ISRN KTH/MSE--02/12--SE+ENERGY/AVH ISBN 91-7283-301-7

Black Liquor Combustion in Kraft Recovery Boilers-Numerical Modelling

Doctoral thesis by

Reza Fakhrai STOCKHOLM DEPARTMENT OF MATERIAL SCIENCE AND ENGINEERING May 2002 DIVISION OF ENERGY AND FURNACE TECHNOLOGY ROYAL INSTITUTE OF TECHNOLOGY SE - 100 44 STOCKHOLM

KUNGL TEKNISKA HÖGSKOLAN

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Black Liquor Combustion in Kraft Recovery

Boilers-Numerical Modelling

Doctoral thesis by

Reza Fakhrai

Department of Material Science and Engineering Division of Energy and Furnace Technology

Royal Institute of Technology SE-100 44 Stockholm

Sweden

Akademisk avhandling som med tillstånd av Kungliga Tekniska Högskolan i Stockholm, framlägges för offentlig granskning för avläggande av teknisk doktorsexamen, onsdagen den 29 maj

2002 kl. 10.00 i E1, Elektroteknik, Kungliga Tekniska Högskolan, Stockholm

ISRN KTH/MSE--02/12--SE+ENERGY/AVH

ISBN 91-7283-301-7

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Black Liquor Combustion in Kraft Recovery Boilers-Numerical Modelling

Reza Fakhrai Dissertation for the degree of doctor of Philosophy in Energy and Furnace Technology (TeknD) 2000 Royal Institute of Technology Department of Material Science and Engineering Division of Energy and Furnace Technology S-100 44 Stockholm, Sweden

Abstract Black liquor is a by-product, which results from digestion of wood chip in alkaline pulping processes. After the evaporation process the solid in black liquor increases up to 80% and it is combustible. Black liquor is conventionally burned in a large unit called Kraft recovery boiler, for the dual purposes of energy production and recovery of the pulping chemicals. Kraft recovery boiler model in the present context refers to the numerical simulation for solving partial differential equations governing the characteristics phenomenon in a Kraft recovery furnace. The model provides an analytical tool and it is best appreciated when the numerical simulations and the measurement techniques are linked to the real industrial problem and the industry that used it. The purpose of this study was to enhance the understanding of the processes involved in a Kraft recovery furnace through mathematical modelling and to keep the code in state-of-art. It is essential to consider the path of every drop in the cavity of a Kraft recovery furnace. In this process the liquor may accumulate on the walls and depending on the gravity or the flow pattern at the wall it periodically sloughs off and falls to the bed. As new components in the general framework of the Kraft recovery boiler model, the interaction of burning drops and walls in a recovery boiler considering the above mentioned was modelled. The importance of a bed model and its effect on the predicted temperature in the furnace cavity was examined. The heterogeneous conditions in a Kraft recovery furnace with significant local variation of concentration of constituent gas components and temperature level/gradient could affect NO production rate. The NOx model developed in this work considers the NO formation from fuel NO and prompt NO. It is assumed that the fuel nitrogen in black liquor is released either via devolatilization or char combustion. Further work focussed on estimation of the temperature level near/on the bed based on the mass distribution on the char bed. The model was also used to examine the effects of changes in black liquor properties (used in Kraft recovery furnace model) namely effect of the swelling and solid content versus the Kraft recovery furnace performance. The results illuminate also the potential of numerical modelling method as a promising tool to deal with the complicated combustion processes even for practical application in the industry.

ISRN KTH/MSE--02/12--SE+ENERGY/AVH

ISBN 91-7283-301-7

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Supplements

The work presented here this thesis is mainly based on the following publications, referred by Roman numerical

I. Theoretical Analysis of Interaction between Fuel Drop and Walls during Black Liquor

Combustion in A Kraft Recovery Furnace. Energy conversion & management- an International Journal, RAN2001 Special Issues, Nagoya, Japan Dec. 14-17, 2001

II. Combustion Performance of the Kraft Recovery Boiler Versus Black Liquor Properties – Numerical Study, Submitted to Energy Conversion & Management

III. Use of a Computer Model for Evaluation of Combustion and NOx Control Alternatives

in a Kraft Recovery Boiler, International Chemical Recovery Conference, Tappi, Tampa Florida June 1-4 1998

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Acknowledgment The work in this thesis has been performed within the Energy and Furnace technology, Material Science department, Royal institute of technology (KTH) Stockholm Sweden, during the years 1997-2002. Energi Myndigheten (STEM), Ångpannföreningen forskningstiftelse (ÅF) and Värmeforsk financed the project. I wish to convey my sincerest thanks to the leader of the Energy and Furnace division, Associate Professor Wlodek Blasiak, for giving me the opportunity to work in the group and to write my thesis under his direction, and for constantly providing me encouragement, guidance and moral support. I appreciate his commitment and utmost professionalism in all regards. I also thank the other members of the Material Science department Prof. Seshadri Seetharaman, Prof. Pär Jönsson who made me feel like a colleague more than a student during the years I have worked at the department. I would also like to take the opportunity to thank my co-workers Jan Bong, Simon Lille who were my friends when I needed one. I would like to thank my family, particularly my father, Abas who passed away 1999 and my mother Golestan, for their support and encouragement throughout my long academic career. Finally, my heartfelt thanks to my dear wife “Neda” for her patience, concern and understanding, my son “Sam” and my daughter “Ella Mina” who inspired me to aim high and motivated me to achieve it.

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Table of Contents Abstract I Supplements II Acknowledgements III Table of Contents 1V Nomenclature VII Thesis Summary 1. Introduction 1

1.1 Black Liquor 2 1.2 Recovery Furnace 2 1.3 Objectives 4

2. Liteature Review 5

2.1 Flow Field 5 2.2 The In-flight Combustion of Balck Liquor Droplets 6 2.3 The Char bed 8 2.4 Pollutant Formation 9 2.5 Gird Generation and Geometry 11 2.6 Model Validation 12 2.7 Concluding Remarks 13

3. Development of Model of Recovery Furnace Used in

This Work 14

3.1 Geometry of the Furnace Used in This work 14 3.1.1 Conventional Firing Strategy (CF) 15 3.1.2 Rotational Firing Strategy (RF) 17

3.2 Black Liquor Combustion 20 3.2.1 Spray 20 3.2.2 In-flight Drop Combustion 23 3.2.3 Char Bed Processes 24 3.2.4 The Interaction Between Drops and the Walls in the Recovery

Furnace 24

3.3 NO Modeling 28 3.3.1 Thermal NO 28 3.3.2 Turbulence Chemistry Interaction 29 3.3.3 Prompt NO 30 3.3.4 Fuel NO 31

4. Result and Disussion 33

4.1 Estimation of Surface Temperature and Mass Distribution on a Char Bed 33

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4.2 Effect of the Physical Properties of Liquor on The Furnace Performance 36 4.2.1 Effect of the Swelling on the Path of a Droplet 36

4.3 Effect of the Wall-burning Model on the Overall Recovery

Furnace Model performance 40 4.4 Prediction of the NO Level in Recovery Furnace 42

5. Conclusions 46 References 48 Appendix

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Nomenclature Aint Internal surface area of carbon in the bed (m2). The value of the internal

surface area of Kraft chars was taken from the experimental data obtained with two experimental chars obtained

Asp Internal specific surface area of char, taken as 11000, m2 /kg CC, bed Molar concentration of carbon in the bed, mol/m3 CSO4 Sulphate concentration, mol SO4/mol Na2 CC Carbon concentration, mol C/mole Na2 CNa2, bed Sodium concentration, moles Na2/bed volume, mol/m3 Md Mass of the drop/parcel, kg M char, Na Total amount of sodium in the drop parcels, kg M char, S Total amount of sulphur in the drop parcels, kg. Md0 Initial mass of the drop/parcel, kg Mi, Mj Molecular weight of corresponding gas species Pg Local pressure of gas mixture at the cell adjacent to the char bed, bar RO2, overall Overall oxidation rate, RCO2, overall Overall CO2 gasification rate RH2O, overall Overall H2O gasification rate V bed Char bed volume per unit surface area (m3/m2), XH20 Mass fraction of water XVM Mass fraction of volatiles XC Mass fraction of char XSmelt Mass fraction of smelt XNa, XS Initial mass fraction of sodium and sulphur in the drop parcel. YCO2, YH2O Molar fractions of carbon dioxide, water vapour, the cell adjacent to the char

bed. YCO, YH2 Molar fractions of carbon monoxide, and hydrogen at the cell adjacent to the

char bed. YO2 Local oxygen mole fraction at the cell adjacent to the char bed. β Multiplier for Cameron-Grace reduction rate η Sumnicht factor, for carbon surface area available for direct oxidation. and has

a value of 0.6, which corresponds to 50% completion of sulphur reduction58 from the Institute of Paper Chemistry’s drop tube furnace, which had a value of about 11 m2/g.

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1. Introduction In the 1970’s the paper industry in Sweden started to make large scales structural changes to the production of paper. These modifications continued during the 1980’s and 1990’s. The number of the paper making units was 47 in 1998, a reduction by 8 units since 1988. Nevertheless, the total energy used by the industry increased because of environmental regulation [1]. The industry is almost independent on oil, but it requires heat for evaporation plant. Higher electricity costs as a result of nuclear phase-out seems to be the real threat to the industry [2]. In order to reduce the energy consumption, the building new production line or modernisation of the existing one, are possible alternatives. The other option would be, to increase production of the required heat by using the spent liquor. The industry has a long tradition in using residual oil. Heat extraction, from compound found in wood as a resulting of the chemical pulping process can supply a major part of the energy needed by the industry. Of the total energy used, its bulk is from the internal fuel (black liquor and bio-mass). 74% of the total energy used, is self-produced with black liquor contributing 63% in 1998. Chemical recovery is a technique used to recover the valuable cooking chemicals, generating large amounts of heat energy by burning the organic material in black liquor and eliminating the black liquor as a danger to the environment. The main objective of the Kraft recovery steps is to minimize, as efficiently as possible, the loss and subsequent makeup of the chemicals used in preparation of the cooking liquor (commonly called white liquor), [Green et al. (3)]. The recovery boiler in this regard is crucial and the bottleneck to the process. In order to improve recovery boiler performance, it is necessary to understand the effects of furnace design furnace operating variables, and liquor quality e.g. fouling and plugging rates, air emissions and combustion stability etc. One of the tools available in improving the desired operation and design conditions is mathematical modelling. Mathematical modelling of a Kraft recovery furnace is the study of fluid dynamics, combustion and has been under constant development for years. It is commonly accepted as referring to the broad topic encompassing the numerical solution, by computational methods, of the governing equations which describe fluid flow, the set of the Navier-Stokes equations, for example continuity, energy /species concentrations. In Karft recovery model, In-flight burning of liquor and bed burning are the other important components, which are included and affect the gas phase. The essence of the subject of Kraft Recovery furnace is that of judicious compromise between theory and experiment. The goal of using the model is to use the model as a tool in order to enhance the understanding of the phenomena in question and in this case, the recovery furnace in a Kraft cycle. In the next section, terms: black liquor, recovery boiler and black liquor combustion are clarified and the relevant processes are introduced.

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1.1 Black Liquor Black liquor is an important resource for steam generation in the pulp and paper industry. It is a substance unique to the pulping process and represents a readily available renewable energy source. It is also one of the few fuels to have been locally produced and used in the countries with a pulping industry. From the point of view of fluid mechanics, black liquor is a liquid. However, it has two states, fluid and solid. Its main components are inorganic cooking chemicals, lignin and other organic constituents removed from the wood during pulping in the digester, and water. These organic constituents are combined chemically with sodium hydroxide (NaOH) in the form of sodium salts such as Na2S, Na2CO3, Na2SO4. Some organics may be recovered in the chemical recovery process, such as the mixture of resin and fatty acids known as tall oil, or the turpentine recovered in a liquid separation sequence. The exact composition of black liquor depends on the wood species, the pulp yield, and the alkali charge used. Considerable differences exist between liquors from different species, and especially between those from hardwood and softwood. An important aspect of black liquor combustion is elemental analysis of the liquor solids, that is, of the percentage by weight of each chemical element in the black liquor solids. Five elements are always present: sodium (Na), sulphur (S), carbon (C), hydrogen (H), and oxygen (O). In some cases, potassium (K) and chorine (Cl) are also present. The approximate composition of black liquor is given in Table 1.

Table 1: Typical elemental composition of black liquor solids and char

Element Present (wt%) Na 19.17% S 4.76% C 35.93% H 3.56% O 35.20% K 1.02% Cl 0.12% Inerts 0.24% Total 100.0%

1.3 Recovery Furnace A recovery boiler is both an ordinary steam boiler with tubes in the walls, bottom and top of the furnace that delivers the steam required by the mill and a chemical reactor where sodium sulphate is reduced to sodium sulphide. A unique characteristic of such boilers is the use of char bed in the lower furnace. These boilers were developed by Tomlinson, in cooperation with Babcock and Wilcox, in the early 1930s and contributed to the predominance of the Kraft process, in which the recovery boiler is a crucial element. The boiler has three critical functions. First, it uses the chemical energy in the organic portion of the liquor to generate steam for the mill; second, it plays a major role in the sulphate process as a chemical reactor;

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and third, it destroys dissolved organic matter and thus eliminates one type of environmental discharge. Concentrated black liquor enters the recovery furnace as droplets fired through liquor guns in the walls. Traditionally, two systems are used: either wall firing using one or two oscillating guns, or suspension firing with many oscillating guns located on two of the four walls around the furnace. The droplets then go through a series of processes involving drying, pyrolysis, char gasification and finally homogeneous combustion. Burning of the solid char residue that remains after pyrolysis of the black liquor occurs largely in the char bed that covers the floor of the furnace. Older recovery boiler designs incorporated tangent tubes or tubes with flat studs backed by refractory. These water wall constructions had thick casing plates on the outside in order to keep the flue gases within the furnace. The walls of modern recovery furnaces are welded walls, also known as membrane walls. They are constructed of vertical tubes, typically 6.4 to 7.6 cm in diameter, set in rows. The tubes are either placed immediately adjacent to one another and continuously welded along the lines of contact or are spaced approximately 1.25 to 2.5 cm apart, connected by a flat fin. In chemical recovery boilers the water wall tube is normally made from ordinary carbon steel. A series of ducts and boxes introduce the combustion air into the furnace. These open though the boiler walls at various levels, and with different numbers of openings at each level. Dampers located in the ports, wind boxes, ducts or near the forced draft fan control the air flow. Primary air nozzles are located approximately 1 m above the furnace floor. Approximately 35 primary air nozzles are located on each of the four walls of a recovery furnace designed for a 1000 ton per day mill. The size of these ports is approximately 5 by 25 cm. Secondary air nozzles are located approximately 2 m above the furnace floor. The secondary air nozzles are generally large and less numerous than the primary ports. There are 4 to 16 ports on each wall of the furnace, ranging in size from 5 by 25 cm to 12.5 by 67.5 cm. Tertiary air nozzles are located above the liquor guns approximately 8 m above the floor of the furnace. Tertiary air nozzles are the largest and are usually located on two opposing walls, normally the front and rear walls. Typically, 3 to 8 tertiary ports ranging between 10 by 45 cm and 15 by 75 cm are used on both walls. The formation of Na2S requires local under-stoichiometric conditions. This reduction step is critical. The only part of the recovery cycle where oxidised sulphur compounds can be converted to sulphide is during combustion. The inorganic compounds melt and flow out of the furnace as a mixture of molten salts called “smelt”.

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Figure 1 Typical recovery furnace

1.4 Objectives The work presented here has the following general objectives: 1. Enhance understanding of recovery furnace behaviour. 2. Develop new models and modify the existing model where necessary to keep the

recovery furnace model state-of-the-art. To achieve these goals, attention has been directed to the following: • Estimating the temperature level near/on the bed based on the mass distribution on the

char bed. • Developing a model that incorporates the interaction of the droplets and the walls in a

recovery furnace • Implementing the NOx production/reduction mechanism in the recovery furnace. • Examining the effects of changes in the properties of black liquor, particularly the effect

of swelling and solid content on the performance of a recovery furnace. It is hoped that the results provided in this thesis illustrate some of the ways CFD can be used to supplement existing engineering tools and that researchers will be encouraged to tackle new problems using modelling and to develop new methods as well as extensions and improvements of the methods presented here.

