Doctor of Philosophy University of Surrey HEAT TRANSFER IN ROTARY … · The measurement of flame...

Thesis .submittea for the. degree of Doctor of Philosophy in Chemical Engineering

at the University of Surrey

HEAT TRANSFER IN ROTARY CEMENT KILNS

by

Barrie George Jenkins

B.Sc. University of Surrey (197i)

Chemical Engineering Department University of Surrey

Guildford Surrey.

September 1976.

ProQuest Number: 10798264

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.SUMMARY .

The results of an extensive series of trials on a 100 tonnes per day cement kiln have provided a fundamental insight into the aerodynamics, burning mechanism and heat transfer in the sintering zone of such kilns. Accurate monitoring of the input and output variables has enabled mass and heat balances to be made on the system, and from these results it has been possible to isolate the areas where major fuel savings canibe achieved. Slurry moisture, excess air and external heat losses are all variables where improvements and better -control would reduce fuel consumption and increase the efficiency of rotary kilns.

It has been shown that the external temperature profile of the kiln shell provides a useful indication of the various reaction regions that exist in the process cycle. From calculations of the heat lost from the shell , it has been shown that half the external heat losses occur from the sintering zone of the kiln.

Specialized instrumentation has been developed to measure gas temperatures and extract combustion gas samples during the normal range of operation of the kiln. An analysis of these results has led to a formula to predict the length of the flame as influenced by the significant operating parameters. The measured gas concentrations have been used to predict the combustion rate within the flame, and a favourable comparison of this rate has been made with published data.

The measurement of flame temperatures in the kiln has shown that the average flame temperature that may be encountered in a cement kiln is approximately 1800°C. Point temperatures of up to 2100°C were measured, and it was observed that increased excess air produced a shorter, hotter flame, but reduced the temperature of the combusted gases, resulting in a poorer quality product.

A mathematical model has been developed to predict the gas and refractory temperature and heat flux profiles oecuring in a rotary kiln sintering zone. The method is based on that of Hottel and Sarofim, but modified to account for the specialised firing conditions necessary for cement production.

The model has been tested against the measured data obtained from the kiln trials., and the degree of agreement found to be encouraging. Use of this model should enable the cement, lime, and refractory industries to comprehend the effect of changes to operational variables, with a resulting improvement in heat utilisation, product quality and plant life.

ACKNOWLEDGEMENTS

The author wishes to express his sincere gratitude to

Mr. F. D. Moles for his guidance and encouragement throughout this research programme., for devoting so much of his time to the experimental trials, and for providing thethe photographic material in this thesis.

The entire staff of the Barnstone Works for theirassistance and cooperation during the trials, and especially to Mr. P Dover and Mr. C Jones for their invaluable advice on the design and construction of modifications to the furnace, without which the results in this thesis would not have been possible.

Dr. Z. A. Syed for his assistance in the design and construction of the integral probe cooling unit.

The Associated Portland Cement Manufacturers Ltd. for so generously making available the entire Barnstone Works and staff, and the analytical section of their Research Department.

The Science Research Council for the award of a Research Studentship to permit the realization of this work.

CONTENTS

SECTION 1. INTRODUCTION' 1

SECTION 2. LITERATURE SURVEY OF PREVIOUS WORK 42.1 Fundamental studies of enclosed turbulent

diffusion flames 42.1.1 Jet entrainment 42.1.2 Recirculation 62.1.3 Combustion of pulverized fuel clouds 10

2.2 Evaluation of kiln operating data 132.2.1 Critical analysis of production

capacity 132.2.2 Significant variables 15

2.3 Practical measurements in rotary kilnsystems 16

2.4 Theoretical studies of rotary kiln systems 24

SECTION 3. THEORETICAL PREDICTION OF HEAT TRANSFER;THE DEVELOPMENT OF A MODEL 29

3.1 General summary of approaches . 293.1.1 Zone methods 303.1.2 Flux methods 303.1.3 Monte Carlo methods 313*1.4 Fluid mixing methods 31

3*2 Summary of prediction of heat transfer incylindrical furnaces by the zone method 323.2.1 Gas data 323.2.2 System zoning 333.2.3 Evaluate direct exchange areas 333.2.4 Evaluate total exchange areas 333.2.5 Evaluation of net radiative exchange 343.2.6 Formulate total energy balance 353.2.7 Heat flux determination ’ 36

3.3 Representation of a real gas by a series ofgrey gases to simulate the emissivity/ absorptivity characteristics 36

3.4 Representation of soot particles by a seriesof grey gases to simulate the emissivity characteristibs 39

3.5 Radiative properties of dust particles andtheir incorporation into the grey gas approximation principle 4l

3.6 Direct exchange factor evaluation 443.7 Accounting for local variation of

concentration of absorbing media 473.8 Convective heat transfer within the furnace 483.9 Heat loss to the atmosphere by conduction

through the furnace walls 513.9.1 Convective heat loss from a tube 513.9.2 Radiative heat loss from a tube 533.9.3 Total heat loss from a tube 53

3.10 Evaluation of gas enthalpy terms 543-11 Prediction of heat release rates for kilns 553.12 Distribution of mass flows in an enclosed

jet- system 56

SECTION 4. DESCRIPTION OF EXPERIMENTAL APPARATUS ANDOPERATIONAL TECHNIQUES 60

4.1 Brief description of the rotary kiln 604.2 Brief description of the cement making

process equipment 6l4.2.1 Wet process 6l4.2.2 Semi-wet process 634.2.3 Dry process 63

4.3 Temperature and mass flow measurements 644.3.1 Mass flows 644.3.2 Heat flows 714.3.3 Material analysis 76

4.4 Flame measurements 784.4.1 Stationary gas sampling 784.4.2 Rotating gas sampling 804.4.3 Gas sample analysis 874.4.4 Flame temperature measurements 91

SECTION. 5. RESULTS OF THE EXPERIMENTAL AND COMPUTATIONALINVESTIGATION OF THE SINTERING ZONE OF A KILN 101

5.1 Experimental results 1015.1.1 General kiln operating conditions 1015.1.2 Heat and mass balances 1055.1.3 External heat loss 1145.1.4 Gas concentration profiles in the

flame region 1205.1*5 Combustion rate analysis 1295.1.6 Temperature measurement in the flame

region 1325.2 Prediction of heat transfer in a rotary kiln 135

5.2.1 Brief description of program 1355*2.2 Model assumptions 1375.2.3 Model testing data 1385.2.4 Convective and conductive heat transfer

coefficients 1485.2.5 Results of the computational simulationi49

5*3 Discussion of predicted and experimentalresults 1545.3.1 Thermodynamic processes ' 1555.3.2 Aerodynamic processes 157

SECTION 6. CONCLUSIONS AND RECOMMENDATIONS FOR FURTHERWORK 158

REFERENCES 160

NOMENCLATURE 16'8

APPENDIX A. Derivation of weighting factors andabsorption coefficients for a grey gassimulation of the radiation characteristicsof a real gas 170

APPENDIX B. Derivation of exchange factors for cylindrical geometry 176

APPENDIX C, Calculation ...of the overall heat of. formation ' of clinker from dry slurry with reference

to Barnstone Trials raw materials 192

APPENDIX D

APPENDIX E

APPENDIX F

APPENDIX"G

Measured kiln operatind data and chemical analyses for Barnstone Trial A (2-6.4.73) 198

Results of combustion gas analyses and temperature measurements taken from the flame region of the Barnstone kiln during Trial. 4 (2-6.4.73) '202

Calculation of dust particle mean diameter in the sintering zone of the kiln 208

Theoretical estimation of the velocity decay of the Barnstone kiln flame jet operating at 17.24# excess air 210

APPENDIX H. University of Surrey program GSECKC1Z compilation listing 216

LIST OF FIGURES

SECTION. 2,FIGURE 2.1 Position of recirculation eddy in a

confined jet FIGURE 2,2 Comparison of measured temperature

data for gas and solids in a cement kiln

SECTION 3.FIGURE 3.1 Representation of gas absorption

coefficients with bands in the same spectral range

FIGURE 3-2 Variation of weighting factor with temperature for q q 2= '1

FIGURE 3*3 Comparison of grey gas fit with data for a range of temperatures

FIGURE 3.^ Comparison of approximation for soot with measured data

FIGURE 3.5 Variation of species concentration a'' with flame temperature in a coal/air flame with 5$ excess air

FIGURE 3*6 Average gas concentration at any planeFIGURE 3.7 Generalized streamlines in an

axisymmetric pntraining jet system with recirculation

FIGURE 3.8 Simplified model of entrainedconfined jet

SECTION 4.FIGURE 4.1 Line diagram of manufacture of cement

by the wet process FIGURE 4.2 Simplified diagram of mass flow meter

and electronic system FIGURE 4.3 Flow diagram of gravimetric feederFIGURE 4.4 Gas sampling probe for stationary

measurements FIGURE 4.5 Gas sampling probe for rotating

measurements

8

22

37

38

40

42

5056

57

57

62

6870

83

83

SECTION

FIGURE 4,6

FIGURE 4.7 FIGURE 4.8

FIGURE If . 9

General arrangement of integral cooling system. _ 84Integral probe cooling unit 85Suction pyrometer for flame temperature measurements 92Land venturi suction pyrometer for flame temperature measurements • 96

5.FIGURE 5.1

FIGURE 5,2 FIGURE 5-3 FIGURE 5.4 FIGURE 5.5

FIGURE 5.6 FIGURE 5.7 FIGURE 5 .8

FIGURE 5*9

FIGURE 5.10

FIGURE 5.11

FIGURE 5.12

FIGURE 5.13

FIGURE 5.14

FIGURE 5.15

FIGURE 5.16

FIGURE 5.17

Distribution of primary;secondary air ratio 103Distribution of primary air percentage 103 Distribution of excess air percentage 103 Destribution of Craya Curtet parameter 104 Distribution of specific fuel comsumption 104Distribution of gas exit temperature 104 Flowsheet for heat and mass balances 112 Sankey diagram of average heat flow over trial periods 115Average external temperature profile of kiln shell over trial periods 117Variation of kiln shell emissivity with temperature 119Variation of external heat loss with shell temperature 121Variation of external heat loss with distance from burner nozzle 122Flame gas concentration profiles for 17.24$ excess air 125Flame gas concentration profiles for 8.62$ excess air 126Flame gas concentration profiles for 2,87$ excess air 127Fuel burnout as a function of distance from the burner nozzle 130Derived isotherms in the burning zone of the Barnstone kiln 133

FIGURE 5.18 Flowsheet for simulation of heat. transfer in a rotary kiln 136

FIGURE 5.19 Dust burden and gas velocity variationin a rotary cement kiln 140

FIGURE 5.20 Variation of dust concentration withdistance from the burner nozzle l4l

FIGURE 5.21 Predicted and measured gas temperatureprofiles for 100 t.p.d. cement kiln 151

FIGURE 5.22 Predicted and measured wall temperatureprofiles for 100 t.p.d. cement kiln 152

FUGURE 5.23 Predicted wall heat flux for 100 t.p.d.cement kiln * 153

LIST OF PLATES

SECTION 4.'PLATE 4.1 Monitoring of cooler* air flow 65PLATE 4.2 Measurement of kiln shell emissivity 75PLATE 4.3 Stationary gas sampling 79PLATE 4.4 Sample porthole 81PLATE 4.5 Integral cooling unit 86PLATE 4.6 Gas sample analysis 90PLATE 4.7 Suction pyrometer 93PLATE 4.8 Pyrometer shields 94PLATE 4.9 Venturi suction pyrometer 97

SECTION !■;

INTRODUCTION. .

The successful application of rotary kilns to the cement making process in the late 1890’s brought about a major revolution to the industry, and by 1906 they were in almost ,universal use as the primary production unit of cement manufacture. This change to a continuous process, from the batteries of batch kilns, first developed by Aspdin in 1824, presented the possibility of substantially larger production rates at lower costs, thus establishing cement as a major constructional material.

The manufacture of cement from limestone or chalk and clay or shale has changed little in the eighty years since the introduction of rotary kilns. The original wet process of manufacture involves the slurrying of a chalk and clay mixture' to yield a kiln feed with a moisture content of approximately forty percent. This wet slurry is fed into one end of the kiln where it is countercurrently contacted with a stream of hot combustion gases, produced by a single burner at the opposite end of the kiln. The slurry is thus dried and heated until calcination and clinkering have occured at a temperature of approximately l450°C. This basic, wet process is still widely used and accounts for over half of the present world production.

The introduction of the Lepol grate, invented in Lithuania in 1928, had the effect of reducing the length of kiln required for equivalent production; and from the improved heat transfer on the grate and reduced moisture of the feed ( 10-15# )3 a reduction in the specific fuel consumption allowed for still lower production costs. ! .This was taken a step further in Germany by the introduction of dry suspension preheater process equipment in the early 1950’s. The process made use of a ’fluidized bed’ cyclone system to carry out the calcination reaction, and a short rotary kiln to complete the clinkering. This process has approximately half the energy consumtion of the traditional wet process, and provides for very high production rates from relatively small units.

Throughout these changes, the rotary kiln has remained the basic unit for the high temperature clinkering process.The capacity of kilns has increased from units producing 100 tonnes/day to present day units producing 5000 tonnes/day, and the world annual production is now approaching 500 million tonnes. For such a major industry to remain a relative research backwater appears inexplicable, yet, until the last decade, there have been only three notable practical investigations 1,2,3 into rotary kiln systems.

In this work, current combustion expertise on measuring techniques and devices has been applied to the difficult problem of continuous monitoring of a rotating furnace in the flame region. Particular emphasis was placed on this

area of the kiln because it is common to all types of process. The results'of this work have provided the basic data to formulate heat, and mass balances, and to investigate the combustion of the fuel under the influence of the critical operating parameters.

The results also provide a practical datum against which to test theoretical postulates regarding heat transfer within the clinkering region. The advent of fast digital computers has allowed the solution of finite element problems to become realistic, and an analysis of the rotary kiln has been carried out using just such a zone analysis method to predict temperature profiles, and heat fluxes in the sintering region. A favourable comparison is exhibited between the measured and predicted values, thus providing a major step towards the analytical assessment of rotary kiln systems.

SECTION 2.

LITERATURE SURVEY OF PREVIOUS WORK.

2.1. Fundamental studies of enclosed turbulent diffusion flames.

The usual system employed in rotary cement kilns to provide the energy for chemical conversion of the products, is that of a partially premixed jet of fuel and air injected axially into the cylindrical kiln tube, where it is allowed to entrain a limited amount of slow moving secondary air. Fundamental aspects of this system have been studied by various workers concerned with both the aerodynamics, and combustion of enclosed turbulent diffusion flames.

2.1.1. Jet entrainment.

A turbulent jet diffusion flame commences by entraining the same mass of surrounding fluid as does a free jet. This surrounding fluid is however composed of a mixture of air and recirculated combustion products. In the case of an enclosed flame, where the walls of the furnace provide a physical limitation on the j.et, the rate at which the air/fuel ratio increases starts as a free jet, but decreases with increased distance from the burnerk.

Several accounts are available 5»6 of entrainment by free turbulent jets issuing into stagnant surroundings. It may be shown by dimensional analysis that the mass entrained, m, at any point, x, along the jet path is given by the

relationship,

i im = K x tfl2 p2 *a (2.1)

where K is a constant, p„ is the density of theCL

surrounding fluid, and M is the excess jet momentum flux, which is given by

where d is the nozzle diameter and p is the jet fluid o odensity at the nozzle.

The numerical value of K'has been determined by a number of workers 5,7,8 to lie within the range 0.22 to 0.404.This range of values has arisen due to the method of determination of the value,m, where it is given by

m =. f* 2tt pu y dy .................... (2.3)

since the determination of pu at large values of ybecomes difficult practically,or depends on the form of anyanalytical velocity profile that may be chosen.

9Hicou and Spalding carried out practical measurements on isothermal jet systems with varying densities of jet and entrained fluids, and from this work they arrived at a value of K = 0.282. From this a general relationship was developed of the form

M = M = 0.25 it d2 p u2O 0 ^ 0 0 (2.2)

mrno

(2.4)

In the presence of buoyancy and chemical reaction they have advanced a modified relationship of equation (2.4) including the Froude number.

within the vicinity of the nozzle is much lower, and this is attributable to the potential core of the jet and the region of jet stabilization.

2.1.2. Recirculation.

Recirculation is the reverse flow of fluid from a point in the downstream section of a jet, which is reentrained by the jet nearer the injection source. This phenomenon occurs naturally under certain conditions in confined jets11, and may be induced by the insertion of bluff bodies or swirlers in a jet stream10 . Confinement of the jet produces external recirculation, where fluid is disentrained from the periphery of the jet, creating a toroid of reverse flow around the jet. Bluff bodies and swirlers produce internal recirculation, where fluid travels in a reverse direction along the axis of the jet. The advantage in in combustion systems is that recirculation provides a

(2.5)

where

12Fr (2.6)

These entrainment relationships have been found tohold true for values of x>6d . The rate of entrainmento

source of heat near to the burner from the returning gaseous products of combustion, which stabilize - the flame on the burner at a higher turbulent intensity than is possible with a simple jet.

Thring and Newby11 developed a theoretical treatmentof the problem, based on the entrainment laws for a freejet. They introduced a parameter,0 *, which may be used todetermine the point in the downstream jet at whichrecirculation entrainment occurs,x , and the core of the3 n 3recirculation eddy,x .O

d* fP " 2 m +m0 Ko a o2L lPaJ m o

(2.7)

if a jet half angle3cij of 9.7° is assumed1 8 , then

x = 6.25 6' L ................ . (2.8)

x = 3.12 ( 6' + 0.94 ) L .......................(2.9)0

The theory is only valid where d^/2L<0.05. Experimental data of Barchiilon and Curtet12 has shown that equations (2.8) and (2.9) give a reasonable fit. See figure 2.1.The point where the jet strikes the confining wall,x„, is given by

x = — ................ (2.10)p tana

>5 cd p •H bOo oO P H O© JG> pH ©

- CO

P

rH GO *H

HkjQ^D Oo(0COT-

X

©i— !ts!NOcoShCm©OC aJ p co •H Q

, Q aoqeureaad ^qMO^ Suiaq,!

FIGURE

2.1

Position

of re

circ

ulat

ion

eddy

in a

confined

jet

For values of do/2L>0»055 Craya and Curtet13 have developed a theory to predict recirculation characteristics. They proposed a similarity parameter,m, of the form

(2.11)

where

R (2.12)

they also found that a relationship

i6m* = f ( 0,d£/2L ) (2.13)

exists between the Thring-Newby approach and their own.If the jet is considered as a point source, d^/2L=0, then equation (2.13) becomes

The value of m ranges from zero to infinity, with recirculation occuring at values greater than.1.5 in axisymmetric jets.

•Becker et al.1*1 have developed a rigorous treatment of enclosed jets considered as a point source, and have demonstrated that the aerodynamic flow pattern is significantly different from that of a free jet. Using the concepts of dynamic,u^, and kinematic,u^', velocities,

i6 = m 2 (2.14)

they developed an expression to define a characteristicstream velocity,uo

(2.15)

where

G + IG o 2 a, upL^ , (2.16)

and

muk rT + UiTpL a (2.17)

They have shown that -this is related to the Craya-Curte parameter3m

m = \ 2-4 - 1k

(2.18)

and they have presented a similarity parameter termed the Curtet number,Ct, where

i u,ct = i = -4

u.(2.19)

Experimental work has shown that recirculation is limited to Ct<0.75.

2.1.3. Combustion of pulverized fuel clouds.

It has been shown by Nusselt15 and Essenhigh16 that

for a monosize suspension of coal particles burning in plug flow, the burning time is given by

(2.20)

where

td ' 964®pg (2.21)

and

(2.22)

The.multiplying factors,F and F^, are dependant on the excess air and volatiles in the fuel. Their values have been determined by Nusselt 15 and Essenhigh16.

Practical application of this work is limited by the assumption of a monosize cloud which considerably underestimate the length of the tail of a polysize cloud of the same mean particle diameter. Calculations on polysize suspentions of coal particles were made by Hottel and Stewart17 , who •• considered the suspension to be composed of a number ofmonosize fractions. This approach has been extended by

18 .Field et al. to take account of the the diffusional andsurface reaction rate coefficients in an isothermal flame.Two alternative forms of equation, depending on the combustionmode, were developed to calculate the change in weight fractionwith time.

For constant size combustion

(2.23)

(2.24)

Thus if there are n fractions, equations (2.23) or (2.24)provide a series of n simultaneous equations, varying in t,u. and p (U), which may be solved by finite difference J S

methods. •

A comparison of polysize and monosize suspensions by Field et al.J8 has shown by calculation that for a typical pulverized fuel with a weight mean size of 50 microns, the time taken to burn 98$ of the residual char is more than ten times that for a monosize suspension of 50 microns. It is thus important to consider particle size distribution when calculating the course of combustion, especially in the case of substantially complete combustion ( i.e. less than 20$ unburnt char. ).

McKenzie, Smith and Donau Szpindler19 used a mathematical model, similar to that proposed by Field, to predict burnout of high and low reactivity pulverized fuels. Their results were compared with experimental'measurements on a semi- anthracite made by the I.F.R.F.20 , and adequate agreement was observed. They concluded that suspensions containing

du. S. p (U) 1 _ JQ gdt " 1_ 1_k , k d s

and for constant density combustion

dui . (V wj),S7sjoPR(u)dt ' " (u,/w",)-r3 + X

kd ki

appreciable quantities of particles above 100 microns, at near stoichiometric conditions, are unlikely to achieve complete combustion.

Hedley and Jackson21 have considered the influence of recirculation of combustion products on the rate of burnout. They proposed that there is an optimum value of the 'recirculation ratio, dependant on the excess air, at which to operate to achieve burnout in the shortest time. Whilst the work was concerned with monosize suspensions, the implications are nevertheless applicable to polysize suspensions,

2.2. Evaluation of kiln operating data.

The need to understand and predict the heat transfer mechanisms occurring within the confines of rotary kilns has long been recognised, but the volume of published work on this aspect of kiln operation, which may be generally applicable, is relatively sparge.

2.2.1. Critical analysis of production capacity.

It has been recognised that the size of a particular rotary kiln, and the input variables, may be directly related to the production capacity of the unit, and with this concept in mind a plethora of capacity formula are available in the literature. In general these are only valid within narrow limits, due to the analysis of an inadequate range of data taken on units of approximately the same physical dimensions.Martin22 produced a formula

which stated that the production capacity of a kiln is directly proportional to the cross sectional area of the unit, arriving at this conclusion from the assumption that there is an optimum terminal velocity for the exit kiln gases, and that this is the major factor influencing kiln performance. Most capacity formulae have been based on the internal volume of the kiln, and the the most commonly used, and widely applicable of these are Gibbs23, Eckel21* and Kannerwurf and Clausen25. Gibbs and Eckel both proposed a relationship of the form

where ki is a constant dependant on the type of process, the value of which is derived from operational experience. Kannerwurf and Clausen proposed a relationship of the form

where k 0 and ki are constants determined from an analysis of almost 600 kilns in the U.S.A. cement production industry.

P = kiV (2.25)

P = k 0 + kiV (2.26)

Yoshii26 proposed a relationship of the form

P = k 0Ve“klV (2.27)

where k 0 and ki are constants dependant on the type of process. This form of relationship indicates that there is an optimum value for the volume, over which any increase

will result in a diminishing production rate. Whilst this is not wholly true, it is a significant fact that larger unit in the cement industry do not produce proportionally as much as their smaller counterparts.

2.2.2. Significant variables.

Such approaches as are outlined in Section 2.2,1. have been attempts to simplify and generalise what is a very complex system, and their conclusions have been that the physical dimensions of the system are the predominant variables. This basic premise is not only true with regard to the thermodynamics of the system, but as Groume-Grima ilo27 pointed out; in no type of furnace does the concept of hydraulics have more significance than in a rotary kiln. Physical modelling by Moles, Watson and Lain28 has since' shown the importance of this statement, and of the geometric variables of the system.

Burke and Field29 summarized the most important variables from the industrial view as a need to know

(a) The maximum rate at which fuel may be economically burnt for a given kiln diameter.

(b) The fineness of fuel grinding or atomization to give the best results.

(c) The quantity and velocity of primary or burner air.

