’ mechanism, which, though still unconfirmed, is believed to involve hydrocarbon radicals in the formation oflarge quantities of O and N radicals in the reaction zone 121 , for example:

Whatever mechanism operates, NO may be oxidized to NO2, either by direct combination with an O atom involving a third body collision, or by oxygen abstraction from speciessuch as HO2 122 :

The foregoing mechanistic summary suggests that formation of NO (and hence totalNOx) is greatest at high temperature and with fuel-rich conditions (i.e. low primary aeration). Figure 6.4 illustrates the effect of primary aeration on NO and NO2 emissions from a Bunsen-type flame. The diagram shows that NO2 emission is virtually independent of primary aeration, despite the variation in flame temperature as aerationchanges. NO varies considerably, decreasing steadily between 0% and 100% primaryaeration and dropping sharply as the flame becomes fully aerated. Although maximumflame temperature is at about 100% primary aeration, the prompt NO mechanism appearsto dominate, resulting in highest NOx emissions at low primary aeration. While absoluteconcentrations will vary according to burner configuration, the general trend observedhere is common to most ‘traditional’ burners. Novel burners, such as fluidized combustors and catalytic heater pads, which operate at much lower temperatures, showdependence on temperature, such that NOx emission with those systems is very low.

This discussion, albeit brief and somewhat simplistic, shows that NOx emissions can be reduced by using lower temperature combustion and high primary aeration (preferablyover 100%). An example of this is the use of fully aerated radiant plaque burners, whichhave very low NOx emissions and are used extensively in flueless radiant heaters both in Great Britain and abroad. Improvements can be made to partially aerated burners byusing higher primary aeration, but attention needs to be paid to the effect on heatdistribution (owing to the change in flame shape and temperature) and noise levels in theburner resulting from the increased flow rate.

6.4 Individual appliance requirements

The very different applications of domestic gas burners (e.g. cookers, boilers) arereflected in the variety of burner geometries and configurations used in gas appliances.All these burners are constructed according to the same design guidelines and are subjectto similar performance Standards.


Figure 6.4 Nitrogen oxides emission from a Bunsen-type flame as a function of primary aeration.

This section covers aspects of designing burners for particular appliances and discussescurrent examples of these designs. Jasper 123 gives a thorough general guide to appliance construction.

6.4.1 Boilers and water heaters

Although boilers and water heaters perform different functions, there are manysimilarities in design and it is appropriate to consider them together, Both appliance typescontain a heat exchanger consisting of water-filled tubes or panels to which heat istransferred from the hot combustion gases. The heat exchanger has often been regardedas the central part of the design, perhaps to the detriment of the burner. Ideally, the burnerand heat exchanger should be designed as one unit in order to maximize efficiency and


compactness, maintain satisfactory combustion, provide an adequate turndown ratio, etc.

6.4.1.1 Choice of heat input

The first consideration in any new design is the heat output that is required.Consequently, the first design stage is to decide the rating of the heat exchanger andburner. Domestic central heating boilers, which provide stored hot water and spaceheating, are generally rated at between about 8 kW and 60 kW heat input. Ifinstantaneous domestic hot water is required, a heat input of at least 20 kW is usuallyrequired in order to provide an acceptable flow rate from bathroom taps. Lower ratedunits (~10 kW) are suitable for a kitchen sink. Attention must be paid to the need, if any,for modulation of the gas rate. Such a feature is becoming increasingly necessary withboilers because of the widening gap between hot water and space heating requirements.Current energy conservation measures are resulting in highly insulated buildings whichhave a much lower space heating demand but the same hot water requirements.Installation of oversized boilers can lead to excessive on/off cycling of the boiler or roomthermostats which in turn may induce premature failure of the primary control system. Ifmodulating controls are incorporated, the burner must be capable of satisfactoryoperation at full and reduced gas rates; this will be taken into account by testingaccording to the appropriate Standards.

6.4.1.2 Choice of burner

Design parameters discussed in previous chapters relate to port dimensions and burnerport loading. The overall physical dimensions of the burner depend to a large extent onthe size and shape of the combustion chamber and heat exchanger. The burner mustsupply a uniform heat distribution to the heat exchanger in order to prevent localoverheating or underheating of the heat exchanger surface. A good temperaturedistribution maximizes thermal efficiency, while local overheating may acceleratedeterioration of the heat exchanger materials.

The most popular types of partially aerated burner for boilers and water heaters are box burners (Figure 6.5), which are inexpensive to manufacture and can be easily adapted tothe dimensions of the heat exchanger. They may take the form either of a single box withone injector and flamestrip, or of an array of smaller burners, usually mounted in parallelbelow the heat exchanger passes. Each unit of the array may be a separate burner with itsown injector and mixing tube (a configuration generally referred to as a blade burner) orcan be part of a composite burner served by a single injector and mixing tube. Burnerarrays are best suited to high heat input appliances 115 because secondary air can pass between the array elements and thereby mix more thoroughly with the individual outerdiffusion flame mantles.

Fully aerated burners are much less common in current boilers and water heaters. They generally consist of gas and air passing into a burner manifold and then through theflamestrip. Because no secondary air is required, the choice between one or severalburners will depend on the heat distribution and manufacturing costs. To date, surface


Figure 6.5 Typical gas burners for domestic boilers: (top) box burner, (bottom) blade burner.

combustion plaques and punched metal burners have been favoured 115 . Various other burners have been used in boilers and water heaters, but have now fallen into disuse. Thejetted burner (Figure 6.6), consisting of a number of small, individual partially aerated burners (‘jets’) screwed into a suitably shaped manifold, was designed at the time of conversion to natural gas. Its chief advantage was that when an appliance was convertedfrom town gas to natural gas, the non-aerated town gas jets could be directly replaced with a new aerated natural gas jet 124 . It became evident, however, that the aerated jetsare especially prone to linting and that the flame picture is not sufficiently compact foruse with heat exchangers of low thermal capacity. With regard to the latter, the heatdensity can be adjusted by varying the jet spacing. However, if the jets are too closetogether, there is excessive flame interference; if the jets are too far apart, cross-lighting


Figure 6.6 A jetted burner for boilers

The matrix burner (Section 2.3.3) has been assessed with boilers and water heatersprimarily in mind. While the matrix burner offers many advantages in terms of short,compact flames and uniform heat release, the deficiencies described in earlier sectionshave precluded its widespread use. on ignition is unreliable.

6.4.1.3 Burner location

In order to make best use of the heat generated by the burner, the flames and combustiongases need to be in close proximity to the heat exchanger surfaces. Two limiting factorsneed to be considered:

1. If the burner is too close to the heat exchanger, then there is a risk of bothincomplete combustion due to flame quenching at the cooler surface, and general or localoverheating of the heat exchanger leading to material degradation, and, with radiantburners, the possibility of nucleate boiling within the tubes.

2. If the burner is too far from the heat exchanger and the combustion chamber is notwell insulated, then heat losses through the appliance wall can reduce thermal efficiency.

Consequently, the distance between the burner and heat exchanger can be critical.Variation of a few millimetres in either direction can greatly affect applianceperformance. In practice, optimum position is very often determined experimentally for aparticular burner/heat exchanger combination.

As well as combustion and efficiency considerations, there are numerous other factorswhich must be taken into account. For example, for partially aerated burners, the air flow


pattern within the combustion chamber must be such that there are no problems with flame instability or unreliable ignition because of draughts or local vitiation. It may benecessary to fit air guides into the chamber in order to shield the flames and to direct airsupplies to the appropriate parts of the burner.

The problem of linting is an important consideration for aerated burners that use combustion air taken from inside the building; floor-standing boilers with low-level air entry are particularly at risk. Linting is discussed in greater detail in Section 3.5.5, together with a range of designs for lint-resistant burners. In general, the use of room-sealed appliances rather than open-flue appliances will minimize lint accumulation.Where an open-flue appliance is used, air intakes should be at as high a level as possible.

6.4.1.4 Future developments

The majority of currently available boilers and water heaters use partially aerated box andblade burners, whether operating by natural draught or fanned draught, and achievethermal efficiencies of nearly 80%. Recent trends in burner development suggest thatfuture appliances will incorporate fanned draught, fully aerated burner systems 115 , 125 . Fanned draught systems allow lower excess air levels by facilitating better control ofaeration, and permit the use of more restrictive heat exchangers which, with theirimproved heat transfer characteristics, can extract more heat from the combustion gases.Appliances with a thermal efficiency of over 90% using partially aerated burners are nowbecoming available in Great Britain and the rest of Europe. Looking further into thefuture, burner developments featuring pulsed combustion will offer another high-efficiency option if the outstanding problems can be overcome.

6.4.2 Individual space heaters

Individual space heaters, which are designed to supply heat to one room or passage way,are generally divided into two classes, convectors (often referred to in Great Britain aswall heaters) and radiant-convectors (gas fires). Convectors provide warm air by use of aheat exchanger through which combustion gases pass and over which room air is drawneither by natural buoyancy or forced convection (e.g. by using a fan). Radiant-convectors supply a combination of convected heat and direct radiant heat, the latter arising from theheating of a ceramic surface to incandescence. Although design of heatexchanger/transmitter is different for each type of appliance, there are many similaritiesin burner design, such that they may be considered together in this section. Decorativegas-log and gas-coal fires, which are designed for appearance rather than heat output, are not included here.

6.4.2.1 Choice of heat input

As with boilers described above, the rating of the heater and heat exchanger needs to bedefined at the initial design stage. In principle, there is no limit to the heat input of aspace heater. In Great Britain, however, individual space heaters generally have a heat


input of less than 7 kW, which obviates the need for a permanent air vent in a room inwhich an open-flue appliance is installed 126 . Such a vent with its associated draught would be highly undesirable in a living room. Space heaters that are currently availablecover a range of applications. For bedrooms or unoccupied areas such as hallways,convector heaters with a low heat input (up to 3 kW) are usually adequate. In GreatBritain, focal point space heating is generally preferred in living rooms, so radiant-convectors are popular. For an average size living room, a heater with 3 to 4 kW output(about 5 kW input) is acceptable; in smaller, highly insulated modern homes, a heat inputof 2 to 3 kW should be sufficient.

Figures quoted above refer to appliances operating at full rate. However, a range of heat inputs is often desirable. Not only can the heat input be set according to prevailingclimatic conditions, but also full rate can be used to heat a room quickly followed by alower rate to keep the room warm. Consequently, most appliances are equipped tooperate at a number of heat settings. The means by which this can be achieved isdiscussed in the following section.

6.4.2.2 Choice of burner

Guidelines concerning flame stability, port dimensions, and so forth have been discussedin previous chapters. The overall burner dimensions will depend on the shape of the heatexchanger. In order to minimize the extent to which an appliance protrudes into a livingarea, space heaters are often designed to be as slim as possible. This results in heatexchangers which are of approximately rectangular horizontal cross-section with a high aspect ratio. Burners therefore must be compatible with this geometry in order that thecombustion products may pass easily through the slot entrance of the heat exchanger.

Before conversion to natural gas, non-aerated burners were used on almost all spaceheaters. Discussions in Chapter 2 have shown why non-aerated burners have generally proved unsuitable for natural gas despite clear advantages in respect of low noise levelsand freedom from linting. Consequently, partially aerated burners are currently usedalmost exclusively by natural gas space heater manufacturers (Figure 6.7). By far the most frequently used is the box burner; ribbon burners are in use, but are more expensiveand much less common. Both burner types are characterized by large width to depthratios. Fully aerated burners using radiant plaques have not found favour in Great Britainexcept in LPG mobile heaters. Natural gas surface-combustion space heaters are available, however, in many other parts of the world, especially in Japan, where thelowCO and NOx emissions and high efficiencies are particularly advantageous in theirlarge flueless heating market.


Figure 6.7 A typical burner for gas fires and convectors.

Figure 6.8 Simplex (top) and duplex (bottom) gas fire burners, showing respectively one and two gas injectors.