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2. Literature Review The combustion of black liquor has been studied for more than 40 years. During the1980s there was extensive research in this area in the USA, Canada, Finland and Sweden. These findings underlie the present work. However, achievements in other fields have also contributed to our present understanding of recovery furnace operation. The rapid development of numerical simulation methods has made it possible to simulate a recovery furnace and accurately predict its performance [Blasiak et al. (4)] and has the potential to enhance our understanding of the processes occurring in a recovery furnace.

2.1 Flow field Among the early achievements using recovery furnace models in literature is the isothermal simulation of the flow field and mixing pattern in a recovery furnace. The gas flow patterns are determined primarily by the furnace geometry, air inlet geometry. In view of this, considerable efforts have been devoted to determine the flow pattern in the recovery furnace. Jonse and Grace (5) studied the general flow pattern in a recovery furnace, comparing the results of their computational model with experimental cold flow. The framework for the computational modelling techniques was FLUENT. The results showed that the flow pattern and the furnace exit gas velocity distributions have a tremendous effect on the deposition rate on the superheater tubes. Grace et al. (6), 1989, described the construction of and preliminary results from a three-dimensional mathematical modelling of a Kraft recovery furnace. Their results indicated that gas flow patterns are primarily determined by the geometry of the air inlets. The bed shape, on the other hand, can be affected by the gas flow patterns. . Llinares and Chapman (7) used flow modelling to examine a three-level air system. The model was used to evaluate designs in order to determine the best location for introducing air into the furnace. The article gives no information regarding the modelling. Perchanok et al. (8), investigated the flow pattern in a recovery boiler the results indicated that the flow is often grossly unstable. This unsteadiness was also observed in the physical model of boiler as well. Bergman and Hjalmarsson. (9) introduced Rotafire, a new firing strategy applied to a number of the secondary air nozzles. Reducing the total open area of these nozzles produced an initial jet velocity of 70–80 m/s. Since the jets work together, this velocity caused vigorous rotation of the gas in the lower furnace. Mathematical modelling was used to support the theory. Jones and Chapman (10) used computational fluid dynamics to describe two models in order to examine a refined secondary air system. They used big/small air nozzles configuration in secondary level. The results predicted a dramatic improvement in combustion behaviour using big/small secondary design Salcudean and Gartshore (11) investigated flow pattern and associated transport phenomena in black liquor recovery boilers and found that aerodynamics is critically important to the performance of these boilers. They concluded that buoyancy effects are most likely secondary over large parts of the boiler.

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Salcudean et al. (12) presented preliminary computations of the cold flow in a simplified geometry of a recovery boiler. They did not take into account the effect of the char bed shape on the flow since the char bed was not modelled. However, the results indicated that the flow in the secondary level interacts strongly with the upwards flow in the central region of the boiler to form the upwards-rising core. All of the research discussed so far used symmetry boundary conditions. Modelling just half of the furnace further reduced the mesh size requirements. The resulting mesh provided the optimum resolution and run-time performance possible at the time. Tao et al. (13) studied Rotafire air system designs. They showed that Rotafire creates a more effective mixing of gases than conventional firing methods, making it possible to achieve higher combustion and thermal efficiency and more evenly distributed gas temperatures and species..

2.2 The In-flight Combustion of Black Liquor Droplets It has long been understood that black liquor behaves in very unconventional fashion during combustion. Its combustion behaviour is more like that of solid fuels, such as coal, than that of oil or other liquid fuels. During the combustion period, the droplet undergoes drying, devolatilisation, char burning and smelt reaction. The amount of literature associated with spray modelling in general is huge, and comprehensive results have been presented for a droplet trajectory problem that involves drying, vaporisation and burning [Dwyer (14), Lintries (15), Stockel, (16), Bousfield et al. 1990]. Spielbauer et al. (17), presented droplet size distribution data for the spray from splash-plate and swirl-cone black liquor nozzles. They showed that the distribution is square root normal. The result also indicated that normalised size distribution does not change, or changes very little, as a function of nozzle geometry, flow conditions and fluid parameters. Horton et al. (18) studied spray variables including mean droplet size, breadth of spray distribution, and the angle at which liquor is sprayed into the furnace. However, little information is given about the parameters used in the model and symmetry plane was used in order to save CPU. Empie et al. (19) evaluated two types of commercial nozzles, splash-plate and swirl-cone, with a number of liquors from different mills and found the same qualitative trends for both nozzles, despite differences in the quantitative dependence of droplet diameters on the parameters studied. The two most important variables were the velocity of the jet and the viscosity of the liquor. Increases in the velocity produced sprays with decreasing diameters and increases in the viscosity produced sprays with increasing diameters. Increasing jet diameters also increased the droplet sizes. Correlations showed that the type of liquor did not appear be important. Droplet diameters in these studies generally ranged between 2 and 3 mm. Helpiö et al. (20) using splash-plate nozzles, focused on the effect of temperature on atomisation, especially above the boiling point of the liquor. They found that even though the

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temperature increase near the boiling point results in an increase in droplet diameters, the phenomenon of flashing, which occurs several degrees above the boiling point, leads to significantly smaller droplet sizes. Droplet diameters in these experiments (liquor with S = 0.698) ranged mostly between 3 and 4 mm, but decreased sharply with temperature to about 2 mm several degrees above the boiling point. Hupa et al. (21) introduced the use of the single droplet burning technique to measure and characterise the swelling of black liquors during combustion. In their study, they observed a droplet swelling factor (d/d0) of 1.3 to 1.8 almost immediately upon contact with the hot gas, which remained constant during the drying period. The average value for a set of 19 determinations was 1.55, with a standard deviation of 0.11. Rapid and much more drastic expansion ensued during volatilisation, reaching maxima exceeding linear ratios of 3 (volumetric up to 35) at the end of that stage. Diameters decreased during char combustion, falling below those of original droplets before burnout. Hupa et al. (22) published an article on black liquor combustion in which they defined the stages of black liquor combustion as drying, devolatilisation and char burning. Their work with a laboratory furnace enabled them to define three time-periods describing the entire droplet combustion process: drying time, ti, from the initial contact with the gas to ignition; devolatilisation time, tv, from the appearance of the flame to maximum expansion; and char-burning time, tc. Good empirical correlations were established for all three: ti is proportional to the initial droplet diameter, and both tv and tc are proportional to the 5/3 power of that diameter. They were also able to generate profiles of the swelling and of the droplets’ temperature as a function of the liquor combustion time. Fricke (23) performed extensive studies of the properties of black liquor and defined some of them by mathematical expressions (Appendix 1). Jones (24) as well as Grace et al. (6), presented models which included some critical features of Kraft recovery boilers, such as the combustion of black liquor droplets in flight and in the char bed. Frederick (25), (26), studied the combustion rate models for each of the combustion stages of black liquor. The drying and devolatilisation were limited by the rate of heat input to the particle. In the char burning stage, the burning rate was controlled by mass transport. These studies provided much needed data on such aspects of black liquor combustion as droplet surface temperature, the yield of volatiles during pyrolysis, and the impact of the temperature of the gas flame on swelling behaviour. They also succeeded in simplifying the swelling process somewhat by considering the change in dimensionless droplet diameter (the ratio between the minimum and maximum diameter) as a function of the dimensionless devolatilisation time (the fraction of pyrolysis time). An empirical expression was developed such that the dimensionless droplet diameter was roughly equal to the dimensionless time to the power of 0.8. Walsh (27) and Hyöty et al. (28), 1989, developed a sub-model for in-flight combustion of black liquor droplets. Horton et al. (29) studied the key parameters of this model using a recovery furnace model. The model traces a droplet’s flight path as it simulates the rates and stages of combustion as well as the locations where water, volatile organics, char carbon, and inorganic ash portions of the black liquor were transformed from one phase to another. In all these studies a symmetrical plan was used to save CPU time. However, Quick et al. (30)

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pointed out that even when a boiler was geometrically symmetric, flow instabilities and jet interactions could cause major asymmetries in the flow patterns. Malmgren (31) used analytical solution of partial differential equations to investigate the movement of droplets. The simulation showed that the trajectory of a droplet was influenced by the drag coefficient, mass transfer, its initial properties and variations in its size or density during flight. Rotation of a droplet was less important and the risk of the droplet breaking up due to aerodynamic forces was small in a recovery boiler. Grace et al. (32) presented a new model of elemental-based black liquor drop burning. The processes of drying, devolatilisation and char burning were treated as occurring in parallel, eliminating the need for arbitrary criteria for transitions between steps. The model gave rate equations for the transfer of individual elements in the black liquor to the gas phase and included a proper treatment of char gasification as well as oxidation.

2.3 The Char Bed A char bed with a wide active layer ensures stability and consistency in the combustion process in a recovery boiler. However, controlling the burning process and combustion rate in and on a char bed is very difficult. One of the critical issues for recovery boilers is the avoidance of blackout. Blackout occurs when • heat released above the bed does not reach the bed surface and the gasification process

stops (thermal feedback) • the shape and height of a char bed are such that primary air nozzles are blocked, resulting

in incomplete combustion of the released gas [Hough et al. (33), 1985]: Sumnicht (34), 1989, presented a sub-model for char bed combustion that incorporates several factors that are important for the design of such a model, including the char combustion rate [Brown et al. (35), 1989] and the surface roughness of the char bed. The results suggested that the bed influences gas flow patterns as well as playing an important role in overall char combustion. Karvinen et al. (36), 1989, developed a char bed model in which the surface of the bed was assumed to be insulated and the air flow of primary jets was uniformly distributed across the whole bed surface. The heat released when droplets hit the bed was also assumed to be uniformly distributed. This model did not take into account the interaction between flowing air, incoming material and burning in the char bed. Frederick and Hupa (37) developed a char bed model that took the chemical reactions into account and gave the temperature profile of an active zone, the rate at which carbon was burned The model required the mass flow rate of char to the bed and its composition and temperature as input. Sutinen et al. (38) presented a two-dimensional char bed model linked to the gas atmosphere above the bed in the furnace. The velocity, temperature and concentrations of gas in the furnace were solved numerically for the vicinity of the char bed and linked to the chemistry of the char bed.

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Wessel et al. (39) extended the work done by Fiveland et al. (40), 1988. Their model included carbon monoxide oxidation kinetics but did not considered the H2 and CH4 in homogeneous chemistry. The char bed surface temperature was given, but no information was supplied regarding the nozzles and conversion rates. Tao et al. (41) examined char bed geometry using the partial three-dimensional approach. Their results showed that aerodynamics in the furnace had a strong impact on the performance of a Kraft recovery boiler. Any changes to the air system configuration and the char bed geometry would significantly influence the flow field in the furnace. Tao et al. (42), 1997, presented a char bed model based on and incorporated into a recovery boiler model. They based their work on the fact that the reaction rate is influenced by the rate of mass transfer of the reacting gases (oxygen, carbon dioxide and water vapour) to the char bed as well as by the kinetics of the reactions of these gases with the carbon in the bed. Grace (43), 1996, reviewed computer models of recovery boilers and identified two factors that currently limit the ability of models to deal with issues concerning char bed shape and inventory. One is a lack of an adequate model for the char bed inventory. The other is the lack of consensus on what constitutes a good bed, especially in a quantitative sense.

2.4 Pollutant Formation Control of pollutant emission is an important factor in the design of a modern recovery boiler. Pollutants of concern include particulate matter, such as soot and fly ash, metal fumes, and various aerosols; the sulphur oxides, SO2 and SO3; unburned and partially burned hydrocarbons, such as aldehydes; oxides of nitrogen, NOx, which consist of NO and NO2; carbon monoxide, CO; TRS (total reduced sulphur), and greenhouse gases such as N2O and CO2. Some of these gaseous emission such as TRS, CO, and volatile organic compounds (VOCs) are destroyed by oxidation. Adjusting the stoichiometry and improving the mixing in the furnace is sufficient to ensure the destruction of these gases. There is thus no need to simulate these species using source/sink methods. Turns (44) however, pointed out that CO is a major species in rich-combustion products, and that substantial amounts of CO will be produced whenever rich mixtures are used. The other source of CO is quenching by cold surfaces. In a recovery boiler the flow is unstable. Consequently, there will be local areas of rich and poor combustion. The CO concentration is highly uneven and therefore CO emission could be locally high. Sricharoenchaikul et al. (45) reported that 30 to 60% of the carbon originally in the black liquor converted to tar. Secondary pyrolysis reactions then produced CO and CO2. These results were conformed by several studies. Sricharoenchaikul et al. (46) investigated sulphur species transformation and sulphate reduction during the pyrolysis of Kraft recovery liquor. He found that sulphur was released from the burning liquor droplets and developed an algorithm for modelling SO2 behaviour in the furnace.

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Frederick et al. (47) proposed an empirical model for predicting the rate and total amount of sulphur releases during devolatilisation of black liquor. The rate of sulphur release is assumed to be proportional to the rate of carbon release, while the total amount of sulphur released is a fraction of the total sulphur in black liquor. Although NOx is a minor species in the combustion process, it is important because of its contribution to air pollution and has a large quantity of literature devoted to it. Someshwar et al. (48) investigated the mechanism of fuel NOx and thermal NOx in relation to NOx formation in a Kraft recovery furnace. The results showed that NOx formation is most likely a result of fuel NOx formation rather than thermal NOx. The impact of an increase in the solid content on NOx emission was unclear in that study. Jonse et al. (49) studied NOx formation and found that it increased above 75% solid content in black liquor, primarily due to additional thermal NOx formation. Nichols et al. (50) reviewed NOx formation mechanisms in a recovery boiler. The limited data available showed an increase in NOx as the solid increased from 62% to 80%. The observed increase was much less than predicted by the temperature dependence of thermal NOx. Nichols et al. (51) studied the formation of fuel NOx during black liquor combustion. They observed that fuel NOx is formed during devolatilisation and char burning. The formation of NOx was found to be moderately sensitive to temperature in the range of 800–1000°C. It was also found that conversion of only 25% of the nitrogen in the black liquor could account for the level of NOx measured in the flue gas from furnaces. Aho et al. (52) studied fuel nitrogen released during black liquor pyrolysis. Their results indicated that ammonia was the main fixed nitrogen species formed and that the rate of fixed nitrogen released increased with increasing temperature. Adams et al. (53) used the results from CFD calculations for a recovery boiler to estimate the maximum contribution of the thermal NOx mechanism to the ultimate NOx emission from the furnace. The results were compared for cases run at 67% and 80% black liquor solid content. In both cases the estimated total thermal NOx was very low, 0.09 ppm and 8.3 ppm respectively. Information regarding the max temperature in the furnace was not given. Forssen et al. (54) studied char nitrogen oxidation in single droplet experiments. At higher gas temperature (>800°C) and higher oxygen concentrations (>1%), the char nitrogen was oxidised at the same time as char carbon, and NO was formed. At low temperatures and low oxygen contents, however, the nitrogen was not released until all of the char carbon was consumed. If oxidation was stopped in the middle of carbon oxidation, the N remained in the char. During gasification in 20% CO2 there was little formation of NO, and approximately two-thirds of the original char nitrogen was retained in the residue. Brink (55) proposed a NO model for black liquor droplets. The model assumed that the fuel N was released in the gaseous phase via devolatilisation and char combustion. During devolatilisation, the amount of nitrogen released depends on temperature and increases as the temperature increases. Laboratory studies have found that approximately 70% of the fuel nitrogen is released during devolatilisation, mainly as NH3 and N2.

11

Iisa et al. (56), 2000, reviewed the NOx formation and reduction mechanisms in a recovery boiler. They proposed algorithms for predicting NOx in a recovery boiler based on the three NOx formation mechanisms; thermal, prompt and fuel NOx. The basic information on reduction of the NOx in a recovery boiler is also available and interacts with the NOx formation mechanism.