(d) The best' type of burner to obtain maximum radiative heat transfer in the kiln.

They note that experience has shown that in smaller unit's control is easier, and burning conditions are more satisfactory. Work at IJmuiden by the I.F.R.F.20 has shown that the degree of. grinding or atomization of the fuel is not significant, and this conclusion is only to be expected, since the system is generally one of a turbulent diffusion flame burning-'in confined surroundings, where mixing time is predominantly more important than the burning time of a particle30 .

As Burke and Field have pointed out, the sizing of a rotary kiln for a given capacity cannot reasonably be based on empirical information concerning volume or area factors, or retention time and loading factors, with any degree of certainty. Unless the heat transfer aspects are understood, the influence of minor variations in fuel, ambient conditions, etc., on the overall process will not be predictable.

Rigorous evaluation of heat transfer mechanisms within the rotary kiln require a knowledge of the various heat transfer coefficients, radiative, convective and conductive, through the various stages of processing.

2.3» Practical measurements in rotary kiln systems.

A number of attempts at practical measurement of heat and mass flows in the rotary kiln have been made.The most difficult problem of physical measurement, and the reason why so little work has been carried out, is

to overcome the rotating function of the kiln, which prevents the simple insertion and withdrawal of measuring probes using the techniques that have been employed . •' successfully in most static industrial furnaces.

Some preliminary work on the measurement of flame temperature was carried out by Martin22 using platinum/ platinum-rhodium thermocouples, the results of which indicated a temperature at some unrecorded point in the flame region of 1400°C.

Gilbert31 carried out an investigation * into the operation of a M.I.A.G. calcinator, to study the effect on kiln performance. During the course of;the investigation he measured material composition, and gas and material temperatures at various points through the kiln system.By analytically dividing the kiln and calcinator into thirteen stages, he constructed predicted profiles for gas and material temperatures, and attempted.'to assign values to the various heat transfer coefficients in the system.

Gygi32 measured average gas and material temperature longitudinally along the kiln and assumed a circumferential temperature distribution for the rotating kiln wall. By assuming complete heat penetration of the charge he was able to arrive at some measure of the thermal efficiency of the system. The assumption of both wall and charge temperatures implies an assumption of the value of heat

transfer coefficients. The true value of Gygi's work lies in the measurement of some of the basic data values required for a complete evaluation of the system.

Yoshii33 measured the external shell temperature of both a kiln and cooler at twelve stations along their lengths in order to determine the radiative and convective heat losses from the unit, and to predict the internal wall temperature.

The first comprehensive series of experiments on a rotary kiln were carried out by Folliot1 in 1955 •Using a production unit manufacturing ordinary Portland cement by the wet process, he studied the radiation ; characteristics of the gas bulk and internal walls of the kiln at eleven stations along its length, using a total radiation pyrometer. He discovered that the emissive power of the gases was higher than that predicted from a knowledge of their composition, and radiative abilities. This increase he contributed to the dust entrained by the gases as they flow over the charge bed. He also carried out measuremen to determine the emissivity of' the.charge at various temperatures through its process cycle. From this data, and a measure of the bulk gas temperatures, he was able to calculate the net energy transferred by radiation from the gas bulk to the walls and charge. By measuring the temperature gradient and thermal cycle of specially constructed firebricks, containing embedded thermocouples, in the kiln, he was able to assess the conductive heat transfer to the charge and through the walls.

concluded that(a) The radiation component of heat transfer

from the gas is much greater than that expected from a knowledge of the gas components. This discrepancy was attributed to .the*dust icohtent of the gases as they flow over the charge.The emissivity is dependant on the kiln diameter and any increase in the diameter over 2,4 meters will not significantly improve the radiative heat transfer coefficient, and particularly if the gases have a high dust concentration.

(b) In the preheating, calcining and sintering zones approximately' 84$ of the heat transferred to^the charge is transmitted by radiation, reradiation and convection ( 90$ of this is radiation and 10$ convection ). The remaining' heat is transmitted by conduction between the refractory walls and the charge, and is dependant on the thermal conductivity of the charge.

(c) Due to its high bulk porosity, the charge material is a good heat insulator, and thus heat penetration is small from the charge surface to its interior. The amount of heat absorbed at the charge surface is reduced with an increase in material bulk- temperature, and a reduction in the grain sizeof the charge. ( For grain sizes greater than 1mm. the reflectivity of the surface o.f the bed is constant, depending on the chemical state of the charge. ).

(d) Since heat penetration is so poor, heat transfer to the bulk of the charge must be carried outby a kind of ’solid convection’ involving a rapid renewal of the exposed surface. Heat transfer between the refractory wall and the charge is proportional to the square root of the speed of rotation of the kiln.

(e) The throughput of a rotary kiln of given dimensions can be increased only by raising the gas temperature. This results in a higher exit gas temperature and heat losses.

* Folliot’s results are not in agreement with Gygi.Heat balances on the product using Gygi’s data indicate that the charge receives apprpximately three times the heat manifested by measured values, and Folliot attributes this to an erronious assumption by Gygi of the gas emissivity which takes no account of the dust burden.

Weber2 carried out the most complete study of the cement making process yet undertaken in 1963* He measured the changes in the major variables using industrial plant of the three process systems most commonly used, i.e. a long wet process rotary kiln,a- Lepol kiln, and a suspension preheater dry process kiln. During his investigations Weber measured the chemical state, and physical condition of the charge material at various stations through each kiln system. He also measured overall mass flows of charge, air and dust in the system; carried out chemical analyses of the fuels used and the exhaust gases, and measured the external

wall temperatures of the equipment to estimate the atmospheric heat losses. The principle objective of Weber's work was to understand more fully the sulphur and. alkali cycles which occur within the kiln, since the vapourization of these compounds affect the overall heat consumption of the system. From his work he concluded that

(a) All the calcium oxide formed in the calcining zone reacts immediately with the alumina, silica and iron oxide, and that all the exothermic reactions take place inthe calcining region.

(b) The cycling of dust, sulphur and alkalis within the kiln, and through the system, cause high grade heat to be taken up and regenerated as low grade heat at the back end of the system. These cycles also make the operation of the kiln more difficult.

(c) In each of the kiln systems studied, 90$ of heat from the hot gases is transferred by direct or indirect radiation, and formulae may be

. developed , based on the internal dimensions of the kiln, to predict the optimum output of any unit.

(d) The criterion for heat transfer in a dry process kiln is the temperature at the boundary between the preheating and calcining zones.

(e) The criterion for heat transfer in a wet process kiln is the exit gas temperature in conjunction with the excess air.

Foll

iot

a c: r- r-a.ina'sjedme.i,0

Dist

ance

’ along

kiln

from

nozzle

( % kiln

length

)FIGURE

2.2

Comparison

of measured

temp

erat

ure

data

for

gas

and

solids

in a

cement

kiln

3In 1965 Ruhland carried out an investigation on the flame length in a wet process rotary cement kiln, by extracting and: analysing gas samples from the flame zone of the kiln. This work was an extension of his isothermal water modelling work using acid/alkali techniques to determine flame lengths. By continuous analysis of the carbon monoxide, carbon dioxide and oxygen concentrations in the gases at fourteen points along the kiln, Ruhland was able to establish a formula to predict the length of the flame in terms'of its dependant parameters. The formula derived was

L = d /K o 3.21 f 2%+B.3 o + 3'.862 n -1 o* ^ 2 (a+b)le v ..(2.28)

where

K =m +m o a

m m o ■ apo pa

(2.29)

a = 2.12 -tm1*245

2L-d0 I °.2L-d

b = 0.1052 o

(2.30)

(2.31)

Thus the flame length,L, is a function of the mass flows of primary,* , and secondary,* , fluids, the nozzle,d .

O cl ‘ O

and kiln,2L, diameters, the densities of'the primary,pQ , secondary, pa , and flame,p^,, fluids, the secondary fluid requirement of the primary jet,B , and the excess air in the secondary fluid,n. Ruhland concluded that the excess airwas the most significant factor influencing flame length.

2.4. Theoretical studies of rotary kiln systems.

Pike 3h divided a kiln system.into*three analytical zones, clinkering, calcining ana preheating, and developed equations to predict the thermal efficiency of each, and the optimum length of each zone for maximum thermal efficiency. Of the three zones, it was considered that the calcining zone was thermally most important.

3 5Lyons et al. used the same divisions as Pike, and with thermodynamic data available in the literature, set up an analog model to predict solids composition and gas and solids temperature variation along the kiln length.The simulation gave a reasonable agreement with available measured data. The values for heat transfer coefficients were arrived at by a trial and error method using approxima data from manufacturing experience. The predicted exit gas temperature was 280°C, as compared to a data value of ~260°C. The maximum predicted gas temperature was 1790°C and solids temperature was 1570°C.

3 6Imber and Paschkis developed a mathematicalmodel to predict kiln operation. They constructed heat balances over elemental sections of the kiln to account for the heat transferred from the gas to the kiln wall, the gas to the charge, the wall to the charge, the reradiated heat\from the wall to the charge, and the external heat losses 'from the element. Prom a"consideration of two limits of heat transfer through the charge, the upper and lower limits of kiln length were arrived at. The lower limit

is given by assuming a ’well mixed’ charge bed in which there is no temperature gradient. The upper limit is given by assuming a ’non mixed’ charge acting as a slug of material travelling along the kiln, within which there are temperature gradients depending on the thermal conductivity and chemical reactions occurring. It was found that the well mixed assumption agreed reasonably with published data for production kilns. The problem of calculating the reradiation from the walls to the charge was discussed, but the calculation procedure made disregard of the effect.

Vaillant37 extended the work of Paschkis to include all the process variables. By the use of a passive network analog computer he developed a model to include the effect of

(a) The regenerative heat action of the rotating wall in the cyclic equilibrium for numerous combinations of temperature, heat transfer coefficient rotational velocity, and volume loading.

(b) Radiation loss to the surroundings as a function of gas and charge temperatures.

(c) Variable flowrate due to mass transfer between the charge and gas. .

(e) Space and temperature dependant heat transfer coefficients.

(f) A partially mixed kiln charge.

The model examined the effect of rotational speed on • the system and heat losses from the kiln shell, but was

not tested against an existing unit, due to a lack of basic data concerning the values of the various heat transfer coefficients.

Bowers and Read38 developed equations to predict the effect of kiln diameter and length on kiln/heat transfer rates for four rotary kiln processes; limestone calcination, cement burning, dolomite calcination, and shale expansion.An analogy of kiln heat transfer rates with those occuring in a fixed slab was made. The authors demonstrated that the bed preheat and particle size of the material have a significant effect on the heat transfer rate. It was concluded that for cement and limestone production the overall heat transfer coefficient is approximately numerically equal to the inside diameter of the kiln ( f.p.s. units ).

Siedel39 has proposed that both the primary air percentage and the rate of dust insufflation in a rotary cement kiln influence flame formation. From available data, he set up relationships for a wet process kiln between flame temperature, and kiln throughput, and has suggested that increases of 50°C in the flame temperature, without the addition of extra fuel, results in a 3 to 8% increase in production capacity.

Schwartzkopfk0 has proposed a computer program flowsheet which may be used to analyse heat transfer in rotary lime kilns, and to determine kiln size for a given production rate. The analysis takes into consideration the process heat requirements, geometric factors, shell heat losses, radiation

from the gases, and conduction from the bricks. Results are presented to show the effect of changing various parameters on the length of the kiln required for lime processing, and the heat required per unit mass of product. Investigations were made into the effectiveness of both soaking pits and pre-calciners.

Rosa41 has developed a computer program to carry out the following calculations from mathematical model developed to describe the heat transfer within various zones of a rotary kiln.

(a) The course of the gas, and raw materials temperature curves in the kiln system. *

(b) The total amount of heat liberated in any particular section of the kiln.

(c) The shortest required residence time of the raw' material in the individual zones.

(d) The highest possible kiln output attainable from a thermodynamic standpoint.

(e) The qualitative and quantitative effect of eachof the variables in the model on the values of the variables it is desired to determine.

No attempt has been made in the paper to present results which may be compared with measured data, or to predict trends.

Maritius et al.42 have recently developed a steady state model of a rotary kiln used for the calcination of basic ammonium aluminium sulphate to aluminium oxide.In this system dust entrainment is significant ( 20-50$

of the feed rate )5 and chokes are used to hold up the material in the kiln. ’The model developed is presented as a series of ordinary differential equations describing the reaction kinetics and mass and heat transfer balances; a set of algebraic equations for space dependant parameters; and a set of two point boundary conditions. An alogrithmic program was developed for the solution of the equations with and without chokes in the system. Results are compared with measured values and good agreement ( ±10% ) is displayed.

SECTION 3.

THEORETICAL PREDICTION OF HEAT TRANSFER;THE DEVELOPMENT OF A MODEL

3»1. General summary of approaches. •

Within the last decade the well established empirical furnace design methods, for example that of Lobo and Evans^3 have been superseded by, what may be termed, ’finite element’ .methods. For years furnace designers have known that such methods would ultimately yield the required answer, but only the advent of large and fast digital computers has made the methods feasible.

A rigorous calculation of heat transfer from a flame or flames to a furnace enclosure requires the simultaneous solution of partial differential equations of fluid flow wîth chemical reaction,.and of energy transfer. Since these are interdependant the problem becomes highly complex and has not yet been resolved, for practical cases. The general approach has therefore been to develop models to simulate the energy transfer problem and superimpose information about flow patterns and heat release onto these. The patterns of heat release and flow are usually obtained from isothermal physical models or plant trials. This approach has been adopted because such models are relatively insensitive to errors in flow but highly sensitive to errors in energy transfer.

There are four major methods of approach to the problem,

but all consider the interaction between various points or zonal elements in the furnace, assuming that there is no change of property of the gases or surface enclosed within a zone or surrounding a point.

3-. 1.1. Zone methods

If ifThis method, due originally to Hottel & Sarofim, depends upon dividing the furnace into a number of zones, either volumes or areas, such that it can be considered that all the properties of the constituents of that zone are constant throughout the zone. The interaction between each zone is then studied using a modified ’view factor’ technique, and a series of simultaneous equations in terms of zone temperatures is set up. These are then solved by iterative procedures. This method is discussed, and' developed in more detail later in this chapter.

3.1.2. Flux method

The method considers a small elemental volume, or surface in the furnace enclosure, and by calculation of the net heat flux to/from that element in one or more directions, generates, a total energy balance on that element. This procedure may be repeated for as many elements as required until the desired degree of accuracy is obtained. Its advantage over other methods is that it is much faster in computational time, but it is not so accurate. Reference can made to Sidall for more detailed information!?5

3»1.3« Monte Carlo methods

This method is much like the zone method, 3*1*1» except that in place of zones the Monte Carlo method considers bundles of energy to simulate the actual physical processes of radiant emission and absorption of energy occurring within the enclosure. These energy bundles are similar to photons in their behaviour, but the energy per bundle is constant and does not vary with the temperature, its spectral region, or its point of emission. The energy per bundle is some fraction of the total or net radiant energy emitted throughout the system per unit time. The history of an energy bundle from its emission until it is finally absorbed is determined by a series of random numbers which are generated every time a decision with respect to position, direction, spectral region, path length, reflection and absorption is required.Reference is made to Steward and Cannon for more detailed information.4 6

3.1.^. Fluid Mixing methods

This approach is fundamentally different to the first three methods as it is based on a mathematical treatment of the flow systems predominant in the furnace enclosure, taking into account entrainment and recirculation, derived from the basic Navier Stokes equations for momentum conservation.The furnace is again divided into a series of zones and ’finite difference equations’ between adjacent zones are set up and solved by iterative procedures. Energy release,

convective heat transfer, and enthalpy changes in the gases are 'taken into account, but the major inaccuracy is that no consideration is taken of radiative heat transfer. Reference c be made to Spalding et. al. for more detailed informations7

3.2. Summary of prediction of heat transfer in cylindricalfurnaces by zonal analysis.

The zone method of Hottel and Sarofim has been developed in this research programme to consider the problems of heat transfer in the cylindrical rotary cement kiln.

The method may be used to predict temperature profiles in the gas stream, and temperature and heat flux distribution along the furnace wall, if the gas composition and flow pattern, within the furnace are first specified or known. The method can be summarized by reference to a-series of steps as follows.

3.2.1. Gas data

Prom a knowledge of the gas concentrations of CO2 and H 20 vapour in the furnace, and the fuel combustion products, emissivity/absorptivity relationships to a real gas system may be calculated by use of the weighted sum of a series of grey

Suitable values for kn are calculated and the variation ofa and a with T are determined, gn sn g

3.2.2. System zoning

The system is zoned into the coarsest structure consistent with the accuracy desired, in a manner dictated by the configuration of the furnace and by prior estimation of regions where steep gradients in the temperature field necessitate fine-scale zoning.

3.2.3. Evaluate direct-exchange areas

We must first evaluate how much area each gas or surface zone can 'see' of each other gas or surface zone. These -areas are derived in the form of integro-differential equations, and the actual effort of calculation is considerably reduced by

4 8reference to pre-evaluated tables such as those of Erkku or Einstein.49 These exchange areas must be evaluated for each of the values of kn in the gas emissivity/absorptivity fit.

3.2.*]. Evaluate total exchange areas

We must now evaluate the total radiation incident on/ reflected from each zone due to direct exchange and reflected exchange. The leaving flux density, W, at any surface zone is equal to the emitted and reflected fluxes. A radiation balance on surface zone i gives:

A .W. = A.(e.E • + R .) = A .e •E_• + p.(Is.s.W. + Eg.s.E .) 1 1 l l si 1/ l l si Hi v . j l j ;6J i gjJ J

orA. A.e.E .

Z( s. s. - 6 • . — ) W. -■ ■■ --- - Z g.s.E, . (3-3)• J i iJPjL J j 1 SJ

Thus we build up a series of equations like (3.3*) with differing values of W for each surface/surface and surface/gas zone pair. Prom these W values we may evaluate total exchange areas between surface/surface, gas/surface, and gas/gas zone pairs.

A . e .(G.S.) = -J— i .W...................................... (3.if)i j n p . gi j

A . e .(S.S.) = —J— ( .W. - 6..£.nv i j ' n p. si j ij i) ..... (3*5)

(W n = gigj + f k gj giWk ....... (3‘6)

We again have to evaluate (3*^)a (3*5) and (3*6) for each of the values of kn in the gas fit.

3.2.5* Evaluation of net radiative exchange

An estimate is made of the temperature field in the system, and, the direct flux areas for each zone pair is calculated; thus for example we have:

S.S. = Z 1 J n

a (T. ) snv l <Si V n .................... (3-7)

3»2.6. Formulate total energy balances

Total energy balances can now be written for all gas and surface zones taking into account convection to surface zones, enthalpy changes in gas zones (Qej_)3 energy release within gas and surface zones (Qc^)3 and transient terms or storage of energy in zones (Qugj_) •

Thus we have for surface zone

Thus we obtain a series of nonlinear simultaneous equations for as many unknown E ’s (or aT^’s) as there are equations. Terms in T which are not large may be rearranged in a form that will make equations linear in T \

Tsi)..(3.9)

And for volume zone V.1

(3.10)

These equations can be solved by iterative procedures using a combination of Gauss-Seid’el and Newton-Raphson methods to converge to the correct solution.

Temperatures obtained from step 3.2.6. are inserted into the surface heat balance equations to evaluate net heat

3.2 .1 . - 3.2.4. give a complete description of the radiation characteristics of the system, (i.e. Fixed furnace shape or size, wall emissivity, and gas compositions.) Changes in steps 3.2.5 - 3 *2 .7 . will show the effects on furnace performance of changes in the gas flowrate, flow pattern, enthalpies of gases, heat generation, and temperature distribution of the controlled parts of the furnace walls.

3.3. Representation of a real gas by a series of grey gâsesto simulate the emissivity/absorptivity characteristics.

In furnace calculations it is most convenient to represent the radiation properties of a real gas by the combination of a weighted sum of a series of grey gases

flux to that surface. It should be pointed out that steps

(3.2)

(3.1)

where

spectral windows

represents the fractionof blackbody energy inthe wavenumber regionassociated with k. about1Ak. ( shaded areas.)

0)

Figure 3.1* Representation of gas absorption coefficients with bands in the same spectral range.

It is usual to have a clear gas component which has a value of k^=0 to account for the spectral windows in the system.A one clear - two grey gas approximation to a real gas has been found to be sufficient for most engineering calculations.

Using data available for the variation of e T with & g gpL for particular gases involved in the system being studied50 (i.e. C02, H 20, and CO), values of k^ and a^ for the system

over the temperature range likely to be encountered in the furnace can be evaluated. It is usual to adopt fixed values of kâ since these vary little over a wide range, of temperature, and to allow the weighting factors a^ to vary so as to simulate the variation of e and « with the gas temperature. Figure 3.2. shows the variation of a^ with temperature calculated for a system where ^2^ P e o 2 = *(see appendix A ) Rationalizing this data, a gas fit was

derived (Table 2.1).

2 5 0 0 01 0 0 0 15 0 0 2 0 0 0 T. c

. 9

e

. 7

6

- 5

- «»

3

2

1

0. 0 0 0 *♦ .0 0 1. 0 0 0 7 . 0 0 0 9. 0 0 0 5 . 0 0 0 6 . 0 0 0 8

FIGURE 3.2 Variation of weighting factor with temperature for

pH20/pC02 =1

Table 3.1 Absorption coeff. and weighting factor values

i k . m 1 atms \l .a.1

0 0 1.0029^ - 735.29T"1&1 .7858 0.01125 + 562.50T-1

2 13.65 -O.OOH877 + 182.93T"1§

Figure 3.3. shows a comparison between values calculated using the data in Table 3.1 and actual results for a range of temperatures.

3. 1. Representation of soot particles by a series of grey gases to simulate the emissivity characteristics ,

5 1Johnson and Beer have shown that a grey gas approximation to the characteristics of soot can be made in the same way as for a real gas3 providing that

asn(T) = agn(T) (3a2)The emissivity of soot may be calculated using the expression

e = L (T) (l-e"ksncsL ) ........................... (3.13)s n gn

Thus combining equations (3*1) and (3.13) we have

em = I*gn(T) d-e-knL )

where

(3.14)

o to r—I

m

«—ico

HP=i

We have now incorporated the soot factor into the general equation describing the radiation characteristics of the medium. By using the values of a^ calculated in Table 3*1 and some suitable experimental data on soot emissivity, values of kgn have been determined to simulate e, in the form of equation-3•13• Soot emitsihover the entire range of wavenumbers, w, and thus no clear component exists. ' Table 3-2 gives the best fit values thus obtained.

Table 3-2 Absorption .coefficient for soot •

i ksi m 2. kg 1.

0 10001 6002 500

Figure 3*4 shows the comparison between the experimental data of Howarth5and the approximation using the data in Table 3.2.

3.3. Radiative properties of dust particles .and theirincorporation into the grey, gas approximation principle

If we have a mass concentration, c^, of spherical dustparticles of diameter d and density p , and each particleP Pabsorbs a fraction A of the radiation falling on it (so

Conc

entr

atio

n -

path

ieng

th1500Temperature

o

1000sO600500s 2

Data of Howarth

Grey gas fit to data

* 0 0 0 0 1

c 1. 2 6 .8Emissivity e

FIGURE 3.^ Comparison of approximation for soot wi'th measured data

that for ’large* black particles d >>!_, A=l), then thep m -fraction dl of radiation I incident on a gas layer of thickness dl which is absorbed is given by 5 3

dl -And2 ,T 1 , v nr.\— ~ “Tj— P Npdl..... .................................. ....(3.15)

where

N =------- 2... ...(3.16)6(ppffdp

Thus for a finite thickness, L, we may integrate equation (3 .15) to obtain

log lo = A--3eL (3.17)I 2d p ■ ,P P

in the case of large ’black’ particles for which A =-E = 1,where E is the emissivity of a single particle then

3cLed = ^d = 1”e 2dPPP ............ (3.18)

In the. rotary cement kiln there is a considerable amount of dust generated,(both from the ’rolling’ action of the charge, and from the cooler) which is entrained by the secondary air and combustion gases in the kiln. In the burning area of the kiln this dust has been considered to be fine particles of clinker having a known density. Prom a consideration of the gas velocities in the system the maximum particles size that can be supported in the gas stream may be calculated and assuming normal particles size distributions a mean particles diameter arrived at. Thus we can reduce equation (3.18) to the following

d d (3.19)

and equation (3.1*0 may be rewritten to incorporate (3 *19) so that

For the rotary cement kiln we know the following information54 (see Appendix F )

p = 1715 kg./m3.