Because of the need for more than one gas rate, means must be provided for thecustomer to vary the heat input to the burner. Most space heaters are provided with anumber of predetermined settings, which are selected manually by the customer. Themanner in which gas rate to the burner is varied has an important influence on burnerdesign. There are two general methods available (Figure 6.8):

1. A simplex burner is one that is fitted with a single injector which supplies the wholeburner. At reduced gas rate, a flow restriction in the gas valve lowers the gas pressure atthe injector and hence gas flow rate and burner port loading. Care must be taken to ensurethat flames remain stable at all gas flow rates.

2. A multiplex burner is divided into sections, each with its own injector. On turndown,the gas valve shuts off the supply to one or more sections, while maintaining the samegas rate to other sections. Consequently, all sections of the burner, when alight, operate atthe same port loading over the complete range of available heat input.

It is possible to combine both burner features to produce a multiplex burner where, for instance, each section may operate at full or half rate. Whether a simplex or multiplexburner is used depends on the application and, to some extent, personal choice of thedesigner. The designer has a number of factors to consider, which are discussed hereseparately for convectors and radiant-convectors.

Convectors rely on good convective heat transfer in the heat exchanger for optimumefficiency. They are also low-cost items, where components need to be as cheap aspossible; the introduction of one expensive component can greatly influence the retailprice. Both of these factors suggest a simplex burner, which is cheaper than a multiplexsystem and maintains an even heat distribution across the heat exchanger on turndown.

With radiant-convectors, consideration must be given to the effect on radiant output and appearance, which have a greater influence on the customer than the convectiveefficiency. If a simplex burner is used, the radiant appearance on turndown deterioratesuniformly across the burner; use of a multiplex burner allows the remaining operativesections to retain a high radiant output and a more pleasing appearance. Consequently,multiplex burners are usually fitted to radiant-convector heaters.

6.4.2.3 Location of burner

For a convector heater, the location of the burner with respect to the heat exchanger issubject to constraints which are very similar to those mentioned in connection withboilers in Section 6.4.1.3, and will not need to be repeated here. For radiant-convectors, attention must be given to the method by which radiant heat is produced.

Radiant heat may be produced either directly by using fully aerated radiant plaques(described in Chapter 4), or indirectly by impingement of a partially aerated flame on arefractory ceramic. Two general designs of the latter are currently in use (Figure 6.9):

1. Box-radiants, comprising a perforated ceramic box, the inner surface of which is heated by the flame. Design of box-radiants is outside the scope of this publication. Ingeneral terms, the cross-section area inside the box must be small enough to allow


Figure 6.9 A radiant convector gas fire fitted with (top) box-radiants and (bottom) radiant plaques.

impingement, but not so small that impingement is so excessive that incompletecombustion may occur. Box-radiants must also be sufficiently tall to permit a degree offlame lengthening (due to gas composition changes, vitiation, overgassing, linting, etc.)without the risk of impingement on the entrance to the heat exchanger.

2. Ceramic coals or logs, comprising a layer of suitably shaped and coloured pieces of


ceramic material, which is heated from below by a horizontally firing or inclined burner.The burner may contain, in addition to a ‘standard’ partially aerated box burner, a numberof low-input flames which are non-aerated or of low primary aeration (up to about 25%). These burn with a yellow flame, and protrude through and above the ceramic, therebyenhancing the live-fuel effect.

Whatever the arrangement for radiant heat output, the heat exchanger entrance ispositioned above the radiating surface, with due regard to any risk of flame impingement.In order to avoid an excessive intake of secondary air (and hence in order to maximizethermal efficiency), a heat-resistant glass panel may be incorporated into the front casing of the appliance.

Linting is of particular concern with space heaters, because many designs (especially radiant-convectors) draw combustion air from within the room. Appliances are usually mounted on or very near the floor and air intakes are generally at a low level. Thus spaceheaters are particularly susceptible to accumulation of lint within the burner assembly.The reader is referred to Section 3.5.5 regarding guidelines for lint-resistant burners.

6.4.2.4 Future developments

Current British burner design for space heaters is firmly rooted in partially aerated boxburners operating by natural draught. Recent boiler developments using fanned draughthave not been taken up by manufacturers of space heaters. This is perhaps becauseradiant-convectors, which form the majority of the space heaters in use, are traditionally fixed to an existing chimney. Future developments may include an assessment of thebenefits of fanned draught and the potential increases in thermal efficiency 79 . In the longer term, fully aerated radiant burners are likely to be a design feature of radiant-convectors. Pulsed combustion is also being considered for individual space heaterapplications 127 .

6.4.3 Cookers

Domestic cooking appliances in general use in Great Britain incorporate three differenttypes of burner (hotplate, oven and grill), but all have the same basic objective: to ensurean efficient, evenly distributed and controllable transfer of heat from the flame to thefood. Each of the three burner types is considered separately here.

6.4.3.1 Hotplate burners

The main requirement for hotplate burners is for an efficient transfer and evendistribution of heat between the flame and the cooking vessel. A number of factors mustbe considered by the hotplate designer:

1. The burner geometry must produce a compact flame picture that gives a uniformdistribution of heat to the base of the pan, without local overheating.

2. Because heat is lost by the flame with increasing distance from the burner, greatestefficiency is obtained by the smallest possible separation between the top of the burner


and the base of the pan. 3. Pan-to-burner separation must not be so small that incomplete combustion is caused

by excessive flame impingement, secondary air starvation or lack of space for the egressof combustion products.

4. Arrangements must be made to prevent excessive flame impingement on the pansupports causing overheating and subsequent material degradation.

5. Hotplates must provide good pan stability, ease of cleaning, and be of attractiveappearance.

Figure 6.10 Typical hotplate burners for a domestic cooker.

Most current British domestic hotplates utilize four circular, partially aerated burnerswhich have an array of vertical flame ports spaced around the periphery (Figure 6.10). Because different cooking loads have different heat demands, hotplate burners must becapable of satisfactory operation with a very large turndown ratio. For example, a burnerof between 3 kW and 4 kW is needed for fast boiling of a kettle, but the same burner willneed to deliver less than 0.3 kW for simmering or for small pans. Cookers maybe fittedeither with pairs of differently rated burners (e.g. two at 3.5 kW and two at 2 kW) or withfour equally rated burners of intermediate size. Most British burners operate at a primaryaeration of about 45% to 50%, with a pan-to-burner separation of 7 mm to 15 mm toensure good combustion 128 . A turndown ratio of up to 10:1 is normal with Britishhotplate burners.

In other countries, particularly in continental Europe, a wider range of hotplate burner ratings is normally provided (e.g. 3 kW, 2 kW and 1 kW for what are known as rapid,


semi-rapid and auxiliary burners). This obviates the need for large turndown ratios, which can be difficult to achieve if appliances are for use in several countries over a range ofdifferent gas compositions. However, such an arrangement provides less flexibility forthe user and can lead to local overheating of large pans when using the smaller burners.

Transfer of heat from a burner flame to the contents of the pan occurs by a combination of conduction, convection and radiation. The latter component is usuallyvery small (perhaps only 2%) for a partially aerated burner, so heat transfer is mainly byconvection at the vessel boundary layer and by conduction through the base of the vessel.The greatest heat transfer is obtained with vertical or inward-firing flames and high port loadings (gas velocities), both of which increase turbulence and hence convective heattransfer in the boundary layer. However, because vertical and inward-firing burners may restrict access for secondary air, outward-firing burners are usually fitted. Completeness of combustion is thereby ensured, but with a slight loss in efficiency.

Thermal efficiency of hotplate burners is difficult to quantify, because the amount of heat transferred is dependent on pan size, shape and material and the way in which the

Figure 6.11 The effect of gas rate on the thermal efficiency of a typical hotplate burner.


Figure 6.12 A high aeration cooker hotplate burner.

customer uses the burner. BSI/European testing procedures 117 define a standard vessel and contents, which can then be used to compare the heating performance of differenthotplate burners. Typical partially aerated burners give thermal efficiencies in the range52% to 60% based on the latest European test method 117 , 128 . The effect of usage pattern on efficiency is shown schematically in Figure 6.11. At low gas rates, convective heat transfer at the base of the vessel is low and radiative losses from the vessel areproportionately high. At very high gas rates, flames may extend beyond the edge of thevessel, and heat is lost directly to atmosphere. Between these two extremes is the regionin which the customer should operate the burner.

Future developments in hotplate design are likely to include more efficient burners. One such option 128 involves an increase in primary aeration to 75% to 85%, which inturn gives hotter, more compact flames (Figure 6.12). This allows much smaller pan-to-burner separation (1 mm to 2 mm) and increased efficiency (up to about 65%). NOxemissions should also be reduced.

Current development work is also aimed at developing hotplates which utilizeconventional partially aerated burners beneath glass-ceramic discs in place of pan supports (Figure 6.13). As well as providing an attractive aesthetic appearance, the discprotects the burner from draughts and spillage of food. The disc also ensures that theflame position remains unaltered when pans are removed or replaced, so that combustionperformance is largely unaffected by the presence of a pan, and flame detection becomesmore reliable than with conventional systems. This contrasts with traditional open


Figure 6.13 A ceramic hob hotplate (partially aerated).

Figure 6.14 A ceramic hob hotplate (fully aerated).


hotplate burners where the flame impinges on the pan and the flame shape is influencedby the pan geometry.

Another novel development uses glass-ceramic hobs that are fired by radiant plaqueburners (Figure 6.14). Such burners are fully aerated, emit very low NOx levels, and air can be supplied by either atmospheric injection 128 or use of a fan 129 . However, there is a loss in both controllability and thermal efficiency compared with conventional openhotplates 128 . In addition, radiant plaque burners offer very limited turndown (seeChapter 4). This can be overcome by the use of sequential burners, which can cycle off and on (at full rate), thereby giving a low rate for simmering. Such a burner is moreefficient than a conventional burner at part rate (cf. Figure 6.11), but attention needs to be paid to component reliability if cycling becomes excessive.

6.4.3.2 Oven burners

Ovens can be divided broadly into three types, depending on how the cooking space isheated:

1. In an internally heated oven, the burner is located (usually in the base at the rear) inthe cooking chamber, such that the combustion products heat the food directly. Thisarrangement has traditionally been used in Great Britain. Heat is transferred by naturalconvection from the hot gases and by radiation from the walls and fittings. Because of thereliance on natural convection, a vertical temperature gradient is a feature of internallyheated ovens, with the hottest part at the top of the chamber. This allows differentcooking conditions depending where the food is placed.

2. Most European cookers use external heating, where the burner is sited below thebase, and combustion products enter the chamber through slots or louvres in the base orside walls. Convected heat is thereby more evenly distributed throughout the chamber.Compared with internally heated ovens, the rate of heat-up from cold is much slower owing to the thermal mass of the chamber walls. Additionally, the base can become veryhot, which increases radiant heat output from the base, resulting in an inversion of thetemperature gradient. Externally heated ovens are generally of lower efficiency thaninternally heated ovens 128 .

3. Some domestic ovens now operate using forced convection, whereby a hot airstream is directed over the food by use of a small fan. The improved heat transfer anduniform temperature distribution allows larger quantities of food to be cooked morequickly and at a lower temperature than in conventional ovens. Therefore less fuel is usedduring the cooking process. Heat can be introduced into the chamber by a number ofmethods. The simplest and cheapest is to use an internally heated oven with a fan fitted inthe rear panel or roof of the cooking chamber. Externally heated ovens may be adaptedsimilarly. In each case, the main problem to be overcome by the designer is to preventhigh velocity air streams interfering with flame stability. An alternative scheme uses aheat exchanger, the outside of which is heated by the combustion gases, and the inside ofwhich forms part of the cooking chamber; this system is inherently less efficient and canbe bulky, although it is widely used in commercial catering appliances.