2.5 Grid Generation and Geometry Very few studies have addressed the issue of grid generation and geometry. A full-scale boiler is approximately 30 m high and has a cross section of 100 m2. It features one or two liquor guns on each wall. The bull nose is located some 30 m above the floor and occupies about one-half of the furnace cross-section. The combustion air system in the boiler has three levels: • Primary air, nozzles located on all four walls, about 1 m above the floor and 100 x 50 mm

in size. • Secondary air, nozzles located on all four walls, about 2–3 m above the floor and 200 x 50

mm in size. • Tertiary air, nozzles located on two walls (the front wall and the rear wall) about 9.1 m

above the floor. In some cases there is a fourth level, at an arbitrary height. Comparing the size of the air inlets and the height of a furnace shows the need for high local grid resolution, particularly in the air nozzle regions. Jones et al. (57) described the first stage in the development of a three-dimensional recovery furnace model, namely, the simulation of cold flow in an existing furnace design. One of the limitations mentioned in the work was the use of staggered nodes to describe a diagonal surface (i.e. the bed surface). Grace et al. (43) listed the limitations in the simulation of a recovery boiler at that time. They included problems in reaching convergence and the use of symmetry plane to limit the number of computational cells. Salcudean et al. (12) addressed two issues in their simulation of transport phenomena in recovery boilers. The first of these is the need for high local resolution, especially at the air nozzle levels, and the second is the slow convergence rate of the solution algorithm for the large domains typical of recovery boilers. They suggested the adoption of a multigrid solution algorithm in order to increase the convergence rate. However, they also used a symmetrical model (only half of the furnace was modelled). Tao et al. (58) investigated the flow field in a recovery boiler and showed that use of an unstructured grid and local mesh refinement made it possible to simulate the complex flow and geometry in a Kraft recovery boiler with optimum resolution and run-time performance. Studies of the isothermal flow field that combined mathematical and physical modelling verified that this approach gives satisfactory accuracy regarding general flow characteristics. It minimises the need for compromises in describing the geometry of the boiler when setting up a problem with a reasonable number of computation cells or elements. Air nozzle arrangements do not have to be adjusted to fit grids and line up with other nozzles and the geometry of the air nozzles can be correctly described.

12

Salcudean et al. (59) examined the numerical solution and convergence. Their results showed that an accurate solution of the flow field in a recovery boiler could only be obtained with very high numerical resolution and excellent convergence characteristics. The computer code developed at the University of British Columbia and used in the work uses a multigrid method and segmentation technique. Yang et al. (60) presented a CFD study of jet flows near the char bed in recovery furnaces. They examined the effects of step and smooth char bed surfaces on gas velocity, shear stress, and heat transfer rates. Comparing a step bed model with a smooth bed model showed that a step boundary could be used to predict main gas phase flow patterns, but that the velocity errors caused by the step boundary result in large errors in shear stress and heat transfer coefficient distributions. A step boundary also resulted in overestimation of the surface area and led to an error in the total heat transfer rate. A smooth boundary was recommended if the surface transport process needed to be accurately calculated. The results indicated that the primary air jets produce high shear stresses on bed surfaces and are therefore important in controlling bed shape. The geometries of the nozzles affect shear stress distributions, and therefore also the mass and heat transfer processes that play important role in char bed combustion.

2.6 Model Validation It has proved very difficult to measure lower furnace gas temperatures, species concentration or bed temperature with any degree of reliability. The basic problems are the extreme dirtiness of the furnace atmosphere and interference from the sodium and water vapour present. Borg et al. (61) measured the temperature inside a Kraft recovery furnace under various operating conditions. Temperatures were measured inside the bed (5–25 cm below the surface). They distinguished cold spots and hot spots on the surface of the char bed. The hottest zones were where the primary (and high primary) air hit the bed surface. They found that the average bed temperature was usually quite stable, as if controlled by a themostat and close to the melting point of the system. This study gave a normal temperature distribution vertically in the furnace. Blackwell (62) used two simple techniques to check the validity of the results of physical flow modelling: a) cold-flow air tests on a full-scale boiler, b) camera observations in an operating boiler. The work aimed to examine a new air system. Blasiak et al. (63) studied the isothermal flow pattern in a recovery furnace model using a water model. Flows calculated using isothermal models were tested against data from water flow models of two different recovery boilers with reasonable success. Vaclavinek et al. (64) studied the stability of the flow field in a recovery furnace using a water flow model. The results were verified by mathematical modelling, which indicated that the overall flow predictions were reasonable. Karidio et al. (65) measured the velocity distributions of cold air flow at the liquor gun level. The velocity boundary conditions for the mathematical modelling were determined from the pressure measurements at the primary and secondary windboxes. Subsequent comparison of the measurements and the computations yielded good agreement, showing that balanced air flow through opposed walls resulted in a central region of high velocity upwards flow.

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Wessel et al. (66) compared the modelling results for flow, combustion and fume formation to the performance test data for an operating Kraft recovery boiler. However, the gas flow and temperature distribution were not measured during performance tests. Visual inspection of furnace brightness through observation doors in the upper furnace was used. It indicated a bias in the temperature distribution towards the left-hand corner of the front wall.

2.7 Concluding Remarks Indeed a great deal of research has been performed on issues related to the combustion of black liquor and mathematical modeling of recovery boiler. While in some instances the results reported by different researchers are consistent with one another, in a surprisingly large number of cases, the results are contradictory. Some of these inconsistencies are undoubtedly due to differences in modeling methods and approaches but many result from the complex nature of recovery furnace and lack of needed CPU power at the time. It highlights the difficulty of studying such a complex phenomena. One of the phenomena that did give consistent results about a recovery furnace was the influence of the operation strategy on flow field i.e. “central chimney“ in a recovery furnace. Central chimney is formed as the primary and secondary air from the four walls converge, usually in the middle of a recovery furnace, displaying a strong upward flow. But influence of the furnace geometry on the flow field deserves further study. Research on the influence of the furnace geometry on the flow field has not provided consistent results and as such the impact of the variables like the relation between the total inlets area and the char bed surface area is still not fully understood. The issue that how the liquor-dependent parameters affect a recovery performance is one of the uncertainties. Dealing with these parameters in the in-flight model requires more study. Despite, a considerable effort to characterise and understand the recovery furnace, there is still no good explanation regarding the furnace performance based on the known properties of the liquor. Swelling is a good example in this regard. At the present swelling factor made in the laboratory under arbitrary conditions can be used as a basis for describing swelling in the furnace. One critical need is to tie the NO forming processes more closely to the combustion model and to define the rate formation and nature of it in the recovery boiler. A step in right direction is to identify the mechanism, dominating the NO production. NO model could answer some of the question regarding the NO forming processes. A wall in a recovery furnace is an important part of the burning surface when it comes to smelt reduction. In an ordinary recovery furnace the surface of the walls is ten times larger than the surface of the char bed. A droplet reaching a wall may continue interacting with gases in the furnace. In order to handle droplets reaching the furnace walls correctly, it is necessary to have a model considering the interaction of the droplets and the walls in a recovery furnace.

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3. Development of Model of Recovery Furnace model Used in This Work

It is clear from the literature survey that several issues are involved in obtaining an accurate representation of the flow field in a recovery boiler. The problems include the high numerical resolution required, especially at the air nozzle levels, and the large amount of computer time required to obtain convergence. Slow convergence of the solution algorithms is generally due to persistence of low frequency errors that are not effectively removed by a grid that is fine relative to the wavelengths of the errors. Kaul et al.’s (67) suggestion of using an unstructured grid system with local mesh refinement was adopted for the recovery furnace model. Several studies done by Tao et al. between 1992 and 1997 [13, 41, 42, 58] at the division of Energy and Furnace Technology, KTH, confirmed that use of an unstructured grid and local mesh refinement makes it possible to simulate the complex flow and geometry in a Kraft recovery boiler with realistic computing power. The code used in this work is a steady state/transient, finite difference, computational fluid dynamics program that can solve three-dimensional fields for pressure, velocity, temperature, kinetic energy of turbulence, dissipation rate of turbulence, and several chemical species. The code operates by solving the governing differential equations of the flow physics by numerical means on a computational mesh and is able to predict gas velocity, temperature profile, and concentration fields. The code is also modified to accommodate the unique characteristics of black liquor combustion. The black liquor combustion models interact with the gas fields and provide source/sink terms to the CFD equations. Black liquor burning rates depend on the gas fields and, in turn, influence them through the source/sink terms (Appendix 3). The boiler model is limited to the furnace cavity itself and terminates at the nose arch. However, it may be possible to carry flow and temperature calculations through the superheater.

3.1 Geometry of the Furnaces Used in This Work Combining the body-fitted meshing capabilities with unstructured non-orthogonal grids, local mesh refinement and arbitrary coupling between mesh blocks gives great flexibility in representing highly complex geometries. However, the geometry of a recovery boiler and recovery furnace in particular is not complicated. Eliminating the boiler bank on top of the furnace from the calculation makes the geometry even simpler. This elimination is reasonable since no information is available regarding the cooling mechanism of the passing gases, the mechanism of deposition of the carried-over materials on the tube and the correct presentation of gas passage area. The main problem in regard to the geometric representation of a recovery boiler is the shape of the char bed. A description of the char bed geometry as a boundary in the model will be given at a later stage. Two furnaces were chosen for study. The first operates with the conventional firing method with three elevations for air delivery, while the second boiler has four elevations for air delivery and uses a Rotafiring strategy on the secondary level.

15

3.1.1 Conventional Firing Strategy (CF) Air is delivered to the studied recovery boiler using the conventional firing method at three elevations. Primary and secondary nozzles are located in all four walls, but the tertiary nozzles are located only in the side walls. The furnace is 10.248 x 9.912 m2 in cross-section. The height to the midpoint of the bull nose is 27.4 m. Combustion air enters the unit through 179 inlets that can be individually set. Figure 2 The computational domain for the CF recovery boiler.

Primary air enters the furnace through 24 air nozzles on the right and left walls, and 26 on the front and back walls. The primary air system on all four walls is located 1.050 m above the floor. 30% of the combustion air enters the furnace through these nozzles. The temperature of the air at this level is 153°C. The velocity of the air entering the combustion chamber through these nozzles is between 18 and 36 m/s. Figure 3 The computational grids illustrating the shape of the primary nozzles.

16

Figure 4 Mesh defining the primary air nozzles.

Secondary air enters the furnace through 16 air nozzles on each wall (64 in total). The secondary air system on all four walls is 2.4 m above the floor. These nozzles introduce 55% of the combustion air. The temperature of the air at this level is 128°C. All the nozzles at this level were open at the time of measuring. The air is delivered to the combustion chamber through these nozzles with a velocity between 63 and 81m/s. Figure 5 Mesh defining the secondary air nozzles.

Tertiary air enters the furnace through 14 air nozzles, 6 on the front wall and 8 on the back wall. These nozzles introduce 15% of the combustion air. The tertiary air system is located 9.1 m above the floor. The temperature of the air at this level is 35°C. The velocity of the air entering the combustion chamber through these air nozzles is between 50 and 60 m/s. The geometry of this CF recovery boiler was represented in a three-dimensional data model consisting of a mesh with 242 000 cells. See Figure 6.

17

Figure 6 The computational grids used in the simulation of CF recovery boiler.

3.1.2 Rotational Firing Strategy (RF) Air is delivered to the studied recovery boiler using the Rotafiring strategy at four elevations. There are primary and secondary nozzles in all four walls but the tertiary and the fourth nozzle levels are located only in the side walls. Rotafiring was done by the secondary air nozzles, a number of which were closed. Reducing the total open nozzle area on the secondary level produced an initial jet velocity of 70 to 80 m/s. Since the jets work together, vigorous rotation of the gas in the lower furnace is achieved. The furnace is 10.5 x 10.5 m2 in cross-section. The height to the midpoint of the bull nose is 28.5 m. Figure 7 RotaFire secondary air nozzles configuration.

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Primary air enters the furnace through 33 rectangular nozzles on each side, 1.2 m above the floor. 38.6% of the combustion air enters the furnace through these nozzles. The temperature of the air at this level is 223°C. The air entering the combustion chamber through these nozzles has a velocity of between 18 and 36 m/s. Secondary air enters the furnace through nozzles arranged as follows: Front and back walls Port 1–3 and 13 are 100% open Port 14–16 are closed Ports 3–12 have their opening areas decreased by 10% Right and left walls Ports 1–3 and 11 are 100% open Ports 12–14 are closed Ports 3–11 have their opening areas decreased by 10% These nozzles introduce 43% of the combustion air. The temperature of the air at this level is 99°C. Figure 8 Secondary air nozzles configuration used in this work.

1 3 11 14

Rig

ht a

nd le

ft ai

r del

iver

y co

nfig

urat

ion

Front and back air delivery configuration 16 12 3 1

19

Tertiary air enters the furnace through 8 nozzles, 4 on the front and 4 on the back wall. These nozzles introduce 15% of the combustion air. The tertiary air system is located 10 m above the floor. The temperature of the air at this level is 36°C. Fourth air enters the furnace through 4 nozzles, 2 on the front and 2 on the back wall. These nozzles introduce 2% of the combustion air. The temperature of the air at this level is 36°C.

Figure 9 Computational grids used in the simulation of the RF recovery boiler.

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3.2 Black Liquor Combustion A model of black liquor combustion was built using STAR-CD. Such a model has to be able to account for what happens to a droplet based on its modelled trajectory. A droplet of liquor burns in a number of different ways in the recovery furnace, all of which must be accounted for. It may burn in flight, on the char bed, or on the walls, or may exit the furnace only partially burned (carry-over). The fully coupled recovery boiler model in this work has several interactive components:

• Built-in models in STAR-CD describing the flow, heat transfer and chemical reactions in the gas phase.

• A droplet trajectory model. • An in-flight droplet combustion model describing the characteristics of liquor sprays,

droplet motion and the physicochemical behaviour of droplets subject to local furnace conditions.

• A char bed model describing the interaction between the char bed reactions and the gas flow above the bed.

• A wall model describing the interaction between the walls and the gas flow. • A NOx emission model, which considers NO formation from thermal NO, fuel NO

and prompt NO mechanisms.

3.2.1 Spray Several types of commercial nozzles are used to inject black liquors into recovery boilers. The splash-plate type is the most common, but swirl-cone and V-jet types are occasionally used. While it would be expected a priori that droplet size would be determined by both properties of liquors and nozzle design, the relative importance of the various parameters has to be determined empirically. A successful black liquor distribution system must control the droplet size in order to minimise carry-over and deliver relatively dry liquor to the bed surface. To do this, liquor guns are arranged in a symmetrical pattern that distributes liquor evenly across the full cross-section of the furnace. The mechanisms for stream break-up exiting the liquor gun nozzle generally fall into two basic regimes: ordered and chaotic. The most common commercial spray nozzles and atomisers fall into the latter category, as do all current black liquor nozzles used in recovery boilers in pulp and paper mills [Green et al., (68)]. The injection velocity, that is, the velocity of the liquid fuel as it exits the nozzle and enters the furnace, and the viscosity of the liquor are the most important parameters in a spray calculation. They strongly influence the atomisation and break-up processes, the spray penetration, the interphase transfer processes, and droplet–droplet and droplet–wall interactions. In general, modelling of the spray can be done in two ways:

• Using the nozzle geometry (e.g., the diameter, D, of the nozzle hole) as a parameter [Adams (69)].

• Setting up the initial conditions obtained under laboratory conditions. Although these two approaches have different starting points, they should produce fairly similar results. The size distribution of the droplets should fall within the initial conditions, for these were set up to simulate the injection of the droplets by liquor guns into the furnace.

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The current model uses the second option. This option was chosen because setting up the initial conditions as the input parameter shortens the CPU time tremendously. In mathematical modelling of spray, groups of droplets with similar initial conditions are commonly assigned to classes or parcels. Each parcel is distinguished by the initial position, velocity, diameter, temperature and density of a typical droplet as well as by the number of droplets within it. A single droplet possessing the properties of that class represents an individual parcel. The trajectory and heat/mass transfer between the phases is then calculated based on the local gas phase conditions as a droplet parcel moves through the flow field [Star-CD (70)]. The characteristics of a black liquor spray are defined by its

• Initial position and velocity. • Temperature and density. • Size distribution.