A complete description of the radiative properties of the medium having thus been derived, taking into account the variation of gas species, soot generation, and dust burden, we can now use this to predict the heat transfer mechanisms taking place within the furnace enclosure.

3.6. Direct exchange factor evaluation

The method of Einstein49 is preferred to the more commonly used one of Erkku4,8because it allows for much larger furnace configurations to be handled without the need to extend the pre-calculated values as is the case with Erkkufs values.

t

(3.20)

d 125 microns P =thus

U = 7.01 m 2/kg

In the case of a cylindrical enclosure axisymmetryis assumed and the enclosure is divided into a series of concentric annuli or rings, these being further divided along the axis at regular intervals.

If we consider a ring of radius R radiating to an infinitesimal volume or surface at a point P within the enclosure, this point is at an axial distance x from the plane of the ring and at some distance, radius r, from the centerline. The distance from P to some point on the ring is given by z, then z is found from the relationship:

z2 = x2 + (r-Rcos0)2 + R2sin20 ........ (3.21)

entire ring to the absorption at P is obtained.

If equation (3.21) is put into (3 *22) and integrated from 0 = to 2tt the contribution of the

6q = 2k 2dvSdR dx /^ P TT (3.22)

Some simplification can be achieved if we define the following

Thus equation (3.21) may be rewritten

z2 + z2 z2 + z2 •Z 2 = -------- :---- +--— ------ COS0 (3.25)

2 2

If we make the further substitution that T=kz then equation (3.22) can be written entirely in terms of the two parameters

and t^. There are six different types of exchange that can take place within the enclosure, gaz zone to gas zone, wall surface to gas zone, end surface to gas zone, wall surface to wall surface, wall surface to end surface, and end surface to end surface. The exchange factors for each of these are calculated as follows

gg = - -— - I-, (3.26)2 2 —7T T 0 '

if = K 3R e_lo (I _ r x } ..........................(3.27)tt2 0 t 3 2 R 030

if = k 3x silo. I .(3 .28)2 3 *-TT T 0

if = K V 2-JL- 1 ....................... ....(3.29)TT2 0 t" 5

0

ie = -*R x e 0 (I, - 2 1 ) ....................... (3.30)ir2 3 t 1* 00

ee = - 11 x2 2-1". 1 .... ........................... (3.31)tt2 t " 1 0

(See Appendix B for the derivation of these expressions.)

1^ to 1 are integrals in terms of x^ and x^, v-alues ofwhich have been calculated for the range of x in the furnace,and by interpolation other values may be derived.

3»7» Accounting for local variation of concentration of absorbing media.

When the concentrations of the components of the absorbing gas media vary locally without any change in their relative proportions, as would be the case where the combustion products would have a higher concentration at the axis of the kiln than they would near the kiln wall, allowance for the variation with regard to the exchange factors can be made.The average value of K(=knp) along a path r is given by

K = (/r Kdr) / r ....... (3-32)a v o

To evaluate K we must know the relationship between r andavp, and as a first approximation is is assumed that they are directly proportional so that

K. + K . p . + p .Kav — ---~ = kn — L .............

2 . 2

We then evaluate the direct exchange factors at the value of Kav and adjust the results as follows

(3 • 3*0

(3.35)

(ss) = (ss)T,/var \ ^

K. J

(SB)var = <S«>K “av Kav

3.8. Convective heat transfer within the furnace.

If in the system being studied Le = 1 then theconvective heat transfer rate can be calculated by making use of the difference in enthalpy as the driving force. The gas composition need not be known provided the total enthalpy of the mixture as a function of the temperature is available.

For combustion gases the assumption that Le = 1

(a) the flame gases are sufficiently turbulent for the contribution to heat transfer by aerodynamics to be very much greater than the contributions of either molecular collisions or diffusion.

(b) the concentration of hydrogen atoms is so low that it has a negligible effect on the overall heat transfer rate.

qconv g* (H - H ) &h g s (3.37)

whereh G St (3-38)Cp

is only valid if either

In practice, most combustion systems involving hydrocarbon fuels produce a significant amount of hydrogen

atoms at high temperatures (>2500°K. see Pig. 3*5)• Under these conditions Le i 1. Several calculation methods have been developed to account for this situation, but these seem to agree only in their orders of magnitude, as shown by Chen and McGrath.55 The method of Altman and Wise5 Appears to give answers in the middle range of results, and is expressed as the ratio of heat transfer rates with and without chemical reaction in a turbulent boundary layer

q ’ o Q. (N.„ - N. )= 1 + X i IS .....:... (3.39)qconv 1 (1 - Nn ) Cp (T - Tg)

where

q = h (T - T ) (3.40)Hconv g ■ s

and h is evaluated from the standard relationship

Nu = 0.42 Re *6 Pr ’ 3 3 ......................... (3 - 4l)

Prom a consideration of the results in Section 5 obtained on a coal fired Barnstone kiln, where the maximum flame temperature measured was 2400°K, a correction for Le i 1 is unnecessary, but a convection correction for chemical reaction is significant since this increases convective heat transfer by a factor of 1.5 - 2 .0 .

In the case of gas and oil fired kilns a further correction for Le / 1 may be required.

CO

NO

.01- OH

.005

.001454025 30 35201510

Temperature °CxlOO FIGURE 3.5 Variation of species concentration with, flame temperature in

a coal/air flame with 5% excess air

3«9» Heat loss to the atmosphere by conduction throughfurnace walls.

The conductive heat loss through the furnace.wall can be directly equated to the radiative and convective heat loss to the atmosphere from the external surfaces of the furnace5 provided that the system is in a steady state condition.

Heat is lost from the shells' of both the kiln and cooler by two major modes of heat transfer3 vz. Convection and Radiation. Using published relationships an expression has been derived to calculate both the radiative and convective heat losses with just a knowledge of the shell and ambient temperatures.

3*9«1« Convective heat loss from a tube

According to the equation of Krausold for natural convection.57

0 * 3 3Nu 0.13(Gr.Pr) f (3-42)

which on rearrangement gives

Nu0 * 3 3

(3 .43)p . k *

or

h 0.13D

.k P2-gg-CP .d !At (3.44)p .k *

t This has been shown to be valid for horizontal tub^s as well as vertical tubes in the turbulent range Gr = 1 0 . “ 10•, r ** »«> 0 , *

McAdams59gives the following data (Table 3.3) for varying values of t.

tTable 3*3 Values of temperature dependant data required

in the calculation of the convective heat transfer coefficient.

t

°K

k*

kcal/hr m 2°C/m

„ _ p?8.g.Cp p.k*

(m3°C)”1

273 0.02080 -

373 0.02635 77 X 106473 0.03330 36.25 x 106573 0.03860 18.08 x 106673 ■ - 11.05 x 106

From this data a straight line relationship between k and t has been derived of the form

k * = t + 72 16670

(range 0-400°C) ....(3.45)

.A similar relationship between y and t has been derived by polynomial curve fitting of the form

8813.4(t-273)

- 9.18 x 106 .... (3.46)

Thus we may rewrite equation (3-46) so that the total convective heat loss (Qc ) is given by

Q, = 0,13 D

(t+72)16670

8813.(t-273)

- 9.18 X i o 6 . d 3 ( t - t )£L

0 . 3 3

(t-t

kc al/m2hr (3.47)

3.9'» 2. Radiative heat loss from a tube.

From experimental'• work carried out by Kuhle5,8 a relationship for the emissivity of heavily dust coated steel shells in terms of surface temperature has been derived.

0.96 - 5.2.x lO"4 (t-373) (3.48)

Q R

The general radiation heat transfer equation gives

= o e ( t ^ t M ..... ....(3.49)d

o = Stefan-Boltzmann constant = 4.88 x 10 8 kcal/m2hr°ktt

thus combining equations (3*48) and (3-49) we have

Q r = 4.88 X l o ' e ( 0 . 9 6 - 5.2 X 1 0 " ( t - 3 7 3 ) ) • ( t " - t & ")

kcal/m2hr (3 - 50)

3.9.3 Total Heat loss from a tube

If we combine equations (3*47) and (3*50) we may say that the total heat loss (Q,p) from 1 square metre of kiln surface is given by

Qrp

(t-t„) + 4.88 x 10

0.13 (t+72)'

[8811-1 . - 9 is]*

.D3 x 106D 16670 [(t-273) J 2

*

0 . 3 3

0.96 - 5.2 X 1 0 “ " ( t - 3 7 3 ) . ' ( t " - t a ")

keal/m2hr .............(3 •51)

Thus if we take values of t at measured distances along the kiln or cooler we can get corresponding values of Qrp . These values of and L (distance along the kiln from origin) can then be related by means of a polynomial using curve fitting procedures, if we call this polynomial U (i.e. Q 1 = aiL + a2L2+ etc.)

Tiien the total shell heat losses are given by

Q = >.D.£L D.dL. keal/hr....... (3.52)

3.10 Evaluation of gas enthalpy terms

Prom McAdams 59the following values,(Table 3**0 of Cp have been taken for the primary combustion products and air.

oTable 3.4 Specific heats’ of combustion gases (kcal./kg. C)

Temp. °C CpC02 CpH20 CpAIR(N2)

1430 0. 3H 0.65 0 • 0

750 O.30 O.38 0.275

Average values of Cp for air and CO2 may be used in all calculations with little loss in accuracy. For H 2O a linear relationship of the form

Cp„ n = .08 + .0004T (3-53)H 2 0 g

has been adopted.

From a knowledge of the stoichemtric air requirement of the fuel and the' excess air level in the furnace, the total combustion gas mass flowrate M_ can beocalculated. The enthalpy of the combustion gases at any point in the furnace is then given by

oH„ = £ p.Cp.T M ............... ...... (3 • 5*0g i i l g g

3.11. Prediction of heat release rates for kilns.

By considering a series of planes at right angles to the burner axis in the flame zone of the kiln, and by using available gas profile data, the time mean average gas concentrations of CO, C02, and 02 in those planes can be estimated. Thus by determining the rate.of CO2 generation and its intermediary CO’, we can gauge the rate of heat release within the flame. The rate of 02 consumption may also be used to verify the heat release rate.

There are two major drawbacks to the use of this method for the rotary kiln.

(a) Recirculation of combustion products in the latter part of the flame will lead to higher values of C02 and CO concentrations than would be expected, and thus overprediction of the heat release from the fendTof the flame is probable.

(b) C02 release from the charge material will alsolead to higher values than would be expected, giving the same results as (a).

If we consider some plane in the burning zone of the furnace i, then the gas concentration of any specie-s at that plane may be calculated as follows.

[Gas]

avvo max

If[Gas] =

Then[Gas] =

av

F(-x) (3.55)xmax

x

[Gas] = avDistance x

Figure 3.6 Average gas concentration at any plane.

/ IIla-x [Gas] dx ... (3.56) o/ m a x I ( x ) dx .(3 .57)

max

Equation (3*57) may be determined at each chosen plane by graphical integration or by fitting an expression to equation (3 .55) and integrating between limits.

3.12. Distribution of mass flows in an enclosed jet system.

In a direct fired rotary kiln system some, or all, of the air required for complete combustion must be entrained into the fuel/air jet stream from the secondary air surrounding the jet. Because the system is closely confined by the kiln shell, thus limiting the availability of secondary air, and depending on the momentum ratio of the two streams, areas of recirculating jet fluid can occur (See Fig. 3.7). A proven criteria6for determining the onset of recirculation in sucha system is the Craya Curtet parameter, m, defined by: 13

^ AREA OF RECIRCULATION

FIGURE 3*7 Generalized streamlines in an axisymmetric entraining jet system with recirculation.

entrainment and recirculation between two plug flow systems

singleunidirectional plug flow

no'ainment ent

FIGURE 3 i8 Simplified model- of entrained confined jet.

m K R' *R

*R2 (3-58)

oR

where*R

Q

7r r-o (uo - V

tt R u + qcl

The value of m varies from zero to infinity, with the onset of recirculation occurring at m>1.5 (for an axisymmetric jet) and increasing with increasing m.

The rate of entrainment of secondary fluid into a fast moving jet stream has been investigated by Ricou and Spalding9 and the effect of density differences between the fluids has been considered, an expression was derived of the following form

mm_ = 0 .32. — .

d_ex .PiJ (3.59)

A simplified model is proposed to describe the mass flow in a rotary kiln system (see Figure 3*8) which makes the following assumptions.

(a) No entrainment occurs within 3 nozzle diameters of the nozzle (Region of jet establishment)

(b) Flows within and without the jet are uniform (but not equal) over any cross-sectional plane at right angles to the kiln axis.

(c) The jet is axisymmetric and expands at a constant jet half angle of 11°.61

(d) Entrainment rates are governed by the expression of Ricou and Spalding (Equation 3-59)•

(e) Recirculation may or may not occur, but sufficient secondary fluid is always available forjet entrainment and no account is taken of disentrainmen from the downstream jet.

(f) When the jet strikes the wall of the furnace, all entrainment ceases and flow is uniform in one direction across the duct (i.e. plug flow).

SECTION 4.

DESCRIPTION OF EXPERIMENTAL APPARATUS AND OPERATIONAL TECHNIQUES

4.1. Brief description of the rotary kiln

The rotary kiln is one of the most important types of high temperature process furnaces and is generally employed when the feedstock to be processed or heated is in a granular or slurry form. The main industrial applications of rotary kilns are in the calcining and sintering industries, such as lime burning, cement manufacture, sintering ore ’fines’ from blast-furnace flues, dolomite burning, basic aluminium oxide calcination, and calcination of alumina and silica/alumina carbonate mixtures for brickmaking.

The kiln used to obtain our experimental results is a hollow steel tube .834 metres internal radius and 45.7 metres in length. The whole tube is mounted in metal tyres at approximately 20 metre intervals, which in turn rest on roller beds. The drive is by a central spur wheel gear producing rotational speeds between 1-1.5 r.p.m. The tube is inclined at an angle of 2° to the horizontal to allow for the passage of material through the kiln, and is lined with-refractories for the major part of its length.

The work carried out in this practical study is , specifically concerned with the rotary kiln as applied in the

cement manufacturing industry, but the problems encountered in this industry are typical of those for any•in which rotary furnaces are used.

4.2. Brief description of the cement making process equipment.

Figure 4.lis a block diagram of the Barnstone cement plant where the experimental program was conducted. Portland Cement is made by three major types of plant, depending on the state of the feedstock, these being wet (slurry), semi-wet (damp nodules) and dry (powder); but in all three, final processing is carried out in a rotary kiln.

4.2.1. Wet process

In this process chalk or limestone is milled with clay and water in open or closed mills to give a slurry with a moisture content between 35-44$ wet basis, this being governed by the free moisture in the starting materials and the slurry viscosity.

This slurry is fed to the cold, upper end of the kiln where it encounters a curtain of chains hanging vertically down in a spiral arrangement to within0.5 metres of the kiln bottom. These chains act as heat exchangers between the hot kiln gases and the wet slurry by increasing the surface area, thus helping to speed the drying process, and also as dust collectors, trapping the airborne dust in the kiln gases on the sticky chains. The whole chained area comprises some 20$

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of the total kiln length, and is lined with hardened steel plates welded to the mild steel shell to reduce wear from the chains.

The kiln shell ranges in thickness from 2 cm. at the cold end .to 4 cm. in the burning zone, and is usually of an all welded construction. The remainder of the kiln shell beyond the chained zone is lined with approximately 20 cm. thick refractory bricks of various grades.

4.2.2. Semi-wet process

In this process a pre-wetted, milled mixture of limestone and clay or shale are fed to a dish noduliser which produces nodules of 1-2 cm. diameter with a moisture content of 10-15$. These nodules are discharged onto a moving metal perforated grate through which the kiln exit gases pass. In this way the nodules are dried and partly decarbonated before dropping into the cold end of a short rotary kiln for final processing.

4.2.3. Dry process

A dry milled mixture of limestone and clay or shale is fed to the top of a series of 3 or 4 cylones where it passes countercurrently to the exit kiln gases and is decarbonated.This decarbonated feed then drops into the cold end of a very short rotary kiln where the final clinkering process■takes place

4.3. Temperature and mass flow measurements

The Barnstone kiln was specially instrumented to enable a continuous monitoring of all the input and output parameters to be made. In addition, specific measurements of gas concentrations and temperatures were made in the flame area of the kiln. Thus overall heat and mass balances were calculated to ensure that the individual flame measurements were valid, and to provide basic data for the testing of theoretical models.

4.3.1. Mass flows

(a) Vane anemometryThe air flow through the system was measured

using a variety of instruments and techniques. The air drawn into the system through the cooler end was metered using an Abbirko Flowmaster, Model PA, vane anemometer. Twelve positions on a rake across the mouth of the cooler were recorded and an arithmetric mean average taken. The instrument consists of a freely rotating ’windmill’ protected by an outer ring. Mounted in the outer ring is a small capacitance transducer. As the vane assembly rotates, so electrical pulses are developed, these being proportional to the speed of rotation. By means of simple electronics this signal is displayed as a direct velocity reading on an attached meter.

The range of the instrument is 0 - 25.0 metres/ sec. Readings were taken hourly so that a daily average-'- figure was obtained.

PLATE 4.1 Monitoring of cooler air flow.

Measurement of cooler air flow by yane anemometry Cright) and argon tracing (.left),

(b) Argon tracingThe techniques of inert gas tracing were

modified to measure total and inleaked air flows in the system. A small metered quantity of argon gas was injected at the cooler air inlet. Gas samples were withdrawn from the system into rubber holding tubes by means of vacuum pumps. These samples were taken from a point in the kiln hood uptake, and at a point at the back-end of the kiln. Since the quantity of gas injected is known, analysis of the samples to find the dilution of argon will give a value for the airflows. All samples were analysed for argon, nitrogen, carbon dioxide and oxygen using a mass spectrometer at the Wolfson Institute of Interfacial Technology,University of Nottingham. Results from vane anemometry were compared with the values obtained by argon tracing to ensure accuracy.

(c) Pitot traversingPitot traverses of the primary air flow ducts

were made at hourly intervals using a standard pitot tube and inclined limb manometer. From these readings the total primary air quantity and the air inleakage through the coal milling system were calculated.

The accuracy of these readings was seriously affected by the coal dust present in the airsteam, as this tended to block up the pitot tube pressure tapping holes and constant clearing was thus essential

(d) B.C.U.R.A. coal and air flow measurementsTests were carried out during trials 4 and 5 using

a B.C.U.R.A. Mass flowmeter kindly loaned to us by the Coal Research Establishment, of the National Coal Board.

This instrument is designed to measure pulverized coal-air flows in ducts, and was used to monitor the primary airflow in the firing pipe duct.The flowmeter consists of an ultrasonic velocity meter, a beta density gauge, and an electronic system for combining the two measurements. The meter has the advantage that none of the component parts need be inserted into the flow so there is no wear or obstruction to the flow.

Figure 4.2 is a diagramatic representation of the instrument set up. Matched magnetostrictive transducers are used to generate, or receive ultrasonic signals. The gas velocity is obtained from the difference in transit time of two simultaneously generated, in phase, acoustic waves, one propagating upstream and one downstream of the flow. The wave directed upstream is advanced, while the downstream wave is retarded. The mass flow unit which measures the attentuation of a stream of 3 particles across the duct was not used.

Difficulties were encountered in the system because the transducers used have a temperature tolerance of ~50°C to +105°C. Duct temperatures rose to 120°C at times which affected permanently the signal frequency of the transducers.

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(e) Coal mass flow by weighout and weighfeederIn Trial 1 the coal rate was measured by

weighing all coal deliveries over a weighbridge, weighing the remainder of coal in the hoppers at the end of a trial, and by subtraction, calculating the total coal consumed during the trial periods.The coal to the milling system was controlled by a gate at the base of the hopper, and unless it was absolutely necessary this was fixed for the duration of the trial. Thus an average coal consumption was worked out.

The weighout system, as can be seen, is not particularly satisfactory for accurate working of results, and thus for trials 4 and 5 a Wallace and Tiernan pneumatic gravimetric feeder was fitted between the coal hopper and Attritor to monitor and control the coal rate continuously. Figure 4.3 shows the construction of the instrument, which operates by weighing a section of coal on a conveyor belt, as it passes. This measured weight is used to govern the belt speed and gate setting so that the set point is maintained. This instrument provided a reliable and instant check on the coal mass flowrate. .

(f) Clinker production by weighoutPrior to discharge to the storage silos, all

the clinker produced by the kiln was emptied into a lorry. The amount of clinker produced over a certain period (usually 2-3 hours) was then ascertained using a weighbridge and average daily

15D

10 13

12 6

1. Feed hopper.2. Weigh belt.3. Synchronous motor4. Gear box.5. Weigh deck.6. Belt rollers.7. Pivot yoke.8 . Scale beam

9. ’Flapper and nozzle’ air relay10. Piston.11. Pneumatic controller.12. Pneumatic gate positioner.13. Stationary deck.14. Feed•delivery recorder.15. Feed control panel.

FIGURE 4.3 Flow diagram of Gravimetric Feeder.

production figures were thus arrived at. A further check on these figures was made by noting the rate of slurry feed to the kiln, and, from a knowledge of the chemistry a clinker production rate was worked out.

(g) Dust loss by weighoutAirbourne dust in the exhaust gases from

the kiln was removed by electrostatic precipitation. This dust was then discharged into a bulk powder lorry which was tared over the weighbridge. Thus the daily average dust loss was calculated.

*1.3.2. Heat flowsc

(a) Clinker temperatureThis was measured at regular (approximately

hourly) intervals over each trial period using a Pyrotenax metal sheathed Ni.Cr./Ni.A 1 . thermocouple and a Comark electronic thermometer. The thermocouple had a maximum operating temperature of 800°C and was of the bonded junction type, giving faster response.

Samples were extracted into a refractory cement lined bucked while the temperature was recorded.

(b ) Secondary air temperatureThe temperature of the air after its passage

through the clinker cooler and prior to its entry into the kiln and kiln hood was measured by a water cooled suction pyrometer positioned in the uptake from the cooler. A metal shielded, metal sheathed Pyrotenax Ni.Cr./Ni.A1. thermocouple was used to measure the temperature, and compensating cable connected this signal to a continuous Kent chart recorder.

(c) Primary air temperatureThe air temperature in the firing pipe duct

was measured, using a mercury in glass thermometer, concurrently with the air flow pitot measurements described earlier.

(d)Clinker nose ring temperatureThe temperature of the clinker as it fell

over the nose ring was measured using a disappearing filament optical pyrometer. This instrument provides a method of manually matching the colour of a calibrated resistance wire against the background of the clinker sample by adjusting the current flowing through the wire. When the temperatures of both are matched the wire effectively ’disappears’ into the background. From the current flowing a direct temperature reading is shown on the instrument.

Errors of - 2q°C are possible with this method. This measurement was recorded hourly during the trials.

(e) Ambient air temperatureThe ambient air temperature was recorded

hourly during the trials using a mercury in glass thermometer placed in a sheltered position just outside the kiln firing platform. .

(f) Kiln and cooler shell temperaturesPositions were marked at approximately two

metre intervals along the length of the kiln andcooler. The temperature of the shell at these pointswas then recorded at two hourly intervals using a KaneMay Infratherm infra red electronic thermometer. Theinstrument operates remotely from the kiln shell by pointing the receiving lens at the shell. Theinfra red energy impinging on the lens is directed to alead sulphide cell, which then generates a small D.C.signal. This is modulated to an A.C. signal by choppingthe energy source. The strength of the signal isgoverned by the energy received, thus a direct readingof temperature can be indicated. The instrument ishand-held and battery operated giving quick and easilyrecorded results. The instrument had a range of 100°Cto 500°C in three spans with an accuracy of - 2%' P.S.D.of the span employed. A knowledge of the emissivityof the surface was required to obtain accurate temperaturreadings.

(g) Kiln and cooler shell emissivitiesThe emissivity of the shell of the kiln and

cooler were measured at the same positional intervals along the kiln using a Land Type S.P. surface pyrometer.