All modern domestic ovens in Great Britain use partially aerated burners, either of the


box type with punched ports, or bar burners made from cast iron with drilled ports. European cookers that are fitted with externally heated ovens use a wider variety ofburner configurations. Whatever design is employed, a simplex burner with one injectoris commonplace. Burner ratings are usually related to the cooking chamber volume.Typical values vary from about 50 kW m−3 for internally heated ovens to 120 kW m−3 for forced convection ovens. Primary aeration is usually about 50%. In order to ensuresatisfactory combustion, secondary air inlets need to be sized and positioned correctly,but must not be so large that recirculation of combustion products into the injector systemcould occur. It should be remembered that, unlike many other appliances, oven burnersare frequently required to operate in a vitiated atmosphere. They must also be stable todraughts, such as are created when the oven door is opened or closed. With internallyheated ovens, there is usually little problem with combustion because the open flame isunlikely to impinge on an adjacent surface. Flame lengthening can be a problem inexternally heated ovens, where the burner is located underneath the baseplate.

Future developments for ovens are most likely to improve design and efficiency of the whole system rather than just the burner. Present studies are concentrating onimprovements in insulation and heat transfer together with the development of a ‘dual-purpose’ oven which will offer the customer a choice between natural and forced convection cooking.

6.4.3.3 Grill burners

The grill is on average the least used burner in the domestic cooker, using about 14% ofthe total fuel consumption 128 . Such figures can disguise the extremes of use: somecustomers never use a grill, while others may use a grill more frequently than the oven.Thus, although most development work has concentrated on hotplate and oven design,the grill should be regarded as an equally important component in the system.

Grilling requires a high intensity radiant heat for a relatively short cooking period.Desirable features include rapid heat-up, some degree of turndown and an evendistribution of heat at all gas rates. British grills are traditionally separate from the oven,usually mounted at high level and rated at about 4.0 to 5.5 kW. European grills areusually mounted in the roof of the oven and are much lower rated (2.0 to 2.5 kW).

Traditional grill burners are comprised of a cast iron or pressed steel simplex bar burner (Figure 6.15). The flames, which are partially aerated, impinge on a metal gauze or refractory surface, which is heated to incandescence. Non-uniformity of heat distribution is the major drawback with this design of grill burner. Because the heatcontent of the flame decreases along the flame axis, the gauze/refractory surface must besuitably inclined or curved in order to improve heat transfer in the tail of the flame.Additionally, owing to heat losses from each end of the burner, equally spaced ports givea greater heat flux at the centre of the burner. This may be compensated for by increasedport spacing at the centre of the burner, but the spacing should not be so large that cross-lighting on ignition is prevented. A second drawback with the bar burner is that, becauseit is of the simplex type, the grilling of small quantities of food results in


Figure 6.15 A grill bar burner.

Figure 6.16 A grill ceramic burner.


wasted fuel at the burner extremities and a lower overall efficiency. Recent developments 128 are designed to overcome both disadvantages of the bar

burner. Surface combustion radiant plaques (Figure 6.16) offer direct radiant heat uniformly over a wide area. Duplex operation also can be provided, such that each plaqueis either fully on or off. Although such a combustion system with associated controls ismore expensive than a simplex bar burner, there are overriding advantages in terms ofconvenience of use and fuel savings which make such a system very attractive.

Future work on grill burners is likely to concentrate on the development of surface combustion plaques. To a large extent, such work will complement similar developmentswith hotplate burners, which were discussed above.

6.4.4 Other appliances

Sections 6.4.1 to 6.4.3 have concentrated on boilers, space heaters, water heaters and cookers, which together constitute 97% of British gas appliances 130 . It is not proposed to consider other appliances (e.g. refrigerators, decorative fires). However, the same basicguidelines apply, and from the discussions so far it should be possible to design a burnerwhich gives satisfactory performance over a wide range of operating regimes.


CHAPTER 7 Burner controls

7.1 General

The information gathered together in previous chapters provides guidelines for theconstruction of a burner that is correctly sized (physically and thermally) for a particularapplication and gives a flame that is stable with respect to lift and lightback. Animportant aspect not yet covered is that of burner controls, which may be fitted to anappliance in order to optimize efficiency, safety and ease of operation. A wide variety ofsuch controls is available both for industrial and for domestic equipment. For example,devices are available that can regulate pressures and flow rates of gas and air and ensureeffective and safe ignition of the gas stream. Controls that maintain combustion qualityover a range of operating conditions are used in industrial applications but have yet to beproven in domestic appliances.

This chapter is intended as an introduction to a subject, the theory and application of which could in themselves occupy several volumes. As well as currently used controls

7.1 General 7.2 Gas controls 7.2.1 Gas valves 7.2.2 Governors 7.2.3 Flame failure devices 7.2.4 Thermostats and timers 7.2.5 Ignition 7.2.6 Full sequence control 7.3 Air controls 7.3.1 A.C. fans 7.3.2 D.C. fans 7.4 Air/gas ratio controls 7.4.1 Zero governors 7.4.2 Flue gas sensors

(e.g. governors, thermostats), possible future developments are discussed, such as theprobable widespread introduction of microelectronic control. Discussion here is confinedto an appreciation of the needs for various controls and a description of a number ofoptions that are or soon will be available. Consequently, no attempt is made to presentcomplete design details of each device. A number of general texts are available to thereader. Jasper 123 and Miles and Pinkess 131 offer detailed, though now in parts somewhat dated, reviews of domestic control systems, while Pritchard et al. 49 and BGSFM 132

discuss industrial controls, which can, to a certain extent, be scaled down for domesticapplications.

Many controls are required to be fitted to appliances for reasons of safety (e.g. flamefailure devices). All such controls must comply with the relevant performance and safetyStandards for the appliance 116 . 117 and, where appropriate, for each component. No discussion of detailed performance requirements is presented here; it is left to the readerto consult the relevant documents.

7.2 Gas controls

Control of the flow of gas to the burner is an essential feature of all gas appliances. Thecheapest and simplest appliance must be fitted with at least a manual tap or valve whichenables the user to turn the gas on and off. More complex controls include gas governors(which regulate pressure or, occasionally, volume flow), methods of igniting the gas, andflame failure devices (which shut down the appliance if the gas supply fails). The moreexpensive appliances (such as central heating boilers) may be fitted with full sequenceautomatic ignition, which controls gas input, ignites the gas and incorporates a flamefailure device. The development of microelectronic control units is currently receivingmuch attention worldwide. Miles and Pinkess 131 give a thorough and detailed review ofmechanical and electromechanical controls; Finch et al. 133 discuss possible future applications of microelectronics in domestic control systems.

7.2.1 Gas valves

Gas control valves can be divided broadly into two types: firstly, devices that requiremanual operation by the customer (as fitted to most gas fires and cooker hotplates);secondly, devices that incorporate a sensor which detects a demand for heat and sets thegas rate accordingly (e.g. boilers, water heaters). The degree of automation provided by amanufacturer depends largely on the requirements of the user and (for cheaper appliancetypes) the acceptability of increased appliance cost. If controls are too complex, theappliance may be prohibitively expensive. Therefore, of the systems described below,space heaters are almost exclusively fitted with manually operated valves; modulatinggas valves tend to be used only in cooker ovens and in the more expensive domesticboilers.


7.2.1.1 Manual gas valves

Manually operated gas valves are generally confined to two appliance types. Individualspace heaters are usually designed as low-budget items where construction costs are kept to a minimum. The use of an intricate control system can add significantly to the price ofthe appliance, thereby reducing its competitive position in the market-place. With cooker hotplate and grill burners, users generally require (or prefer) complete manual control ofgas rate in order to cook using a range of temperature regimes. Although adding to thecost, suitable customer controls are considered a necessary part of a cooking appliance.

Manual gas valves can be divided loosely into two types, those which offer acontinuously variable gas rate up to the burner rating, and those which supply gas at anumber of fixed gas rates. Continuously variable taps are generally fitted to all cooker burners and simplex gas fire burners. Such taps provide a variable restriction in the gasline which lowers the gas pressure at the injector and consequently reduces heat input(equation 3.22). Gas fire multiplex burners are usually fitted with a control tap that isprovided with one inlet and several outlets. Each outlet feeds a proportion of the totalflow to a section of the burner (Figure 7.1), such that the user can select the number of sections operating at any given time. For example, a three-section burner may be fitted with a control tap which provides three heat inputs, corresponding to one, two and threeburner sections operating at full rate.

Figure 7.1 A gas fire control tap for a multiplex burner.

Burner controls 153

Figure 7.2 Four-valve push-button control for a gas fire. (From Miles and Pinkess 131 .)

Push-button valves, using the same principle of flow splitting, have been used to a limitedextent with space heaters, but have not generally found favour. The four-button valve illustrated in Figure 7.2 contains four outlet pipes, each of which supplies gas to one ofthe four burner sections. By depressing the correct button, the user can select operation ofone, two, three and four sections.

7.2.1.2 Automatic gas valves

Automatic control of gas rate to an appliance is intended to simulate the manual processof adjusting heat input to the burner. The simplest automatic control is the ‘two-position’ switch, which is usually designed to turn the burner on and off with no intermediatesettings. Such a control reacts to a mechanical or electrical signal from a sensor, e.g.thermostat, pressure switch or timer. Examples of appliances that use this system areboilers and instantaneous water heaters. Boilers use a thermostat and a timer, whichtrigger the gas valve when a demand for heat is detected; instantaneous water heaters usea water pressure sensor, which detects flow of water through the appliance and opens the


gas valve accordingly. A variant of the on/off device is a high/low valve, which may be fitted in systems that

have to operate at two different gas rates. Examples of this mode of operation arethermostatic control of gas fires and refrigerators, which work at a reduced heat settingwhen demand for heat is satisfied. Initial activation of the appliance is manual; thehigh/low sensor obviates the need for a full reignition sequence whenever there is arenewed heat demand.

Figure 7.3 Schematic response of a modulating gas valve.

The alternative to the two-position switch with predetermined gas rates is a device which gradually reduces the gas flow rate as the required temperature is approached, i.e.the gas rate is modulated. This allows the gas rate to vary according to the demand forheat. Typical valve performance is shown in Figure 7.3. The upper temperature limit corresponds with the desired temperature. Below the lower limit, the valve is fully open(i.e. full gas rate) while above the upper limit, the valve is fully closed or at a low by-pass rate (i.e. no heat required). Between these two limits (within what is termed the

Burner controls 155

proportional band), the valve opens partially to a predetermined position for a giventemperature. The temperature/gas rate relationship need not be at all linear within theproportional band. Valve response is determined by the valve design and can be tailoredto the design and requirements of the appliance. An appliance with low thermal mass(such as an oven) reacts quickly to changes in heat input. Consequently, a wideproportional band is normally required in order that sudden changes in gas rate (whichwould produce sudden changes in heat output) can be avoided. At the other extreme, ahigh thermal mass appliance (such as a boiler with a cast iron heat exchanger) should beequipped with a valve with a narrow proportional band. In this case, a slight drop intemperature will fully open the valve, thus avoiding the otherwise long delays in heatingup.

7.2.2 Governors

Since conversion from town gas to natural gas, domestic gas meters have normally beenfitted with a pressure regulator (or governor) just upstream from the meter inlet. Thefunction of a governor is to reduce a higher (and possibly variable) gas pressure to alower constant outlet pressure for a range of volume flow rates. In Great Britain, thenominal outlet pressure is 20 mbar. Consequently, all natural gas appliances distributed inGreat Britain are designed to accept G20 gas pressure of 20 mbar at the appliance inlet.

The full available supply pressure need not be used in the appliance. With partially aerated cooker and water heater burners, the full 20 mbar pressure is generally used inorder to maximize the energy available for entraining primary air. Unfortunately, highinjector pressures can produce noisy burners. Noise may be tolerated on cookers andwater heaters, but can cause a nuisance with space heaters fitted in living rooms. Thus anappliance governor may be provided, or else sufficient pressure losses built into thesystem (e.g. a restrictive gas valve). Boilers, like water heaters, benefit from a highinjector pressure. However, if a governor is fitted (and the heat exchanger is designedaccordingly), the installer can vary the gas rate in order to match boiler output with thedesign heat loss of a dwelling. Boilers provided with such adjustment are known asrange-rated appliances.