The two first points can be can be obtained from known plant operation conditions and are given in Table 2. Table 2: Initial liquor conditions

Recovery Furnace Initial position

(m) Above the floor Initial velocity (m/s)

Temperature (K)

Load mtotal (kg) ds/s

CF 6,2 8 120 24,67

RF

6 7,4 115 19,67

In the present model, the velocity and size distributions of a black liquor spray are determined using the following relations. The mass carried by an individual droplet is:

d

totald n

mm = (1)

A Rosin-Rammler size distribution is assumed for the diameters of droplet parcels. The size distribution function has the following form [Tao (41), 1997]:

d dCVd dm

q

=−

ln

/100100

1

(2)

where ddm is the mean droplet diameter for a given nozzle, CV is the cumulative volume and q is a parameter that describes how wide the distribution of diameter is for the nozzle. Two parameters must be specified: the mean diameter and the distribution parameter. A splash-plate nozzle has the following spray distribution parameters: ddm = 3 mm and q = 2.7. These

22

parameters have been applied in the present work. Ten droplet diameters are used to represent the size distribution in the calculations. These diameters i

dd are calculated by the following relations:

q

jdmj CV

dd/

ln

1

100100

−= (3)

jCVj 10= 9993210 .,,....,,,,=j

)(. 150 ++= jjid ddd 10321 ,....,,,=i 999210 .,,...,,,=j

The number of droplets in a droplet parcel is given by:

d

id

pid

d

MN

ρπ 3

6

= (4)

The initial velocity vector of a droplet parcel is determined from known spray angles (horizontal and vertical) and known mass distribution within the spray angles. In the present work, a group consisting of ten droplet parcels with the diameter distribution described above is assigned a unique initial velocity vector. The magnitude of the relative velocity between the droplet and the gas implies that the Reynolds number is less than 103 and thus that the drag force is important in this case. The model uses the average gas velocity in a given computational cell to determine the drag on a liquor droplet passing through the cell. The trajectory used in this work takes into account the changes in mass and physical size of the droplets as they burn (Appendix 3). Knowledge of droplet velocity makes it possible to determine the position of a droplet at any moment, when it could be on the hearth, at a furnace wall, on the char bed or be being carried out of the furnace by flue gas. One current method for calculating droplet trajectories is based on the force balance. This method is not particular to recovery boilers and has yielded reliable models in other applications. The initial velocity and droplet size are liquor gun–dependent variables and should be considered carefully. The trajectory model should also account for the swelling/deswelling of a droplet during combustion. In this work the swelling/deswelling of a droplet is treated empirically. Generally there is a need to find a swelling/deswelling model that is dependent on the liquor properties and furnace environment. Trajectory calculations directly determine the amount of carry-over [Fakhrai (71)].

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Figure 10 Droplets trajectories and cloud of the drops injected by a liquor gun located on the back wall.

3.2.2 In-flight Drop Combustion There is still some uncertainty with respect to the model just described, particularly as regards modelling and accounting for the behaviour of the black liquor. The liquor-dependent parameters are among the uncertainties in the in-flight model. Despite considerable efforts to characterise and understand swelling, there is still no model available that can adequately estimate swelling based on the known properties of the liquor [Whitty (72)]. However, models of swelling produced in a laboratory under arbitrary conditions can be used as a basis for describing swelling in the furnace [Frederick (73)]. One of the most critical points in mathematical modelling is verification of the implemented models and, in the case of the recovery boiler, of the droplet burning models for the specific existing furnace conditions. Existing droplet burning models are based on laboratory experiments. These experiments have been carried out in the following conditions [Noopila (74)]:

• Black liquor droplets have been suspended in a laboratory furnace. • Droplets have been examined in laminar entrained flow reactors.

One problem is obtaining a proper description of the ratio processes that depend on the particular environment in which the droplet is present. Another is defining the stoichiometric partition factors and swelling factors that do not have an obvious counterpart in the furnace. There is a need for an experimentally validated droplet-burning model that involves only those variables that are meaningful in a CFD furnace model

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Appendix presents a complete description of the current in-flight model and the gas phase combustion reactions.

3.2.3 Char Bed Processes The char bed is one of the main differences between recovery boilers and boilers for virtually every other fuel. This difference is important because of the role of the char bed in achieving the key steps of chemical recovery and organic combustion in recovery boilers. The char bed is composed of the residual material from in-flight combustion that arrives at the hearth and accumulates in a pile. Char bed combustion is the second major mode of burning in a recovery furnace. The char bed provides a fuel reservoir that helps to stabilise the combustion process in the furnace. It also plays a significant role in maintaining high reduction efficiencies. It usually covers the entire floor area and rises in the middle to form a peak 1 to 3 m high. The char bed shape is the net result of material landing on the bed, physical entrainment of material along and off the bed surface, and combustion and smelting on the bed surface. A model of a the char bed must be able to accommodate the following variables: • The burning rates, considering both exothermic combustion with oxygen on the surface of

the bed and endothermic gasification of the carbon inside the char bed by CO2. • Interactions with the gas fields above the bed. • Accumulation and depletion of material and physical transport of char along and off the

bed surface in arriving at the ultimate size and shape. • Surface temperature of the char bed. • Temperature inside the bed. • Bed shape and stability There are two types of zones on the surface of a char bed, cold and hot spots. The hot zones are generated when the primary air hits the bed surface [Borg et al. (75)]. The average bed temperature is usually quite stable, and is close to the melting point of the system.

3.2.4 The Interaction Between Drops and the Walls in the Recovery Furnace The landing place and burning scheme are an extremely important part of any mathematical model of the burning of droplets of black liquor in a recovery furnace. The way in which wall burning is handled has a considerable effect on carry-over. A wall in a recovery furnace is an important part of the burning surface when it comes to smelt reduction. In an ordinary recovery furnace the surface of the walls is ten times larger than the surface of the char bed. A droplet reaching a wall may continue interacting with gases in the furnace. In order to handle droplets reaching the furnace walls correctly, it is necessary to have answers to the following questions: a) What happens to a droplet’s trajectory when it hits a wall? b) In the case of burning at the wall, which burning model should be chosen? In answering the first question, the following options are available. • Ignore the droplets that hit the walls. • Assume that liquor droplets will bounce off.

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• Assume that droplets stick to the wall and burn on the wall. Ignoring the droplets that hit the walls of the furnace means that one cannot account for the physical state of the droplets reaching the heat transfer surfaces in the upper part of the recovery furnace. The physical state of the droplet is a function of the temperature in the boiler, which would be affected if this approach were adopted. A second concern with this approach is that the temperature gradient near the furnace walls could not be correct if one ignored combustion of black liquor droplets at the walls. The second approach, which has the droplets bouncing off the furnace walls, is unrealistic since more than 85% of the droplets will reach the char bed in a different physical state. One consequence of this approach would be the creation of high temperature zones (up to 2000 K) in the lower part of the furnace close to the char bed. The existence of such high temperatures has neither been proved experimentally nor observed. From a mathematical point of view, the third approach also has some disadvantages. If all droplets hitting the furnace walls burn at the walls, the proper amount of fuel will not reach the bed and the temperature profile in the boiler will be affected (cold furnace). Furthermore, there is the risk of overflow at the walls (stopped calculation). Walls was previously considered to be an unimportant aspect of the burning process and generated no studies. However, it accounts, at least in part, for such observations as “high peak temperature”. The high peak temperature phenomenon shown by mathematical modelling in the lower part of a recovery furnace is complex and involves a number of factors. To address these concerns, the following assumptions have been made: • The droplets are considered single reactors working independently of each other. • The mass of the droplets is not spread uniformly but is induced at each droplet’s landing

place. • If a droplet sticks to a wall it is “alive” and reacts, and therefore it is also considered

during computations. Two burning models are available for use:

• Expand the char bed to the walls and use a burning rate similar to that on the char bed (char bed model).

• Assume that a particle reaching the wall continues to burn at the same rate as when it reached the wall (in-flight burning model).

The internal surface of the walls is divided into two areas, defined in terms of the local fuel and air injection and the furnace temperature distribution as shown in Figure 11. In the lower part of the furnace, where the temperature is high enough to sustain the reduction process at the walls, the char bed model is expanded to cover the entire wall surface area given as:

R at wall = R overall char bed (5)

where R at wall is the char combustion rate at the walls and R overall char bed is the same rate at the char bed. In the upper part of the furnace, the droplets hitting the walls will adopt the second

26

burning regime, which is continuation of combustion at the same rate as in-flight combustion given as:

R at wall= R combustion stage (6)

where R combustion stage is the combustion rate in devolatilisation or char burning stages.

Figure 11 The adaptive wall burning model

A realistic wall-burning model must also allow for the fact that liquor accumulates on the walls in the furnace and then periodically sloughs off and falls to the bed. This can, in fact, be a very significant way that fuel and inorganic materials reach the bed. In reality, the only factor affecting these processes is the force of gravity on the accumulated material. At the present time, insufficient computing power and the risk of overflow make it impossible to account for every droplet in a recovery furnace. In order to overcome this problem, the following approach is suggested: 0 mdrop > C

mcell= ( 7)

mdrop 0<mdrop<C

When a droplet reaches a cell at a wall and the cell is empty, the droplet will stick to the wall. The mass carried by the droplet, mdrop, will be assumed as mass in the cell, mcell. This mass will be introduced as appropriate gas species into the gas phase according to the adapted wall-burning model. In the second case, where another droplet, is present the droplet will slough off and re-enter the gas phase. C is a constant, calculated based on the volume of the cell in

The primary air nozzles

The Char bed Burning model

The In-flight burning model

The secondary air nozzles

The tertiary air nozzles

27

question, which prevents overflow in calculation (calculation stops). Figure 12 illustrates a) accumulation of droplets on the wall b) sloughing off of droplets

Figure 12 Schematic representation of the wall in case of a) accumulation of the droplets on the wall b) sloughs off

a) b)

28

3.3 NO modelling There are several main points to control a combustion process in a recover boiler. The first is the selection of a proper, clever, operation strategy with intention to ensure the safety, efficiency and continuity in operation. The second main point to a combustion process is that to reducing the air pollution down to a level which meets the current and forthcoming regulations. This is due to the fact that the products of combustion are distinctly identified as a severe source of environmental damage. Less obvious products of combustion are NO and NO2, collectively called NOx. Now, it has become apparent that NOx is a major contributor of photochemical smog and ozone in the urban air and participates in a chain reaction removing ozone from the stratosphere. The consequence of this removal is increased ultraviolet radiation to the earth’s surface. Three different routes are now identified in the formation of NO. These are thermal, the prompt, and the fuel-bound nitrogen route. The mean NO concentration is obtained by solving is transport equation based on the flow field and combustion solution.

3.3.1 Thermal NO The conditions in the furnace of a recovery boiler are extremely heterogeneous, with significant local variation. The concentration of constituent gas components, temperature level and gradient, and flow field in the furnace of e recovery boiler are affected by these variations. Although at this stage the level of these variations and their effect on the NO productions rate can’t be measured directly, they can be calculated by mathematical modelling. If the temperature inside a furnace is sufficiently high, thermal NO will be the formed by the there elementary reaction known as Zeldovich mechanism:

O +N2 1+

↔k

N+NO (8)

N+O2

2+

↔k

O+NO (9)

N+OH3+

↔k

H+NO (10)

Based on the quasi-steady state assumption for N, d[N]/dt=0, it can be shown that the, maximum NO formation rate is given by:

[ ]=

dtNOd 2k1[O2][N2] (11)

ki is the forward reaction rate constant for reaction(1), and it is equals:

k1= 1.8.108exp(-38400 /(RT)) m3/(mol.s) (12)

An approximation for [O] based on the Partial Equilibrium results in

29

[O]=36.64 T0.5[O2]0.5 exp (-27123/T) (13)

3.3.2 Turbulence Chemistry Interaction One fundamental feature of the combustion process in a Kraft recovery boiler is that the flow is fully three-dimensional and turbulent. There are two consequences from the interaction between turbulence and chemistry. Turbulence can enhance the mixing of the reactants but also influences the mean reaction rate. Several investigators [Bai, (76), Weber et al.(77)] have shown that the mean reaction rate, R( )φ , is sensitive to the method used to calculate R( )φ . Note that R R( ) ( )φ φ≠ . Additional effort is therefore necessary in order to calculate the mean reaction rate correctly. Many simulations of industrial boilers have been based on the Eddy-Dissipation Concept (EDC) proposed by Magnussen and Hjertager (78). The EDC model assumes that the chemical reactions are fast in comparison with the mixing of the reactants and the heat transfer, thus the mean chemical reaction rate in turbulent combustion will depend on the slow process, i.e. turbulent mixing and heat transfer. With certain calibration of model constants, the EDC model has been very successful in the simulation of industrial combustion processes, since the hydrocarbon combustion is normally much faster than the mixing process. The Damköhler number (the ratio between the time scale of mixing to the time scale of chemical reaction) is typically in the order of tens of thousands to millions. The EDC model has also been used for computing fuel-NO formation by several investigators [Lindsjo (79) as the fuel-NO path involves the chemical reactions in which the Damköhler numbers is high. In this study, the EDC model is applied to the calculation of the main combustion reactions and the fuel-NO formation. Bai, Xue-Song (80) indicated that the Damköhler number shows that thermal-NO formation is essentially a slow reaction compared with turbulent mixing. Our experience shows that the thermal-NO formation calculated using the mean gas temperature is very low ( less than 2 ppm), which suggests that the PDF model should be considered. The PDF approach used in this study is a presumed single-variable PDF model by Hand et al. (81) with the gas temperature being the only fluctuating variable. This approach gives:

∫−−=−

bT

uT uTbTdT

baTPTNOtRNOtR ),,()( (14)

where a and b are parameters of the PDF function. It is assumed that the distribution of the Probability Density Function follows the β function:

∫−−−

−−−= 1

0111

111

dfbfaf

bfafbaTP

)(

)(),,( (15)

uTbTuTT

f−−

= , uTbT

uTTs

sa−

−−=

1 , uTbT

TbT

ssb

−

−−=

1 (16)

30

“s” can vary from 0.5 to 1. A value of 0.6 has been recommended by Missaghi (82). The other important parameters in the PDF model are the lowest possible value, Tu and the highest possible value, Tb of the fluctuating temperatures. The choice of the unburned temperature, Tu, is rather straightforward. It is taken as the inlet temperature for the combustion air. The value of Tb is locally calculated throughout the furnace. For this purpose, the local stoichiometry λ is defined as:

2244

2

HmHsCOmCOsCHmCHsOm

++=λ (17)

where mO2, mCH4, mCO and mH2 are the mass fractions for oxygen, methane, carbon monoxide and hydrogen while sCH4, sCO and sH2 are their stoichiometric oxygen requirements (kg/kg) respectively. An enthalpy increase ∆h which occur if all the combustibles are burned, can be calculated as:

if λ ≥ 1, 2244 HmHHCOmCOHCHmCHHh ++=∆ (18)

if λ < 1, ( )2244 HmHHCOmCOHCHmCHHh ++=∆ λ (19)

where HCH4, HCO and HH2 stand for the lower calorific values for methane, carbon monoxide and hydrogen gases. The resulted temperature increase can therefore be determined by:

pCh

T∆

=∆ (20)

The burned temperature, Tb, is then calculated as: TTbT ∆+=

3.3.2 Prompt NO A roughly estimated prompt-NO formation rate for methane combustion was given by De Soete 1975 and in terms of concentration it reads (in gmole/cm3s).

[ ] [ ] [ ][ ] )exp(

RTE

FuelNOCfdt

NOd abP −××= 22 (21)

C=6.4*106 s-1 According De Soete, the coefficient f depends on the number of carbon atoms in a fuel molecule and the equivalence ratio φ(=stoichiometric air demand divided by actual air amount).

31

32 21203222308190754 ϕϕϕ ..... −+−+= mf pr (22)

where m is the number of carbon atoms in the hydrocarbon. These values are valid for aliphatic hydrocarbons and for equivalence ratios between 0.62 and 1.43. The reaction rate coefficient and activation energy depend on the fuel species. And b varies with the oxygen mole fraction [O2] as follows,

b=1.0 [O2]<0.0041 b=-3.95-0.9ln[O2] 0.0041 ≤ [O2]<0.0111 b=-0.35-ln[O2] 0.0111 ≤ [O2] ≤ 0.03

b=0.0 [O2]>0.03 The effect of fluctuating temperature is accounted for by Beta-pdf function in the same way as for thermal-NO.

3.3.3 Fuel NO Brink (83) proposed a NO model for black liquor droplets. The model assumed that the fuel N was released in the gaseous phase via devolatilisation and char combustion. During devolatilisation, the amount of nitrogen released depends on temperature and increases as the temperature increases. Laboratory studies have found that approximately 70% of the fuel nitrogen is released during devolatilisation, mainly as NH3 and N2.

NBL

Nvol

Nchar

N2

NH3

NO

N2NO

Nsmelt

NO

Ngreen liquor

Figure 13 Fuel Nitrogen Pathways for Black Liquor Combustion [Brink 1995].

The nitrogen remaining in the char after devolatilisation is either oxidised to NO or captured by the smelt. It was assumed that the fuel N is released in the gaseous phase and rapidly converted into an intermediate nitrogen radical, often identified as NH3, which may then follow two alternative reaction paths: an oxidation to NO or a recombination with NO to form N2. The formation rate of NH3 is assumed to be proportional to the release rate of fuel N.