A hemispherical gold cup with a thermopile, protected by a flourite window, at its apex is placed over the surface and the D.C. signal generated is translated to a temperature reading (T ) using a Cambridge Spot galvonometer. The gold surface is then replaced by a matt black surface and a second temperature reading (T^) taken. From the difference T^ - T^, using calibration charts provided by Land Pyrometer Ltd., the emissivity of the surface is calculated.

Readings were taken at 3” hourly intervals during Trial 1 and daily during Trial 4 and 5. Thus an emissivity vs. surface temperature relationship was derived for the kiln and cooler shells.

(h) Back-end Gas TemperatureThe back-end gas temperature was metered

continuously using a Pyrotenax metal sheathed Fe./Con. thermocouple (maximum operating temperature = 750°C) inserted through the back-end chamber into the gas flow. The generated e.m.f. was translated to a temperature reading using a Kent continuous chart recorder.

PLATE 4.2 Measurement of kiln shell, emissivity

Land surface pyrometer measuring shell emissivity sintering region of the Barnstone kiln.

4.3.3* Material analysis

Laboratory analysis was carrried out on the slurry feed, clinker, dust and.coal to determine the percentage composition of some or all of their chemical constituents.

(a) Clinker analysisA cumulative daily sample of clinker (every

two hours) was analysed to determine the percentage free lime (CaO) using standard A.P.C.M. analysis methods, thus giving an indication of the degree of -clinker formation taking place in the kiln. A cumulative trial sample was analysed by A.P.C.M.Research Department to determine the chemical composition of the clinker for each trial.

(b) Slurry feed analysisA cumulative daily sample (every two hours)

of slurry was analysed to determine -its moisture and carbonate content. The moisture percentage was found by drying a measured weight of slurry. The carbonate percentage was determined by back titration using AnalaR hydrochloric acid and sodium hydroxide.A dried cumulatived trial sample was analysed by A.P.C.M. Research Department to determine its chemical composition for each trial. The heat of formation of clinker from slurry for the Trials 1, 4 and 5 has been calculated as 432.20 Kcal/kg. ( See Appendix C ).

(c) Coal analysisSamples of pulverized coal were taken every

two hours and analysed for moisture content by drying, and for fineness between 90y and 300y sieves. A cumulative pulverized coal sample was taken over each trial period and analyses for calorific value using a Bomb Calorimeter, moisture by drying,, ash by ignition, and volatile matter by controlled combustion at low temperature. These tests were carried out by A.P.C.M. Research Department. A proximate analysis of these samples was carried out using a mass spectrometer in the Department of Chemistry, University of Surrey, to determine carbon, hydrogen and nitrogen, and in the Department of Chemical Engineering using Eshka’s Mixture to determine sulphur, oxygen was found by difference.

An ultimate analysis of the coal, which was delivered from Pleasley colliery throughout the trials, was supplied by the National Coal Board.

(d) Dust analysisSamples of dust from the electrostatic

precipitators were taken every two hours and analysed for loss on ignition to determine the carbonate percentage and volatile mineral matter in the sample. A cumulative trial sample was analysed for carbonate and carbon percentage by A.P.C.M. Research Department.

4.4. Flame measurements

4.4.1. Stationary gas sampling

The first flame gas sampling system to be employed at Barnstone was one in which the kiln was stopped for brief periods; approximately 3 minutes. Samples were withdrawn through water cooled, stainless steel probes (see Figure 4.4 ) which were inserted vertically through portholes in the kiln shell from a platform above the kiln. This platform extended for the full length of the burning zone of the kiln.

Flexible reinforced rubber tubing was used to supply cooling water from a permanent water supply to the probes and to discharge the hot water into a surge tank and drain. This surge tank also provided a visual check to ensure that cooling water was circulating at all times to the probes. Gas samples were extracted from the flame by means of a vacuum pump, connected to the probe by rigid plastic tubing, with a sintered bronze filter placed in the sample line to protect the pump.

The sample line was purged for approximately one minute before samples were collected. Samples were pumped into rubber inner tyre tubes for subsequent analysis (see Section4.4.3.),

Samples were only taken at a predetermined series of excess air levels as indicated by the back-end oxygen reading. Whilst the kiln was stationary the kiln burner controlled the back-end oxygen reading at a constant value by adjusting .

PLATE 4,3 Stationary gas sampling,

Clearing a porthole from the sampling platform prior probe insertion.

%

the back-end damper.

This method of sampling was used to determine flame gas compositions in Trial 1.

During Trial 1 samples were taken from five portholes in the flame zone, three sample positions being chosen at each porthole; axial and .407 metres above and below the axis. From a consideration of the data extracted from Trial 1, two extra portholes were inserted in the kiln and samples were taken only from the axis position at each porthole during Trials 4 and 5•

\

For safety reasons no more than two probes were inserted into the kiln flame zone during any one stop. Each probe was attended by two people, with a third person extracting the gas sample.

TABLE. 4.1. Porthole positions along kiln measured fromthe nose ring.

PortholeNo. ± 2 3 • 4 5 6 7

Trial No.1 4.67m. 7.46m. 10.25m. 13.15m. 15.9m. - -

4/5 2.69m. 4.67m. 6.12m. 7.46m. 10.25m. 13.15m. 15.9m.

4.4.2. Rotating gas sampling

In order to eliminate the effect of rotation of the kiln on flame characteristics, a gas sampling system was devise to operate integrally with the kiln.

PLATE 4.4 Sample porthole.

View of a continuous sampling porthole, showing quick release clamps and suction pyrometer fan.

Figure 4.5 shows the construction of the sampling probe used for continuous operation. Figure 4.6 is a diagramatic representation of the cooling water and gas sampling arrangements on the kiln and Figure 4.7 shows the actual arrangements of the forced convection water cooler and sampling unit, which was attached to the cool end of the kiln. The radiator units and fan units were sufficient to cool 52 lits./min of water against a 20 metre head from 80°C to '50°C or less. Booster pumps were fitted to the unit to supplement the main Mono pump in order'to maintain a good supply water pressure to the probes. Pressure, temperature and flow indicators were fitted to the supply line, and a temperature indicator was fitted to the return line from the probe, which activated an audio alarm if the temperature was over 85°C, since this would imply that steam was being generated at the probe. A pressure relief valve was also fitted to the unit as a precaution against steam generation. A vacuum pump was also mounted in the unit, connected to a galvanized steel/flexible plastic pipeline from the probe. The whole unit was supplied with power through a slip ring round the kiln, giving sufficient electrical input to drive an equivalent total load of 4KW.

During Trial 4 samples were taken from the flame zone of the kiln in the same manner as for the stationary system, except that the gas sampling line was purged for a longer

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period since it was more remote. No filter was placed in the sample line, but a compressed air purge was available to clear any blockage. Only one probe was inserted in the flame zone at any one time, and all samples were taken on the axis of the kiln, since off-axis samples would only provide a mean about a circumference, and would not help in the determination of the flame shape. The-same series of excess air levels were used in Trials 4 and 5 as in the stationary system, but the control of these levels was much easier, since the kiln was operating under steady state, normal running conditions. In all cases three samples were taken at a given excess air level and their arithmetic mean values used in the determination of flame gas composition.

Gas sample analysis

All flame gas samples extracted from the kiln into rubber holding tubes were analysed for carbon monoxide and carbon dioxide using a I.R.G.A. infra-red analyser, and for oxygen using a D.C.L. Servomex paramagnetic meter. The operation and accuracy of these instruments are summarized in the following paragraphs.

Oxygen Analyser

The D.C.L. Servomex oxygen analyser depends for itsoperation on the paramagnetic susceptibility of oxygen. Itsrange was 0 -'100# oxygen on 3a600° (ten turn) linear dial,

+graduated to 0.1% and its accuracy - 0.1#. The analyser

owould function within the range -10 to +^0 C and the•hO .accuracy was maintained over ~5 C. variation of ambient

temperature. With the sintered glass disc filter removed and the sample flow rate at 100 ml/min.(optimum working flow rate), distance velocity lag did not exceed 0.1# oxygen. As the reading of volume percentage dry gas was required, the sample had to be - dried before being analysed and the conditions were maintained such that condensation could not occur within the analyser’s measuring cell. Pure nitrogen was required for gas check and dry air for adjusting the span. Readings were taken at the same pressure (usually atmospheric) at which the span was adjusted.

Carbon Monoxide and Carbon Dioxide Analyser

The standard Poly I.R.G.A. depends upon the selective absorption of infra red radiation by CO or C02 . Both these gases have a unique absorption spectrum and therefore the absorption of one gas in a complex combustion mixture could be selectively measured. The amount of absorption was a function of the concentration of the gas and the length of the absorption path. For measurements of a wide range of concentrations, a multisection analysis cell was fitted, consisting of three sections with cell lengths in decade steps.

The detector was sensitised to the gases to be measured and selectivity was obtained by interposing matched pairs of filters in the optical paths. The filter system was fitted with trimming devices so that the optical balance

could be adjusted for individual channels. The gas to be measured was selected manually by a simple filter changeover valve which directed purge gas (Nitrogen) to the two other range cells not being used at any time.

Analyser Specification

Reproducibility. + 1# of the full scale readingStability. + 1# of the full scale reading

over 2k hoursStandard Measurements Carbon Monoxide/Carbon DioxideStandard ranges 0-1#, 0-10# and 0-100#.

All samples to both these, instruments were filtered through glass wool and dried over silica gel prior to analysis. During Trial 1 standard Orsat gas analysis apparatus was used to check the accuracy of the analysers for oxygen and carbon dioxide readings, but this was abandoned in subsequent trials when it was found that the instruments were giving reliable, repeatable results.

The control of the excess air levels in the actual sampling operation from the kiln was effected by measuring the oxygen gas concentration at the kiln exit. A probe was placed at a representative sample point in- the backend of the kiln, and a continuous iso kinetic sample was drawn through a D.C.L. paramagnetic meter, similar to the one previously described, by a peristaltic pump.. The measured reading was recorded visually both in the kiln control room and at the back-end of the kiln, near the

PLATE 4.6 Gas sample analysis.

Analysis of combustion gas samples by Orsat Cleft) and infra-red absorption (right) apparatus. A rubber holding tube for gas samples is shown in the left foreground.

flame gas sampling apparatus, so that the burner could control the oxygen level, and the sampler could see when the kiln was in a steady state.

4.4.4. Flame temperature measurements

Measurements were obtained in Trial 5 of flame gas temperatures to supply basic data for a study of the shape, heat release and buoyancy of the flame in the hot kiln.

Figure 4.8 shows the construction of the suction pyrometer used to obtain flame gas temperatures of up to 1550°C. The water cooled probe was constructed of stainless steel with an o.d. of 63*5 mm and a central gas suction port of 25*4 mm i.d. The probes were supplied with cooling water from the cooling unit previously described in Section 4.4.2. Platinum/Platinum, 13#Rhodium thermocouples with a temperature limit of 1770°C were used to measure the gas temperature.The thermocouple junction was protected by a recrystalized .alumina sheath which has a temperature tolerance of 1700°C.The sheath was fixed to a narrow bore stainless steel pipe which carried the sheathed thermocouple wires through the centre of the probe to a junction box at the rear of the unit. Special, high temperature, radiation shields made from recrystalized aluminia and coated with aluminia cement were used to ensure an accurate gas temperature reading, since the thermocouple must be isolated from any radiation from the surrounding walls if a true measure of the gas temperature is required. The shield was fixed externally

9 2 ex.o

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FIGURE

4.8

Suction

pyrometer

for

flame

temp

erat

ure

meas

urem

ents

PLATE 4'. 7 Suction pyrometer.

Water cooled suction pyrometer probe? assembled prior insertion into the kiln.

PLATE' 4.8 Pyrometer shields.

Porthole No.7

Porthole No. 6

Porthole No, 4

Unused

Flame temperature = 1085°C.

Flame temperature - 1305ftC,

Flame temperature ^ 1540°C,

Suction pyrometer shields showing the results of approximately four hours operation in the burning zone of the kiln.

■•‘-.V-v . « . *'< v.

to the probe, thus reducing its life, but significantly increasing the accuracy of the readings.

The convective heat transfer coefficient to the thermocouple was increased by aspirating the gases through the shield at about 150 metres/secusing a blower mounted on the kiln shell next to the probe, and connected by a flexible metal tube to the suction connection. The thermocouple signal was recorded on a Honeywell continuous chart recorder which was fixed to the kiln shell at the rear, ’cool’ end of the kiln next to the water cooling unit.Compensating cable carried the signal from the junction box of the probe to the recorder.

■ (For the measurement of temperatures in excess of 1550°C

two Land Venturi pneumatic suction pyrometers were used,Figure 4.9 is a sketch of one of these probes showing its detailed construction.’ Since thermocouples and their ancillary shielding cannot operate with any reasonable limits of accuracy or life above l600°C, some other temperature dependant property of the gases must be measured. The density of the flame gases is employed in the Venturi system to measure their temperature.. The hot gases are aspirated into a Venturi restriction (’hot Venturi’) at the end of a water- cooled probe. Pressure tappings at the Venturi measure the pressure drop,AP^ of the hot gases. As these gases pass down the probe they are cooled by the water jacket. The rate at which the gases are sucked from the system is controlled so that they are above their dev; point prior to exit from the probe. Here they are

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PLATE 4.9 Venturi suction pyrometer.

Venturi, pneumatic suction pyrometer operating in porthole No. '3. Flame, .temperature recorded as 1950ffC ,

drawn over a second Venturi restriction (’cold Venturi') and pressure tappings measure the pressure drop,AP of the. cooled gases. A resistance thermometer fixed in the gas stream after the ’cold’ Venturi measures the temperature of the cooled gases.

Since the mass flow rate of gas is the same over both the 'hot' and 'cold' Venturis we may write, for subsonic conditions.

m = A, £, a, /p. A P. h h h Kh h A e a / p A P c c c ^c c (4.1)

The ideal gas equation gives the following results

_h>h

RTh andMh

RT (M (4.2)

Thus combining equations (1) and (2) we may say

Th T Apht

P^M.h h A, e, a,h h h2

AP P M A e ac c c o o o

4

(4.3)

or shortly

T.hAPV hAP .F (4.4)

The calibration factor F can be simply determined byaspirating cold air through the instrument so that T^ = Tc ,and measuring AP, and AP .® h c

The shorter of the Venturi probes (probe length = 1.83 metres) was installed in the kiln in the same manner as the suction pyrometers. The pressure tapping connections were coupled by rigid plastic pressure tubing to a small analogue computing unit and amplifier mounted remotely on the kiln shell next to the Honeywell recorder used for the suction pyrometry described previously. The larger of the Venturiprobes (probe length = 3*66 metres) was used to investigate flame temperatures near the jet nozzle, and was inserted through the kiln hood. Since this does not rotate, the supply of cooling water and the other service pipes to the probe were installed using normal service mains and equipment.

All rotational probe sampling was carried out at three positions through six of the seven portholes, (porthole 1 was not used as the longer Venturi provided data in this region of the furnace). Readings of flame gas temperature, at two excess air levels were taken axially, and at .152 and .457 metres from the axis. The off-axis readings were a mean value about a circumference because the response time of an ordinary suction pyrometer is of the order of 2-3 minutes, and the rotational speed of the probe was one revolution per minute. Also fluctuations with a periodicity of less than one minute are damped out. This was not considered a major disadvantage since the suction pyrometers were used only in relatively ’cold’ areas of the flame, where the temperature gradients are consequently less steep.

The Venturi pneumatic probes were used in the ’hot’ region

of the flame to provide a comprehensive ’map' of this area since they have almost instantaneous response times. Some difficulty was however encountered in the rotation' of these instruments and ancillary equipment due to the complex nature of the probe and computer, with the readings tending to ’’drift’'from the initial set points. It has been estimated that all;suction gas temperature readings taken in the trials are accurate to + 50°C.

SECTION 5.

RESULTS OF TEE EXPERIMENTAL AND COMPUTATIONAL INVESTIGATION OF THE SINTERING ZONE OF A KILN

Two programmes of work have been carried out concurrently in this dissertation.

I. An experimental investigation has been conducted on a small cement kiln, during which measurements of flame gas composition and temperature have been taken, and heat balances worked out from input/output mass flows and temperatures.

II. A computational study has been undertaken, using the- theoretical analysis developed previously ( see Section 3 )l in an attempt to simulate the measured results of the experimental investigation, and hence disclose any theoretical inadequacies.

5*1. Experimental results * ^

The kiln which was made specially available for this work by the Associated Portland Cement Manufacturers Ltd. is situated at Barnstone in Nottinghamshire, and is described fully in Section *1.

5.1.1. General kiln operating conditions

From data on the operation of kilns within the Blue Circle Group of Companies during 19713 Histograms have been constructed to show the variation of the significant parameters affecting wet kiln operation. ( see Figs, 5-1. to 5.6. ).

a

These factors are:-

(i) Velocity ratio of primary to secondary air at the nozzle plane.

This governs the rate of entrainment of secondary air into the primary jet stream.

(ii) Primary air percentage.This governs the rate of combustion of the volatiies

in the fuel, and the amount of preheated secondary air that is available for combustion.

(iii) Excess air percentage.This governs the total air available for combustion,

and hence the length of the flame and its temperature.

(iv) Craya-Curtet parameter.(m)This quantifies the degree of recirculating combustion

gases in the flame.

(v) Specific fuel consumption.This demonstrates the overall efficiency of heat

transfer and recuperation in the system.

(vi) Exit gas temperature.This measures the efficiency of the chained drying

section in the recovery of low grade heat.

From this data, the suitability of Barnstone as a represents!: i\

CO£i—l • H

O55

3 015 2 55 2010u /u o a

3 0 -

10-

Fercentagi

FIGURE 5.1Distribution of Primary:

Secondary Air Velocity Ratio.

50,

4 0-

£ 3 0,I—I • H

Cm / O. 20-

10

1 1.5 2 2.5

Percentage 02.

FIGURE 5.2Distribution of Primary Air

Percentage.

FIGURE 5.3Distribution of Excess Air

Percentage.

40

30 -

DOc 20_i—i •HCh OO 10-

—

0 .4 . 8 1.2 1.6 2.0 2.4

Craya Curtet Parameter.FIGURE 5-4Distribution of Craya Curtet Parameter

4 0 .

0 -j---- ■; — — > f—— f—— j.—~p13 14 15 16 17 18 13 20 2.1 22 23 2 4 2 5

3 0 -

mCrH 20-] • H

OO 10

0 w _I

kcal./kg. x 10 - 7175 225 275 32 5 375 425

Degrees Cent igrade

FIGURE 5.5Distribution of Specific Fuel

Consumption.

FIGURE 5.6Distribution of Gas Exit

Temoerature.

kiln has been studied, and dashed lines have been superimposed on Figs. 5*1* to 5-6. showing the average values of each parameter for the Barnstone kiln taken from measurements during the kiln trials. Fig. 5*2. indicates that the Barnstone kiln has a high primary air percentage, this being due to the type of coal milling system used; and this has in turn led to a high velocity ratio and Crasâ-Curtet parameter. (The value of m'at Barnstone was 4.15 which does not appear on Figure 5*4.)

The kiln does have a better than average thermal efficiency and operates with typical excess air and back end. gas temperature values. Thus, although the kiln is not an ideal average operating unit, it may be assumed that any values obtained at Barnstone can be translated with reasonable confidence to other kiln systems.

5.1.2. Heat and mass balances__

Appendix D contains tables of daily average values of mass flows, temperatures, and other operating data calculated from periodic sampling during the course of the 5th .kiln flame trial on April 1973- The sampling methods used are described fully in Section 4. This data provides a basis for the calculation of the heat and mass balances shown in Tables 5.1. to 5*6.Figure 5*7. is a flowsheet of the calculation procedure used for the heat and mass balances (University of Surrey program GSECKCIH)

4-The mass balances for trial 5 agree to within -• 0 .5% • The exit gas mass is calculated from a knowledge of the mass flows and compositions of the fuel and feed combined with- the cooler

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FIGURE 5.7 Flowsheet for heat and mass balances

and inleaked air flow, and the excess air measured as that not required for combustion. This was felt to be a more reliable calculation procedure than the use of the exit gas rate from argon tracing, which proved to be extremely variable. Table 5*7 shows a comparison between the volumetric concentration of the excess oxygen measured, and that calculated in the mass balance (converted to a volumetric basis). •

TABLE 5.7’ Comparison of measured and calculated oxygen concentrations.

Date 02calc. % 02 measured %

3.4.73 4.43 , ' 3.375.4.73 3.53 2.696.4.73 2.34 1.70

These values correspond to mass balance errors of up to+ --3.0%, which is felt to be highly acceptable, considering the number of measuring devices employed and the large and insensitive nature of the plant.

The values of the mass flows, together with measured temperatures and heats of reaction, and a knowledge of the specific heats of the various compounds and elements, has enabled the calculation of overall heat balances to be made, (see Tables 5.4 to 5.6). The balances for 5 - - 7 3 and 6. 4.73 agree to within - 3.0%, whilst that for the 3- * 73 to within -.10%. The large

discrepancy in the latter may be accounted for to some extent by the fact that the unit was started up on that day, and perfect steady state may not have been achieved for a major part of the day. (This is reflected in the lower feed rate compared with other days). Pig. 5*8. is a Sankey diagram showing the average heat flows over the trial periods. It becomes obvious from a study of this diagram that major fuel saving may be brought about by a saving in the exhaust gas enthalpy; this may be achieved by:

(a) Reducing the moisture content of the feed. A 1% reduction of moisture content will produce a 1.1% heat saving in a wet process kiln.

(b) Controlling the excess air rate to a minimum,concommitant with efficient combustion and product quality. A 1% reduction of oxygen in the back endgases (i.e. approximately 5# reduction of excess air)will produce a 0.75$ heat saving in a -wet process kiln.

(c) Savings may also be made by reducing the external surface losses. This is discussed in more detail in Section 5.1*3.

In a dry process kiln, the effect on fuel saving of excess air is doubled and becomes a very significant factor.

5.1.3. External heat loss

The temperature profile of the kiln shell provides an

Fuel

OrganicMatter

Product

External 10.2# Surface Losses

25.9# Heat of ReactionExhaust Gases 62.1#0.7# Dust 2.6# Excess Air

12.6# Combustion Products

46.2# Proces Gases

FIGURE 5,8 Sankey diagram of average heat flow over trial periods

immediate, rough guide to the operation and efficiency of the unit. It indicates areas of thin brickwork, characterized by Thot spots’ which lead to unnecessary heat losses, and also the length of the sintering zone, as defined by the internal coating of clinker on the walls. Fig. 5 -9 . shows the average temperature profile of the kiln during the trial period, with 98# confidence limits on each point.

The peak of highest kiln shell temperature, at approximately 10# of the kiln length from the nose ring, corresponds to the area of maximum flame temperature as shown in Fig.5.17; furthermore, there is a region at about 30# of the kiln length where the shell temperature varies by up to 100°C and after which the shell temperature rises. This region corresponds to the end of the coating, or the start of the sintering zone, the temperature_fluctuations being due to the building up, and breaking away of coating. From the work of Moles, Lain and Shaw62on flame lengths, it can be shown that the end of the flame is also about this point. Thus from measurement of the shell temperature in the sintering region, we can find the area of maximum temperature in the flame, and the length of the flame.

The region from 30 to 60# of the kiln length is the decarbonating zone and the depression in the profile at about 50# is probably due to the decarbonation reaction, which is highly endothermic. The region from 60 to 75# is the preheating zone as mentioned by both Weber and Foliot* , during which the feed is heated to about 800°C, when decarbonation

_ o

o

tn

O0 aanq.UtioduiojjPIGURE 5.9 Average external temperature profile of kiln

shell over trial periods

-paO£0)1—I

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takes place. The increase in shell temperature in the drying zone from 75 to 100$ is due to the unlined kiln shell in the chained section, which, since there is no refractory lining, allows more heat to be lost to the atmosphere.Fig. 5-9* compares favourably with the measurements taken

2 . .by Weber on a larger wet process kiln, the same regions beingdefined, and being of the same equivalent lengths.-

Fig. 5.10. shows the measurements of kiln and coolershell emissivity along their length taken using a Land surfacepyrometer. (See Section 4.3*2. (g) For a description of theapparatus). There is a large scatter of points, due to theuneven nature of the surface, but a regression analysis (shownby the solid black line) was carried out on the data in order

s'sto compare the results with those of Kuhle quoted m Section 3.9*2. (shown by the broken line). Regression analysis led to equation(5•1•)

e = 0.95 “ 4.0 x lCf1* (t - 273) ............... (5.1)

which may be compared to equation(3•48)due to Kuhle

e = O .96 - 5.2 x 10” (t - 373) ................ (3.48)

These equations compare reasonably well, and for dusty oxidized steel shell kilns either may be used to predict the emissivity with a high degree of confidence.