Governor design is beyond the scope of this chapter; it is, however, worth whiledescribing the principles of operation. Full details of the numerous design features aregiven by Pritchard et al. 49 and by Miles and Pinkess 131 . The special applications of zero pressure governors are discussed in Section 7.4.1. Figure 7.4 shows schematically the operation of a constant pressure governor. It contains a flexible diaphragm, which isloaded with a weight (as shown) or by a spring, and it is linked to a valve which moveswith the diaphragm to restrict the flow. The pressure above the diaphragm is maintainedat atmospheric by a small breather hole. The diaphragm itself is kept in position by theupward force due to the outlet pressure (p o) being balanced by the downward force dueto the weight. If the inlet pressure (p i) increases, p o also increases initially. This exerts a greater upward force on the diaphragm, which itself moves upwards taking the valvelinkage with it. Flow through the valve is progressively restricted and the pressure dropacross the valve increases until p o is restored and the forces balance. The converse is true


if p i decreases, i.e. the diaphragm moves downwards until the outlet pressure is restored.

Figure 7.4 A simple constant pressure governor showing (left) the schematic layout and (right) the pressure and force distribution.

Most modern governors are fitted with an adjustable spring (essentially a variableweight). This permits the governor outlet pressure to be set by the installer, as would berequired for the installation of range-rated boilers.

7.2.3 Flame failure devices

A flame failure device (FFD), also known as a flame protection or flame supervisiondevice (FSD), is fitted to many gas appliances in order to prevent unburnt gas issuingfrom the burner. Such a circumstance can occur if, for any reason, the ignition sequencefails to ignite the gas or the flames are extinguished by draughts or vitiation duringoperation of the appliance. Any consequent build-up of gas is potentially explosive and could put the user at serious risk of injury. An FFD must be able to maintain the gassupply if a flame is present, and must shut down the gas supply promptly if the flame isextinguished. Three general methods of flame detection are discussed here:

1. Sensing heat generated by the flame. 2. Using electrical conduction or rectification by the flame. 3. Sensing infra-red or ultra-violet radiative emissions.

The first is the most common, mainly because it is cheap and requires no externalelectrical supply. More recently, flame rectification has found increasing favour,especially with more expensive appliances that are provided with an electric mainssupply. The third method is presented here for completeness; it is currently used only inindustrial applications.

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7.2.3.1 Heat sensing FFDs

1. Bimetallic strips. Figure 7.5 shows the operation of a bimetallic strip FFD. A curvedstrip is fixed to the burner at one end, while the other is connected by a fixed mechanicallink to the main gas valve. The FFD is heated either by an independently fed pilot (oftendesigned to be more susceptible to lift than the main burner) or by a low ‘by-pass’ gas rate to the whole burner. In either case, differential expansion causes the main valve toopen and the full gas rate to be delivered to the burner. If the gas supply fails, the stripcools, thereby closing the main gas valve. The device is simple and relativelyinexpensive, and has been used on a number of cheaper domestic appliances, such aswater heaters and cooker ovens. Disadvantages are that it does not shut off the pilot/by-pass supply, and it cools very slowly owing to its large thermal mass; both effects cancause a build-up of gas in the combustion chamber. Long-term reliability is also questionable; metal fatigue or corrosion can distort the strip such that the valve becomes ineffective. Because of these problems, the popularity of bimetallic strips has decreasedin recent years.

Figure 7.5 Schematic operation of a bimetallic strip flame failure device.

2. Mercury vapour devices. Figure 7.6 shows schematically a typical vapour pressure detection system. A sensing phial is filled with liquid mercury and is connected via acapillary tube to a diaphragm, which in turn is linked mechanically to the gas valve. If thephial is heated by a flame from a pilot or from the burner by-pass, the mercury vaporizes


and expands. The resulting pressure increase extends the diaphragm and the levermechanism opens the valve. If the gas supply fails, the mercury condenses, and the springforces the valve to close. Like the bimetallic strip, the device is simple and inexpensive,and is currently in use in a number of space heaters, water heaters and cooker ovens.Although there remains the risk of a build-up of gas from not shutting down the pilot/by-pass supply, the device has two major advantages compared with bimetallic strips: firstly,response time is much faster owing to the lower thermal mass; secondly, in the event ofphial damage (e.g. breakage) insufficient pressure is available to open the gas valve(i.e.the device is fail-safe).

Figure 7.6 Schematic operation of a vapour pressure flame failure device.

3. Thermoelectric FFDs. Thermoelectric devices work according to the well knownprinciple that an electromotive force (e.m.f.) is produced when a junction of twodissimilar metals is heated. Thermocouples operating on this principle can be used formeasuring temperature and form the basis of a frequently used method of flameprotection. By allowing the main burner or pilot flame to impinge on a thermocouple, theoutput voltage (usually 20 to 30 mV) can be used to energize an electromagnet and holdopen a springloaded gas valve. If the flame is extinguished, the thermocouple cools andno longer generates an e.m.f., such that the valve is released and the gas supply shuts off.Figure 7.7 shows the operating cycle of a typical device. In the off position, no gas can pass (Figure 7.7b). The ignition cycle must be initiated by manually pushing the valve against the magnet. This admits gas to the pilot/by-pass (Figure 7.7c). The gas is ignited and the resulting flame heats the thermocouple, which energizes the electromagnet. Thepush-button must be held in until sufficient e.m.f. is generated (normally about 30 seconds). Release of the button leaves the valve open and also allows full gas rate to passto the main burner (Figure 7.7a). Thermoelectric devices are now in common use onboilers, water heaters and space heaters. Their main drawback is that they require manualoperation for every ignition sequence unless a permanent pilot is fitted. Although manual

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ignition is acceptable for individual space heaters, it is unsuitable for, say, oven burnersfitted with automatic timers. Response time can be lengthy (up to 30 seconds) if the flameextinguishes; however, thermoelectric FFDs do shut down both the main burner and thepilot/by-pass.

Figure 7.7 Valve mechanism of a thermoelectric flame failure device: (a) operating position, (b) off position, (c) start condition. (From Jasper 123 .)

7.2.3.2 Conduction/ rectification FFDs

The chemical reactions occurring within a flame generate small concentrations (about 1ppm) of electrically charged particles (ions) and free electrons. Consequently, a flamewill conduct an electric current if a d.c. potential is applied between a pair of electrodes.It is common practice 49 , 131 for the burner to form an earthed electrode while a thin rodof heat-resistant metal protruding into the flame acts as a live electrode (Figure 7.8). The


detected electrical signal can be amplified and made to operate an appropriate gas valve.If the flame extinguishes, the current falls to zero and the valve closes. The chiefdisadvantage of flame conduction is that the device fails to warn of danger in the event ofan electrical short circuit.

Flame rectification devices do not suffer from this disadvantage. The technique relies on the ability of a flame to rectify an applied a.c. voltage, owing to the very much lowermobility of positive ions compared with the lighter flame electrons. An a.c. voltage isapplied between two electrodes of different area. (An area ratio of at least 4:1 is usual 49

with the burner head at earth forming the larger electrode.) The unequal transfer ofcharge carriers results in a small rectified current flowing between the two electrodes.The signal can be amplified and can be made to operate the gas valve. Becauserectification only occurs if a flame is present, the drawbacks of flame conduction do notapply.

Figure 7.8 A general layout for flame conduction or rectification flame failure devices.

The great advantage of conduction/rectification FFDs is that the response to flame failure is almost instantaneous. They respond to the slightest hint of flame lift and are notsubject to the operational time delays of heat-sensing FFDs. One drawback is that anexternal electrical supply is required. This does not pose a problem with most modernhigh-budget ovens and boilers. However, many low-budget space heaters and water heaters operate without the need for electrical power; the economic viability of providingan electrical supply solely for an FFD is questionable.

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7.2.3.3 Ultra-violet/ infra-red FFDs

Flames radiate over a wide range of wavelengths, including the infra-red (IR), visible and ultra-violet (UV) parts of the electromagnetic spectrum. With suitable UV and IR sensors(e.g. a photoelectric tube for UV, zinc or cadmium sulphide coated cell for IR), thepresence of a flame may be detected and the generated signal may be used to operate anelectrically operated valve. These systems are in use in industrial gas-fired plant 49 , 131 ,132 ; cost and physical size currently preclude their use in domestic appliances.

7.2.4 Thermostats and timers

Thermostats and timers are both designed to give users a degree of flexibility andautomation in the operation of their appliances. A thermostat maintains a steady presettemperature for the environment in which it is positioned (e.g. room air, oven air) byindirectly controlling the flow of gas through the gas valve. Both methods of automaticgas valve operation described in Section 7.2.1.2 can be used in conjunction with athermostat, i.e. thermostats can control gas rate either by a two-position switch (on/off or high/low) or by a modulating valve. Refrigerators, space heaters and boilers tend to usetwo-position switches, while ovens generally use modulating valves.

Methods of operation of thermostats are very similar to heat-sensing FFDs (Section 7.2.3.1), except that they are required to function in a much lower temperature range. Bimetallic devices have been used extensively in ovens, but have become less popular inrecent years. Vapour pressure and liquid expansion thermostats, which work in a similarmanner to the mercury vapour FFD, are frequently used. Adjustment of the setting knobaffects the distance that the valve must travel and therefore the temperature/gas ratecharacteristic of the device (Figure 7.9). Fuller construction details are given by Jasper123 and Miles and Pinkess 131 .

The provision of a timer enables the user to arrange for automatic ignition and shutdown of a gas appliance without the need for the user to be present. Timers are nowconsidered an integral part of any central heating system, and are commonplace oncookers. Early timer devices were mechanical, and therefore did not require an electricitysupply. With the increasing acceptability of electrical control of appliance components,electric timers are now used extensively on cookers and boilers.

Future developments in thermostats and timers appear likely to include the incorporation of microelectronic control 133 . Digital display timers are already available as part of integrated packages for cookers and boilers. Further advances inmicroelectronics may see the introduction of electronic temperature control, for instanceusing thermistors in place of existing electromechanical devices.


Figure 7.9 Schematic operation of a liquid expansion or vapour pressure thermostat

7.2.5 Ignition

Some form of ignition system is generally fitted to every gas appliance. As outlined inChapter 1, an ignition system must supply at least the minimum ignition energy to the gas/air mixture. In addition, ignition must be non-explosive, safe and reliable. Ignition bythe user employing, for instance, a match or taper is no longer regarded as acceptable.Three general ignition methods can be identified and are discussed in more detail below:

1. Permanent pilots. 2. Hot filaments (glow-coils). 3. Spark ignition.

All three methods have been fashionable during the last few decades. Spark ignition iscurrently most popular and can be incorporated with comparative ease into an integratedcontrol system. Further details of ignition systems for domestic applications are given byJasper 123 and Miles and Pinkess 131 together with a number of research papers 134 − 138 .

7.2.5.1 Permanent pilots

Permanent pilots, as their name suggests, are small burners (gas rate up to 250 W) whichare continuously alight and located in such a way that gas to the main burner can beignited when the main gas valve is opened (manually or automatically). In most

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applications, the pilot is placed immediately adjacent to the main burner. An exception is the cooker hotplate, where a pilot is positioned in the centre of the hob, with flashtubesextending outwards to each hotplate burner 123 , 131 , 136 . Care must be taken by the burner designer to ensure that the pilot/burner arrangement gives non-explosive and reliable ignition.

Permanent pilots offer advantages of cheapness and simplicity, and do not require another energy source. There are, however, many disadvantages. Pilot jets are very smalland are therefore particularly susceptible to draughts, vitiation and blockage by lint.Careful design may minimize, but cannot eliminate these problems. Additionally, whilean appliance is unused, a permanent pilot is wasting fuel; hotplate permanent pilots, forexample, consume almost half the total gas used on a hotplate 128 .

The use of a permanent pilot is currently confined mainly to water heaters and boilers. In recognition of the need for fuel economy, micropilots (gas rate about 60 W) havenowbeen developed specifically for these appliances. Where remote operation of a gasvalve is necessary (as with water heaters and boilers), a permanent pilot remains the bestignition source unless electronic control can be incorporated.