32

The release rate of the NH3 is proportional to the total devolatilisation rate. The fuel NO chemistry in the gas phase is difficult to model since the reaction rates are highly dependent on radical concentrations. De Soete [84] and Mitchell et al. [85], proposed the following overall reaction scheme for fuel NO formation:

2223 50 HOHNOONH Ak .++→+ (23)

2223 50 HOHNNONH Bk .++→+ (24)

The overall reactions for fuel NO formation indicate that NH3 is both a NO producer and a NO reducer. The rate of fuel NO formation is dependent upon the local oxygen concentration. The eddy-dissipation concept (EDC) is used to calculate the rate of fuel-NO formation. By EDC model, the mean reaction rate is:

κερ

=

SRmmCR r

fEDC ,min (25)

where fm is mass fraction of the fuel , rm is the mass fraction of reactant SR is stoichiometric ratio. C is EDC constant, and ε and Κ are the local kinetic energy of turbulence and its dissipation rate respectively. Iisa et al. (86), 2000, reviewed the NOx formation and reduction mechanisms in a recovery boiler. They proposed algorithms for predicting NOx in a recovery boiler based on the three NOx formation mechanisms; thermal, prompt and fuel NOx. The basic information on reduction of the NOx in a recovery boiler is also available and interacts with the NOx formation mechanism.

33

4. Results and Discussion

4.1 Estimation of Surface Temperature and Mass Distribution on a Char Bed As described in the previous sections, a substantial fraction of the total combustion occurs on the char bed, and therefore the corresponding amount of the total air should be supplied to the bed. Overall bed control (shape of the bed, burning rate, reduction efficiency, etc.) is of the utmost importance for stable operation of a recovery furnace and is associated with char bed inventory and surface temperature levels. The bed surface temperature, in turn, is critical to sulphate reduction fume formation and combustion stability. Mass distribution on the char bed. The new model used in this work enables the visualisation and calculation of the mass distribution on the bed. Figure 144 shows the droplets’ distribution on the char bed surface in the CF recovery furnace. The amount of mass (droplets) reaching the bed depends on the flow field, the temperature gradient in the furnace and the diameter size distribution. The trajectory of the droplets indicates the following: • Droplets swell considerably during devolatilisation, which may cause them to be carried

into the upper part of the boiler, and then fall back to the char bed during the char combustion process when they shrink (with a resultant increase in density) (Figure 15a).

• Droplets with an initial diameter of 5 to 6 mm fall directly to the bed (Figure 15b). • Droplets hit the wall in the lower part of the furnace and, after accumulation, slough off

the wall.

Figure 14 Predicted contours of surface temperature (Kelvin) on the char bed and black liquor droplets distribution on the char bed.

34

Figure 15 a) A droplets diameter and trajectory b) The dropletss trajectories

Using results obtained from the model, it is possible to estimate the initial size of the droplets that hit the char bed (Appendix 2). The majority of the droplets that went directly to the bed had an initial diameter between 5 and 6 mm. In mathematical modelling of a bed in this work, the size of a bed is characterised by its height. Here the height of the bed was assumed to be 2 meters above the floor. In general, the adoption of a constant bed height is required as a starting point in a steady-state furnace model. At the current state of CFD, it is very difficult to calculate the shape of the char bed because of the CPU demand. The growth rate of the char bed reflects the dynamic balance between the combustion of the material on the bed and the smelt flowing out. The growth rate of the bed could be calculated in a time-dependent simulation of a recovery furnace. In time-dependent simulation case the main steps to study growth of a char bed would be as follows: • Calculate the fall-out sites by direct numerical simulation (mass distribution on the bed). • Eliminate the droplets based on their mass content and the temperature of areas of the char

bed surface The shape of the bed is highly dependent on the flow field in the lower part of boiler.

a) b)

35

Char bed surface temperature. The present model reveals the relation between the high temperature areas and the mass distribution on the surface of the bed. This makes it possible to link the bed surface temperature to the properties of the liquor (i.e. to swelling). Having done that, the reduction efficiency can also be defined as a function of the area of the active zone and indirectly linked to the liquor properties. As shown in Figure 14 the hottest zones on the char bed surface are droplet fall-out regions. There are areas at the corner of the boiler that the droplets do not hit, and consequently these areas have a lower surface temperature. The temperature is highest on the surface of the bed and decreases towards the interior due to the endothermic reduction and sodium volatilisation reactions occurring within the bed. The most critical thermal parameter for sustainable combustion in a recovery furnace is the average bed surface temperature, which determines surface reaction rates as well as the amount of heat being carried into the bed, and hence the extent of reduction and fume formation within the bed. In this work, the highest calculated surface temperature is about 1400 K and the lowest temperature is about 550 K. The temperature of the bed is highly affected by the primary and secondary air input, the liquor injection method (angle of injection) and the flow field close to the bed. .

Figure 16 Predicted contours of surface temperature (Kelvin) on the char bed

36

4.2 Effect of the Physical Properties of the Liquor on The Furnace Performance.

The chemical composition of black liquor depends on the raw material, pulping conditions, and the process parameters used in the transformation from weak to strong liquor. The physical properties of the liquor are a function of its chemical composition. The chemical composition and the physical properties govern the behaviour of the liquor when it is processed in the mill recovery unit. It is of interest to know how the properties of black liquor change the conditions prevailing in a recovery furnace. Studying how swelling and the solid content of liquor affect the flow pattern, temperature profile and the like contributes to an understanding of associated operation strategies. As regards evaporation and combustion, the most important physical properties used in the model are viscosity, solid content and swelling tendency of a black liquor droplet. There have been many attempts to model these physical properties and most of them have been verified on the basis of the conditions prevailing in a laboratory. Therefore the relationship between the temperature and the solid content of a droplet could be used with confidence in the numerical simulation. Viscosity is the physical property that plays an important role, changing the initial droplet size. It shows both Newtonian and non-Newtonian behaviour dependent on the solid content of the droplet. It is affected by liquor temperature, which alters the viscosity and thereby changes the liquor composition, particularly the solid content and NaOH concentration [Blackwell et al. (87)]. A factor that currently limits the ability of the model to deal with the viscosity of a black liquor droplet is the lack of an adequate model linking viscosity to the burning model in a recovery furnace. The main approach in this work has been to simplify the model to a limit that a modern computer is able to handle. One option is to eliminate the spray model using appropriate droplet diameter distribution.

4.2.1 Effect of the Swelling on the Path of a Droplet In heterogeneous reactions, such as the combustion of a black liquor droplet, in which several processes take place simultaneously, the mass transfer of the reacting gas species from the droplet to the bulk gas is controlled by the surface of the droplet and its swelling. In short, the temperature increase in the droplet is the driving force in the drying and devolatilisation stages. The temperature also affects the swelling. While it has no effect on the maximum volume attained by a black liquor droplet, it changes the rate of swelling. This change in the rate of swelling also affects the timing of drying and pyrolysis and the path a droplet may take depending on the existing condition in the relevant part of a furnace. Each of these processes requires a certain amount of information as input to the model. As a first step, the following analysis was done in order to understand the impact of swelling on a droplet’s path. A black liquor droplet with swelling factors of 3 during drying and 6 during devolatilisation (drdry = 3 and drmax = 6) was defined. Once the swelling factors have been defined, it is possible to estimate the drag, gravity or other forces acting on the droplet during each combustion step and their contribution to the trajectory of the droplet. It is particularly interesting to know the location of an individual combustion stage in relation to the flow pattern, which can overwrite or reinforce the effect of the swelling and carry the droplet no matter what the level of swelling. The controlling forces that determines the path of a droplet are:

37

1) the initial velocity of the droplet. 2) gravity. 3) drag Other forces that affect a droplet’s path, such as pressure, are assumed to be negligible. Near the walls of the furnace, where most of the drying takes place, the flue gas velocity is low. The gravity force is downwards and the initial velocity gives the droplet momentum in the horizontal direction. The drag coefficient, which is used to calculate the drag force, is obtained by using the Reynolds number and the relative velocity of the droplet. The relative velocity is calculated using the velocity of the droplets and the gas velocity. Figure 17 shows that the Reynolds number in a recovery furnace is less than 103. Other forces that affect a drops path, e.g. pressure force are assumed to be negligible. Near the walls of the furnace where the most of the drying takes place the flue gas velocity is low. The gravity force in downward and the initial velocity give the droplet momentum in the horizontal direction. The drag coefficient, which is used to calculate the drag forced, is obtained based on the Reynolds number using the two velocities, namely the droplets and the gas velocities to estimate the dropels relative velocity. Figure 17 shows that the Reynolds number in a recovery furnace is less than 103.

Figure 17 Predicted diameter for a droplets (drdry=3 and drmax=6) and calculated Reynolds number

As it is shown in Figure 18, an increase in the droplet diameter in a recovery furnace does not automatically result in an increase in the drag coefficient. This is largely due to the fact that there are a number of slow recirculation areas around the liquor guns in a recovery boiler using CF strategy. The average gas velocity in these areas is around 4 m/s. Therefore a droplet’s trajectory will be dominated by gravity. According to the model, the devolatilisation process begins when all water has been removed. Since drying is a slow process, a droplet has enough time to reach the plug flow area before devolatilisation begins. In the plug flow, the gas velocity is the dominant force on the trajectory of the droplet. The effect of the gravity force diminishes as volatiles enter the gas phase and the mass of the droplet decreases rapidly. When the swelling increases from 3 to 6 times the diameter of the initial droplet, the only noticeable effect is the shorter time it takes a droplet to change its path during devolatilisation. This is due to the dependence of the duration of devolatilisation on the surface area of the droplet.

38

Figure 18 Predicted diameter for a droplets (drdry=3 and drmax=6) and calculated drag coefficenet In the second step, the effect of swelling on the overall performance of a recovery furnace was investigated by simulating a CF recovery furnace using liquors with different swelling factors. The swelling factors (drmax) used in this work were 2.5 (case a) and 5.5 (case b). The burning rate and duration of a liquor droplet is a function of swelling of the droplet. The duration of the combustion stages is also affected by the swelling of the liquor, as shown in Figure 19 Figure 19 Predicted diameter of a droplet with drmax=2.5 (blue line) and drmax=6 (yellow line)

Dro

p di

amet

er (m

m)

Time step (S)

39

Figure 20 Predicted velocity (m/s) vectors plot in the front-rear central vertical plane of a CF recovery furnace, case a left, case b right

As an example, of the effect of the flow pattern on a droplets path in a boiler with a “plug flow pattern” in the middle the effect of swelling on a flow pattern is relatively limited. In this study, the swelling of black liquor droplets during drying is quite low: 1.5 (case a) and 1.75 (case b) of the initial diameter. The time required for the complete evaporation of water is longer than the time required for pyrolysis and char burning. Drying takes place near the walls where the liquor sprays into the boiler. Figure 2020 shows that recirculation zones with relatively low gas velocities usually dominate these areas. The temperature profiles for the recovery boiler in the two simulations are presented in Figure 21. The results from the two simulations are somewhat different. It is clear that swelling affects the overall performance of a recovery furnace. The swelling during drying affects the trajectories of the droplets and therefore has a direct effect on the location of the combustion zones and the temperature profile in a recovery furnace. The swelling during devolatilisation affects the burning process, not because of the change in the trajectory of the droplets, but in terms of where along the trajectory the volatile evolves. Thus the swelling during devolatilisation has an indirect effect on the trajectory. By changing the duration of the devolatilisation stage, it changes the temperature profile and the flow pattern as a result of the density variation. It is clear from the figure that the swelling of the liquors controls the combustion zones. Where the swelling is greater, the combustion zone extends beyond the tertiary air nozzles. The high temperature area (flame volume) is also larger and more symmetrical, and consequently the thermal field is more uniform in the combustion chamber. Low swelling of the liquor droplets makes the combustion somewhat unstable.

40

Figure 21 Predicted contours of gas temperature (K) in the front-rear central vertical plane case a left, case b right

4.3 Effect of the Wall-Burning Model on the Overall Recovery Furnace Model Performance

If a model is to contribute to an understanding of thermal behaviour, overall emission levels, and solid phase entrainment (carry-over), it needs to take into account the interaction between black liquor droplets and the walls of the recovery furnace. The model presented here allows material and energy balances to include the liquor burning on the furnace walls by taking into account both the accumulation of droplets on the walls and sloughing from the walls. It does this by comparing the droplets’ burning rate (the mass left to burn at the wall), the mass already attached to the walls. Development and implementation of the wall model allows material and energy balances including the liquor burning on the furnace walls. The model considers: • Accumulation of the droplets on the walls • Sloughing from the walls t would be of interest to study the temperature profiles of the recovery furnace. Unfortunately, measurement limitations have made such investigation very difficult. In order to generate enough information for accurate analysis, the best we can do at this point is refer to parametric studies. The CF recovery furnace was used in this part of the study. Two simulations were carried out, one with and the other without the adaptive wall model. In case a) the wall model was modified allowing the droplets to bounce off and lose 70% of their momentum. In case B, the

41

adaptive wall model was employed allowing the droplets to stick to or bounce off the wall. Several parameters were considered for each condition and assumption. Parameters of interest were • temperature profile in the furnace cavity • the amount of the black liquor mass burned on the walls • estimation of the carry-over in the heat transfer surfaces In the spray model with an initial velocity of 6–8 m/s, the droplets reach the wall relatively fast. This is because the liquor guns are usually located in the area where the slow (4–5 m/s) and large recirculation dominates the flow pattern. The droplets reaching a wall are usually still in the drying stage (incomplete). In case A, the results indicate that more than 85% of the droplets reach the char bed, mostly after colliding with the walls and returning to the calculation domain. The bed is therefore growing and the liquor-firing rate has to be reduced to get the char supply and depletion to match. Of the remaining droplets, 14% burn in flight and 1% exit the furnace as carry-over. The fact that 85% of the droplets hit the char bed could be a very significant contributor to the high temperature spot near the char bed calculated by the code. In case B, only 72% of the droplets entering the furnace reach the bed. Of the remaining droplets, 14% burn in flight, and 0.5% exit the furnace as carry-over. Despite fewer droplets hitting the char bed, the amount of carry-over decreases by 50% compared to the base case. The remaining 13% of the droplets stick to the walls and continue to combust there. These results suggest that the combustion of black liquor on the wall is a significant consumer of black liquor. The max temperature in the recovery furnace was also lower in this case. Figure 22 shows the results of the simulations. The second result obtained from this model is that 88% of the droplets that hit the walls, hit the walls in the lower part of the furnace. In this part of the furnace the temperature of gases is relatively high, 1000 to 1400 K. This may favour the use of the wall model based on the char bed combustion model.

Figure 22 The faith of the droplets in case of using the bounced-off model and wall-burning model

0%10%20%30%40%50%60%70%80%90%

100%

C harbed W all In -fligh t C arry-over

Perc

enta

g of

the

drop

s

B ounced O ffW all B urn ing M odel

42

4.4 Prediction of the NO Level in Recovery Furnace Substantial resources were allocated to modelling NOx production/reduction. The results enable some preliminary comments on this issue. As explained in Appendix 3, correct prediction of flue gas temperature and temperature distribution in the furnace is of the utmost important. Any model of temperature distribution over the furnace cavity should include the regions of high gas temperature, indicating the intensity of combustion and the rate of temperature changes across the furnace cavity. In the lower furnace, the gas temperature distribution is characterised by a relatively cold region (the drying zone) and a relatively high temperature zone (devolatilisation region). The distribution of oxygen concentration in the furnace is also critical to the NOx model. The contours of oxygen concentration indicate at which level the combustion of combustible gases is completed, and the predicted final level of oxygen concentration should agree with the measurements in the boiler. In this study, the individual contributions of the fuel NO, thermal NO and prompt NO mechanisms have been investigated (Appendix 3). The nitrogen content in the combustion air and the high temperature cause the emission of thermal NO. The temperature in a recovery boiler is lower than that in other types of boiler. The result shows that the individual NO emissions due to thermal NO mechanisms are 5 ppm in the high temperature area therefore the thermal NO mechanisms is could be irrelevant. Figure 28 shows the predicted contours of fuel NO concentration (mass fraction) in the front–rear central plane. This figure shows that the region of high NO concentration is generally located in the centre of the lower furnace. The emission of fuel NO is caused by the small nitrogen content in the black liquor during the devolatilization. This content is approximately one order of magnitude lower than what is normal for other types of fuel, such as biomass and coal. The result shows that individual NO emissions in the lower part of the furnace due to fuel NO mechanisms are 70 ppm at an average of 4% O2. Figure 27 shows the predicted contours of prompt NO concentration (mass fraction) in the front–rear central plane. This figure shows that the region of high NO concentration is generally located in the lower furnace and near the char bed. The emission of prompt NO is caused by the CH radical and in this case is based on CH4. CH4 is produced by the devolatilisation process and char gasification in the char bed. The amount of CH4 from the devolatilisation process is very low compared with the CH4 from the char gasification at the bed. The results show that the individual NO emissions at the char bed level due to prompt NO mechanisms are 5 ppm at an average of 4% O2. The NO concentration is reduced as the gases move up to the bull-nose of the furnace. The predicted average NO concentration at the bull-nose level is about 75 ppm at an average of 4% O2, which differs from the measurements from the actual boiler. This is due to the fact that the complete fuel-NO mechanism is not implemented. At this time and the current NO model, the results are not conclusive and more must be done in this area. Fuel NO seems to be the only contributor to the NO production and therefore deserves more attention. The complete path of NO production should be considered in future.