Using the analytical approach outlined in Section 3.9*

1.0

w coCD > 3 <D -P

r—I i— I CdbO cd H (L) £ 13 Q) M K

jJ— o co

oo

Oo<L>U2

-PcdU<Dft6(DEh

iC^TATSSTUia FIGURE

5*10

Variation

of kiln

shell

emissivity

with

temp

erat

ure

the radiative and convective heat losses from the surfaces of the kiln and cooler may be calculated, and Pig. 5*11 shows the relative contributions of each for varing shell temperature. Since the heat loss by radiation is proportional, to the fourth power of the temperature, it is important that the shell temperature is kept as low as possible in the sintering zone. This may be achieved by insulating firebrick, and ensuring a sound, thick coating of clinker on the brick surface. A decrease in the average shell temperature from 400° to 300°C in this region will produce a 3.5% fuel saving, whilst a decrease of 300 C to 200°C will give a fuel saving of 2.5%.

Pig. 5.12 shows that half the heat losses occur within the first third of the kiln length (i.e. the sintering region). This figure has been used later in this section to derive the amount of heat transferred by conduction and convection to the bricks between any two points along the kiln length. This is important in the simulation process described in Section 5.2.

5.1.4. Gas concentration profiles in the flame region

Measurements were made of the active combustion gas species in the burning zone of the kiln. These results were used to study three aspects of flame formation in the kiln.

(a) The length of the flame for various operational conditions, to determine the optimum flame length for efficient combustion.

120C

0

o

SSOI Q.'B0H

FIGURE 5.11 Variation of external heat loss with shell temperature

Temp

erat

ure

_ o

_ o

_ o

_ oin

T-(ssoi 1^03 %) ^20i ^roq oSequoouaj

FIGURE 5.12 Variation of external heat loss with distance from burner nozzle

Distance

along

kiln

from

nozzle

{% Kiln

leng

th)

(b) The shape of a non~isothermal, buoyant flame as compared with an isothermal air model, to verify the applicability of cold modelling.

(c) Determination of the combustion rate within the flame under various conditions, to predict the heat release rates from the flame.

This dissertation is concerned primarily with the third part of the study, the other sections being investigated by other workers 61,6 2

Appendix E contains tables of results of gas analys for three excess air levels, taken at various points in the flame region. The concentrations of carbon dioxide, carbon monoxide and oxygen were determined at each point using the analysis equipment described in Section 4.4.3* The flame length was determined by considering the carbon monoxide extinction curve, and the end of the flame was defined as that point at which combustion is 99*5$ complete, (i.e. CO concentration is 0.05$). The oxygen and carbon dioxide concentration curves will not give a true indication of flame length, due to recirculation, and generation of carbon dioxide from the charge.

Moles, Lain and Shaw62 derived a formula to determine flame length, from the results, which contains the significant operating parameters.

Lf— = 20.02 + 2.89 1 . + 0.39' D 2 -d 2'o - 0.32

fuo*o n-1

V. J . d 2 k o •°a -

Figures 5.13 to 5*15 are gas concentration profiles in the flame zone for different excess air levels, constructed from cross plots of data in Appendix E • All the plots show.the non- axisymmetric nature of the flame jet in the vertical plane, showing an initial downward deflection with a later buoyant lift. Both these effects have been predicted from cold model work by Lain60and Smith61.

The profiles for carbon dioxide show unusually high concentrations near the base of the kiln, this being due to the decarbonization of the charge. The. region at which this decarbonization occurs varies with the excess air, as does the degree. The decarbonization reaction is slow, and the decomposition of calcium carbonate commences at approximately 800°C, and is not complete until a material temperature of 1100-1200°C is attained. It is therefore probable that the high carbon dioxide concentrations indicate the region in which the majority of calcium carbonate is decomposed. The plot Figure 5*13 with the highest excess air rate shows an absence of any significantly high concentrations of carbon dioxide, and, along with the knowledge that a poor quality, underburnt product is produced, it is reasonable to conclude that the calcium carbonate decarbonization reaction does not occur sufficiently fast enough due to a failure to attain an adequate material temperature.

Perc

enta

ge

CO

-p

FIGURE

5‘13

Flame

gas

conc

entr

atio

n profiles

for

17.24$

excess

air

Percentage

C0

2CD

IGURE

5*1

Flame

gas

conc

entr

atio

n profiles

for

8.62%

excess

air

Percentage

C02

-p-p

FIGURE

5*15

Flame

gas

conc

entr

atio

n profiles

for

2.87$

excess

air

The profiles for oxygen extinction are the most accurately indicative of the true flame jet path, although dilution by charge carbon dioxide will distort the correct values. There is an initial decay of oxygen concentration along the jet axis, due to the reaction of the primary air oxygen with the volatiles in the fuel, followed by a steady increase to a constant value, indicating the localised mixing and combustion of the entrained secondary air within the spreading flame jet. This agrees with the observations of I.F.R.F. for a similar confined turbulent diffusion flame system.63 The aerodynamic jet deflection, as postulated by Lain, and the later buoyant jet lift are both shown to exist in the hot jet system. The buoyancy effect is most manifest at the 8.62$ excess air level and this is shown from later experimentation, to be accountable to a higher flame temperature.

The carbon monoxide extinction profiles show a matched relationship to the oxygen profiles during the initial stages of jet development and entrainment. But in the later region of the jet, where recirculation and disentrainment occur, local areas of higher carbon monoxide concentrations are observed near to the kiln charge. These are probably due to a combination of the combustion of both the organic matter in the fuel, and of any oversized, partially burnt, coal particles which have droped out from the slower moving flame jet, onto the surface of the kiln charge. This phenomena becomes more significant, as would be expected, in the case of the lowest excess air level.

5.1.5• Combustion rate analysis

The gas analysis-data in Appendix E has also been used to calculate the rate of combustion within the flame using the method outlined in Section 3*11. The inaccuracies introduced by recirculation are minimal, due to the: insignificant amount of combustion (<10/0 that occurs in the recirculation /region. Figure 5.16 shows the combustion rate- curve calculated from the data of Figure 5.1^ assuming plug flow in the jet. (This assumption has been made in the absence of any suitable mass or velocity flow d a t a ) . I t becomes apparent that approximately 75$ of the combustion takes place within the first 30% of the flame, and that the latter half of the flame is concerned writh the combustion of the oversize coal residue.

Calculations have been made to compare the combustioncurve with data on the burnout rates of volatiles and residue

2 8in coal clouds, using the flow model advanced in Section 3.12. and equation (5*2) to determine the 99*5$ combustion flame length.

From size analysis data on the pulverized fuel used in the flame trials, it was found that 90% of the coal was ground to less than 90 microns. It is thus reasonable to assume that the coal particles remaining at the end of the flame will be the oversize, having a mean particle diameter of approximately .100 microns.

ooCMO5J-o(0OCOoo

-phOc(D i—i(D£cdi—lCm

0 i—1tSlNO£oCm0£cdi—icmbOCOi—1cd0oCcd-pm•HQ

^no uanq FIGURE

5»l6

Fuel

burnout

as a

function

of distance

from

the

burner

nozz

le

From equations (2.20- 21), the diffusion controlled burning time for coal residue is given by

% = ,K V < V .... '••(5.3)

Taking as an example the condition, for 17.24$excess air and 40.0$ primary air* the flame length is given by equation (5*2) as 8.63 metres. Further, using the model in Section 3.12, the time for a particle, travelling at the same speed as the jet gases, to attain a distance of 8.63 metres from the jet nozzle is calculated to be 0.562 secs, (see Appendix G ). Substituting these values into equation 5.3- gives

K F d = 5620

But for the conditions given; from

1+E . F(1+E - V/100)

where F2 E l j [ (E+l)

(5-5)

we can calculate that Fd 4.003. Thus

K = 1404 sec, cm 2 (Dry mineral matter free C = 81.

This value compares well with a value of 1424 sec. cm 2 for a High-Hazel (Durham ) coal residue obtained by Essenhigh.16 (Dry, mineral matter free *C = 81.9$) •

If we consider the volatiles in the system, using the same data, but with a mean particle diameter of 70 microns and a value for K of 130' 16 we can calculate that the combustion time for volatiles i s . t = 0.025 sec. From Appendix G this corresponds to a distance travelled of 17.4$ of the flame length. From Figure 5*17 this corresponds to 46$ combustion (compared with 34$ volatile matter from analysis).

From these calculations it is reasonable to assume that the combustion curve, Figure 5.16 agrees within the limits of experimental error to that which may be predicted from theory.

5.1.6. Temperature measurement in the flame region

From measurements of the temperature'of the flame gases taken by suction pyrometry (See Appendix E) Figure 5*17 has been constructed to show isotherms in the flame region for two levels of excess air, namely 8.62$ and 17.24$. The plots are shown to be axisymmetric, which is not the true case, but since the response time for a sheathed thermocouple in a simple suction pyrometer is of the order of 1 minute, and the rotational speed of the kiln is of the same order of 1 rpm, off axis continuous readings will only provide a mean value around the circle of rotation, and not individual point values. The venturi- pneumatic suction pyrometer used to determine flame, gas

Degrees

Cent

igra

1 00 0

----

rH

rH COo

cv

o

£rH•HxCDCO

-PCO£u■ cd PQ

0X-pUO0£ON50C

•H£U£X0,£

•P

•H

COfi.£0X-poCO•H

0>•H£0Q

,i—IinWoHfH

temperatures near to the burner (up to 3 kiln diameters away from the nozzle plane.) has almost instantaneous response, and since it was not rotating, provided point temperature data. However in the region near to the nozzle, the jet is approximately axisymmetric due to its high excess momentum over its surroundings, which renders the ’geometric’ and ’buoyancy’ forces discussed in Section inconsequential.

Figure 5*1.7 shows the expected result that a higher excess air, which provides a shorter flame, gives rise to an increase in maximum flame temperature and a higher combustion intensity near to the burner. However, for any point downstream of the maximum flame temperature, the bulk gas temperature is lower in the case of the higher excess air.

The maximum flame temperatures measured were higher than expected, since the adiabatic flame temperature was calculated to be of the order of 2300°C. • If the contributionsof charge heating and heat losses in the flame region are taken into account, the mean adiabatic gas bulk temperature is calculated to be of the order of l400°C, which closely corresponds to the end of the flame as given by equation 5*2. and the measured 1*J00°C isotherms in Figure 5*17* Thus we may alternatively define the flame length in terms of the axial temperature profile, as that distance from the nozzle at which the axial temperature is l400°C.

5.2. Prediction of heat transfer in a rotary kiln.

A mathematical model to predict heat transfer in an axially fired kiln tube has been developed in Section 3*This model has been realised in the form of an Algol program, the flowsheet of which is given in Figure 5*18.

!5.2.1. Brief description of program.

Basic data concerning the operation and geometry of the furnace are read into the program, along with values for the calculation of direct exchange coefficients as determined by Einsteink9 . The zoning pattern is specified, and matrices

are set up to accomodate the individual data concerning each zone. Data on the distribution of gas concentrations are read in, and the partial pressures of the radiative gases in each zone are calculated.

For each gas in the grey gas approximation ( see Table 3*1. ) direct exchange coefficients are evaluated between all surfaces and volumes in the zoned region. From these values total exchange factors for the zoned regions are calculated, taking into account reflection and reradiation. At this stage of the program we have complete data concerning the radiative properties of the furnace. Heat balance equations are now set up for each zone to include the effect of convection at surfaces, heat liberation and enthalpy changes in the gas, gas mass flow within each zone, enthalpy changes in the feed, and external heat losses. With this given pattern,

H EA D b a s ic d a ta end

P R IN T in p u t d a ta

C A LC U L A TE z o n in g

‘G A S P R E S ’ c a lc u la t e p a r t ia l p re s s u re s of ra d ia t in g g a s s p e c ie s

R E A D in it ia f va lu e s fo r i t e r a t io n

a r e a /v o lu m e of ro n e s

'E X F A C T ' in te rp o la te

7 C A LL E X F A C T in te g ra l from o p tic a l d is ta n c e an d o p ac ity

C A LC U L A TE d ire c t e x c h a n g e c o e tfs . fo r

'B U M F ' p r in t ou t of d a ta ra n g e re p o rt

'F 0 3 A A A ' (N A G lib ra ry ) e v a lu a te d e te r m in a n t

C A LL F 0 3 A A A

C A L C U L A T E re sp o n se

o f m a tr ix d e term in an ts

C A LC U LA TE t o te l ex c h a n g e c o o f fs . fo r e l l to n e s

g re y gas In re a l g a s f i t ?

*H EA TO U T* c a lc u la te ra d ia t io n h e a t t ra n s fe r b e tw e e n

zones for g iv en te m p e ra tu re f ie ld

C A LL H E A T O U T

'M A S S FLO W ' c a lc u la te

m ass of g a s in zoneC A LL M A S S F L O W

'S E N S H E A T* c a lc u la t e h e a t lo a d per zone fro m g as c o n v e c tio n , c h e m ic a l r e a c t io n , en th a lp y c h a n g e and

C A L L S E N S H E A T

i t e r a t io n l im it

u t in Q te m p e ra tu re s d e r iv e d fro m prev io u s ite r a t io n

A D J U S T te m p e ra tu re s of u n c o n tro lle d zon es

C o n v e rg e n c e

c o n v e rg e n c e re p o rtP R IN T t e m p e ra tu re )

end t e a t f lu x e s

FIGURE 5*18 Flowsheet for simulation of heat transfer in a rotarv kiln

iteration to a temperature field which will provide a satisfactory heat balance is carried out, and the final temperature field and heat fluxes to the surfaces reported on a line printer.

A complete account of the mechanics of the program i developed ( University of Surrey program GSECKC1Z.) ;is given in Appendix H , together with a listing. The approximate storage requirements for the program are (19*6 + O.Ol4n2) K, where n is the total number of zones in the system.

9.2.2. Model assumptions.

(a) No entrainment of secondary fluid occurs within three nozzle diameters of the nozzle plane.

(b) Plug flow is assumed in both the jet and secondary fluid.

(c) The jet expands along the axis at a constant jet half angle of 11D- until it strikes the tube wall.

(d) Entrainment into the jet occurs until the jet has reached a width of two thirds of the kiln diameter. Disentrainment from the jet is then assumed until the jet strikes the tube wall ( assumption (c) ), at which point plug flow over the whole tube is assumed.

(e) Recirculation may or may not occur, but sufficient secondary fluid is always available for jet entrainment. Any recirculated fluid is supplied from the last third of the jet.

(f) All flows are assumed to be in the direction of jet flow.

(g) The three term grey gas approximation assumes that the ratio of partial pressures of water vapour to carbon dioxide is one in all gas zones.

(h) Carbon monoxide is assumed as equivalent carbon dioxide for the purpose of evaluating gas radiation.

(i) Combustion proceeds at a uniform rate at any plane in the jet, parallel to the nozzle plane.

5.2.3« Model testing data.

The model was tested against the measured results of theBarnstone trials using the data for an 8.62% excess air flame.Table 5*8. summarizes the input conditions.

TABLE 5*8. Measured input data used in model testing.

Fuel rate 14.98 kg/min.Calorific value of fuel 7417 kcal/kg.Air requirement of fuel 8.998 kg/kg.Excess air percentage 8.62 %

Primary air percentage 40.0 %

Primary air temperature 100 0 CSecondary air temperature 700 °CAmbient air temperature 5 °cKiln shell temperature 250 °cProduct exit temperature 1400 °cProduct flowrate 70.0 kg/min.Kiln diameter 1.668 m.Nozzle diameter 0.1952 m. '

From equation (5.2), the length of flame produced using the data in Table 5*8 was calculated as 11.1 metres, and model testing was carried out over this length of the kiln.

The system was zoned using a characteristic zoning dimention-of 0.278 meters. This gave three annular rings and forty,longitudinal divisions. From this grid, using Figure 5.1' * the mean volumetric concentrations of carbon monoxide, carbon dioxide and oxygen within each zone were calculated. These values are listed in Table 5*9. Using Figure 5*16, and assuming a jet half angle of 11°, the rate of heat release due to fuel combustion was calculated for each zone. These values arc listed in Table 5.9*

Dust concentrations within the flame zone of the kiln have been measured by Weber2 . Figure 5*19 is a replot of the data of Weber, showing the variation of dust burden and gas velocity within the kiln. From this data and a^knowledge of the kiln diameter ( 3*5 meters ) and production rate ( 30 tonnes/hour ) of Weber’s kiln, Figure 5.20 has been developed to show the variation of dust concentration( kg./m? with kiln length for the region under investigation. This data may then be used directly in equation (3.19)

-Uc ,Led = 1 - e a ............. (3.19)

to incorporate the emissive contribution of dust to the radiative characteristics of the gas zones. Table 5*9 lists the values of c^ used for each zone in the model. The value of U used is that calculated in Section 3*5 as 7*01 m?/kg.

( rs/*ui ) Xq.iD0X3A

o00

o

co

V

CM

CO Oin

( AOûixo *8q/*3>[ ) uoiXTmiuoouoo qsnp oiuoquiv

FIGURE 5.19 Dust burden and gas velocity variation in a rotary cement kiln

Distance

along

kiln

from

nozzle

( %

kiln

length

)

Airborne

dust

conc

entr

atio

n ( kg

./m?

t— \

A.

5030 4020100Distance along kiln from nozzle ( % kiln length )

FIGURE 5*20 Variation of dust concentration with distance from the burner nozzle '

TABLE 5•9• 6aszone input data.

Gas concentrations in.volume fraction

Heat rate in Kcal./min.Dust concentration in kg./m!Soot concentration in kg./m!

Legend for method of zone identification.

Zone No. 1 2 3 4 5 6 7 8 9 10 11 12 et<

Axis of Symmetry

Ring p No.

IN v-/ • _L- •

Zone Gas Cone. Heat Dust SootNo. CO C02 02 Rate Cone. Cone.1 0.0005 0.0100 0.1500 8881 0.0279 0.000012 0.0015 0.0300 0.0600 9341 0.0331 0.000073 0.0020 0.0350 0.0500 9452 0.0418 0.000254 o .0050 0.0400' 0.0400 9230 0.0756 0.000625 0.0150 0.0600 0.0450 9341 0.0959 0.000776 0.0200 0.0700 0.0400 9674 0.1073 0.000877 0.0250 0.0800 0.0300 7784 0.1129 0.000698 0.0400 0.0900 0.0350 6005 0.1150 0.000209 0.0420 0.1000 0.0350 4559 0.1100 0.00005

10 0.0450 0.1100 0.0350 3892 0.1000 0.0000111 0.0480 0.1100 0.0350 1501 0.0992 012 0.0450 0.1150 0.0350 1334 0.0992 013 0.0420 0.1150 0.0350 1390 0.0992 014 0.0400 0.1200 0.0350 1168 0.0992 015 0.0350 0.1250 0.0350 1056 0.0992 016 0.0320 0.1280 0.0350 945 0.0992 017 0.0280 0.1350 0.0350 612 0.0992 018 0.0220 0.1500 0.0350 66 7 0.0992 019 0.0180 0.1650 0.0350 556 0.0992 020 0.0150 0.1800 0.0350 500 0.0992 021 0.0130 0.1900 0.0320 445 0.0992 022 0.0120 0.2000 0.0320 334 0.0992 023 0.0110 0.2020 0.0320 388 0.0992 024 0.0100 0.2050 0.0320 445 0.0992 025 0.0092 0.2050 0.0320 445 0.0992 026 0.0087 0.2080 0.0320 297 0.0992 027 0.0080 0.2080 0.0320 260 O.O992 028 0.0075 0.2100 0.0320 297 0.0992 029 0.0070 0.2100 0.0300 260 0.0992 030 0.0065 0.2120 0.0300 260 0.0992 031 0.0060 0.2150 0.0300 260 0.0992 032 0.0055 0.2150 0.0300 222 0.0992 033 0.0050 0.2180 0.0250 185 0.0992 034 0.0045 0.2180 0.0250 185 0.0992 035 0.0040 0.2200 0.0250 185 0.0992 036 0.0057 0.2200 0.0250 185 0.0992 037 0.0033 0.2200 0.0200 185 0.0992 038 0.0030 0.2200 0.0200 185 0.0992 039 0.0028 0.2200 0.0200 222 0.0992 040 0.0026 0.2200 0.0200 185 0.0992 0

ZoneNo.

Gas Cone. HeatRate

Dust Cone.

Soot Cone.CO CO2 0 2

1 0 0.0050 0.2000 5659 0.0065 02 0.0005 0.0100 0.1800 0 0.0129 03 0.0010 0.0150 0.1200 0 0.0216 04 0.0015 0.0180 0.1000 0 0.0496 0.000105 0.0015 0.0200 0.0900 0 0.0712 0.000256 0.0018 0.0300 0.0800 0 0.0863 0.000337 0.0019 0.0400 0.0800 0 0.0949 0.000158 0.0030 0.0400 0.0700 0 0.0992 09 0.0030 0.0400 0.0700 0 0.0992 0

10 0.0050 0.0400 0.0700 0 0.0992 011 0.0050 0.0400 0.0650 1501 0.0992 012 0.0050 0.0450 0.0620 1334 0.0992 013 0.0048 0.0480 0.0600 1390 0.0992 014 0.0045 0.0500 0.0600 1168 0.0992 015 0.0035 0.0550 0.0580 1056 0.0992 016 0.0032 0.0650 0.0550 9^5 0.0992 017 0.0032 0.0750 0.0550 612 0.0992 018 0.0032 0.0820 0.0520 667 0.0992 019 0.0032 0.0900 0.0520 556 0.0992 020 0.0030 0.1000 0.0500 500 0.0992 021 0.0030 0.1100 0.0500 445 0.0992 022 0.0029 0.1200 0.0500 334 0.0992 023 0.0029 0.1280 0.0480 388 0.0992 024 0.0028 0.1350 0.0480 445 0.0992 025 0.0028 0.1420 0.0480 445 0.0992 026 0.0027 0.1500 0.0450 297 0.0992 027 0.0027 0.1550 0.0450 260 0.0992 028 0.0026 0.1600 0.0420 297 0.0992 029 0.0026 0.1620 0.0420 260 0.0992 030 0.0025 0.1650 0.0420 260 O.O992 031 0.0025 0.1680 0.0400- 260 0.0992 032 0.0024 0.1700 0.0400 222 0.0992 033 0.0024 0.1700 0.0400 185 0.0992 034 0.0023 0,1720 0.0380 185 0.0992 035 0.0023 0.1720 0.0380 185 0.0992 05.6 0.0022 0.1750 0.0350 185 0.0992 037 0.0022 0.1750 0.0350 185 0.0992 038 0.0021 0.1780 0.0320 185 0.0992 039 0.0021 0.1780 0.0320 222 0.0992 040 0.0020 0.1800 0.0320 185 0.0992 0

ZoneNo.

Gas Cone. HeatRate

Dust Cone.