7.2.5.2 Filament ignition

Filament igniters and glow-coils consist of an electrically heated filament made from asmall coil of wire (usually made from a precious metal) and positioned adjacent to a pilotburner. Power is supplied to the filament by battery or mains electricity. Operation isgenerally manual: depression of a control knob simultaneously completes the electricalcircuit, energizes the filament and opens the gas valve (Figure 7.10). The pilot gas supply is ignited, which in turn ignites the main burner. Unlike the permanent pilot, this systemrequires the ignition pilot to be alight only during the ignition sequence. Although there isno waste of fuel while an appliance is out of operation, a filament igniter normallyrequires manual start-up by the customer.

Hot filament ignition was popular on British town gas appliances. Natural gasespossess much higher minimum ignition energies than town gases 135 , as a result of which filaments need to pass much higher currents in order to ignite a hydrocarbon fuel.Ignition of a given gas/air mixture also becomes less reliable, and tolerances inpositioning the igniter with respect to the burner and gas/air mixture become far tighter134 . Such problems with igniting the gas/air mixture have become the major disadvantage of a hot filament ignition system for natural gas appliances. Other minorproblems include the need for an electrical supply (batteries need regular replacement, amains supply is inconvenient for many space heaters), and longer term reliability (e.g. aswith a light-bulb, filaments are fragile and prone to breakage, burning out). For thiscombination of reasons, and owing to developments in spark ignition, filament ignition isnow little used in British domestic gas appliances.


Figure 7.10 Push-button switch for a hot filament ignition system

7.2.5.3 Spark ignition

Ignition of a gas stream by electric spark is currently the most common method employedin British domestic gas appliances. A high voltage (up to 30 kV) is produced on anelectrode positioned in the unburnt gas/air stream in close proximity to an earthed surface(either a second electrode or the burner itself). Three methods of producing a suitablevoltage are described here. Fuller discussions on the development of spark ignition aregiven in references 135, 137 and 138. 1. Piezo-electric igniters. Piezo-electric crystals possess the property that whenmechanically stressed or deformed, a voltage is produced between the two ends of thecrystal. By using a suitable connection to an electrode mounted on the burner, the crystalmay be discharged as a spark in the unburnt gas stream, thereby igniting the gas. Figure 7.11 shows a typical system: rotation of a cam (often part of the gas control tap) movesthe lever so that the crystal is pressurized. A spark is produced when the potentialdifference between the electrodes is sufficient to overcome the electrical resistance of thegas/air mixture. The advantage of piezo-electric igniters is that no external electrical supply is required; consequently they are popular with such appliances as space heatersand water heaters, both for igniting the main burner or a permanent pilot.

2. Mains-operated transformer igniter. If an electrical supply is already connected to an appliance, a simple step-up transformer may be used to generate the required voltage (Figure 7.12). Mains-operated spark generators are not common in domestic gas appliances, primarily because the component cost is high and the spark energy is far inexcess of what is needed.

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Figure 7.11 General layout of a cam-operated piezo-electric ignition system

Figure 7.12 Typical electrical for a mains step-up transformar ignition system: (A) one electrode earthed, (B) electrodes attached to each end of the secondary wiring.

3. Electronic pulse igniter. Developments in solid state electronics have now madepossible production of a series of sparks in rapid succession (up to eight per second) froma single generator (Figure 7.13). When the ignition sequence commences (either by


manual operation as in cookers, or automatically as in boilers), the input current (mainsor battery) charges up a capacitor through a resistor and diode rectifier. A solid stateoscillator switch then allows discharge at the electrode via a step-up transformer. Modern electronics permits rapid cycling of the charge/discharge sequence until ignition hasoccurred.

Figure 7.13 Electrical circuit for a pulsed-spark ignition system.

With all spark igniters, the position of the high voltage electrode with respect to earthis critical 135 , 137 . The spark gap must be precisely located in order that the spark is of the correct energy and occurs where the gas/air mixture is within its flammability limits.With the high voltages produced, insulation resistance must be high enough to preventleakage. Despite these tight tolerances, spark ignition is currently a relatively cheap,reliable system and is therefore extremely popular with appliance designers.

7.2.6 Full sequence control

Depending on the application, one or more of the devices described above will bearranged as part of the complete system. On some appliances (e.g. cookers, spaceheaters), some degree of manual control is acceptable or desirable; on central heatingboilers, a high degree of automation is necessary. In connection with the latter inparticular, control packages (known as full sequence controllers) are now available towhich may be connected the igniter, FFD, and electrical outputs from timers andthermostats. On the basis of information received, the appropriate signal can be sent to anelectrically operated gas valve. If either the timer or thermostat senses a demand for heat,an electrical signal is sent to the controller and the ignition cycle is initiated. When heatdemand is satisfied or if the flame is extinguished, another electrical signal closes the gasvalve.

Until the last five to ten years, the relatively high costs of sophisticated automatic

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controls have limited their use to the industrial market 139 , 140 . Developments in solid state circuitry, particularly the microprocessor, have made possible the introduction ofelectronic control into domestic appliances (e.g. see references 133, 141–143).

7.3 Air controls

Section 7.2 has described a number of devices, all of which can affect the gas flow rate.These included on-off devices (FFDs, timers) and a number of components (e.g.modulating gas valves) which permit more than one gas flow rate. With modern heat exchangers, in order to maintain high efficiency at low gas rates, means must be providedfor reducing the air flow rate as well as the gas in those systems where some form ofturndown is required. With natural draught burners that use atmospheric injection forprimary air, the physical process of air entrainment ensures that the primary aeration isvirtually unchanged as the gas rate varies, i.e. if the gas rate decreases, the entrained airrate automatically decreases (see Figure 3.4). Secondary air flow rate will also decrease,though not to the same extent, such that the excess air level will rise as gas rate drops.

Many modern appliances are equipped with fans, which operate at a fixed rate and donot respond automatically to changes in gas rate. Consequently, an appliance operatingwith 40% excess air at full gas rate will operate with nearly 200% excess air at just 2:1turndown (i.e. half-rate). This reduces thermal efficiency and, for fully aerated burners, will lead to major problems of poor combustion due to flame lift (see Section 4.3; Figures 4.13 and 6.1). Indeed, if any degree of turndown is required for a fanned draught fullyaerated system, then some form of control is necessary for both gas and air flows.

This section considers a number of methods for controlling the speed of (hence flow rate through) a.c. and d.c. fans. Although presented here as an independent control, thesesystems in practice will be connected electronically to the gas valve or full sequencecontroller.

7.3.1 A.C. fans

Fans fitted to domestic gas appliances are generally driven by two-pole a.c. brushless motors. The running speed is dependent on supply voltage and frequency, such that240V, 50 Hz gives a rotational speed of approximately 2800 rpm 144 Three methods of fan speed control are described here (see also references 144, 145).

7.3.1.1 Variable voltage control

The simplest and cheapest method of speed control is to vary the supply voltage to themotor by placing a resistor in series with the fan (Figure 7.14A ). There are a number of disadvantages with the arrangement:

1. Each motor speed requires a different resistor. If only two speeds, e.g. part and fullrate, are needed, then a single resistor may be switched in or out of circuit accordingly. If


a number of speeds are required, a variable resistor or a bank of individual resistors withassociated switching would have to be fitted.

2. With a resistor in circuit, unwanted electrical power is dissipated by the seriesresistor. In view of other available methods, this represents a waste of energy.

3. Energy is lost by the resistor to the surroundings as heat. Because the resistor mayneed to dissipate up to 20 W (depending on motor type and flow rate), the resistor mayneed to be quite large (physically) and will need to be adequately ventilated in order toprevent overheating.

Figure 7.14 Three method of controlling the rotational speed of a.c. fan motors.

Alternative methods of voltage variation are provided by a series inductance coil (choke)or by a step-down transformer. The latter may use fixed tappings taken from thesecondary winding or may be of a continuously variable type (variac). The drawbackwith these options is that the components tend to be somewhat bulky and expensive.

7.3.1.2 Phase angle control

All the voltage control options just described change the r.m.s. voltage by varying theamplitude of the sinusoidal a.c. supply. An alternative method of reducing r.m.s. voltageis to use what is known as a ‘phase angle controller’ 145 . A semi-conductor switch (triac)

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with its associated control circuit cuts out the same part of each half-cycle of the a.c. signal (Figure 7.14B, cf. light dimmer switch). Suitable electronic circuitry determinesthe extent to which the signal is altered, thereby enabling the voltage applied to the motorto be varied between 0% and 100% of nominal. Because the device works by breakingthe electrical circuit, no power is wasted (cf. a simple voltage dropping resistor) andhence there is no need for ventilation of the components. Controllers are fairlyinexpensive and are therefore suited to the domestic appliance market. One disadvantageis that certain motor types cannot be controlled effectively using this method. Forexample, with some capacitor-start capacitor-run motors, the capacitor will discharge through the coil when the circuit is broken, thereby perturbing the required waveform andadversely affecting motor performance.

7.3.1.3 Variable frequency control

Fan rotational speed is linearly dependent on the a.c. signal frequency 144 ; consequently, if, by using suitable electronic circuitry, the signal frequency can be controlled, then theair flow rate may be varied according to requirements. At the time of writing, variablefrequency controllers (e.g. as in Figure 7.14C) with a sufficiently low power delivery (less than 50 W) are only at a developmental stage and are substantially more expensiveand physically much larger than an equivalent phase angle controller. An advantage,however, of this device compared with all variable voltage systems is that full power isavailable to the motor at all fan speeds*. Consequently, motor torque is not affected to the same degree as when voltage is reduced (whether by dropping resistor or phase anglecontrol). Additionally, because a frequency controller can increase as well as decrease thesupply frequency, fan speed can be increased such that a small fan is able to perform thenominal duty of a larger fan.

7.3.2 D.C. fans

Although only fans with a.c. motors are currently employed in domestic gas appliances,the use of d.c. motors needs to be mentioned, if only to point out why they have notgenerally found favour. One major advantage of d.c. fans is that they are not subject tothe supply voltage/frequency (hence rotational speed) constraints of a.c. fans. Variation inthe d.c. signal (e.g. by using dropping resistors or a form of chopping device) permitsselection of a number of fan speeds. With the voltages generally used (up to about 25V),rotational speeds of up to 5000 rpm are feasible. This increases the static pressure headdeveloped by the fan and enables more restrictive heat exchangers or flue systems to beused.

Despite the relative ease with which fan speed maybe varied, d.c. motors possess anumber of disadvantages:

1. Domestic installations usually only have available a 240 V, 50 Hz a.c. electricalsupply. In order to provide d.c., a transformer and rectifier must be supplied in addition tothe speed controller; inevitably, this adds to appliance costs.

2. Simple d.c. motors rely on contact being made between two carbon brushes and a


commutator. There are two reasons why this may lead to problems. Firstly, suchmechanical contact must result in wear of the components; if such wear is excessive, thenpremature failure of the motor can ensue. Secondly, mechanical contact may produceisolated arcing between brush and commutator. In a gas appliance this is clearly

*Electrical power is dependent on the signal peak voltage and resistance; it is independent of signal frequency 144 .

hazardous, especially if the appliance is arranged in such a way that unburnt gas may passthrough the fan.

3. Brushless d.c. motors are available, but, including associated circuitry and controls,a complete package is currently much more expensive than a standard a.c. fan. Unlessthere is further development (coupled with adequate demand) for a brushless system, thecost seems likely to remain prohibitively high.

7.4 Air/gas ratio controls

The previous two sections have outlined methods and devices used in the control of gasand air flow rates. Although presented as separate controls, the two flows are linked, evenif indirectly, in that a satisfactory fan speed control must be able to act in response to asignal from the gas valve. However, all methods described will control the air only to apredetermined rate dependent on the position of the gas valve; they do not controlaccording to the actual gas flow, and are therefore insensitive to any transient variation ingas flow rate, gas composition, etc. This section outlines current developments in controlsystems that can respond to such variations and are able to maintain a preset air/gas ratioby use of a modulating gas valve and/or fan speed control. Forms of air/gas ratio havebeen used for many years in industrial systems (e.g. references 146, 147). Transfer of thattechnology to smaller domestic systems has been hampered by difficulties withminiaturization and the cost of any systems that have been developed. Recent work,particularly on microelectronic control, suggests that these drawbacks can now beovercome; it seems probable that within the next few years, sophisticated control systemswill be fitted to many appliances, particularly at the high-budget end of the market.