43

Figure 23 Computational grids for the recovery boiler and contours of gas temperatures (Kelvin) in the front-rear central plane

Figure 24 Predicted contours of gas temperature (Kelvin) on the char bed

44

Figure 25 Predicted contours of oxygen concentration

Figure 26 Predicted contours of NH3 concentration

45

Figure 27 Predicted contours of prompt NO concentration

Figure 28 Predicted contours of fuel NO concentration

46

5. CONCLUSIONS This thesis concerns the mathematical modelling of a recovery furnace with the overall goal of developing new models and modifying existing ones in order to keep the code state-of-the-art. The conclusions that can be drawn from this work are presented below.

• As a starting point of the simulation, char bed size, shape and surface temperature, which are critical for stable operation of a recovery furnace, were assumed to be constant. Based on the number of droplets hitting the bed and the amount of the mass carried by the droplets, a limited growth rate of the bed was predicted. The surface temperature of the char bed was also obtained, based on the mass distribution on the bed.

• The physical properties of black liquor used in the model were investigated in a

parametric study focusing on the effect of swelling on the performance of a recovery boiler. The results show that swelling affects the combustion process in the recovery boiler. The swelling during drying affects the trajectories of the droplets and therefore has a direct effect on the location of the combustion zones and the temperature profile in the boiler. However, the swelling during devolatilisation has an indirect effect on the trajectory by changing the duration of the devolatilisation stage. Swelling changes the temperature profile and the flow pattern as a result of the density variation. It is clear that the swelling of the liquors controls the combustion zones. Where there is higher swelling, the combustion zone extends to the tertiary air nozzles, and the high temperature area (flame volume) is larger and more symmetrical, resulting in a more uniform thermal field in the combustion chamber. Low swelling of the liquor droplet makes the combustion somewhat unstable. The flow has a tendency to move from the right to the left side of the furnace. The effect of the solid content of the droplet on the temperature profile was not clear. Controlling the effect of swelling involves the operation strategy (air system strategy) adopted. In the case of a conventional, three-level, combustion air system (the plug flow in a recovery boiler), the swelling of a droplet during devolatilisation has a limited direct effect on the trajectory of that droplet since the velocity of the surrounding gas predominates over other forces.

• A model of the interaction of the droplets and walls (the burning process at the walls)

is proposed and implemented. The results suggest that at-wall combustion is important when calculating the peak temperature in the furnace and the amount of carry-over.

• Modelling of NOx formation in a recovery furnace showed that thermal and prompt

NO have a marginal effect on the overall NO formation. The fuel NO level showed that the formation of the NO in a recovery furnace is liquor-dependent and the level of N in the liquor is important in this regard.

• The major conclusion of this work is that a recovery furnace model based on solving

the equations describing flow reactions is a strong and flexible way of obtaining numerical solutions. The model maximises physical insight into an actual recovery furnace and is capable of accommodating a large variety of numerical methods and different systems of equations. These characteristics make it one of the first choices an investigator should consider when choosing a numerical method for solving simple and complex problems. At this point, the model should be considered a tool in

47

improving recovery boiler performance. Information dealing with modifications of the model or introducing new models is being developed.

The results from each step of the process have been reported here. It is hoped that the work will encourage researchers to tackle new problems using models and to develop new methods as well as extensions and improvements of the methods presented here.

48

References: 1 “ Energiförbrukning i Massa och Pappersindustrin “ ÅF-IPK, 1994. 2 Brännland, R., Hartler, N., “Environmental Protection in Pulp and Paper Industry,“ Department of Cellos Technology, Royal institute of Technology, KTH, 1992 3 Green, R. P., Hough G., “ Chemical Recovery in the Alkaline Pulping Processes,” ISBN 0-8985-255-2, TP B-046, 1992 4 Blasiak, W., Lixin, T., Vaclavinek J., “Recovery Furnace Modelling”, Preliminary study, Nutek/IEA. Project no 92-09979, Department of Energy and Furnace Technology Stockholm, Sweden,1992 5 Jones, A. K, Grace, T., “ A Comparison of Computational and Experimental Method for Determining the Gas Flow Patterns in the Kraft Recovery Boiler,” Tappi Engineering Conference, 1988 6 Grace, T., Walsh, A., Jonse, A., Sumnicht D, ”A Three- Dimensional Mathematical Model of the Kraft Recovery Furnace” US Department of Energy, Black Liquor Recovery research Nov, 1989 7 Llinares, V. Jr, Chapman, P. J., “Combustion Engineering Update Stationary Firing, Three Level Air System Retrofit Experience” Engineering Conference, page 629, 1989 8 Perchanok, M.S., Bruce, D.M., Gartshore, I.S., ”Velocity Measurements in an Isothermal Scale Model of Hog Fuel Boiler Furnace,” J. Pulp and Paper Science, Vol. 15, no. 6, November 1989 9 Bergman, J., Hjalmarsson, L., “Rotafire: A New Recovery Boiler Operation Concept”, Inter. Chemical Recovery Conference, 1992 10 Jones, A. K, Chapman P. J., “Computational Fluid Dynamics Combustion Modelling: A Comparison of Secondary Air System Designs”, Tappi Journal vol. 79, no. 7, 1993 11 Salcudean, M., Gartshore, I. S., “Modelling of Black Liquor Recovery Boiler”, DOE Program Review, April 8–10, 1991 12 Salcudean, M., Abdullah, Z., “Mathematical Modelling of Recovery Furnace”, International Chemical Recovery Conference, Seattle WA, June 7–11, 1992 13 Tao L. “Three-Dimensional Computer Simulation of the Flow and Combustion in a Kraft Recovery Boiler”, 3rd Asian-Pacific International Symposium on Combustion and Energy Utilisation, Hong Kong, Dec 11–15, 1995 14 Dwyer, H., “Calculation of Droplet Dynamics in High Temperature Environments”, Progress in Energy and Combustion Science, 1989, pp. 131–158, 15 Lintries, G. T., “Droplet Dynamics in a Non-Uniform Flow Field”, Combustion Science and Technology, 1991, pp. 319–335

49

16 Stockel, I. H. US Department of Energy Tech. Rep. CE/40626 T4, 1988 17 Spielbauer, T. M., Adams T. N., Monacelli, J. E., Bailey, R. T., “Droplet Size Distribution of Black Liquor Spray”, International Chemical Recovery Conference, 1988 18 Horton, R. R., Grace, T. M., Adams, T., “The Effects of Black Liquor Spray Parameters on Combustion Behaviour in Recovery Furnace Simulation”, International Chemical Recovery Conference, Seattle WA, June 7–11, 1992 19 Empie HJ, Lien SJ, Yang W, Samuels DB. J., Pulp Paper Sci.;21(2):j63–j67,1995 20 Helpiö, T. E. J., Kankkunen, A. E. P., Tappi Journal; 79(9):158–163,1996 21 Hupa, M., Backman, R., Hyöty, P., “Investigation of Fireside Deposition and Corrosion in Sulphate and Sodium Sulphite Recovery Boiler”, Proceedings Black Liquor Recovery Boiler Symposium, Helsinki, Finland, 31 Aug.–1 Sep. 1982, pp. D1–17 22 Hupa, M., Solina, P., Hyöty, P., “Combustion Behaviour of Black Liquor Droplets”, J. Pulp Paper Sci. 13(2): J67–72, 1987 23 Fricke, A. L., “Physical Properties of Kraft Black Liquor”, Summary Reports: Phases I and II. DOE/CD/406006-T5 (DE88002991), Sept. 1987 24 Jones, A. K ., “A Model of the Kraft Recovery Furnace”, Institute of Paper Chemistry, Appleton, Wisconsin, Jan. 1989 25 Frederick, W. J., “Combustion Processes in Black Liquor Recovery: Analysis and Interpretation of Combustion Rate Data and Engineering Design Model”, US Report DOE/CE/40637-T8 (DE90012712), 1990 26 Frederick, W. J., Noopila, T., Hupa, M., “Modelling of Black Liquor Droplets Combustion”, Dept. of Chemical Engineering, Åbo akademi, Combustion Chemistry Research Group, Report 89-15, Oct. 1989 27 Walsh, A. R. “A Computer Model for In-Flight Black Liquor Combustion in a Kraft Recovery Furnace”, Institute of Paper Chemistry, Appleton, Wisconsin, Jan. 1989 28 Hyöty P., Uppstu E., “Alternative Air Register Arrangement for Recovery Furnace”, Proceedings of 1989 International Chemical Recovery Conference, Ottawa, 3–6 April 1989 29 Horton R. R. , Slayton D. “In-Flight Black Liquor Combustion Simulations Using a Three-dimensional Computer: Sensitivity Studies”, Forest Products Symposium, Boston Massachusetts 1992 30 Quick, J. W., Gartshore, I. S., Salcudean, M., “The Bifurcation of Steady States for Opposing Jets in Cross Flow: Application to Recovery Boiler”, Proceedings of the Thirteenth Canadian Congress of Applied Mechanics, CANCAM, June 2–June 6 1991

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31 Malmgren A.,“ Investigation of the Flow Conditions in a Recovery Boiler with respect to Gases, Droplets And Dust, part 3: Mathematical Modelling of Non-Isothermal Jet Behaviour Under Boiler Conditions”, Internal report, Division of Heat and Furnace Technology, Royal Institute of Technology, Stockholm, Sweden, July 1992 32 Grace, T., Frederick, J. W., Iisa, K., Wåg, K., “New Black Liquor Drop Burning Model”, Int. Chemical Recovery Conference, Tampa, Florida, vol. 1 p. 257, June 1998 33 Hough G., Grace T., “Chemical recovery in the alkaline pulping process” Tappi Press, ISBN 0-8985-046-0, 1985 34 Sumnicht D. W. “A Computer Model of a Kraft Char Bed”, Ph. D Thesis, Institute of Paper Chemistry, Appleton, Wisconsin, May 1989 35 Brown, C. A., Grace, T., “Char Bed Burning Rates: Experimental Results”, Chemical Recovery Conference, Tampa, Florida 1988 36 Karvinen, R., Siiskonen, P., “Problems and Difficulties in the Numerical Modelling of a Recovery Boiler Performance”, Midnight Sun Colloquium on Recovery Research, Mariehamn, Finland, June 15–16, 1989 37 Frederick, W., J., Hupa, M., “Steady State Kraft Char Bed Model”, Combustion Chemistry Research Group, Report 91-9, Åbo Akademi, Turku, Finland, P.24 Finland 1991 38 Sutinen, J., E., Karvinen, R., “Numerical Modelling of Char Bed Phenomena”, Proceedings of the International Chemical Recovery Conference, Tappi Press,1992 39 Wessel, R. A., Parker L. K. Akan-Etuk, A.”,Three-Dimensional Flow and Combustion Model of a Kraft recovery Furnace”, Tappi Proceedings Engineering Conference, pp. 651, 1993 40 Fiveland, W., A., Wessel, R., A., “Numerical Model for Predicting the Performance of Three-Dimensional Pulverized Fire Furnace”, Engineering for Gas Turbines and Power, vol. 110, no. 1, p. 117, 1988 41 Tao, L., “Mathematical Modelling of Kraft Recovery Boiler”, NUTEK/IEA project no.92-03379, Recovery Furnace Modelling, report no. 3, June 1997 42 Tao, L., Blasiak, W., “Numerical Simulation of a Kraft Recovery Boiler Using Rotation Firing Method”, Finnish–Swedish Flame Days, Naantali, Finland, 3–4 Sep. 1996 43 Grace, T., “A Critical Review of Computer Modeling of Kraft Recovery Boilers”, Tappi Journal, July 1996 44 Turns, S.R., “An Introduction to Combustion”, ISBN 007-116910-5, 45 Sricharoenchaikul, V., Frederick, W. J., Grace, T. M., “Thermal Conversion of Tar to Light Gases During Black Liquor Combustion”, International. Chemical Recovery Conference, Torento, p. A209, 1995

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46 Sricharoenchaikul, V., Kymäläinen, M., Grace, T. M., “Sulfur Species Transformation and Sulfate Reduction During Pyrolysis of Kraft Black Liquor”, International Chemical Recovery Conference, Torento, p. A227, 1995 47 Frederick, W. J., Iisa, K., Wåg, K., Reis, V. V., Boonsondsup, L., Forssen, M., Hupa, M., “Sodium and Sulfur Releases and Recapture during Black Liquor Burning”, US DOE Report DOE/CE/40936-T2 (DE960006558), 1996 48 Someshwar, A. V., “An Analysis of Kraft Recovery Furnace NOx Emission and Related Parameters,” Technical Bulletin, NCASI, no. 636, 1992 49 Jones, A. K. Stewart, R. I., “The High Solids Breakpoint: A Trade-off Between SO2 and NOx”, Pulp and Paper Canada, page 149 1993 50 Nichols, K. M., Tompson, M., Empie, J. H., “A Review of NOx Formation Mechanisms in Recovery Boiler”, Tappi Journal vol.76 no.1, 1993 51 Nichols, K. M., Lien, S. J., “Formation of Fuel NOx During Black Liquor Combustion”, Tappi Journal vol.76 no.3, 1993 52 Aho, K., Hupa, M., Vakkilainen, E., “Fuel Nitrogen Release During Black Liquor Pyrolysis,” Tappi Journal vol. 77, no. 5, 1994 53 Adams, T., N., Stewart, R., I., Jones, A. K.“Using CFD Calculation to Estimate Thermal-NOx from Recovery Boilers at 67% and 80% Dry Solid”, Tappi Proceedings Engineering Conference, pp. 625, 1993 54 Forssen, M. Hupa M., Pettersson R. and Matina, D. “Nitrogen Oxide Release During Black Liquor Char Combustion and Gasification” International Chemical Recovery Conference, Toronto Canada, B231-239 1995 55 Brink, A., “Evolution of Particle Composition and Gas Phase Chemistry in Recovery Boilers”, Internal Report, Combustion Chemistry Research Group, Åbo Akademi, Finland, Dec1995 56 Iisa, K., Jing, Q., “Model for NO Formation in Recovery Boiler”, Journal of Pulp and Paper Science, vol 26 no. 1 Jan 2000 57 Jones, A. K., Grace, T. M., Monacelli, J. E., “A Comparison and Experimental Methods for determining Gas Flow Patterns in the Kraft Recovery Boiler”, Tappi Proceedings, Engineering Conference, 1988 58Tao L., “Investigation of the Flow Conditions in a Recovery boiler with respect to Gases, Droplets and Dust, Part 4: Mathematical Modelling of a Non-Isothermal Jet Behaviour under Boiler Conditions”, Internal Report, Division of Heat and Furnace Technology, Royal Institute of Technology, Stockholm, Sweden, Sept. 1992 59 Salcudean, M., Nowak, P., Abdullah, Z., “Cold Flow Computational Model of a Recovery Boiler”, Journal of Pulp and Paper Science, vol. 19 no. 5, Sept1993