Soot Cone.CO- c o 2 0 2

1 0 0.0001 0.2100 11318 0.0065 02 0 0.0050 0 .2C00 0 0.0129 03 0 0.0060 0.2000 0 0.0216 04 0 0.0080 0.1700 0 0.0496 05 0 0.0100 0.1600 0 0.0712 06 0 0.0100 0.1600 0 O.O863 07 0 0.0150 0.1500 0 0.0949 08 0 0.0180 0.1500 0 0.0992 09 0 0.0180 0.1500 0 0.0992 0

10 0.0010 0.0200 0.1450 0 0.0992 011 0.0010 0.0200 0.1420 0 0.0992 012 0.0010 0.0200 0.1400 0 0.0992 013 0.0010 0.0200 0.1370 0 0.0992 014 0.0010 0.0200 0.1350 0 0.0992 015 0.0010 0.0220 0.1320 0 0.0992 016 0.0010 0.0250 0.1250 0 0.0992 017 0.0010 0.0280 0.1200 0 0.0992 018 0 0.0400 0.1150 0 0.0992 019 0 0.0500 0.1100 0 0.0992 020 0 0.0600 0.1020 0 0.0992 021 0 0.0700 0.0950 0 0.0992 022 0 0.0850 0.0850 0 0.0992 023 ' 0 0.1000 0.0770 0 0.0992 024 0 0.1100 0.0700 0 0.0992 025 0 0.1200 0.0650 0 0.0992 026 0 0.1250 0.0600 297 0.0992 027 .0 0.1280 0.0580 260 0.0992 028 0 0.1300 0.0550 297 0.0992 029 0 0.1320 0.0550 260 0.0992 030 0 0.1320 0.0520 260 0.0992 031 0 0.1320 0.0520 260 0.0992 032 0 0.1350 0.0500 222 0.0992 033 0 0.1350 0.0500 f—j CO 0.0992 034 0 0.1350 0.0500 185 0.0992 035 0 0.1350 0.0480 185 0.0992 036 0 0.1370 0.0480 185 0.0992 037 0 0.1370 0.0450 185 0.0992 038 0 0.1370 0.0450 185 0.0992 039 0 0.1400 0.0420 222 0.0992 040 0 0,1400 0.0420 185 0.0992 0

No data is currently available for the concentration of soot in a pulverised fuel flame, but it is reasonable to assume that its contribution to the overall emissivity of the flame will be negligable when compared to the dust and pulverised fuel contributions, which raise the flame emissivity to between 0.8 and 0.9* Nominal values have been included in the test data, to.ensure that this facility is operational in the program, but variation of these values during program testing proved to have no significant effect on the final results.

Folliot 1 investigated the variation of emissivity of the charge material as it is heated from wet feed through to dead burnt clinker. He observed three distinct stages which corresponded with the regions observed in the results from Figure 5*9> these being the drying, decarbonating and sintering stages of the process. Table 5*10 summarises Folliot1 s results.

TABLE 5*10* Emissivity of kiln material.

Region EmissivityDrying/Preheating 0.80Decarbonating 0.70Sintering 0.88

In this study we are concerned with the sintering region of the kiln, where the kiln walls are covered with a thick, permanent coating of clinker, thus we may assume that the emissivity ofall the internal surfaces of the kiln is 0.88.

The end of the kiln, which is effectively the kiln hood, is constructed of a high quality firebrick, which,from manufacturers specifications, has a mean emissivity of 0 .72.

From the data in Table 5.8, the enthalpic contribution of the preheated primary and secondary air may be calculated. Assuming plug flow in the primary and secondary air streams at the nozzle plane, the enthalpy increase due to preheating for each ring is given in Table 5.11.

TABLE 5.11. Enthalpy of preheated air.

Ring No. Air Heat input rate kcal./min.

1 Primary & Secondary 23262 Secondary 56593 Secondary 11318

Total - 19303

The calculated values in Table 5*11 have been included in the value of the heat release rate for each ring in zone one in Table 5.9.

A similar enthalpic contribution from the charge material is made in wall zone number forty. It is assumed that the charge is at a bulk temperature of 900°C as it enters the region of the system under investigation.Folliot 1 has measured the specific heat of -cement clinker as 0.26 kcal./kg.°C, thus the heat input rate due to preheated charge is 16290 kcal./inin.

5.2.4, Convective and Conductive heat transfer coefficients.

Using equation (3*41), an average value for the convective heat transfer coefficient, assuming air at 1000°C, has been evaluated and incorporated into the simulation process.

Nu = .42 Re*6 Pr*33 -.(3.41)

From the data in Table 5*8, and standard values for the properties of air at 1000°C, normal kiln operation at 8.62% excess air gives

Re = 44630Pr = 0.72k = O.O69 kcal. /hr .in? °C/m.

thus

h „ = 9*6 kcal./hr.m?°Cconv

From Figure 2.2, the average internal wall surface temperature has been evaluated as 1150°C. From the data on the Barnstone kiln in Figure 5* 9* the average outside kiln shell temperature was 200°C, and Figure 5*12 gives a heat loss of 3150 kcal./hr.m? at this temperature. Thus the average conductive heat transfer coefficient through

the furnace walls- is given by

= 3.3 kcal./hr .m? °Ccond

5.2.5. Results of the computational simulation.

Using the data in Tables 5*8 to 5.11* together with other values quoted in Section 5*2, the temperatures for each zone, and the heat flux values to the wall zones were predicted using the mathematical model outlined in section 5.2.1. The results of the simulation are listed in Table 5.12. The convergence level was set at ±5% agreement with the preceeding value on each individual zonal heat balance, which is equivalent to ±10°C in most cases. Convergence was achieved on all zones within forty iterations.

Temperature profiles have been constructed from these results, and they are compared with measured values in Figure 5*21. Comparison of these gas temperature results shows reasonable agreement along the axis of the kiln, but at radial distances of greater than 0.5 meters in the region near the nozzle plane some discrepancies are apparent. This is probably attributable to the inability of the model to account for recirculating flows which would be expected to increase the temperature of the gases in this region.

Figure 5.22 compares predicted and measured values of the wall temperature in the flame region. These again exhibit a good agreement ( ±10$ ) in all areas except that near the nozzle plane. This discrepancy may be accounted for by the lower predicted gas temperatures in this region, giving a reduced rate of heat transfer to the wall. Figure 5*23 shows the predicted wall heat flux variation in the flame

°c ! °c---

°c °C KW/m21 751 702 700 676 25.722 823 710 710 709 24.663 938 724 710 717 28.924 1055 736 709 785 29.575 1208 742 714 821 44.196 1410 738 719 923 45.367 1583 742 725 1087 112.328 1722 758 736 1119 112.769 1819 764 73 6 1351 111.87

10 1872 781 742 1626 114.2211 1908 781 744 1709 117.0412 1922 799 - 751 1792 120.7513 1922 803 755 1806 124.0814 1917 844 755 1787 125-6415 1904 897 760 1751 126.0616 1892 1080 768 1690 129.2317 1872 1243 772 1558 127.1918 1857 1351 779 1460 118.3819 1844 1410 781 1355 122.9020 1814 1468 797 1301 125.^921 1798 1515 821 1208 128.7422 1777 1536 866 1122 116.7223 1754 1547 923 1058 107.0424 1729 1547 1002 1008 102.0325 1705 1536 1081 .937 103.6626 1662 1524 1128 940 100.7727 1618 1500 1172 928 70.3028 1569 1476 1198 917 64.7929 1526 1452 1216 892 62.4530 1491 1438 1222 888 55.6831 1465 1413 1222 887 54.5432 1446 1401 1222 889 54.7633 1429 1380 1217 862 53.4034 1414 1365 1206 857 51.9135 1402 1361 1199 855 51.7036 1395 1340 1189 861 52.2637 1387 1328 1182 850 51.7738 1372 1311 1168 846 51.9139 1361 1307 1157 836 51.3940 1350 1298 1152 1—

i -=

r

CO 50.88

TABLE

5*12

Temp

erat

ure

and

heat

flux

values

from

zonal

simu

lati

on

Isotherms

in °C

xlOO

Pk

cvi

CM

coi

00,

FIGURE

5i21

Predicted

and

measured

gas

temp

erat

ure

profiles

for

100

tpd

cement

kiln

.

MEAS

URED

CM

0

FIGURE

5.22

Predicted

and

measured

wall

temp

erat

ure

profiles

for

100

tpd

cement

ki

o o o o

0 I—I Cd tslo£OUCm£

r H• H

hO£Oi—Icd

0o£cd

‘-p02• HQ

jW/MM xnu QB0H

FIGURE

5.23

Predicted

wall

heat

flux

for

100

tpd

cement

kiln

region. The total heat transferred to the walls in this region is 4.842 MW. ; this is equivalent to 53*3$ of the total energy input to the system ( i.e. fuel plus preheated air ). There is a region of maximum flux density between 2 and 6 meters from the nozzle plane where the average flux is 117*^ KW./m?^ In this region the wall temperature passes through a cycle of 700°C from 1100°C to a maximum of l800°C and back to 1100°C.

5.3. Discussion of predicted and experimental results.

5«3'1» Thermodynamic processes.

From the results of Section 5»l»5a it is apparent that the heat release pattern from the flame is substantially the same for any excess air level within the interval of the measured range ( 8 - 18% excess air ). The curve generated in Figure 5*16 can be divided into three distinct sections: A constant combustion rate section for the first 30$ of the flame, releasing approximately 60% of the total combustion energy. This may be accounted to the steady, chemically controlled, combustion of the fuel volatiles and some residual carbon in the presence of an excess of combustion air, supplied by the premixed primary air and some entrained secondary air. A falling combustion rate section from approximately 30% to 50% of the flame length during which a further 25% of the total combustion energy is released. This is accounted to the burning of residual carbon in the presence of a decreasing air supply, concurrent with the diluting effect of entrained recirculating

combustion products. The tail of the flame is a slow, steady, mixing controlled combustion, releasing the remaining combustion energy, and may be attributed to the burnout of char and oversize particles in the presence of a small amount of excess air.

It would be anticipated that the effect of a reduction in the primary air percentage, and hence jet momentum ( assuming a fixed nozzle velocity ), would be to reduce the length of the initial constant rate combustion section due to a reduced volume of premixed air, and to elongate the falling combustion rate section, due to the increased entrainment requirement’ of the j et and its reduced turbulent intensity. There would, however, be a reduction in the dilution effect of combustion products due to a lower recirculation ratio.

From the good agreement between the measured values, and the equations for the combustion of turbulent diffusion, pulverised fuel flames ( See Sections 2.1.3 and 5»1«5 )>

it would appear that it is possible to predict the shape of Figure 5-16 for varying excess and primary air, coal fineness and fuel volatiles; and to calculate the flame length. Some work along these lines has been carried out at Sheffield University6 ,f on the combustion of polysize, coal clouds, but the difficulty of accounting for the recirculating combustion products tended to predict flame lengths that were shorter, than the measured values.

The heat release pattern given by Figure 5.16, provides a good correlation between the measured and computed values of the gas and wall temperatures, and the heat flux from the flame shows a good correlation with the calculated heat flux to the walls. The measured shell temperature profile in Figure 5*9 exhibits a peak in the same region as the predicted maximum heat flux to the walls in Figure 5.23. Figure 5.21 shows that the measured and predicted l400°C isotherms lie within 10% of each other on the'flame axis. Thus the mathematical model may be expected to give a reasonable estimate of the flame length, using the definition proposed in Section 5*1.6.

From a consideration of the measured values of product quality related to the excess air level in the trials, discussed in Section 5.1.4, it would appear that the heat flux to the walls must be distributed over a reasonable length of the sintering zone at a sufficiently high rate, since an average flux of 117.4 KW/m. over *4 metres ( 8.62% excess air ), will give a better quality product than either a heat flux of 156.4KW/m. over 3 metres ( 17.24% excess air ), or 78.6 KW/m. over 6 metres ( 2 .87% excess air ). .

Of the assumptions made in the mathematical model, listed in Section 5.2.2, those regarding the even combustion within the flame, (i), and the grey gas approximation, (g) and (h), appear to be justified by the agreement between the measured and calculated gas isotherms in the flame envelope.

" - The values of the convective and conductive heat transfer coefficients used in the simulation appear to give a good agreement to the measured and calculated wall temperature, but it must be noted that in program testing it was observed that the value of the wall temperature was significantly influenced by large changes ( ±100% ) in the convective heat transfer coefficent.

5.3.2. Aerodynamic processes.

Of the assumptions listed in Section 5.2.2, concerning the aerodynamics of the system, only assumption (f) appears dubious. The inability of the model to account for the physical reversal of flow of recirculating combustion products has produced lower than measured values for the gas temperature outside the flame envelope. In the analysis of the kiln data in Section 5.1.1* Figure 5*4 shows that in most kiln systems the recirculation ratio is much lower than that in the Barnstone kiln, thus the effect of the inaccuacies introduced by assumption (f) will be less, and it was felt that' the problems involved in reprogramming did not justify the result at the present time.

An overall examination of the computed results indicates that a coarser zone grid may be used in the longitudinal direction with little loss of accuracy, and with a considerable saving in computing time and storage requirement.

SECTION 6.

CONCLUSIONS AND RECOMMENDATIONS FOR.FURTHER WORK

1. Accurate heat balances on the kiln system indicate that there are substantial advantages to be gained by reducing the heat lost in the exit gases and from the kiln shell.

(a), A reduction of 1% moisture in the feed will produce a 1.1% heat saving.

(b). A reduction of approximately 5% excess air will produce a 0.75% heat saving in a wet process system, and a 1.5% heat saving in a dry process system.

(c). A reduction of the average shell temperature of the kiln over the sintering zone of 300°C to 200°C will produce a 2.5% heat saving.

2. The external temperature profile of the kiln indicates the position and length of four regions of feed processing, i.e. drying,preheating,calcining, and sintering. The lengthof the sintering zone as indicated by the external temperature profile corresponds closely with the length of the internal coating of the kiln, which is a good indication of the length of the flame.

3. Gas concentration profiles in the burning zone of the kiln indicate that the phenomena of jet deflection and buoyancy observed in cold model tests are present in the non-isothermal situation. Such profiles may be used to study the mixing of the air and the fuel in the flame jet, and to observe the

effect of excess air on the rate and position of decarbonization in the feed,

4. Measured temperature profiles indicate that average flame temperatures of approximately 1800°C are usual in cement kilns. Hotter, shorter flames are produced with higher excess air levels, but the resulting lower combustion gas temperature leads to a poor quality product due to insufficient calcination.

5. Computational simulation of the sintering region of the kiln has produced comparable results to the measured temperature profiles, and has indicated that the conditions for optimum kiln working ( excess air level-of 17.24$ ) are related to an area, approximately 4 metres long, where the heat flux to the walls is about 120 kW/m?

The areas where further investigation would help to advance the knowledge gained in this dissertation are

(i). A practical study of the emmisive properties of the flame, to elucidate the relative contributions of dust and soot to the heat transfer in the sintering zone.

(ii). The development of a mathematical model to predict the burning rate of different fuels and fuel properties under the conditions of varying excess air, recirculation, primary air and excess jet momentum ratio.

(iii).The development or refinement of the semi-empirical model used to define the aerodynamics of the system,in order that the mixing and gas concentration profiles may be predicted, for changes in the major process

variables.

REFERENCES

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WEBER P.Heat transfer in rotary kilns with due regard to cyclic processes and phase formation.Zement-Kalk-Gips. Spec. English Ed., ( Wiesbaden, Bauverlag GmbH.) 1963.

RUHLAND W.Uber die lange von kohlenstaubflammen in drehofen schriffenreihe zer zementindustrie.Heft, 32, C Dusseldorf, Betonverlag GmbH.) 1965.

HUBBARD E.H.Comparison between turbulent jet diffusion flames of gaseous, liquid and pulverised fuels.Sheffld. Univ. Fuel Soc. Jnl.,'9* 63 to 80, 1958.

SCHLICHTING H.Boundary layer theory.( McGraw Hill : New York.) -1955.

PAI S.I.Fluid dynamics of jets.( Van Nostrand : New York,). 1954.

DONALD M.B. and SINGER H.Entrainment in turbulent fluid jets.Trans. Inst. Chem. Eng., 37* 255 to 264, 1959.

PAIZIS S.T. and SCHWARTZ W.H.Entrainment rates in turbulent shear flows. Jnl. Fluid Mech., 68, 297 to 308, 1975.

9. RICOU F.P. and SPALDING D.B.Measurement of entrainment by axisymmetric turbulent jets. Jnl.' Fluid Mech., 11, 21 to 32, 1961,

10. WATSON E ,A , and CLARKE J.S.Combustion and combustion equipment for aero gas turbines. Jnl. Inst. Fuel, 21, 1 to 34, 1947.

11. THRING M.W. and NEWBY M.P.Combustion length of enclosed turbulent jet flames.4th. Symp. (Int.) on Combustion, 789 to 796*( Williams ana Wilkins Co. Inc. : Baltimore.) 1953.

12. BARCHILLON M.W. and CURTET R.Some details of the structure of an axisymmetric confined jet with backflow.Jnl. Basic Engng., 860, 777 to , 1964.

13. CRAYA A. and CURTET R.

C. R. Acad. Sci. , 24l, 621 to 622, 1955.

14. BECKER H.A .Concentration fluctuations in ducted jet mixing.Sc.D. Thesis, ( Mass. Inst. Tech., Cambridge.) 1961.

15. NUSSELT W.

Z. Ver. dt. Ing., 68, 124 to 128, 1924.

16. ESSENHIGH R.H.Predicted burning times of solid particles in an idealised dust flame.2nd. Pulverized Fuel Conf., B.I., Paper 2,(Inst. Fuel.: London.) 1957. alsoJnl. Inst. Fuel, 34, 239 to 244, 1961.

17. HOTTEL H.C. and STEWART I, McC,Space requirement for the combustion of pulverized coal. Ind. ErigngY Chem., 32, 719 to 730, 1940.

18. FIELD M.A., GILL D.W., MORGAN B.B., and HAWKSLEY P.G.W. Combustion of Pulverised Coal, 216 to 220,( Br. Coal Utilisation Res. Assoc. ; Leatherhead, Surrey.) 1967.

19. McKENZIE A., SMITH I.W. and SZPINDLER G.a Donau. Calculation of burnout of polydisperse suspensions of low and high reactivity pulverized fuels.4th. Symp. on Flames in Industry, Paper 13.( Inst. Fuel : London.) 1972.

20. SMITH I.W.Studies of the combustion rate and stucture of pulverized fuels burned in the C.15. trials ,.(1969) of the I.F.R.F. Div. Min. Chem. Invest. Rept., No. 86, CSIRO, 1970.

21. HEDLEY A.B. and JACKSON E.W.A simplified mathematical model of a pulverised coal flame showing the effect of recirculation on combustion rate.Jnl. Inst. Fuel, 39* 208 to 218, 1966.

22. MARTIN G.Chemical engineering and thermodynamics applied to the cement rotary kiln.( The Technical Press : London.) 1932.

23. GIBBS R.Controlling factors for efficient rotary kiln operation. Rock Products, 45, 58,86 to 90, 1942.

24. ECKEL E.C.Cements, Limes and Plasters. 2nd. Ed.,( John Wiley and Sons ; New York.) 1922.

25. KANNERWURF A,S. and CLAUSEN C.F,What do you know about cement kiln?Rock Products, 65, ' 128,130,132,136,138,142,144, 1962..

26. YOSHII T.On the capacity of the cement rotary kiln.Jnl. Soc. Chem. Ind. Japan, 4l, 330B to 334b , 1938.

27. GROUME-GRJIMAILO W.E.The flow of gases in furnaces.( John Wiley and Sons : New York.) 1923.

28. MOLES P.D., WATSON D. and LAIN P.B.The aerodynamics of the rotary cement kiln.4th . Symp. on Flames in-'Industry, Paper 6.(Inst. Fuel : London.) 1972. alsoJnl. Inst. Fuel, 46, 353 to 362, 1973.

29. BURKE E. and FIELD G.Problems of heat transfer from flames in the cement industry.One day Symp. on Flames.in Industry, Paper 6, B15 to B17. (Inst. Fuel ; London.)

30. TRAUSTEL S. and RUHLAND W.The objectives and results of flame research for the construction of industrial furnaces,Tonind. Ztg., 86, 564 to , 1962.

31. GILBERT W.Investigations on a slurry drier or calcinator.Cement and Cement Manf., 9* (6)115 to 128,(7)139 to 154, (10)207 to 220, 1936.

32. GYGI H.Thermodynamics of the cement kiln.Proc. 3rd. Symp. (Int.) Chemistry of Cement,( London.) 1952.

33. YOSHII T.On burning mechanism in the rotary kiln.Rock Products, 56, 106 to 108,129, 1953.

34. PIKE R.D.Fuel economy in the rotary kiln burning Portland cement clinker.Ind. Engng. Chem., 21,307 to 310, 1929.

35. LYONS J.W., MIN U.S., PARISOT P.E. and PAUL J.E. Experimentation with a wet process rotary cement kiln via the analog computer.Ind. Engng. Chem. Process Design and Development, 1,29 to 33, 1962.

36. IMBER M. and PASCHKIS V.A new theory for a rotary kiln heat exchanger.Int. Jnl. Heat and Mass Transfer, 5, 623 to 638, 1962.

37. VALLANT A.Thermal analysis of the direct fired rotary kiln.Ph.D. Thesis, ( Columbia Univ. : New York.) 1965.

38. BOWERS T.G. and READ H.L,Heat transfer in rotary kilns.Chem. Engng Progress. Symp. Series, ( Heat Transfer : Boston.) 61, 340 to 345,

39. SEIDEL G.Einfluss und gestaltung von drehofenfeuerungen auf verhrennungs und warmeaustauschvorgange in zementdrehofen, Silikattechnik, 19, 1 to 5, 1968.

40. SCHWARZKOPF F.Heat transfer in rotary lime kilns.Pit and Quarry, , 128 to 131,139, 1969.

41. ROSA J.Mathematische formulierung des warmeuberganges im drehofen. Zement-Kalk-Gips, 23, 368 to 377, 1970

42. MANITUS A., KURCYUSZ E. and KAWECKI W.Mathematical model of the aluminium oxide rotary kiln. Ind. ErignK. Chem. Process Design.'and DeveTpoment, 13,132 to 142, 1974.

43. LOBO W.E. and EVANS J.E.

Trans. Am.' Inst. Chem. Eng., 35, 743 to , 1939*

44. HOTTEL H.C. and SAROPIM A.F.Radiative transfer.(McGraw Hill : New York.) 1967.

45. SIDDALL R.G.Flux methods for analysis of radiant heat transfer.4th. Symp. on Flames in Industry, Paper 16.( Inst. Fuel : London.) 1972.

46. STEWARD F.R. and CANNON R.The calculation of radiative heat flux in a cylindrical furnace using the Monte Carlo method,Int. Jnl. Heat and Mass Transfer, 14, 245 to 262, 1971.

47. GOSMAN A.D., PUN W.M., RUNCHAL A.K., SPALDING D.B. and WOLFSHTEIN M,Heat and mass transfer in recirculating flows.( Academic Press : London.) 1969.

48. ERKKU H.Radiant heat transfer in gas filled filled slabs and cylinders. .Sc.D. Thesis, ( Mass. Inst. Tech., Cambridge.) 1959.

49. EINSTEIN T.H.Radiant heat transfer to absorbing gases enclosed in a circular pipe with conduction, gas flow and internal heat generation.Tech. Rept., R-156, ( N.A.S.A. : Cleveland.) 1962.

a50. HADVIG S.

Gas emissivity and absorptivity, a thermodynamic study. Jnl'. Inst. Fuels 43, 129 to 155, 1970.

51. JOHNSON-T.R. 'and BEER J.M.The zone method of analysis of radiant heat transfer: a model for luminous radiation.4th. Symp. on Flames in Industry, Paper 4.( Inst. Fuel .- London.) 1972.

52. HOWARTH C.R. and FOSTER. P.J.Optical constants of carbons and coals in the infra-red. Carbon, 6, 719 to 729, 1968.-

53. THRING M.W.Luminous radiation from flames.Chem. Process Engng., 46, 544 to 551, 1965.

54. ANON.Private communication.Assoc. Portland Cement Manf. Ltd.- Engng. R. & D. Dept. 1974.

55. CHEN D.D.C. and McGRATH I.A.Convective heat transfer in chemically reacting systems. Jnl. Inst. Fuel, 42, 12 to 17, 1969.

56. ALTMAN D. and WISE H.Effect of chemical reaction in the boundary layer on convective heat transfer.Jet Propulfeion, 26, 256 to , 1956.

57. KRAUSOLD H. •

Forsch. Geblete Ingenieurw., 5, 186 to ,1934.

58. KUIiLE W.Untersuchungen uber die au$ere warmeabgabe von drehofen durch strahlung und konvektion.Zement-Kalk-Gips, 23, 263 to £68, 1970.

59. McADAMS W.H,Heat' transmis s ion. 3rd. Ed,( McGraw Hill : New York.) 1954.

60. LAIN P.B.Plow patterns in rotary cement kiln models.Ph.D. Thesis, ( Univ. Surrey, Guildford.) 1972.

61. SMITH R.-M.The effect of buoyancy in enclosed turbulent flames.Ph.D. Thesis, ( Univ. Surrey, Guildford.) 1976.

62. MOLES F.D., LAIN P.B. and SHAW J.F.G.The determination of flame lengths and firing pipe

. diameters in rotary cement kilns.Rept. CE-RD-32,( Univ. Surrey : Guildford.) 1974.