7.4.1 Zero governors

Figure 7.15 shows a schematic arrangement for a zero pressure governor in its simplestform. This method of control, which has been used successfully in industrial burners formany years 49 , 132 , 146 , supplies air and gas at atmospheric (i.e. zero gauge) pressure. Gas is supplied via a standard pressure governor, the zero governor and a control orificeinto a mixing tube or venturi, which is maintained at a negative pressure by the fan. Air isalso drawn into the manifold through a control orifice. Application of equation (1.6) toeach of the gas and air supplies shows that if the respective gas and air pressures

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upstream of the control orifices are equal, then the ratio of air flow rate to gas flow rate isdependent solely on the density of the gas and the area ratio of the two orifices. In orderto ensure that air and gas pressures are the same, an equalizing tube connects the airintake duct to the diaphragm chamber of the zero governor.

Figure 7.15 Basic layout for a zero governor air/gas ratio control.

Because of the dependence on orifice area ratio, a zero governed system will keep theair/gas ratio constant, even if the total flow rate changes, either intentionally or otherwise.(A zero governor will not maintain a constant air/gas ratio in the event of variation in gas composition.) For example, if a reduced heat demand is detected, a fan speed controller(as in Section 7.3) will decrease the total flow rate through the fan, but the air/gas ratioshould remain constant If pressure variations within the system (e.g. wind pressure at theflue terminal) affect fan performance, the total flow rate will change, but air/gas ratio willremain unchanged. An important by-product of the performance characteristic of a zero governor is that if the fan fails to operate, the gas rate automatically falls to zero. Thisprovides a built-in safety device in addition to any other flow-proving controls that may be fitted.

A number of different arrangements are possible; Hancock 146 has described sophisticated examples for industrial applications. Zero governed systems are now wellestablished in industry but have not as yet been scaled down satisfactorily for domesticapplications. One reason for this is cost; more importantly, though, at a technical level,the low pressure differential (less than 20 mbar) across the diaphragm of a domestic zero


governor means that the forces involved are considerably lower than at higher pressuresused industrially. As a result, unless large diaphragms are used, zero governors are notsensitive enough for domestic applications 148 .

7.4.2 Flue gas sensors

In order to maximize thermal efficiency, the excess air should be kept as low as ispossible without combustion performance deteriorating. Excess air is determined by theratio of the air and gas flow rates through an appliance. Consequently, any control ofexcess air level will also control the air/gas ratio in the unburnt gases. In other words, ifthe flue gases can be monitored and a processable signal obtained, then it is possible, inprinciple, to set up a control system whereby satisfactory combustion and high efficiencyare maintained by direct sensing of the combustion product gases.

Within the normal operating range of gas appliances, there are three components of theflue gases which are potentially suitable for monitoring, CO, CO2 and O2. Of these, only CO2 and O2 concentrations vary sufficiently with excess air, since carbon monoxideconcentration is virtually independent of excess air in the critical 20 to 60% excess airrange (Figure 6.1). Carbon dioxide detectors tend to be expensive. They can also giveambiguous results, since CO2 concentration decreases both with increasing excess air infuel-lean combustion and with decreasing air in fuel-rich combustion. An additional CO monitor could distinguish between fuel-rich and fuel-lean conditions, but the complete package would be prohibitively expensive. This leaves oxygen monitoring as the bestoption for domestic systems. O2 concentration is effectively zero for all fuel-rich combustion and increases steadily with excess air in fuel-lean systems (Figure 7.16). Additionally, the excess air/oxygen concentration relationship is virtually independent ofgas composition. This means that an oxygen sensor will react to changes in gascomposition and maintain the required efficiency and safety.

The oxygen sensor that has found most favour is an electrochemical system based on zirconia ZrO2, which at high temperatures behaves as a solid electrolytic solvent for oxygen anions. A review of electrochemical oxygen sensors is given by Maskell 149 . Figure 7.17 shows the most common form, viz., as a potentiometric device where twoplatinum electrodes are fixed either side of a slab of zirconia. If there is an oxygenconcentration difference between the two halves of the cell, an e.m.f. will be generated.One half-cell is usually designed as a reference (usually atmosphere), while the otherhalf-cell is exposed to the flue gases. The e.m.f. generated is dependent on temperature aswell as concentration difference (Figure 7.18). Consequently, a heating element is usuallyincorporated into the sensor in order to maintain a constant cell temperature (usuallyabout 700 °C). The output signal then becomes a function of O2 concentration (hence excess air) in the flue gases. A microprocessor control unit uses the signal to adjust eithera modulating gas valve or the air supply in order to maintain a preset excess air level.

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Figure 7.16 Oxygen concentration in wet flue gases as a function of excess air for the combustion of methane.

Figure 7.17 General layout for an oxygen sensor based on a zirconia solid electrolytic cell.


Figure 7.18 Temperature dependence of zirconia cell voltage on oxygen concentration.

One disadvantage of the potentiometric device is that the output signal is not a linearfunction of oxygen concentration; it can be used as a switch but is not so convenient foruse as a level control. Better in this respect is the amperometric sensor which comprisestwo zirconia discs sealed together and heated to a preset constant temperature (usually700 °C). The void within the device is pumped electrochemically such that the oxygen partial pressure in the internal void is reduced. This sets up a concentration differencebetween the void and the flue gases which provides a driving force for diffusion into thedevice, the rate of which is directly proportional to the concentration difference and thepumping current. If the concentration difference is sufficiently large, the current islinearly dependent on flue gas oxygen concentration.

With fully aerated burners, attention must be paid to the effect of gas composition onappliance performance (cf. Section 4.3.1.6). For example, consider a change from G20 toG21 (i.e. an increase in Wobbe number and TAR). If the sensor causes adjustment to theair rate (an ‘air-led’ system), an increase in air flow rate is needed in order to return to thedesign excess air. If the sensor causes adjustment to the gas supply (‘gas-led’), a decrease in gas rate is required. It can be seen that although air/gas ratio is kept constant, the totalflow rate and heat input differ. On changing to a higher Wobbe number gas, an air-led system will increase flow rate through the burner, while a gas-led system will decrease

Burner controls 175

flow. Burner port loadings and

Figure 7.19 Schematic representation of a ‘profiled’ respons to gas rate variation for a fully aerated burner with a tempreture limitation of 300 ºC. (From Patterson and Dann 148 .)

heat input will also be different, which may have a considerable effect on flame stabilityand flamestrip temperature. Thus the type of control used may influence the design of theburner (and vice versa). It should prove possible to ‘profile’ the response of the controller to the sensor signal. For instance, a higher excess air may be permitted on turndown inorder to compensate for higher burner temperatures at low port loadings (Figure 7.19). Each burner would then have a unique programmed response.

Until the last few years, the use of zirconia oxygen sensors has been confined toindustrial and commercial installations, owing to the cost of the sensors and their controlequipment. The fuel crisis of the mid-1970s stimulated an awareness in energy conservation and provided a stimulus for development of smaller, cheaper devices (e.g.references 141–143, 150–152). Latest work 143 , 145 suggests that oxygen sensors can, in principle, give extremely close control of excess air and that their use may soon beeconomically feasible when incorporated into a microelectronic control package. Little is


known, however, at this stage regarding long-term stability and reliability. Figure 7.20illustrates a control scheme for a boiler and shows the level of complexity that is possiblein future appliances. A microprocessor forms the heart of the system. To this areconnected thermostats, a modulating gas valve, a fan speed controller, spark ignition, aflame rectification FFD and an oxygen sensor in the flue gases. Assuming that currentdevelopments achieve their objective, microelectronic control packages such as thisshould be introduced into domestic appliances within the next few years.

Figure 7.20 Hypothetical control scheme for future domestic boiler system.

References Subject index Author index Index of Symbols

Burner controls 177

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Combust. p. 613 (1953). 55 Reed, S.B., Combust. Flame 11 , 177 (1967). 56 Reed, S.B., J. Inst. Gas Eng. 8 , 157 (1968). 57 Edmondson, H. and Heap, M.P., 12th Int. Symp. Combust. p.1007 (1969). 58 Edmondson, H. and Heap, M.P., J. Inst. Gas Eng. 11 , 305 (1971). 59 Melvin, A. and Moss, J.B., Combust. Sci. Technol. 7 , 189 (1973). 60 Gunther, R. and Janisch, G., Gas Warme Int. , 24 , 489 (1975). 61 Haniff, M.S. and Melvin, A., 18th Int. Symp. Combust. p. 657 (1981). 62 Haniff, M.S. and Melvin, A., J. Inst. Energy 57 , 432 (1984). 63 Potter, A.E., Progress in Combustion Science and Technology , Vol. I (ed. Ducarme,

J., Gerstein, M. and Lefebvre, A.H.) p.145, Pergamon (1960). 64 Friedmann, R., 3rd Int. Symp. Combust. p. 110 (1949). 65 Berlad, A.L. and Potter, A.E., 5th Int. Symp. Combust. p. 728 (1955). 66 France, D.H., J. Inst. Fuel 50 , 147 (1977). 67 Kay, J.M. and Nedderman, R.M., Fluid Mechanics and Transfer Processes ,

Cambridge Univ. Press (1985). 68 Dutton, B.C. and Gimzewski, E., J. Inst. Energy 56 , 107 (1983). 69 Dutton, B.C. and Wood, S.W., J. Inst. Energy 57 , 381 (1984). 70 Reed, S.B. and Wakefield, R.P., J. Inst. Gas Eng. 10 , 77 (1970). 71 Reed, S.B., Datta, P. and Mineur, J., J. Inst. Fuel 44 , 1 (1971). 72 Reed, S.B., Mineur, J. and McNaughton, J.P., J. Inst. Fuel 44 , 149 (1971). 73 Smith, D.J. and White, I.F., J. Inst. Fuel 49 , 30 (1976). 74 South, R., Gas Eng. Manag. 14 , 225 (1974). 75 Weber, E.J., Amer. Gas Assoc. Res. Bull. 79 (1960). 76 Anon., Amer. Gas Assoc. Res. Bull. 20 (1944). 77 Anon., Amer. Gas Assoc. Res. Bull. 38 (1946). 78 Francis, W.E. and Jackson, B., Trans. Jnst. Gas Eng. 107 , 55 (1957/8).

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79 Heslop, D.T. and Sussex, A.D., Inst. Gas Eng. Comm. 1235 (1984). 80 Pearson, E.J., Saunders, T.G. and Hargreaves, K.J.A, 3rd Int. Gas Res. Conf. , Paper

IGRC/C24–83 (1983). 81 Minchin, L.T., Gas World , 125 , 148 (1946). 82 Kilham, J.K. and Lanigan, E.P., J. Inst. Gas Eng. 10 , 700 (1970). 83 DeWerth, D.W., Amer. Gas Assoc. Res. Bull. 92 (1962). 84 Weil, S.A., Inst. Gas Tech. Res. Bull. 35 (1964). 85 Griffiths, J.C., Thompson, C.W. and Weber, E.J., Amer. Gas Assoc. Res. Bull. 96

(1963). 86 Schreiber, R., Krill, W., Kesselring, J., Vogt, R. and Lukasiewicz, M., 3rd Int. Gas

Res. Conf. , Paper IGRC/D06–83 (1983). 87 Krill, W. and Kesselring, J., Proc. 4th Int. Gas Res. Conf. p. 719 (1984). 88 Coles, K.F. and Bagge, L.P., J. Inst. Gas Eng. 11 , 387 (1971). 89 Coles, K.F. and Wilbraham, K.J., Inst. Gas Eng. Comm. 953 (1974). 90 Coles, K.F. and Shirvill, L.C., Inst. Gas Eng. Comm. 1225 (1983). 91 Anon., Amer. Gas Assoc. Res. Bull. 25 (1944). 92 Mugridge, B.D., Hughes, C. and Roberts, C.A., Inst. Gas Eng. Commun. 1048 (1977). 93 Putnam, A.A. and Brown, D.J., Combustion Technology: Some Modern Developments

(ed. Palmer, H.B. and Beer, J.M.) Academic Press (1974). 94 Tsuji, H. and Takeno, T., 10th Int. Symp. Combust. p.1327 (1965). 95 Schimmer, H. and Vortmeyer, D., Combust Flame 28 , 17 (1977). 96 Radcliffe, S.W. and Hickman, R.G., J. Inst. Fuel. 48 , 208 (1975). 97 Sadamori, H., Chikazawa, A., Okamura, S. and Noda, T., Proc. 2nd Int. Gas Res.