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60 Yang, W., Horton, R. R., Adams T. N., “Effect of Boundary Geometries on CFD Simulations of Recovery Furnace Char Beds” Tappi Journal, vol. 77, no. 8, August 1994 61 Borg A., Teder A. “Inside a Kraft Recovery Furnace: Studies on the Origins of Sulphur and Sodium Emission”, Tappi Journal vol. 57, no.1 1974 62 Blackwell, B., “Validity of Physical Flow Modelling of Kraft Recovery Boilers”, Tappi Journal, Sep., 1992 63 Blasiak, W., Tao, L., Vaclavinek, J., “Recovery Furnace Modelling”, NUTEK/IEA Project (no. 92-03379) report, no. 1, Division of Heat and Furnace Technology, Royal Institute of Technology, Stockholm, Sweden, 30 June 1994 64 Vaclavinek, J., Tao, L., Hellen, C., “Modellering av Sodapanna – Preliminär Studie av Luftstrålarnas Areodynamik i en Partiell Fysisk och Matematisk Modell”, NUTEK/IEA Project (no. 92-03379) report, Division of Heat and Furnace Technology, Stockholm, Sweden, November, 1993 65 Karidio, I., Markovic, C., Uloth, V., Thorn, P., Abdullah, Z., Salcudean, M., “Cold Flow Velocity Measurements and Computations for CFD Validation. Paper I: Interaction of Primary and Secondary Air Flow”, Tappi Proceedings Engineering Conference, pp. 853, 1995 66 Wessel, R. A., Verrill, C. L., “Validation of Combustion and Fume Formation Models for an Operating Kraft Recovery Boiler”, Tappi Proceedings Engineering Conference, pp. 775, 1996 67 Kaul, V., Öman, L., “Sodapanneprocessforskning: Dagsläget, Pågående Forskning och Förväntade Resultat” STFI, Stockholm, Dec. 1992 68 Green, R. P., Hough, G., “Chemical Recovery in the Alkaline Pulping Processes”, Tappi Press ISBN 0-89852-255-2, 1992 69 Adams, T., “Kraft Recovery Boilers”, Tappi Press 0102B064 70 Star-CD manual “User Guide”1999 71 Fakhrai, R., “Modelling of Carry-Over in Recovery Furnace”, Thesis for degree of Licentiate of Technology, Stockholm, Royal Institute of Technology, 1999 72 Whitty, K., “Pyrolysis and Gasification Behaviour of Black Liquor under Pressurized Condition”, Åbo Akademi, Rep. 97-3 1977

73 Frederick, W. J., Hupa, M., “The Effects of Temperature and Gas Composition on Swelling of Black Liquor Droplets During Devolatilization”, Journal of Pulp and Paper Science, 10 pp. j274–j280, 20, 1994 74 Noopila, T., “Laboratory Studies on Combustibility of Black Liquors”, Åbo Akademi Rep 91-11, 1991 75 Borg A., Teder A. “Inside a Kraft Recovery Furnace: Studies on the Origins of Sulphur and

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Sodium Emission”, Tappi vol. 57, no. 1, 1974 76 Bai, Xue-Song, “ Modelling of Turbulent Combustion at Low Mach Numbers,” Doctoral Thesis, Department of Mechanics, Royal Institute of Technology, Stockholm, Sweden, (1994). 77 Weber, R., A. A. F. Peters, P. P. Breithaupt and M. B. Visser, Mathematical Modelling of Swirling Pulverized Coal Flames: What Can Combustion Engineers Expect From Modelling ?, The 1993 International Joint Power Generation Conference Proceeding, Kansas City, Missouri, p.71, (1993). 78 Magnussen, B. F. and B. H. Hjertager, On Mathematical Modeling of Turbulent Combustion with Special Emphasis on Soot Formation and Combustion, 16th Symposium (Int.) on Combustion, The Combustion Institute, p.719, (1976). 79 Lindsjo, H., X. S. Bai and L. Fuchs, Numerical and Experimental Studies of NOx Emissions in a Biomass Furnace, Proceeding of 4th International Conference on Combustion Technologies for a Clean Environment, Lisbon, Portugal, (1997). 80 Bai, Xue-Song, On the Modeling of Turbulent Combustion at Low Mach Numbers, Doctoral Thesis, Department of Mechanics, Royal Institute of Technology, Stockholm, Sweden, (1994). 81 . Lindsjo, H., X. S. Bai and L. Fuchs, Numerical and Experimental Studies of NOx Emissions in a Biomass Furnace, Proceeding of 4th International Conference on Combustion Technologies for a Clean Environment, Lisbon, Portugal, (1997). 82 Missaghi, M., “Mathematical Modeling of Chemical Sources in Turbulent Combustion,” Ph.D. Thesis, University of Leeds, England, (1987). 83 Brink, A., “Evolution of Particle Composition and Gas Phase Chemistry in Recovery Boilers”, Internal Report, Combustion Chemistry Research Group, Åbo Akademi University, Finland, 1995 84 De Soete, G. G., “Overall Reaction Rate of NO and NO2 Formation from Fuel Nitrogen”, 15th International Symposium on Combustion, The Combustion Institute, 1974 85 Mitchell, J. W., Tarbell, J. M., “A Kinetic Model of Nitric Oxide Formation During Pulverised Coal Combustion”, AIChE Journal, vol. 28, no. 2, p. 320, 1982 86 Iisa, K., Jing, Q., “Model for NO Formation in Recovery Boiler”, Journal of Pulp and Paper Science, vol 26 no. 1 Jan 2000 87 Blackwell and King “Chemical Reaction in Kraft Recovery Boiler”, ISBN 0-9692261-0-1

Appendix

Numerical Model of Recovery Furnace

1

Appendix Table of Contents

Numerical Model of Recovery Furnace 2 Flow Field 2 Heat Transfer 2 Turbulent 4 Mass Transfer 5

Black Liquor Combustion 6 Drying 6 Devolatilization 7 Char Burning 8 Swelling 9 Gas Phase Combustion 12 Char Bed Combustion Model 13 Bed Reaction Models 16 Diffusion of reacting gases to char bed 16 Chemical Kinetics in the Char Bed 17 Overall rate of Char Bed reactions 18 Mass Exchange between Char Bed and Over-Bed Gases 19 References 22

2

Numerical Model of Recovery Furnace The starting point for any soundly based description of numerical simulation resides in the conservation equations of fluid mechanics. Since these equations and their derivation is well discussed in most of the fluid mechanics books and in all the texts on the subject [1], we can simply state some of them, with a little elaboration. Supplementing these equations are those arising from species conversation, which are considered in detail in combustion texts. It is convenient to employ Cartesian tensor notation.

Flow Field We start with conservation of mass, namely the equation of continuity because it requires no assumption except that the density and velocity are continuum function. It is expressed as:

( ) mSt =∇∇+∂∂ ρρ . (1)

Conservation of momentum yields

dtdV

iSijPg ρτρ =+∇+∇− . (2)

This is the differential momentum equation in its full glory, and is valid for a fluid in any general motion, particular fluids being characterized by particular viscous-stress terms. For turbulent flows, ui, p, and other dependent variables including ijτ , assume their ensemble-averaged values given, for Newtonian fluids.

jiijk

kijij uu

xu

s '322 ρδµµτ =

∂∂

−= (3)

∂

∂+

∂∂

=i

j

j

iij x

uxu

s21

(4)

Heat Transfer Heat transfer model is handled by the energy equation. The form of the enthalpy conservation equation for a general fluid mixture is given as:

( ) ( ) ( ) hj

iij

jjjhj

j

sxu

xpup

tFhu

xh

t+

∂∂

+∂∂

+∂∂

=−∂∂

+∂∂ τρρ , (5)

3

Here, the static enthalpy h is defined as:

mmpp HmTcTch ∑+−= 00 (6)

jxmm

mDm mhhju

jxTKjhF ∂

∂∑+−∂

∂= ρρ '', (7)

Where K is the thermal conductivity, mD is the molecular diffusivity of species m , and

mh is its static enthalpy. The middle term containing static enthalpy fluctuations 'h represents the turbulent diffusional flux of energy. This turbulent scalar flux is linked

to the time averaged static enthalpy h as:

jth

tj x

hhu∂∂

−=,

''σµ

ρ (8)

where th,σ is the turbulent Prandtl number. Its default value is equal to 0.9. The energy source term hs considers primarily the energy sources or sinks due to exothermic or endothermic chemical reactions, gaseous thermal radiation, and the inter-phase heat exchange in two-phase flow cases. Each species m of a fluid mixture, whose local concentration is expressed as mass fraction mm , is governed by a species conservation equation of the form:

( ) ( ) mjmmjj

m sFmux

mt

=−∂∂

+∂∂

,ρρ (9)

For turbulent reacting flow, the diffusional flux component is given by:

'', mjj

mmjm mu

xm

DF ρρ −∂∂

= (10)

Where the rightmost term, containing the concentration fluctuation mm ’, represents the turbulent mass flux. This turbulent scalar flux is related to the time averaged concentration mm as:

j

m

tm

tmj x

mmu

∂∂

−=,

''σµ

ρ (11)

4

Turbulent The ‘standard’ ε−k model, in which the effects of turbulence are represented by an isotropic “eddy” or turbulent viscosity, is evaluated using two quantities: Turbulent kinetic energy k and its rate of dissipation ε. Turbulence Energy k

( )

( )i

i

i

itBt

jk

effj

j

xu

kxu

PP

xkku

xk

t

∂∂

+

∂∂

−−+

=

∂∂

−∂∂

+∂∂

ρµρεµ

σµ

ρρ

32

(12)

teff µµµ += , (13)

j

iij x

usP

∂∂

= 2 (14)

ith

iB x

gP

∂∂

−=ρ

ρσ1

,

(15)

The first term on the right-hand side of this equation represents turbulence generation by shear and normal stresses and buoyancy forces, the second viscous dissipation, and the third amplification or attenuation due to compressibility effects. Turbulence Dissapation Rate ε

( )

( )i

i

i

i

i

itBtl

j

effj

j

xu

Ck

Cxu

kxu

PCPk

C

xu

xt

∂∂

−−

∂∂

+

∂∂

−+

=

∂∂

−∂∂

+∂∂

ρεερρµµε

εσµ

ερρε

εεεε

ε

4

2

23 32

(16)

εσ , 1εC , 2εC , 3εC and 4εC are the empirical coefficients.

The right-hand side terms represent analogous effects to those described above for the k equation. The empirical coefficients used in the standard k-ε model are given in the following table:

5

µC kσ εσ 1εC 2εC 3εC 4εC 0.09 1 1.22 1.44 1.92 0 0r 1.44* -0.33

* 3εC =1.44 for BP >o and is zero otherwise.

Mass Transfer Each species m of a fluid mixture, whose local concentration is expressed as mass fraction mm , is governed by a species conservation equation of the form:

( ) ( ) mjmmjj

m sFmux

mt

=−∂∂

+∂∂

,ρρ (17)

For turbulent reacting flow, the diffusional flux component is given by:

'', mjj

mmjm mu

xm

DF ρρ −∂∂

= (18)

Where the rightmost term, containing the concentration fluctuation mm ’, represents the turbulent mass flux. This turbulent scalar flux is related to the time averaged concentration mm as:

j

m

tm

tmj x

mmu

∂∂

−=,

''σµ

ρ (19)

6

Black Liquor Combustion Despite being a liquid, the combustion behaviour of black liquor is more like that of a solid fuel, such as coal or liquid fuels. Rather than evaporating and burning in the gas phase, like a typical liquid fuel drop, the black liquor undergoes the same stage of combustion as a solid fuel [Whitty, 2, 1997]. Hupa et al. 3, 1987 successfully identified the burning stages and divided them into four stages; drying, devolatilization, char burning and smelt reaction. In coming sections a qualitative description of a previously mentioned process relevant is given. Regarding the mathematical description of the stages, a complete modelling is given in the references considered for example Tao [31].

Drying The term “drying” in this section is used to refer to the process in which heat is transferred to and through a drop from the surrounding environment and water content of the drop enters the surrounding gas phase. Under this classification the subject of evaporation is not only water, even small amount of solid in the drop will be devolatilized. Black liquor at 65% solids has 35% water content. This is relatively low by the standards of wood waste and other biomass, which typically have 45%-60% moisture contents. However, sine black liquor consists of about 38% inorganics, this is equivalent to 46% moisture on an ash-free basis. This represents a very substantial evaporation load during black liquor combustion. The importance of drying process also stems from the fact that a significant portion of burning time is drying time. There is no flame around the drop indicating that there is not significant loss of volatile matter or sulphur gases. Predicting a drop’s drying rate is a complex process. However, in the case of a pure liquids or an externally controlled heat transfer the task is easier. The specific thermal and transport properties of the material, and its pore structure are important to the drying rate. Water can diffuse through this material or, if heating is rapid, evaporate and cause swelling. During drying, the droplet swells and collapses in rapid succession, expanding its surface area and mixing its contents. Swelling characteristics of black liquor increase the complexity of the transport processes. The trapped water vapour continues to make the particle swell until the surface ruptures releasing the vapour and allowing the particle to collapse. There are two classical periods of particle drying. Black liquor droplets swell almost instantaneously, typically by a factor of 1.5 in diameter, once the droplet has entered the hot furnace environment. Therefore it can be treated as being of constant diameter. The droplet temperature rises rapidly to about 150°C early during the drying stage to increase more slowly after approaching 300°C at ignition. By the end of the drying stage, all visible signs of boiling cease. The droplet is extremely viscous or looks like a solid particle. Frederick [4, 1991] called the elapsed time between the droplet entering the furnace and its ignition the drying time or the time to ignition, ti. Drops initial diameter, solid content, viscosity [Clay, 5, 1988] and swelling are important properties during this stage. Convective heat transfer rates from the gas, both natural and forced, depend upon the drop diameter. As long as the surface of a drop is dampened with water, evaporation rate is constant and external heat transfer limits the rate of evaporation. However when the surface is dried, temperature gradients within the drop are significant.

7

Devolatilization If water content of a black liquor drop decreases enough, either locally on the surface of the drop a bright yellow flame surrounding the drops will be formed at some distance from the drop surface. Appearance of the flame indicates the initiation of devolatilization process and disappearance of the flame indicates the end of the stage. Devolatilization processes are much dependent on the rate at which the drop heat and on the maximum drop temperature. The processes include: (a) heating the black liquor drop from its temperature at ignition to that at the end of devolatilization (i.e., beginning of char combustion), (b) vaporization of the residual water, and (c) supplying the heat for pyrolysis reactions. Devolatilization time tv was taken as the duration of the visible flame and time needed to complete above mentioned processes. The point of maximum swelling of the drop also indicates the end of devolatilization stage. Frederick and Hupa (6, 1993) investigated the influence of liquor type, droplet size dry solid content, furnace temperature and pyrolysis time on the volatiles yield of single liquor drops undergoing pyrolysis. The volatiles yield for different Kraft liquor ranged from 35% to 47 %of the initial mass of the liquor. Feuerstin et al. (7, 1967) and Brink et al. (8, 1967) investigated the composition of the product of the devolatilization. They found that CO2, CO, H2, light hydrocarbons, H2O, H2S and mercapatans are the main products. The experiment was carried out using black liquor pyrolysis at a very low heating rate. They also found that the concentration of CO and CO2 increased with increasing temperature. Fredrick and Hupa (9, 1993) studied the influence of furnace temperature, oxygen content, droplet size, and dry solid content on the yield of volatiles and the carbon content of the char produced during devolatilization. They found that the amount of volatiles formed increased and the carbon content of the char decreased with increasing reaction time at a constant furnace temperature. Carbon, sodium and total mass loss from the char particles continued beyond devolatilization. This was assumed to be due to sodium carbonate decomposition. MaKeough et al. (10, 1994) measured the concentration of gases produced during the pyrolysis of black liquor in a grid heater at different temperatures and pressures. They found that the release of H2 and CO and light hydrocarbons increases with temperature whilst the release of CO2 and sulphur-containing species decreases. A higher pressure results in the decreased production of H2 and CO but an increased in CO2 and light hydrocarbons. The remaining solid material contains the residual non-volatile organic substance, mostly carbon, along with most of the inorganic material. Within the pyrolysing particle, the species may reconvene onto the solid surface and remain in the char. Frederick [6, 1993] took the initial conditions for a base-line calculation of tv to be the values of the relevant variables at the end of the drying stage, average particle temperature of 1508°C (estimated limit of the boiling point as the BL drop dries); d/d0 =1.5; and Si=0.8.

8

The model also requires several other input parameters, three of which affect the computation of tv in a major way:

(a) Swelling during devolatilization. (b) Maximum particle temperature (which, according to the laboratory-furnace

experiments, appeared to occur at maximum swelling). (c) Gas temperature throughout devolatilization.