63. THRING M.W.The science of flames and furnaces. 2nd. Ed.( Chapman and Hall : London.) 1962.

64. HEDLEY I.Private communication.Ph.D. Thesis, ( Univ. Sheffield, Sheffield.) To be published.

aNOMENCLATURE

a - weighting factorA - areaB - air requirement of primary j etc - concentrationC - specific heatCp - " "d - diameterD - kiln diameter(2), characteristic dimention(3)t) - diffusion coefficientE . — excess air fraction(2), aTÔ)g - acceleration due to gravityG - mass flowrateh - heat transfer coefficientH - heat of combustion(2), enthalpy(3)k - constant(2), absorption coefficient(3)k* - thermal conductivityK - constant reaction rate coeff.(2), mean absorption coeff.(3)L - kiln radius( 2 ) mean beam length(3)m - mass entrained(2.1.1), Craya Curtet parameter(2.1.2)M - momentum flux(2), mass flowrate(3)n ~ excess air in secondary flow N - mole fraction p - partial pressure P - production rate q - energy fluxQ - " "r - radius of zone ringR - radius of furnace, reflected flux densityR T - gas constantss - direct exchange areaS - surface areaSS - total exchange areaS§ - total exchange area when system properties are temperature

dependent t - time(2), temperature(3)T - temperature u - velocity U - polynomial value

V - furnace volume(2), volatile matt er(5)w - weight fractionW leaving flux densityX - axial distancey - radial distancez - separation distanceAt - temperature differenceCt - Curtet NumberLe - Lewis NumberFr - Froude NumberGr - Grashoff NumberNu - Nusselt NumberPr - Prandtl NumberRe - Reynolds NumberSt - Stanton Numbera - jet half anglecc - absorptivity3 - coefficient of cubical expansion6 - Kronecker delta6* - boundary layer thicknesse emissivityP - reflectance5 densitye - polar anglee ! - Thring Newby parametercr - Stefan Boltzmann constant

- mechanism factory - viscosity

Subscripts

a - secondary(2)a ambient(3) m meanav - average max- maximumc - combustion, cpldCÔ 0 nozzle plane(2), overalld - dust P - particlee - enthalpy R - radiationex - entrained at distance x s surface, sootf - flame T - totalfu - fuel u storageg - gas var- variableh ~ hot H fc.0- water1 - log mean C02- carbon dioxide

APPENDIX Ai

Derivation of weighting factors and absorption coefficients for a grey gas simulation of the radiation characteristics of a real gas.

It has been shown4 that the radiation properties of a real gas or gas mixture may be simulated by the weighted sum of a series of grey gases.

i =b -k.pLeg = g V ( 1 " 6 ..... ...... .......(A.l)

where

1 "z a . = 1 ( a>0 ) ............ ........ (A.2)n gx •

A reasonable fit to gas data for most engineering calculations may be achieved using one clear ( k 0=O) and two grey gases.

-kopL -kipLe = a o ( 1 - e ) + a , ( 1 - e )s s s -k2pL

* a 2C 1 - e ).. .... .(A.3)o

The values ofâ . and k. are temperature dependent3gi lbut to facilitate calculations, it is usual to ascribea mean value to k. and to express a . in a temperature

&dependent form.

For carbon dioxide / water vapour mixtures, Hadvighas developed a simplified emissivity / temperaturerelationship chart. The data from this chart for equi-molarpartial pressures of each component has been used toevaluate a .and k. in the temperature range 1000 - 2500°K, gi xand pL range 0 - 2 m. atm.

The following method has been used to evaluate a .g1and k^.

(1) The sum of a . of grey gases exclusive of a og §is guessed. *

1 =1Z a . = 1 - a o .... ................. (A.4)n gi g u

i =i(2) Z a . - e versus pL is plotted on semi-logrithmic n gi g ^ ^ ecoordinates for a range of temperature values. Equation (A.3) may be rewritten

± = i -kipL • -k2pLZ a . - = a ie + a ^ e . A.5)n gi g g g

If k >>k -i k2>>ki. then all the terms onn n-1 * .the right hand side of equation (A.5) excludingthe first will be negligable at large values ofpL, thus the ordinate of the plot will represent

-kipL i =ia je . A good estimate of Z a .will, atg n gilarge pLTs 3 give a straight line fit with a slopeof -ki;3 and an intercept of a i ( Figure A.lg •shows a typical plot for a i:and ki at 1100°K ).

-kipL s(3) Subtraction of a ie from both sides ofg

equation (A.5) allows the next term to be calculated in a similar manner.

The results from a simple computer program to evaluate a - and k. for the temperature range 1000 - 2500°K are givengi lin Table A.l.

or=

CM

lO

l—I I •s-paji—I I •6

ccLL!H-Z

FIGURE

A.l

Dete

rmin

atio

n of

weighting

factor

and

absorption

coef

ficient

at 1100

K

TABLE A.l. Calculated weighting factors and absorptioncoefficients in the temperature range 1000-2500°K.

T °K ip 1 a ng a 1g a 2g ki k21000 .001000 .270 .520 .210 .7858 13.351100 .000909 .340 .480 .180 .8384 13.911200 .000833 .390 .435 .175 .7479 13.281300 .000769 .435 .420 .145 .770 6 13.351400 .000714 .480 .390 .180 .7858 13.651500 .000667 .520 .355 .125 .8083 13.521600 .000625 .540 .340 .120 .7706 13.281700 .000588 .570 .320 .110 .7781 13.911800 .000556 .590 .305 .105 .7630 14.121900 .000526 .615 .285 .100 .7781 13.652000 ..000500 .635 .275 .090 .7858 13.912100 .000476 .655 .257 .0.88 .8008 13.652200 .000455 .670 .245 .085 .7932 13.522300 .000435 .685 .232 .083 .8008 13.912400 .000417 .695 .225 .080 .7630 13.282500 . .000400 .710 .210 .080 .8008 14.12

.7858 13.6.5

A good fit to the data for the variation of a ^ with T is given by

a . = b±0 + bîT-1 ...................... . (A.6)

The values for b^0 and bî are given in Table A.2.

TABLE A.2. Constants in weighting factor equation.Gas No. bpo bi>

0 1.002940 -735.291 0.011250 562.50 ■2 -0.004877 182.93

The values in Table A.2 give a 1% error on the summationoZ a^p = 1 over the temperature range. Finally we must evaluate the mean value of k2 over the range. Rearrangingequation A.l,

-k2 e a i -ki= 1 - ( 1 - e ) (pL=l)-- ------ (A.7)V V

Thus, from substitution of values of a ., and k.* gi g iin the range 1000 -2500 K., the mean value of k2 is 13.65

APPENDIX B.

Derivation of exchange factors for cylindrical geometry.

B .1. Gas to gas exchange factors.

Since the conditions in the cylinder are axisymmetric, the conditions in a ring of gas at a given radiusfrom the centre of the cylinder will be uniform. Thus, it is sufficient to consider the exchange between such a ring of gas of infinitesimal cross section and an infinitesimal volume at some other point in the cylinder. This situation is shown in Figure B.l. The radius of the ring of gas is R and the infinitesimal volume at P, to which the ring radiates, is at axial distance x from the plane of the ring and at some radial distance r from the centreline. The distance from P to some point on the ring is given by z. Let the emissive power be unity and let dY be the infinitesimal volume at P. Consider the radiation from an infinitesimal arc of the ring that is transmitted to and absorbed at P. Since the volume of this infinitesimal arc is RdRdGdx, the radiation that is absorbed at P from this arc is

dq = todRdBdx e-kz k cW ........ .............(B.l)^ 4tt.z 2

Let P T be the projection of P on the plane of the ring and let 0 be the angle between the radius vector through P ’ and the radius vector to the radiating arc on the ring. The distance z can then be found from the relationship

z2 = x2 + (r - Rcos6)2 + R 2sin20 .... ...(B.2)

If equation B.l is integrated from 0=0 to 2tt, the contribution of the entire ring to the absorption at P is

\C

C\J

CQ

Q> <t>

i-HPQ

WKtDC5H

FIGU

RE

obtained. Note that from equation B.2 the integrand given by equation B.l is symmetrical about 0=tt; thus

.p -kz6qp = 2 k2 dV dR It <19 .....................(B.3)

where the relationship between z and 0 is given by equation B.2.As the situation stands^ z, and thus the integral in equation B.3? is a function of the three parameters x 5 r, and R. Since the integral in equation B.3 must be evaluated numerically and then tabulated for use in the zone exchange-integral computations5 it is important to reduce its functional dependence to as few distinc variables as possible. The expansion of equation B.2 results in

z 2 = x 2 + R 2 + r 2 - 2Rrcos0 ........ (B.4)

Define

z o. - x2 + R 2 + r 2 - 2Rr ............. (B.5)z 2 = x2 +. f t 2 + r2 + 2Rr ............ (B.6)TT

thus

•Zq + .Z2x 2 * R2 + r2 =. ...... (B.7)

2

and

2 2 Z”.. - .Zq2Rr = — ..... (B.8)

2

Substitution of these relations into equation B.4 results in

2 2

Thus, as given by equation B.93 z ia now a function of only two parameters, zq .and z^. If, in addition, the substitution t = kz is made, equation B.3 may now also be expressed in terms of the two parameters To and t . The expression for d0 in equation B.3 is found in terms of dx as follows: Write equation B.9 in terms of x and differentiate to obtain

2 2 To ~ X2xdx = ------- sin0d0 ......... ......... ...........(B.10)

2

solve equation B.9 for cos0 to obtain

C O S 02t 2 - (t § + T2)

then substitute

sin0 = /Cl ~ eos0)(l + cos0) (B.12)

into equation B.ll to obtain

sin0 -2 /(x2 -”t o )(tJ - t 2) (B.13)

Substitute this result into equation B.10 to obtain

2xdx/ ( t 2 - x o ) ( t * - x 2 )

Define

TTo

' T L

-(t

2 f >T 2 e >T 2*- 1 'II —TO T 0 10t J J w - k.

. ........(B. 15)

then

d0 = 6 TodrTo

and

ToSq = 2 k" dV ^dR dx /Tir —•P TT I T

-T f yT d \UoJ [t oJ (B.17)

By examining equation B.4 and converting to a function of t , it becomes obvious that To and t are precisely the minimum and maximum optical distances, respectively, from the ring to point P. If gg is defined as the exchange factor between the gas ring of infinitesimal volume 2TrRdRdx and the infinitesimal volume dV at P, then the radiant exchange between those two volumes is given by

6q = (27rRdRdx) (kdV)gg ............... (B.18)

Then, in terms of equation B.17, gg becomes

T7TTo’?7 since thisThe parameterization here is changed to

is the most convenient form for tabular representation of the integral in equation B.19. The function

Ii To, _TTTo - rT «n- e- J 7T —T o

To To(B.20)

is plotted in figure B.4. ^

B.2. Cylindrical surface to gas exchange factors.

The problem now is to determine the radiative exchange between a ring of infinitesimal width dx on the interior cylinder surface and an infinitesimal volume dV at some point P in the gas This situation is exploded in figure B.2.

The radius of the the surface ring is Ro and the point P is at an axial distance x from the plane of the ring at a radius r from the centreline. The radiation emitted at unit emissive power from an infinitesimal arc of the ring that is transmitted to and absorbed at P is given by

, RodBdx . edq = — — -- cos(|) —-kz

k dV .... (B. 21)TT

where <{> is the angle between the ray from the surface element R od0dx to P and the normal to that surface element. The relation between z and 0 is the same as before and is given in equation B.4 The law of cosines is used to find cos$ from the triangle R 0zs formed by the infinitesimal surface element, the centre of the surface ring, and point P. Prom the law of cosines

Ro + z2 - 2R0zcos<f) = s2 (B.22)

and from figure B.2, s 2 = r2 + x2; also z2 may be found in terms of r, Ro, and x from equation B.4. Combine these relations and solve for cos(f> to obtain

j. _ Ro ~ rcosBcos9 = — --------- (B . 2 3)

Now, substitute equation B.23 into B.21 to eliminate cos(f>j then equation B.21 may be integrated from 0 to 2ir to obtain the contribution from the entire surface ring to the absorption at point P. When the result is reduced to terms of To and t^, as before, the following equation is obtained

<Sq = 2 dV BjlTTTo

dx /ttt (1 - r ecos9) ---no _ 3TTo d To

(B.24)

where cos© is a function of t^/to and t/to and is given by equation B.ll. This situation is a little more complex, since the resulting integral is npw a function of three parameters, t , To, and the radius ratio r/R0. Thus the exchange factor sg between the ring surface 27rRodx and the volume dV is best given as the sum of two integrals, eachof which is a function of only the two parameters t^ and t 0.

f T -Again, the parameterization is changed to TO, ITTo for the

same reasons given earlier. Thus, the result is

o oo

CO O

CJoH

I

< 5<0.

O O O — CM

o

CM

U S

'►h*'«a

^ I

I

►U*

h t •f-I K•

I *

HjT048

< n

(VI l O Ifto o o

o o

►Ik

►Ik- = r

CQWKCDaH

M »»

048H k«x

The integrals

TO, TO

T.,*-T 0 J

r T e = f TT -t «T d

* 'TIjoJ LT o J

10. ^ - ( T ~ T o )- e - J TT — (eosG) 3

To

> < T df \ T

[t oJ T 0 W J

are plotted in figures B.5 and B.7 respectively.

B.3 End surface to gas exchange factors.

(B.26)

..(B.26)

Here, it is desired to derive the relation for the radiative exchange between an annular ring surface element of radius R and width dR on one of the ends of the cylinder and an infinitesimal volume at some point P in the gas. Figure B.3 demonstrates this situation, and the nomenclature is the same as that used in previous diagrams. The radiation emitted at unit emissive power from an infinitesimal arc of the annular ring that is transmitted to and absorbed at P is given by

A RdGdR . e ^dq = — -— cosip ----P ^ rr 2 k dV (B.27)

where ip is the angle between the normal to the surface element and the ray from the element to the point P. In this case, cosi|> is very simply given by

cosip = — r z (B.28)

Substitute this relation for cosip in equation B.27 and

integrate from 0 to 2tt to obtain the contribution from the entire annular ring to the absorption at P in terms of To.and t :TT

6q , = P

•Tpn , it ,,r RdR rT e2 k dV —— x / tt —

“T7T L'T 0 J

T[To ( B . 2 9)

Thus, the exchange factor between the annular ring surface element on the end of area 27rRdR and the volume dV is given by

eg = f t o , TTTo

k 3 e-T° X —IT2 To 1

rT e / TT —To

f \T (1' *T[toJ [toJ(B.30)

The integral

( -T11T o ,--I T°J

lA . -.(T-Xo)r T e

= f TT - JoJ lT0j ......(B.26)

has been described earlier.

B.4 Surface to surface exchange factors.

There are three distinct surface to surface exchange factors: ss", the cylindrcal ring surface to cylindrical ring surface, exchange factor j se", the cylindrical ring surface to end annular ring surface exchange factor; and ee, the end annular ring surface to end annular ring surface exchange factor. Most of the basic relations that are needed in the derivations of these factors have already been discussed and

will not be repeated. The graphics of the situation are similar to those given in figures B.2 and B.3.

Thus for the case of ss*5 consider an infinitesimal area dA at a point Q on the cylinder surface that receives radiation from a ring surface of infinitesimal width dx, and let x be the axial distance from Q to the plane of the ring. Thus the radiation exchange at unit emissive power from an infinitesimal arc of the ring to dA at Q is given by

“~kz, R o d G d x ,, e mq Q = — cos <J> — — ...... .............(B.31)z

Using the relation for cos<j) from equation B.23a noting that here r=Roa and integrating over the whole ring the following results in terms of To and t^:

I± - T

x n R o k ^ d x rT e o6qn = 2 — ---- dA / tt (1 - cos0) . 3y ir ' 1 T b

* *

T df NT

ToV v J lToJ(B.32)

Equation B.32 can then be integrated over the ring of which dA is the surface element to obtain the total exchange between the two surface rings. Thus3 the exchange factor between two surface rings on the cylinder wall, each of infinitesimal area 27rRodx and seperated by a distance x is

The function

— - (t -T 0)eJ TT --------1 (1 - cos0)2 3To

p %T d TJo, Joj

(B.3*0

is plotted in figure B.8. Equation B.ll is used to express cos0 in terms of t , to and t .TT 5

The exchange factor se is derived in an almost identical manner. Let the emitting surface ring be on the cylindrical wall as before, and let dA be at a point Q on an end surface at radius r. Then the transmission at unit emissive power from an infinitesimal arc of the wall surface to Q is given by

-kz_ R od0dx , . edqQ = — ~ dA cos$ cosi|> — - (B.35)Z'

Using the relations for coscj) and cosij; derived earlier and integrating over the ring as before

6<1q =

To1 - 5 — COS1 no

-TTo

T[To (B.36)

Again, this expression may be integrated over dA to obtain the exchange between a ring on the cylindrical wall of area 27rR0dx and an annular ring on the end surface of area 27rrdr. Because of the dependence of the result on the radius ratio r/RQa it again makes sense to split the expression for the exchange factor into two integrals, each of which isdependent only on fr ^ 1

'’T\Thus, the exchange factor between

the two infinitesimal areas given previously becomes

se = f To, TTTo

H 0x e“To71“ To

f 7r ---------1 d0To

r fx er — / tt -no 1

r-2- -(t -To )cosO d0

To

.(B.37)

and

The functions

t o , TTTO - / TT -------------

TO

f «

T d TIt o J [t o J (B.38)

/ NT — “ ( T - T n ) f yTTT 0 rT e - I TT ---- -- cos0 3 T d[|-|I 5t 0J 1 f yT Jo; 1 oj

Jo.(B.39)

are plotted in figures B.7 and B.9 respectively.

Finally, in deriving ee, the exchange factor between annular rings at oppsite ends of the cylinder, the same procedure is again used. The radiation transferred per unit emissive power from an infinitesimal arc on one end of area 27rrdr to an area dA at a point Q at radius r f on the other end is

dq rdrd0 ___2... e-kzQ TT dA coszi|; (B.40)

Integration over the annular ring yields

where x is the axial distance between the two surfaces, and is, in this case, merely the length of the cylinder. Then, in the same manner as for the previous cases, the exchange factor between the two annular surface rings on each end of area 27rrdr and 2Trr,dr’, respectively, becomes

— - n tee = f t o , - tt-■ v TO.

(B.42)

The integral

Ii» To,- (B.38)

has been discussed previously.

APPENDIX C.

•Calculation of the overall heat of clinker from dry slurry with Barnstone Trials raw materials.

of formation reference to

Heat of Formation of Barnstone Clinker.

The reactions occuring in the kiln are listed below, together with the accepted heats of reactions.

Endothermic.

Dehydration of clay compounds to liquid water, a-alumina,- and 3-quartz (i.e. heats of reaction corrected for the exothermic crystallisation of the dehydrated clay.)

(a) kaolinite(b) montmorillonite.(c) illite

Heat of Reaction kcal/kg.

(referred to compound, notclinker)

+ 141.0 + 39.4+ 41.7

2. Dissociation of carbonates(a) calcium carbonate(b) magnesium carbonate

+ 422.0 + 310.0

Exothermic.

3. Heat of formation of cement compounds(a) Formation of 3-d.icalcium silicate C2S(b) Formation of tricalcium cilicate C3S(c) Formation of tricalcium aluminate C 3A(d) Formation of tetracalcium alumina

ferrite Ci+AF

- 171.0- 126.2

3.7

- 20.1

(C=Ca; S=Si02; A=A120 3; F=Fe20 3)

Considering first the dehydration of clay.The composition of the clay content of Barnstone secondary

material is estimated to be 251. kaolinite and 751 illite.The weight of shale used is .82 kg/kg clinker, and the totalclay content is 441.

Therefore wt. kaolinite /kg. clinker = .82 x .44 x .25 = .09k and wt. illite /kg.clinker = .82 x .44 x .75 = .27 kg.

The CaC03 content of the raw material may be obtained from the % CaO in the clinker:

CaCO 3 = CaO + CO 2100 56 44

CaCOs content CaO56 x Jq o“ kg/kg clinker

and similarlyMgC03 content = % MgO

40.3 100

Finally to determine the exothermic heat of reaction from clinker formation, the compound composition needs to be calculated.

Compound Composition

Clinker analysis:- (2nd Flame Trial)

CaOSi02AI2O 3MgOTiONa20K20S03

L„0.1.Fe 2 O 3 Mn 2 0 3Free CaO = 2.6%

65.2,

19.3,

6.9,

' 2 ,0 ,

0.3. 0.23 0.93 1.0. 0 .2 .

3.9. 0 .1 .

Compounds formed:- CaO + SO3 = CaSOit4CaO + AI2O 3 + Fe203 = 4CaO.Al203.Fe2033CaO + AI 2 O 3 = 3CaO.Al2 0 3

2CaO + Si02 = 2CaO.Si02

3CaO + Si0 2 - 3 Ca0 .Si0 2

3CaO.Si02 = 2Ca0.Si02 + CaO.

1.01 S0 3 = 1.0 x 0.701 CaOFree CaO Total

= 0,7% CaO= 2 .6 %

= 3.31 CaO

Remaining CaO = 65.2-3.3. =3 .9 !Fe. 2 0 3 = 3.9 x 0.64% AI 2 O 3

= 3.9 x 1.4% CaO= 3.9 x 3.04% 4CaO.AI2 O 3 .Fe2 0 3

61.9%2.50% AI 2 O 3

5.46% CaO.

11.86% 4 CaO.Al2 O 3 .Fe2 O

Remaining AI 2 O 3 = 6.9 - 2.50 4.4% AI 2 O 3 = 4.4 x 1.65%CaO

= 4.4 x 2.65% 3 CaO.Al2 0 3

4.4% AI 2 O 3

7.26% CaO.11.6% 3CaO.A1 2 0 3

Reamaining CaO = 61.9 - 5.46 - 7.26 = 49.18%

The calculation of the 3CaO. Si02 and 2CaO. Si02 contents maynow be calculated from

% 3 CaO. Si02 = 4.07P - 7.60Q% 2 CaO. Si02 = 8 .6OQ - 3.07P.

where P is % CaO remaining for silicate formation and Q is the % silica content.

% 3 CaO. Si02 = 4.07 x 48.18 - 7.60 x 19.3 = 53.48%% 2 CaO. Si02 = 8.60 x 19.3 - 3.07 x 49.18 = 15.0%

The compound composition is thus:4 CaO. AI2O 3. -be203

(Ci*AF)C 3A C 3S C 2S

11.9%11.7%53.5%15.0%

The net heat of formation is now calculated as in the following table:

Reaction kcal/kg kg/kgclinker

kcal/kgclinker

CaCOa = CaO + C02 + 422.0 1.164 + 491.2MgCO 3 = MgO *3- CO 2 +310.0 .042 T 13.0kaolinite = 01-AI-2O 3 +

Si02 + H 20 + 141.0 .09 + 12.7illite + 41.7 .27 + 11.3

+ 528.2

c 3s ‘ - 126.2 .535 - 67.5C2S - 171.0 .150 - 25.7c 3a 3.7 .117 - 0.4CifAF - 20.1 .119 2.4

Total 432.2

APPENDIX D.

Measured kiln operating data and chemical analyses for Barnstone Trial 4 (2-6.4.73)

TABLE

D.l

KILN

OPERATING

AND

TEMP

ERAT

URE

MEAS

UREM

ENTS

(12

HOURLY

AVER

AGE)

1 Ambi

ent

Air

TemP*

°c '4.8

7.8

10.0 9*0

Kiln

Entr

y Se

cond

ary

Air

Temp

..°c O r A O A

r H o o o l a-=T VO Lf\Kiln

Exit

Gas

Te

mp.

°ct

VO LA CTi 0'\-n- c— c—CM CM CM CM

Cool

erEx

itPr

oduc

tTemp'.

C— D— C— C— VO VO VO voi— 1 i— 1 i—1 i— 1

Kiln

Exit

Pr

oduc

t Te

mp.