Conf. p.1467 (1981). 98 Dongworth, M.R. and Melvin, A., 16th Int. Symp. Combust. 255 (1977). 99 Pearson, E.J., Gas Eng. Manag. 20 , 247 (1980). 100 Thring, M.W. (ed.) Pulsating Combustion, The Collected Works of F.H.Reynst

Pergamon (1961). 101 Griffiths, J.C. and Weber, E.J., Amer. Gas Assoc. Res. Bull. 107 (1969). 102 Davis, D.D., Handbook of Noise Control (ed. Harris, C.M.), p.21–1, McGrawHill

(1957). 103 Chiu, H.H., Clinch, J.M., Blomquist, C.A. and Croke, E.J. Proc. 2nd Int. Gas Res.

Conf. p.1530 (1981). 104 Proceedings from Symposium on Pulse-Combustion Applications, Gas Research

Institute Report GRI-82/0009.2 (1982). 105 Vishwanath, P.S., Proc. 4th Int. Gas Res. Conf. p. 912 (1984). 106 Francis, W.E., Hoggarth, M.L. and Reay, D., J. Inst. Gas Eng. 3 , 301 (1963). 107 Reay, D., J. Inst. Fuel. 42 , 135 (1969). 108 Hanby, V.I., Heat. Vent. Eng. 53(2), 6 (1979). 109 Ahrens, F.W., Kim, C. and Tam, S.W., Trans. ASHRAE 84(1), 488 (1978). 110 Lee, J.H. and Soedel, W., 3rd Int. Gas Res. Conf. . Paper IGRC/C23–83 (1983). 111 Briffa, F.E.J., Staddon, P.W., Phillips, R.N. and Romaine, D.R., Inst. Gas Eng.

Commun. 860 (1971). 112 Chaplin, B., Chart. Mech. Eng. 30(1), 41 (1983). 113 Davidson, J.F. and Harrison, D., Fluidised Particles , Cambridge Univ. Press (1963). 114 Davidson, J.F. and Harrison, D. (ed.) Fluidisation Academic Press , (1971). 115 Searle, M. and White, M., Inst. Gas Eng. Comm. 1223 (1983). 116 BS 5258, Specification for Safety of Domestic Gas Appliances , BSI.

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117 BS 5386, Specification for Gas-burning Appliances BSI. 118 BS 6332, Thermal Performance of Domestic Gas Appliances BSI. 119 Melia, R.J.W., Florey, C. du V., Sittampalam, Y. and Watkins, C., Proc. 6th World

Congress Air Qual. p.263 (1983). 120 Ogston, S.A., Florey, C. du V., Walker, C.H.M., Br. Med. J. 290 , 957 (1985). 121 Fenimore, C.P., 13th Int. Symp. Combust. p. 373 (1971). 122 Hargreaves, K.J.A., Harvey, R., Roper, F.G. and Smith, D.B., 18th Int. Symp.

Combust. p. 133 (1981). 123 Jasper, G. (ed.), Gas Service Technology (3 volumes), Ernest Benn (1978). 124 Henshilwood, C.P. and Prigg, J.A., J. Inst. Gas Eng. 6 , 106 (1966). 125 Curran, A.H. and Green, M.B., 16th World Gas Conf. , Paper IGU/E4–85 (1985). 126 BS 5440, Code of Practice for Flues and Air Supply for Gas Appliances of Rated

Input not exceeding 60 kW , BSI. 127 Anon., Gas Res. Int. Dig. 8(4), 33 (1985). 128 Flood, J.J.F. and Enga, T.G., 3rd Int. Gas Res. Conf. , Paper IGRC/C09–83 (1983). 129 Shukla, K.C., Hurley, J.R. and Lockwood, J.W., Proc. 4th Int. Gas Res. Conf. p.1078

(1984). 130 Curran, A.H., Newcombe, J. and Price, M.D., Inst. Gas Eng. Comm. 1261 (1985). 131 Miles, V.C. and Pinkess, L.H., The Principles and Practice of Gas Appliance

Controls , Walter King (1970). 132 British Gas School of Fuel Management, Gas Controls (Industrial Gas Development

Committee Report 763/69) (1969). 133 Finch, P.J., Miles, A.J. and Moore, N.G., Inst. Gas Eng. Commun. 1197 (1982). 134 Ekins, B. and Brown, A.M., J. Inst Gas Eng. 8 , 223 (1968). 135 Sayers, J.F., Tewari, G.P., Wilson, J.R. and Jessen, P.F., J. Inst. Gas Eng. 11 , 322

(1971). 136 Miles, A.J., J. Inst. Gas Eng. 11 , 339 (1971). 137 Ekins, B. and Wilson, J.R., Inst. Gas Eng. Gas Council Comm. GC185 (1971). 138 Curran, A.H., Ross, N.C. and Sayers, J.F., Inst, Gas Eng. Comm. 1163 (1981). 139 Atkinson, P.G. and Hancock, R.A., J. Inst. Gas Eng. 1 , 550 (1961). 140 Atkinson, P.G., Grimsey, R.M. and Hancock, R.A., J. Inst. Gas Eng. 8 , 341 (1968). 141 Stuurgroep Herontwikkeling Aardgastoestellen met Kontrolerende Electronica

(SHAKE), Microelectronics and Gas Applications (Symposium Report) (1981). 142 Franx, C., Inst. Gas Eng. Comm. 1198 (1982). 143 Bergman, A.P. and Franx, C., Proc. 4th Int. Gas Res. Conf. p.1053 (1984). 144 Osborne, W.C. and Turner, C.G., Woods Practical Guide to Fan Engineering (2nd

edn), Woods (1960). 145 Hargreaves, K.J.A. Jones, H.R.N. and Smith, D.B., Inst.GasEng.Comm. 1309 (1986) 146 Hancock, R.A., J. Inst. Gas Eng. 5 , 470 (1965). 147 Churchill, D.A., Inst. Gas Eng. Comm. 1108 (1979). 148 Patterson, M.C. and Dann, R.G., Inst. Gas Eng. Comm. 1270 (1985). 149 Maskell, W.C., J. Phys. E. Sci. Instrum. 20 , 1156 (1987). 150 Pegler, S.M., Jones, G.E. and Weall, P., Inst. Gas Eng. Comm. 1161 (1981). 151 Bergman, A.P., Proc. 2nd Int. Gas Res. Conf. p. 987 (1981). 152 Sakurai, K., 3rd Int. Gas Res. Conf. , Paper IGRC/D02–83 (1983).

References 181

Subject index

aerated flames 1 –2 air controls 177 –81 air entrainment 39–51, 179 air/gas ratio control 181 –7 appliances

British Standards requirements for 136 –7 classification of 12

autoignition 54 auxiliary flames see retention flames

bar burner 34, 47, 48–9, 157 –9 blade burner 143–4, 146 blow-off see flame lift boilers

burner controls for 162, 165 169, 172, 187 burners for 90, 106, 140 –6 range-rated 165, 166

box burner 34, 83, 143–4, 146, 147, 157 Bray burner see wedge-cavity burner Bunsen burner 2–3, 7 burn-back sooting 28 burner controls see controls burner overheat 28, 29, 96, 100, 106 –7 burner temperature

influence on flame lift 61 on flame quenching 64

of radiant burners 113 of ribbon burners 95–100, 104 –6

burning velocity 6 –8 burnt gases (in flame) 2

carbon monoxide 9, 137 –40 catalytic combustion 118 –23 ceramic burners see radiant burners combustion diagrams 29, 52, 56, 100–1, 104 combustion (CO/CO2) ratio 136, 137 –40 controls 161 –87

air 177 –81

air/gas ratio 181 –7 flame failure devices 166 –71 governors 165–6, 181 –2 ignition 173 –7 thermostats 164, 171 –3 timers 164, 171 –3 valves 162 –5

convector heaters see under space heaters cookers

grill burners for 34, 90, 158–60, 162 hotplate burners for 34, 61, 151–7, 162 oven burners for 34, 157–8, 162

diffusion flames see non-aerated flames draughts 79–80, 115 –7

excess air

definition of 9 effect on combustion ratio of 137 –40

extinction, lightback on 67, 68–9, 85

fans

a.c. 178 –80 d.c. 180 –1 selection of 88 –90 speed control of 178 –90

flame failure devices (FFDs) 166 –71 bimetallic strip 168 flame conduction 171 flame rectification 171 infra-red/ultra-violet 171 mercury vapour 168 thermocouple 168 –9

flame front see reaction zone of flame flame height 18, 19 –21 flame lift 7–8, 112, 136 –8

fully aerated burners 91–5, 106–7, 112 partially aerated burners 52, 53 –61

effect of gas composition on 73 effect of vitiation on 74 –8

flame stretch 53 flame structure 1 flammability limits 4 –61 flue gas sensors 182 –7 fluidized bed combustion 129 –32 frequency control 179, 180

Subject index 183

fuel cells 135 fully aerated burners 86 –

flame stability 91–5, 101–4, 106–7, 111 influence of gas composition 104–6, 184 –6 temperature of 95–9, 104 –7 using atmospheric injection 86–8, 138 using fanned draught 88–91, 117, 138, 146, 151

fully aerated flames 1, 2, fully premixed burners see fully aerated burners fully premixed flames see fully aerated flames

gas composition

influence on burner temperature 104–6, 114 on flame stability 72 –3

gas engines 133 –5 gases

classification of 12 first family 12, 14 interchangeability of 12 –5 physical properties of 11 second family 12, 14 –5 test 16 –7 third family 12

Gilbert-Prigg diagrams 13 –4 governors 165 –6

zero-pressure 166, 181 –2 grills see under cookers

heat pumps 135 hotplates see under cookers

ignition 173 –7

full sequence control 177 hot filament 174 –5 lightback on 67 –8 minimum energy for 62 –6 pilot 173 –4 spark 175 –7

incomplete combustion 52, 69–72, 76, 138 injectors 35 –9

flow rate from 35 manufacturing tolerances for 61 –2 multihole 62 –39 noise from 39

jetted burners 34, 83, 144 –5

Subject index 184

lightback 8, 52, 62–9, 95, 106–7, 112, 114

on extinction or ignition 67–9, 85 lint 18, 69, 80–5, 117, 144, 146, 151 lint-resistant burners 82 –3 live-fuel effect heaters see under space heaters

manufacturing tolerances, influence on burner performance 29, 73 –4 matrix burners 28–9, 145 minimum ignition energy 61 –6 mixing tubes 34, 48, 49, 143 multiplex burners 148 –9

natural gas, composition of 12 nitrogen oxides (NOx) 9, 138, 140 33 noise

burner 90, 114–5, 127–31, 165 injector 165 –39

non-aerated burners 18 –32 non-aerated flames 1, 2–3, 18 –21

height of 19 –21 stability of 21 structure of 18 –9

ovens see under cookers oxygen sensors 183 –7

partially aerated burners 34–85, 133, 147

air entrainment design equations 39 –51 flame lift 52, 53 –61 lightback 62 –9 port design 51 –72

partially aerated flames 1, 2 stability of 51 –72

phase angle control 179 –80 piezo-electric ignition see ignition, spark pilots see under ignition pinhole burners 26 –8 port geometry

influence on flame height 21 on flame lift 55, 56 –60 on flame quenching 63 –4 on yellow tipping 69 –72

preheat zone of flame 2 pressure regulators see governors

Subject index 185

primary aeration calculation of 44 –6 definition of 9 effect of air temperature on 46 –7

primary air 1, 2 pulsed combustion 124–9, 151

noise suppression 124, 127 –9 punched metal burners 106–7, 115, 143

quenching 62 –4 quenching diameter 62 –3

radiant (surface-combustion) burners 90, 107– , 117, 147

flame stability 112, 114 influence of gas composition on 114 porous medium 107 radiant plaque 88, 107– , 141, 143, 150, 156–7, 159 –60

radiant-convector heaters see under space heaters

radiant efficiency 109 – reaction zone of flame 2 regulators see governors resonance see noise, burner retention flames 26, 60–1, 101 –5 ribbon burners 91–106, 115, 147

burner temperature 95 –100 flame stability 91 –5

rod-stabilized burners 30, 32

secondary air 1, 2, 86 sensors

flue gas 182 –7 oxygen 183 –7

simplex burners 148–9, 158 soot 18–9, 27, 28, 30, 32, 69, 73 space heaters 118, 146 –51

burner controls for 162, 163, 168, 169, 172 convector 146, 147, 149 live-fuel effect 151 radiant-convector 88, 146, 147, 149 –51

spark ignition see under ignition spiral burner 133 standards 136–7, 161 stoichiometric combustion 2, 6 submerged combustion 133 substitute natural gas (SNG) 14

Subject index 186

sulphur oxides 10 surface-combustion burners see radiant burners

theoretical air requirement (TAR) 8 –9 thermostats 164, 171 –3 timers 164, 171 –3

Uniplane burner 30 –2

valves, gas 162 –5

modulating 162, 164 –5 vitiation, influence on flame stability 74–8, 115 voltage control 178 –9

wall attachment 101, 103 –4 water heaters

burner controls for 165, 168, 169, 172 burners for 106 –46

Weaver coefficient 11 Weaver flame speed factor 10 wedge-cavity burners 21, 23 –5 Wobbe number 10 –1

yellow tipping 52, 69 –72

zero-pressure governors 166, 181 –2 zirconia sensors see oxygen sensors

Subject index 187

Author index

Ahrens, F.W. 125 Andrews, G.E. 7 Atkinson, P.G. 176 (2)

Bagge, L.P 112 (2), 112 (2) Barnard, J.A. 1 Barr, 131 Bergman, A.P. 176, 186 (3) Berlad, A.L. 63 Berry, W.M. 49 Blomquist, C.A. 123, 125, 127 Bradley, D. 7 Bradley, J.N. 1 Briffa, F.E.J. 126, 128 Brown, A.M. 123, 171, 174 Brown, D.J. 114 Brumbaugh, I.V. 49 Burke, S.P. 20

Chaplin, B. 128 Chikazawa, A. 117 (2), 120, 122 Chiu, H.H. 123, 125, 127 Churchill, D.A. 180 Clinch, J.M. 123, 125, 127 Coles, K.F. 112 (4), 112 (2) Connor, N.E. 39 (2), 55, 113, 114, 132 (2), 160, 164, 169 (2), 171, 180 Cooper, J.R. 7, 11 Croke, E.J. 123, 125, 127 Culshaw, G.W. 25, 31, 67 Cunningham, A.C. 21 Curran, A.H. 145, 159, 171, 174

Dance, E.W.G. 38, 113 Dann, R.G. 181, 186 Datta, P. 74, 75 Davidson, J.F. 130 (2) Davis, D.D. 123, 127

Delbourg, P. 11 Denniston, D.W. 52 Desty, D.H. 28 De Werth, D.W. 106, 107 Dongworth, M.R. 118, 121 Dutton, B.C. 16 (3), 25 (2), 72 (5)

Edmondson, H. 21, 23, 25, 53 (2) Ekins, B. 171 (2), 174 (2), 175 Enga, T.G. 151, 153 (2), 156 (3), 157, 158, 173

Fenimore, C.P. 140 Finch, P.J. 161, 171, 176 Flood, J.J.F 151, 153 (2), 156 (3), 157, 158, 173 Florey, C. du V. 136 (2) Forsyth, J.S. 21, 114 France, D.H. 65 Francis, W.E. 39, 46, 49, 89 (2), 126 (3), 126 (2) Franx, C. 176 (2), 186 (3) Friedmann, R. 62

Garside, J.C. 21, 114 Gaydon, A.G. 1, 9, 21, 60, 140 Gazley, C. 21 Gilbert, M.G. 11, 13, Gimzewski, E. 72 Glassman, I. 1 Goodwin, C.J. 39, 44, 47 (2), 48 Green, M.B. 145 Griffiths, J.C. 51, 56 (3), 57, 69 (3), 69 (2), 70, 107, 117, 118, 123 (4), 124, 125, 126 (2), 132 Grimsey, R.M. 176 Grumer, J. 176 Gunther, R. 52 Guy, J.J. 39 (2), 55, 113, 114, 132 (2), 160, 164, 169 (2), 171, 180

Hanby, V.I. 125 Hancock, R.A. 176 (2), 180 (2), 182 Haniff, M.S. 52 (2) Hargreaves, K.J.A. 90 (2), 91 (2), 93, 95, 100, 140, 177, 178, 186 Harris, J.A. 14, 15, 25 (2), 27, 28, 29, 46, 51, 53, 55 (3), 56, 58, 59, 61 (2), 63, 66 (2), 67, 72 (2), 73, 74, 76 (3), 76 (2), 79, 80, 81, 82 Harrison, D. 130 (2) Harvey, R. 140 Hawthorne, W.R. 21 Heap, M.P. 53 (2) Henshilwood, C.P. 143

Author index 189

Heslop, D.T. 89, 150 Hickman, R.G. 118 (2), 118 (2), 120 (3) Hoggarth, M.L. 39, 44, 47 (2), 48, 125 (3), 126 (2) Hottel, H.C. 21 Hughes, C. 113 Hurley, J.R. 156

Jackson, B. 89 Janisch, G. 53 Jasper, G. 142, 161, 169, 171, 173 Jessen, P.F. 14, 171, 174 (2), 175 Jones, G.E. 186 Jones, H.R.N. 177, 178, 186

Kapp, N. 21 Karlovitz, B. 53 Kavarana, B.J. 25 (2) Kay, J.M. 64, 130 Kesselring, J. 111 (6), 112 Kilham, J.K. 107 (3), 107, 111 Kim, C. 125 Knapschaefer, D. 52 Krill, W. 111 (6), 112 Kurz, P.F. 21, 75

Lambert, G.M.S. 14 Lanigan, E.P. 107 (3), 107, 111 Lee, J.H. 125 Lewis, B. 1, 63, 6, 53, 65 Lockwood, J.W. 156 Lovelace, D.E. 14 Lukasiewicz, M. 111 (3), 112

Maskell, W.C. 184 McNaughton, J.P. 74 Melia, R.J.W. 136 Melvin, A. 53 (3), 118, 121 Miles, A.J. 161, 171, 173, 176 Miles, V.C. 161 (2), 163, 165, 171 (2), 171, 173 Minchin, L.T. 23, 106 Mineur, J. 74 (2), 75 Moore, N.G. 160, 171, 176 Moss, J.B. 53 Moulton, G.F. 49 Mugridge, B.D. 113

Author index 190

Nedderman, R.M. 64, 130 Newcombe, J. 159 Noda, T. 117 (2), 120, 122

Ogston, S.A. 136 Okamura, S. 117 (2), 120, 122 Osborne, W.C. 177 (2), 179 (2)

Patterson, M.C. 181, 186 Pearson, E.J. 90 (2), 91 (2), 93, 95, 100, 119, 121, 129 Pegler, S.M. 186 Phillips, R.N. 126, 127 Pinkess, L.H. 161 (2), 163, 166, 171 (2), 171, 173 Potter, A.E. 62 Price, M.D. 159 Prigg, J.A. 11, 13 (2), 13, 23, 25, 31, 36, 37 (2), 39, 44, 46, 48, 49, 51, 55 (2), 67, 143 Pritchard, R. 39 (2), 55, 114, 132 (2), 161, 165, 169 (2), 171, 180 Purkis, C.H. 11 Putnam, A.A. 114, 123

Radcliffe, S.W. 118 (2), 118 (2), 120 (3) Reay, D. 39, 44, 47 (2), 48, 125 (4), 126 (2) Reed, S.B. 53 (2), 73, 74 (2), 75 Roberts, C.A. 113 Romaine, D.R. 126, 127 Rooke, D.E. 13 Roper, F.G. 21 (3), 140 Rose, J.W. 7, 11 Ross, N.C. 171, 174

Sadamori, H. 117 (2), 122 Sakurai, K. 186 Saunders, T.G. 90 (2), 91 (2), 93, 95, 100 Sayers, J.F. 171 (2), 174 (3), 175 Schimmer, H. 114 Schlichting, H. 23 Schreiber, R. 112 (3), 112 Schumann, T.E.W. 20 Searle, M. 132 (2), 145 Shah Jahan 23 Shawn, G.B. 49 Shirvill, L.C. 111 Shukla, K.C. 156 Silver, R.S. 156

Author index 191

Simmonds, W.A. 39 (3) Sittampalam, Y. 136 Smith, C. 21 Smith, D.B. 140, 177, 178, 186 Smith, D.J. 75 Soedel, W. 125 South, R. 15, 29, 51, 53, 55 (2), 56, 58, 59, 61 (2), 63, 66 (2), 67, 72 (2), 73, 74, 76 (3), 76 (2), 78, 79, 80 Staddon, P.W. 126, 127 Sussex, A.D. 89, 150 Sutherland, A. 38 (2), 38, 113 (2)

Takeno, T. 114 Tam, S.W. 125 Tewari, G.P. 171, 174 (2), 175 Thompson, C.W. 107, 117, 118, 123 (2), 124, 125, 132 Thring, M.W. 123 Townend, D.T.A. 21, 114 Tsuji, H. 114 Turner, C.G. 177 (2), 179 (2)

van der Linden, A. 24, 25 Vishwanath, P.S. 123, 125, 127 Vogt, R. 111 (3), 112 vonElbe, G. 1, 54, 6, 52, 53, 62, 64 Vortmeyer, D. 114

Waight, J.F. 114 Wakefield, R.P. 73 Walker, C.H.M. 136 Watkins, C. 136 Weall, P. 186 Weaver, E.R. 11 Weber, E.J. 51, 56 (3), 57, 69 (3), 69 (2), 70, 81, 107, 117, 118, 123 (4), 124, 125, 126 (2), 132 Weil, S.A. 107 Wells, F.E. 52 Westbrook, C.K. 9 Westwood, C.R. 29 White, I.F. 75 White, M. 132 (2), 145 Whitehead, D.M. 28 Wilbraham, K.J. 111 Williams, F.A. 1, 54, 7, 19 Wilson, J.R. 27, 28, 29, 81, 82, 171 (2), 174 (3), 175 (2) Wohl, K. 21 Wolfhard, H.G. 1, 9, 21, 60, 140

Author index 192

Wood, S.W. 72

Zabetakis, M.G. 4, 54, 63

Author index 193

Index of symbols

Greek symbols

A Area; Primary aeration CV Calorific value C L Loss coefficient c Concentration c d Discharge coefficient D Diffusion coefficient D, d Diameter F Weaver flame speed coefficient; F Bar burner correction factor f Friction factor g Acceleration due to gravity h Height K Pressure correction factor; Proportionality constant L Length n Aspect ratio of rectangular burner port n c Critical number of ribbon elements P. P Pressure V Volume flow rate q Heat input R Ratio of entrained air flow rate to gas flow rate Re Reynolds number r Radial coordinate S Interport separation; Burning velocity t Time u, v Velocity W Wobbe number x Mole fractions-coordinate y y-coordinate

Subscripts

Thermal diffusivity ε Combustion efficiency η Pressure efficiency µ Dynamic viscosity ρ Density σ Relative density (specific gravity)

a Air g Gas i i th component

j Injector m Air/gas mixture o Ambient p Burner port t Venturi throat

Index of symbols 195

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H.r.n. Jones-Application of Combustion Principles to Domestic Gas Burner Design (1990)

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