°C

o o o oCM CTi CM C7\rA -=r arH rH i— 1 i— 1

Prim

ary

Air

Temp

.°c 87 103 98 101

Kiln

Exit

Su

ctio

n

in.W

g. rA CM CM A rH A O - -=T • • » •O i- 1 i— 1 i—1

Kiln

Hood

Su

ctio

n

in. Wg

0.12

0.17

0.07

0.23

COCti •o oc4-5 O ^ • H O X Nw o

o Ch o c— a vo r—O A CM rH

Date

a A A A C— C" C" c— • • • • -=r -=r • • • • CM a a vo

WCD<ttW>tt<HttEHEhOttQOttttQSQWWfinPhOcoHC O>HJ<53tt<oHso

Hott

LA CM

oo

CM

ctit t

0.1

7

-=rCM

oo

CM

t t

CT\COO

t tooo

CMooACM

1

CO

D—VO

o 1

COoCO 0

.30 CO

o

MgO -=3*

i—1

ACM

CaO

-=r

i—l^ r

A

•=rvo

CMo•HEH

OCM

O

oI A

oin

OCM

t t

90

*0'

o 1— 1

oCOoCM£S 0

.07

0.0

9

COoCM

CDfHA

. CM

COI A

COoCM1—1 <

i—1 V O

voCMo

• HCO

A

CMi—1

CM

OCM

Slur

ry

Clin

ker

Comp

onen

t

Weig

htPe

rcen

t

TABLE

D.3

AIR

FLOW

MEAS

UREM

ENTS

(12

HOURLY

AVER

AGES

. KG

./MI

N.

£oI—1pHu• H< -=r inOJ OJ0 .... • . Ix vo o\ c--cd in m CM0HCH

£Oi—1PHu• H CXD H rn voC cr\ -=r vo vou vo oo o m0 m rn •3"1—1 i—i i—i i—1 i—looo

0P0• s<3:• £ o Hpa o m OO• i—1 i . I .P Ph H 1—1# OO oo£ O P.o1—1 CQ <pH

u• H 0< P3►>9 Ehu m OO Hcd -P O OJ VO 1—1e O . . .• H P i—i CT\ -=r oU •rH oo C"- C"— ooPh Ph

C fnO 0 -=3- C"- oohO O . . . 1Jh cd VO OJ i—i< k CT\ OO ooEh

K\ m m m0 C'- r— t— o-P . • . .cd -=r ■=r •=rQ • . •m in vo c-

/P O o\ oo i—Ivo. 0 in O ■=T CT\ m3 • » • .H Q i—irH o o oS

O'WCOpaC5<PPpa><>HP >spa Php Pow CO O -=rp- -=3*i—1 in-=roo m rnOJ cd • . « . •i—l o ■=r in r in ino i—irH i—Ii—Ii—iCOpaswpapCO<pa,spop PhPh 0 t>-OO CT\CO

X vo O vo -=r oCO a . • • • .CO •H 1—1 OO OJ-=3"< i—1 vo vo 0- t— voS O

•QP3P m m m m rnCQ 0 • c— D— c— c--<a P . • • • •Eh cd -=r •=3"■=r -=rQ • • • • •OJ m invo

TABLE

D.5

DAILY

CHEMICAL

ANALYSIS

(12H0URLY

AVERAGES

PERC

ENTA

GE)

Dust

Loss

on Ig

niti

cL

CMCO

t>—i—1

CMrH

-=ri— i

r nCM

CM i—i

COn

CMrH 16

.71

0e

-p • Ho PIdTS CD CO n -=ro CD 1—1 t— nPh U • • • 1 t

Ph Pm 1—1 o H

0-pcdCO ln ■=3* r n -=3"X • • • • •Ph in n n n ncd t - - m M- t—O

0Q) 0Pm Ph2

-P CM co C M o o-m • • • • ••rH oo C M oo o ooO r n r n r n -=r r ns

0Ph3-P

• CO• HOs

i— 1 'dcd 0 i n C O O CM t - -o Ph CM MO n o CMo • H • • • • •Pm i—I o i—I 1—1 o

co<

rn rn rn rn r nt" - C-- t—(L) • • • • •

-p -=r .rr .=r ■=rcd • • • • •Q CM r n n M O n

Wo<Jpmw><.

H«Eh

COHCO>h<<XdjOowEhsHOPMQ£HwOwPMco<

MOQWPQ<Eh

oS •• ■=r> rn

0Phd i—!-P CM»C OOS

. -=rX nCO< m

O JC M

CO O

rnCMo •C M

O O-=r55 •i—I

C MrnK •- = r

mo CM.C OM O

Pc0oPHp 0C Ph0

Pd Po Xa bOe • Ho 0o &

hON i—Icdo0-1—l-=rt—

APPENDIX E.

Results of combustion gas analyses and temperature measurements taken from the flame region of the Barnstone kiln during Trial 4 (2-6.4.73)

TABLE

E.l

GAS

SAMPLE

ANALYSIS

FROM

FLAME

ZONE

AT 2.87%

EXCESS

AIR.

(VOL

UME$

)

o O J o . o o o o A

(

o

O o OJ A o C7\ A o c o oCT\ A o CT\ o ^ o A o A G "\

• • • . . • • » • •

L A GO o OJ A o CM c— o r H

i—1 1—1 O J O J

o o o o v o O o o A

A O o o O i—1 o o o CM

i—1 v o c— o o o r H o A o A« • • . . . . • • *

r A A o CM A o A -=T o o or H i—1 O J v o O J

s o o o O o OA o v o O O -= r O

w O J v o O A Crv O J A

3 * • • ■ • • • ■ • 1 1 1< o -= r o -= r CT\ o CM1— I i— I i—1P h

WJ(S3IS ]O - O o O o O J o o o O ■

3 O J O A A o o \ O o o o. -= r -= r 1—1 O O A A o A o - A

s . . . • . . • • » •

o IS - r H o s - v o o A CM v o i— :

I—1 r H V O CVJ

P h

p qo>1 a*3<C O o O

EH OJ o o OCO i—j s - t s - A

H * 1 1 1 ♦ • • 1 1 i

Q v o -^ r r H AI—1

■

o A o o o o o o oC "- o A A o A A o o A

v o o o i—1 -= r A tr— o o o o o O Y• • • • • • • ■ • • •

-= r o o O o A A 1—1 A r H OJI—1 r H i—1

O O oO n O V O o ov o O J O A

• 1 1 1 • • • 1 1 1O J -= r o o

i— l r H

CM CM CM

o o CM o O CM o o CM

o o o o O o o o O

Z 0 t7 * 0 + 0 *0 iotj * 0 - '

• U I . S f i x v ;

auzzoN wona aoNvisia UVIiNOZIHOH

TABLE

E. 2

GAS

SAMPLE

ANALYSIS

FOR

FLAME

ZONE

AT 8.62#

EXCESS

AIR

(VOLUME

%)

■w < 0-1 w1-3IS3IS1o

oPSpHwoJS<EHCO

O CM o o o o o CM oo o CM LTV o —M A o CM oCT\ CM O rH C O o V O o O •=r.• . • • . • • a a •in c - o —i~ CM O A CM o —i— i 1— i CM CM

o CM o O CTV O O o oA o CM o O IA O O o orH o o L A o O A A o o• • • a • • • • aA o- o ■=T -=r o V O CM ArH i—! CM CM

O O o O L A o O o Oi n o CM o O L A o o o OCM A O o O O . o - A A i— 1• • • . a • • a • ao A O E— 1- 1 o -=T o o *— I CMrH i— i CM A

O o o O CM o O o OCM O 1— 1 o o o O -=r A^r A CM v o L A CM A O C O C O. • . . a • • ■ a • a

A o L A CM • O V O o V O CMi— 1 i— 1 rH

o o oCM o A oi— 1 o O A

1 1 1 • • . • 1 1 1v o L A 1- 1 A

i—1 •

O o o o o O o o ot>- o V O L A o £>- A o C O ovo 1— 1 CM CT\ O A A E— c—a . ♦ . • • • a • • ■ a-=r CM O o\ -=r -cr 1- 1 A o A

1— 1 i— I i— 1

o o Oo \ o o AV O L A A• 1 1 1 • 0 • 1 1 1CM IA C O A

rH

CM CM CM

O o CM O o CM O o CMo o o O o O o o O

LQ\l' 0+ 0* 0 ZOtj •0-

•in SIXV3UZZ0N W0H3 aONVISICI TCINOZIHOH

TABLE

E.3

GAS

SAMPLE

ANALYSIS

PROM

FLAME

ZONE

AT 17.24$

EXCESS

AIR

(VOLUME

%)

O O o o O o o V D oO O CM i n o ■=r o o m oCT\ CM o O O o O r H o o o

• • • • • • * • • •

i n i n O -= r C\J o i n o \ oi- i i— i O J t—1

o LTv O o o \ O o o oL H o C O i n o m o o 1—1 oi— I CM i—i L O o o O J o 1—1 c ^ -

• • • . • • . » • •m LTV o ^ r o o i n - = r o mi— ! r H O J O J

O o o o ■— i O O -= r i nLTY O OJ L T \ o O J o O c o £—CM m o i—1 OO O CM i n o i— i

• • • • • • • • . » •

s o O i n i n o V D C O o CM1—1 i— I r H 1—1w

c

P m

O O o O t > - o o ■ =T QM CM O -= r o o t - i n o C O O

-= r CM O CM i n o O J o o ^ r

[S3 . » * • . • • . . •tS] c— L T V o i n CM o V O V D o ■=r

o i—1 r H i—1

So«

O o oC\J O V D O

W i— I o ^=r CT\o • 1 1 i • • . 1 1 I

s V O ^ r o V Oi—1

EHCOHQ

o CM o O o o o o OE * - o OJ i n o m i n o C O t > -

A D C O o C O o m m o O J C "-• » • ■ • • • . . •

-= r i—1 o C O i n r H m V D o m

i—1 i— i i - l

o o oa \ o t s - o

V D i n i—1• 1 1 1 • » • 1 1 i

CM i n i n r Hr H

CM cm

•

CM

o o CM O o CM o o 'CM

o o O O o o O o o

Z 0t7" 0+ 0* 0 Z0t7 • o -

•ro SIXVa a z z o N w o a a

................. - -----1

aoNVisia UViINOZIHOH

TA

BL

E

t.*

SU

CT

ION

P

YR

OM

ET

ER

M

EA

SU

RE

ME

NT

S

IN

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AM

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A

T

8.62$

EXCESS AIR

(°C)

0cr\

•LTVi—1

OJinCM 1—1 1

12

2 inCOOrH

EH■ < -

WSOCS]

W

0CS

•inrH

0LA 1—! rH

CM CS O 1—I 1

06

0

s S f i .<

w i-3 WPk

< LO O <u mi_3 rH la O CO s O h3 rH m O OJPk • O 1—1 S - H Ph • MO IS- 0 -

m m m O m 1— 1 1— 1 0W 1—! rH rH 1— 1 CO P3 . P3 1—1 rH rH rH

E h H i-3tsi S < CS]IS] W CSl0 S CO O

W CO S« • P3

s ln O . S LAO OJ CM 0 OJ CO « O CM LA O OOP3 • CS C— 1— 1 < W P3 • 1— 1 CM C—Pk O OJ m (—I Pk O CM CM O

1—1 rH 1—1 rH s ■fes. i— 1 1— 1 1—I 1— 1w »=r W0 P3 OJ Os w • *=k< Eh £>- <Eh 1— 1 EhCO COH OJ O 0 la 0 H OJ LA CM CMQ -=r -=r -=r cs Pd Q " CS LA m

• ln LA ■=r >H • •=r LAC'- 1—1 1— 1 rH PH t— rH rH 1— 1

SOH

-=r OJEHO -=r CM

CM C— CM c—

in LTV CO L A LAO 1—1 ^ r la O 1—1 -=r

0 O•

MO O

• m a n z z o N a oWi-3PQ

•u i a u z z o N a o

S I X V W O H t I a o N V i s i a<EH s i x v w o n a - a o N v i s i a

TABLE

E.6

VENT

URI-

PNEU

MATI

C PYROMETER

MEAS

UREM

ENTS

IN

FLAM

E

o

e

P 3

6.

12

oLTN 1 I I 1O N

i— 1

ow <P d f d

H V O O

< p a O J O O

■-a • 1 1 1 1 O N

CO t s i O J i— 1

c o IS ]pa oo aX

p a so

P dO J PhV O H

. p a O N o o o

O O o . 1 O J p - 1 O N

S 3 1— 1 •=r Ol O O

Eh < *—l . i—i< Eh

CO

p a H

s PoCSJ

O NO N O- . I I i O O 1

o p -

O J O J ■=r V O

V O o KN O'.O P - KN KN o

o KN LT* V O

o O O o

■•in auzzoN aosixv won*! aoNVisia

w<PPH&H

COEhs. oM OS — PdPd Pd P H CO < <Pd CO S CO papd opa xs eh pa pa s o tsS.r Pd c\j >HPh p -I—IoI—I Eh Eh ' <<SDpapasoS N

Ph IHPdpEhSS>

pawppaEh

WS<PPh

papN1tSlosoPdPh

paoJ3EHCOHQ

OJ I—iVO

VOOJ

OJ

I—I O'*

a\O n

Olt*ooi—l

o oUN ^T ON P- i—! i—I

Oooo

•oh auzzoN *10 sixv woHd aoNvisia

o-=roOJ

Oini voi—I

OJ OJ ■=r voVO o KN ONp - KN KN oo KN LT* vo

• • • •

o o O o

APPENDIX F .

Calculation of dust particle mean diameter in the sintering zone of the kiln.

The mean diameter of dust particles that may be supported in*a gas stream may be calculated by equating the drag and gravitational forces on any particle

where d is the maximum particle size that may be maxsupported by a gas stream flowing at a velocity u. From the data of Weber2, the mean gas velocity in the burning zone of a kiln is 14 m/s. From measurement, it is known that the density of cement dust is 1713 kg./m? Thus

d = 847 microns (gases assumed as air at 1000°C) max

If we assume a linear relationship between the mass fraction and size distribution of the dust particles in suspension then the mean particle diameter, d, is given

3irydu = ^Trd3 (ps-p& )g (P.l)

thus

l8yu (F .2)max

by

d 1 (F.3)ji dx0 dmaxx+1

thus

d = 125 microns.

APPENDIX G.

'Theoretical estimation of the velocity decay of the Barnstone kiln flame jet operating at 17.2h% excess air..

From the measured input/output data of the Barnstone kiln trials, and the measured temperature profiles an attempt has been made to calculate the jet velocity decay profile, and the time/distance relationship for a particle in the jet, assuming it to travel at the same velocity as the gas stream.

If we divide the system into rings. 0.5 m. in length, we may assume mean values for the temperature of the primary (jet) and secondary gas flows. From this we may calculate the mass of secondary gas entrained into the jet by the use of Ricou and Spalding.1 s9 form&la quoted in Section 2 .

- = 0.32 ^mo do

papo

(2.4)

From the calculated value of the Craya Curtet parameter in the kiln (m=4.15).a a recirculation ratio of 1.87:1 is predicted. Thus the calculation of jet entrainment in each ring ceases when the mass, m, of the jet is.l.87 times that of the total mass input. From this point on, disentrainment is assumed to occur from the jet at a steady rate so that at the end of the flame (as defined by the flame length formula, equation (5*2)) the mass of the jet is equal to the total mass input. ( i.e. plug flow is assumed across the duct beyond the end of the flame.). The jet is assumed to expand at a constant jet half angle of 11° until the onset of disentrainment, where it presumes a constant jet width.

Table G.l is a summary of the calculations for each ring for the results of trials on a 17.24$ excess air flame From these calculations, Figure G.l has been developed to show the jet velocity decay profile, and Figure G.2 to show the time/distance relationship for a particle in the gas stream travelling at the same velocity as the gas stream.

The ledgend for Table G.l is

x = Distance from nozzle to end of ring. m.nip = Mass of gas in jet entering ring. kg/s.m = Mass of secondary gas entering ring kg/s.s

= Mass of fuel kg/s.Tp = Temperature of jet in ring °K.T =. Temperature of secondary gas in ring °K.s ~p = Density of gas jet in ring kg/m?P 9p = Density of secondary gas in ring kg/m:su.. = Velocity of jet entering ring m/s.JUjQ =•Velocity of jet leaving ring m/s. m = Mass of gas entrained in ring kg/s.dj = Diameter of jet in ring m.

o*r-3

G

O o o o vo m m o v o oo oo rH cr\ oo m m m p - c o rH m p - p - p - p - o v cm 00 oo vo o i i n h - h - v o o

# 0 0 0 . 0 • • • • • • • • • • • . •OO t— £— f — VO O i n r l VD ^ m CM O O'* OO C— VO VO -=3* C— VO P~ tm m CM CM i—1 i—1 rH . rH rH

*r-3'G

' n s h L n o \ n s H L n L n i n i n L n i n i n i n i n i nm cr\ p - oo cm h h v o o o o o o o o o o o c r \ o o oo p - p - v o v o m l h i n m m m i r \ i n m m m H m m c— o v rH m m p - p - p - p - p - p - p - p - p - p -0 0 0 0 0 • 9 9 • • • • • • • • • •

O O O O O r H i H i H r H i H rH r H r H r H r H r H r H i H

s+

a6

cm m m m cr\ v o v o o v cm m v o o v cm m o o p p - o (j\ £— O P - O m v o CT\ OO VO P * CM rH CT\ C— VO P " P -m o cr\ p - cm o v v o m o p - p * h c o ^ p oo m cm

H cm cm m p - p - m v o v o m m m p p p m m m

06

m m o p - p - o m p - c— c— c— c— c— c— o— ^ rCO OO P - m CM m m P p rH ■ • rH iH rH rH P P O

o v o m m p - p - c— c— m m m m . m m m m m o # # # • • • • • • • • • • • • • •

o o o o o o o o o o o o o o o o o1 1 1 I 1 1 1 1 1 1

coQ.

P P P P P P P P P P P P r H P P P . P P O O O O P P P P P P P ^ ' ^ - ^ - ^ ^ ’ - H " ^j z r P ’ P ' P ’ C M C M C M C M C M C M C M C M C M C M C M C M C M C M

• 0 ' 0 0 • • • • • • • • • # # # * •o o o o o ; 0 0 0 0 . 0 0 0 0 0 0 0 0 0

aCl

p v o p p p p cm m v o P m ct\ cm m v o v o cm cm m cm v o m m m m m m v o v o v o o o cr\ o p cm cm CT\ CM i—| i—I i—1 i—1 i—1 i—1 i—1 i—1 i—l i —1 i—1 i 1 CM CM CM CM

# # * • • • ' • • • • • • • • • • • • •o o o o o o o o o o o o o o o o o o

t oEh

m m m m m m m m m m m m m m m m m m t— — C— t>— C— t — C"— C"— D— t — E— t-— C— C— tr— CS- P- P- oo o o c o c o ^ ^ ^ h 4 ^ 4 - -5- 4

P P P P P P P P P P H P P P

aEH

m m o o o o o o m o m o o m m m o op - p - o o o m p c M P - O P - o m c M C M P O O m m cm m m m m m c M cm p p ct\ oo p - • v o v o vo

i—I C M C M C M C M C M C M C M C M C M C M i—I t —l i —l i —l i —l i —1

•Htr - j

G

m o o o o vo m m o v o oo c o p c r v o o m m m m o o p m p p - p p c v cm oo oo, v o o\ m p p v o

O oo p - p - p - v o o m p v o p m cm o o v oo p - vo . t— v o p m m cm cm p p i—I p <—I

co6

m m o p - p - p - p p p - o m v o crv cm m o o p p O o cm m v o v o o \ cm m p cm o co p - m m c M o o a o v c m v o o v o m p o o m cm ct\ m cm ov v o m o

• • • • • • • • • • • • • • • # • •P P P O O O P C M C M C M C M P P P O O O O

1 1 1 1 1 1 1 1 1 1 1 1 1

ea

cm cm cm m m o v v o v o v o cm m v o cr\ cm m co co p c M C M m m o m v o a v C M P O N P - m P ' C M o m P -

i P P O V O C M O V V O m P O O p - p O O m C M O V V O C M # # # # • • • • • • • • • • • • • •

p p cm cm m m p - m v o m m m p p p m m m

Xm o m o m o m o m o m o m o m o m o

O P p cm oj m m p - p * m m v o v o p - p - oo c o ov

0 OG cd -p co • H

Qh OG o• H-P Og

<hcd

co cd

w o• H-P M « HG 0 P O cdrHcd ,G op G 0 S G • H cd Ph P G 0

p 0 •r~D<PloG O • HP 0* r HGO g P O 0 G O <H

owPiPQ<Eh

0r HNNOG

the

burner

omCQ E

Y ~ OTq.Bt.1 ^ ^ T O O I 0 A S S 0 O X 3• n

FIGURE G.l Jet velocity decay profile

Distance

from

jet

source

in equivalent

nozzle

diam

eter

s

. O

* SO0S 0UIT,LFIGURE G.2 Time/distance relationship for a particle

travelling in the jet stream

Distance

from

jet

source

in equivalent

nozzle

diam

eter

s

APPENDIX H.

Ûniversity of Surrey program GSECKC1Z compilation listing.

PR

OC

ED

UR

E*

GA

SP

RE

S(A

#B

#C

/D)

INT

EG

ER

* C

,D;

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B;

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N*

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o

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—icDCOUJh-Z

• «OO<\i"ft./ft*

OXUJ03z<cDCUJ

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k i • »

< <\jo + zcs r—i +ft “3 (\J■* %o

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ftj U-l +o *—« < cvi.h- + o

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• ft CL ft “3 —3 «c 0Jcvj K ~ UJ ft* ft ft ■It ■*z K- ft ftft *—If\J<\J

ft 00 «— ft ft It• ft a > UI rv NO 1 • •X O UJ < UI Ua> «r* r-ift * H- r- II < c V/ — i■~3 (\l 00 II • • II II II •>ft C • • Oft • • • • • • u

M OJ — > ft o o r\j ruft . Z 0J 0J O z CO

ft O II »—tDC OJ • • QC 05UJ M » O UJ03 ft z Li CDUJ _l ft *—•ft ft1— <c CC oz UJ o UJ•—• a u. CDm ft «

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ft-« Io

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ft o cc ftft. 0J OJ UJ _J ftfta U- 3 «c oc o r—■ «r03z o a UJ ft • a Z UJ UJUJ UJft o >-z a »—• cc CDft ft OJ _1 UJ c *— • •• * ft ”

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X >cK— UJ Z. OJ •—< <C X Q .

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UJ II• o <£•-J UJ z *—< H > - I L ' C

• UJ 2 U . X UJ2 OC ih ct—t m m m m me?UJCO

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m » • *■ “

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I -2 hh u

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JNCOr—i us • % •v : T~ T” U . o o

« «“ L II 1“ • «.M » II » rs • • o *% « -

u UJ HIH 2 CO c o +2 o 2 HI • H *—1U i H a» UJ o o IIU • U . X UJ 2 « •'■uo •> HIH- cc UJ H-lCO 2 - * • m

II » UJO• •u . x : UJO *•**

mh -• 12

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C“ I %I u -> OHI «— u M -J• Ic— * 22 HI u Ou-i % o • • »O 2 Ui u ul JU|I—i<_> z: HI oCO O 1 » I— oVJ | < »— 2 ■ItVJ CO UJ to 15 o

UiUJ UJ c m 1+ UJ * UI m -J

r—iUJ JU oc * — UJXT' +f\J c • to ¥■

|r~iOC UJ X • -S'*“5 1 3: u CL sfU uc- • H O • « X • fc UJ OHIHI CC 1— _i o u H-U u UJ o u_• « X -a to • « m

2 2 UJ < CO < • • to UJm a •f »—UJUi + UL. to UJ o I— HH o ISJo IIO O (V X c oc 2 jr u x ~ CO H*II 11 tX UJ E <r u u “ UI II H- II• • • •II u. u -J O oc X • « II • •“T—OC r\ioc• •o O HUI Xu 2 is; UIo • H»l 2 « «_1

UJc.u.k«k

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1— 1—12: 2:

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r - r—1X *“ >li X II X#• 1—1 • • 1—1X 21 X 2:

coixi— i

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>>

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+1 ;

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slip

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X II X II X II oc XU I m u< • • l_ l • • l_J • • • « UJ • 0 t-J

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