Char Burning The next stage in black liquor combustion is char burning. Char is the residual solid product from black liquor pyrolysis. It is a black, porous, material containing essentially all of the sodium and about one-half the carbon in the incoming black liquor [Grace 11, 1989]. For most purposes, char can be considered to consist of three inorganic salts: sodium carbonate (Na2CO3), sodium sulphide (Na2S), and sodium sulphate (Na2SO4), along with carbon and bound hydrogen. At the completion of pyrolysis, the char is typically about 75% inorganic and 25% carbon. Typically char composition is given in the table. Table 1 Char Composition. (Source: Chemical recovery in the alkaline pulping process” Tappi 1992)

Moles (chemical equivalents) wt. % Na2S 0.15 Na 31.9 Na2SO4 0.15 S 6.6 Na2CO3 0.7 C 30.8 Fixed carbon 3 H 0.7 Bound hydrogen 1 O 30.0

Char burning occurs as a heterogeneous reaction between the char particle and the oxygen in the combustion air (primary air and secondary air, if present). There is no visible flame, but a rather intense glow on the reacting surface. The residual carbon burns away, and the inorganic salts are reduced and smelted out. Glowing char combustion occurs when the production of volatiles ceases and oxygen diffuses from the environment to the char particle and reacts heterogeneously with it. The burning time of black liquor droplets is strongly influenced by the size of the droplet, and the contact between air and the drop. Burning time in air for droplet of the size range normally encountered in the furnace is about 5-10 s. Drying and char burning stages tend to be the slowest. There are several potential reactions for burning the char and it is not clear which is the most important. If the carbon could access oxygen directly from the combustion air above the char bed the char would be gasified to CO and CO2 by:

C + ½ O2 → CO C + O2 → CO2 CO2 + C → 2CO

Similarly, if the carbon could access water vapour, gasification reactions occurs such as: CO + H2O (g) → CO2 + H2 C + H2O (g) → CO + H2

9

C + ½ Na2SO4 → CO2 + ½ Na2S C + ¼ Na2SO4 → CO2 + ¼ Na2S The most significant parameter governing the stoichiometric air requirement for char is the relative proportion of carbon monoxide (CO) and carbon dioxide (CO2) in the product gas. About twice as much air is needed if the char carbon is burned to CO2 rather than CO. The air requirement for char burning is about 25-45% of the air required for burning all of the black liquor, depending on the CO:CO2 split. The amount of hydrogen and oxygen remaining in char carbon is not sufficient enough to have a major effect on the char burnout. The organic carbon and inorganic salts are intimately mixed in a porous structure provided by the carbon. The temperature of a black liquor droplet burned in air is well above the furnace temperature for the entire char burning stage. The two-colour optical pyrometer measurement shows that the surface temperature for droplets burned in air is typically 220-500 °C above the furnace temperature during char combustion. The molten smelt aids the char combustion and contributes to Kraft black liquor char’s inherently high reactivity. There is no relationship between the carbon content of the char and the initial mass drop. The carbon content changes very rapidly during the first five seconds. Milanova and Kubes (12, 1992) measured the ignition temperatures and the burning times for char particles from various Kraft liquors under laboratory conditions using thermo analytical methods. The ignition temperature was found to be strongly dependent on the sodium chloride content of black liquor solids; ranging from 780 °C to 580 °C, decreasing sharply with increasing content of salt up to about 2 % NaCl, and then levelling off. Burning times for char particles were found to depend on the swelling properties of liquor solids during the pyrolysis stage. Grace et al. (13, 1998) suggested that the char burns via the sulphate - sulphide cycle. In this cycle, the carbon in the char reacts with sulphate, reducing it to sulphide and forming CO2 and CO. The sulphide, in turn, reacts with oxygen from the combustion air, reforming sulphate and completing the cycle. The function of cycle is to carry oxygen to the carbon, which is burned off. Figure 54 illustrates the concept of the sulphate - sulphide cycle. Oxygen comes in from the combustion air and reacts with sulphide (Na2S), oxidizing sulphide to sulphate (Na2SO4). The sulphate carries the oxygen over to the carbon, where it reacts to form CO2 and CO. The sulphate, in turn, is reduced back to sulphide, completing the cycle.

Swelling Swelling the single, most important, specific variable, which complicates the behaviour of the black liquor droplet during combustion. It changes the rate of internal and external heat and mass transport, resulting in a different combustion time. It affects droplet trajectories and the entrainment characteristics. Particle swelling increases the surface area available for transport processes and increases porosity, and thus enhances apparent reactivity. Despite relatively intensive and extensive work in this area, the phenomenon is not well understood. Experimental results and empirical correlation should thus be relied upon in order to model the swelling of black liquor droplets.

10

Balkien (14, 1960) indicated a relation between what he called “burning quality“ and the tendency of a black liquor droplet to swell. He heated liquors in a dish at 420°C until swelling was completed. Based on the amount of liquor swelling, he computed an index for burning quality. Two decades later, Hupa et al. (15, 1982) at Åbo Akademi used a single droplet system to study the combustion behaviour of the sulphate and sulphite liquors. In this study pure air was used and the furnace temperature was 700 °C. They found that both liquors behaved in a fairly similar manner during the drying period. On the other hand, there was a dramatic difference between the two liquors during the escape and burning of the volatiles. The sulphate liquor droplet swelled several hundred percent, while the sulphite liquor droplet hardly shows any swelling during this stage. Miller and Clay (16, 1985) investigated the influence of heating rate, particle size, pyrolysis temperature, and moisture content on the swelling of Kraft liquor. Particle size (1-4 mm in diameter) had no effect on the maximum particle volume. An increase in the heat flux shortened the time of pyrolysis but had no effect on the maximum volume attained; thus the heating rate effectively changed the rate of swelling. The maximum volume of black liquor particles was reached at 500°C. The moisture in black liquor increased swelling during pyrolysis from 400°C to 600°C. The effect of moisture on swelling was not additive as 65 and 80% solid black liquor behaved similarly. Hupa et al. (3, 1985) classified the different burning stages of black liquor. They distinguished four different burning stages and were able to measure the maximum swelling during burning. They also generated swelling and droplet temperature profiles as a function of the liquor combustion time and showed that liquors swell primarily during the pyrolysis stage, and that the extent of swelling varies widely between liquors. Swelling measured in this way gave more consistent data than the corresponding measurement using other methods. Miller et al. (17, 1986) studied the influence of liquor composition on the swelling of laboratory-prepared black liquors in a nitrogen atmosphere at 500°C. They found that the ratio of Kraft liquor lignin to sugar acids dramatically affected the swelling behaviour during pyrolysis, and that a ratio of roughly 1:1 of these components yielded maximum swelling. They also reported that swelling decreased with increasing molecular weight of lignin. Furthermore a higher concentration of extractive resulted in decreased swelling at 500°C. Hupa et al. (18, 1987) examined several specific factors that may affect the swelling tendency and char burning rates, by changing the composition of strongly swelling liquors. Additions of Na2SO4 to increase the “dead load“ slightly reduced the maximum relative swelling. This was explained by a simple dilution effect: the amount of combustible liquor is decreased as the Na2SO4 was added. Addition of tall oil reduced the relative swelling significantly more than expected just by the dilution effect. The reduced swelling was reflected in the longer char burning time. The relationship between combustion behaviour and viscosity was studied by Söderhjelm et al. (19, 1989) Combustion behaviour was determined by single droplet laboratory combustion tests at 700°C and 800°C. The swelling of droplets at 700°C was found to

11

correlate with the viscosity and swelling decreased when viscosity increased. However, no such correlation was found at 800°C. Fredrick and Hupa (20, 1991) examined the effects of a solid content of 55 to 80% on the combustion behaviour. The study was carried out in a laboratory furnace at two temperatures 700°C and 800°C. They found that combustion behaviour of black liquor droplets did not change greatly as dry solid content was increased. There was no effect on swelling for two of the liquors during devolatilization. Swelling increased with increasing dry solid content for third liquor. Fredrick et al.. (21, 1991) studied the influence of liquor type on swelling. They examined softwood Kraft, hardwood Kraft, and sulphite liquors and found that sulphite liquor swelled less than sulphate liquors. Noopila et al. (22, 1991) used a single droplet burning technique to determine the influence of cooking time on pyrolysis swelling when burning at 700°C. As cooking time increased, the droplets swelled more during the pyrolysis phase, with the exception of birch liquor which was from the longest cook. The swelling of the liquor was about the same when cooking time was 180 and 200 min. Fredrick and Hupa et al..(23, 1994) studied the effect of gas composition on swelling and found that the presence of CO2 did not affect the maximum swollen volume as long as some oxygen was present. The results indicate that the relative change in swelling for Kraft liquor with temperature is liquor independent. It also indicates that the effect of gas composition on Svmax in the presence of oxygen can be accounted for as a temperature. The maximum swollen volume reached during devolatilization under combustion conditions (e.g. with oxygen present in the surrounding gases) depends upon the furnace temperature and oxygen concentration. The droplet contains very large fractions of low melting point inorganics. Inorganics dilute of the combustible organic portion of black liquor solids but most salts are more or less catalytic. The inorganic materials enhance the reactivity of the black liquor char. The catalytic activity of an inorganic depends on (1) chemical form; (2) amount; and (3) inclusion size [Walker (24, 1954)].

12

Gas Phase Combustion The gas phase properties of interest are the three velocity components, pressure, temperature, density, and species concentration. These properties are found by a finite-volume solution of the mass, momentum, energy and species conservation equations in a three-dimensional geometry. In practice, the furnace volume is divided into a large number of numerical cells and the difference equations are solved numerically. Interaction between the black liquor droplets and the gas phase properties is handled through the impact of local gas conditions on the droplets and the source terms in the gas phase conservation equations. The gas species considered in the simulation are methane CH4, carbon monoxide CO, hydrogen H2, oxygen O2, carbon dioxide CO2, water vapour H2O and inert (nitrogen) N2. Sulphur is not handled in the present model. Figure below shows the black liquor combustion scheme used in the calculations.

Char bed reactionsC + 0.5O2 => COC + CO2 => 2 CO

C + H2O => H2 + CO4C + Na2SO4 => 4CO + Na2S

Na2S + 2O2 => Na2SO4

Primary air

Secondary air

Black liquor

Tertiary air

Gas-phase reactionsCH4 + 0.5 O2 => CO + 2 H2

CO + 0.5 O2 => CO2H2 + 0.5 O2 => H2O

Char Combustion+CO, +H2-C, -O2,

-H2O, -CO2

Devolatilisation+CH4

+CO, +H2

Drying+H2O

SmeltNa2S+Na2SO4

+Na2CO3

+CO, +H2

Figure 1Graphical presentation of the black liquor combustion scheme

As illustrated in this figure, a three-reaction scheme in the gas phase is assumed, i.e.

CH4 + 0.5O2 => CO + H2

CO + 0.5O2 => CO2 H2 + 0.5O2 => H2O

The volatiles released from the droplets consist of three combustible gases: CH4, CO and H2. Experimental data from black liquor devolatilization25 are used to determine their proportions. In this study, turbulence is handled by the effective gas viscosity determined

13

from the two-equation k-ε turbulence model. Gas radiation heat transfer is calculated by the discrete transfer method [26].

Char Bed Combustion Model The implementation and description of the char bed model was done by Tao [27]. In this section a short of the model is given The drops landing on the char bed supply the char bed with solids (char) and molten inorganic salts (smelt). Measured temperatures inside the bed (5-25 cm below the surface) showed the cold spots, and hot spots on the surface of the char bed. The hottest zones were generated when the primary (and high primary) air hit the bed surface indicating that the mass transfer of reacting gases to the bed dominates this process. The average bed temperature is usually quite stable, close to the melting point of the system. On the surface of the char bed, carbon is converted o CO and/or CO2, and sodium sulphate Na2SO4 is reduced to, sodium sulphide Na2S (the active pulping chemical). The char bed model developed in this project is steady-state model, which is based on the work of Frederick and Hupa [30], considering:

COOC →+ 250. (20)

COCOC 22 →+ (21)

22 HCOOHC +→+ (22)

SNaCOSONaC 242 44 +→+ (23)

The mass input from the landing drop is calculated based on the mass balance.

Mchar = Md (XH20 + XVM + XC +XSmelt ) (24)

∑ =4

1i i

X (25)

Xsmelt = 1- XH20 +XVM +XC (26)

where, Md (kg/s) represent the amount of drop mass and Xi, is the mass fraction of water, volatiles, char carbon, and inorganic salts (smelt) falling on a char bed region resp. (kg/kg). Since no loss of the smelt in the in-flight drop combustion model have been considered the mass flow of sodium and sulphur entering the bed is equal to:

odnaNachar MXM ,,, 0= (27)

odSSchar MXM ,,, 0= (28)

14

Where NacharM , and ScharM , represent the total amount of sodium and sulphur in the drop

falling on a char bed region, 0,NaX and 0,SX the initial mass fraction of sodium and

sulphur in the drop respectively, and 0,dM the initial mass of the drop. From the experiment, it indicates that the depth of the active zone is about a few centimetres. In the present work the density of the bed bedρ is 375 kg/m3, and the depth of the active zone Hbed= 5 cm. In order to simplify the chemical composition of the char bed, the following assumptions were made [Goreg, 28, 1986]. 1. All sodium and sulphur are present as three inorganic compound: carbonate, sulphate,

and sulphide; 2. The relative amount of sulphate and sulphide define the average oxidation state of the

sulphur; 3. Hydrogen present as condensed aromatics is treated as “bound” hydrogen; 4. Carbon not present as sodium carbonate is treated as solid carbon; 5. All oxygen is accounted for by the inorganics; 6. The S/Na2 ratio is 1 to 3 and the initial state of reduction is equivalent to 50%

reduction efficiency (Na2S/(Na2S+Na2SO4)); 7. The relative amounts of carbon, hydrogen, and alkali are close to those found in the

actual char samples. Char burning consist of two essential steps: (1) conversion of the fix carbon to CO and CO2 to permit melting and coalescence of the smelt, and (2) changes in the oxidation state of the inorganic sulphur compounds. In this regards the mina chemical reactions involved in carbon burnoff are as follow:

SNaCOSONaC 242 44 +→+ (29)

SNaCOSONaC 242 250250 .. +→+ (30)

The conditions during char burning are such that the predominant reaction by which carbon is burned are the reaction with molten sulphate. The sulphate and sulphide present can act as a catalyst for carbon burnup, with the sulphate reduced to sulphide by carbon and sulphide subsequently reoxidized to sulphate by reaction with oxygen [Grace 29 1986]. The key factor in this concept is that sulphate reduction can take place simultaneously with carbon burnup as long as the reaction between sulphate and carbon is faster than between sulphide and oxygen [30]. The modelling of sulphate/sulphide cycle requires a time-dependent reduction rate calculation, which is a little bit complicated for a global recovery boiler model at present stage. A simplified approach is to treat the degree of sulphate reduction in the bed as a user-defined parameter. In this work, the sulphur reduction in the char reaching char bed, Rin and the sulphur reduction in smelt leaving char bed, Rout are taken as 0.5 and 0.9 respectively [27].

15

According to the conditions described above, the chemical composition of the char bed can be calculated as the following. The overall mass fractions of carbon, sulphur, and sodium in the liquor droplets/particles falling on a char bed region are given by:

totchar

CcharcharC M

MX

,

,, = (31)

totchar

ScharcharS M

MX

,

,, = (32)

totchar

NacharcharNa M

MX

,

,, = (33)

The overall mass fraction of solid char in the liquor droplets/particles falling on a char bed region is determined through:

−+

−++=

3246

46106

321421

3278

charscharNainincharscharCchar XXRRXXX ,,;, )( (34)

Thus, the mass fraction of carbon, sulphur, sodium, and sulphate of solid char in a char bed region can be calculated by:

char

charcbedc X

XX ,

, = (35)

char

charSbedS X

XX ,

, = (36)

char

charNabedNa X

XX ,

, = (37)

)(,, inbedSbedSO RXX −= 14 (38)

16

Bed Reaction Models Heterogeneous gas-solid reactions involving the char bed are governed by coupling of transport phenomena and chemical kinetics. The overall reaction scheme can be described by the following basic events:

• Diffusion of mass (reactant and product gases) and heat across the boundary layer surrounding the bed;

• Reaction of gases with bed surface. Thus, the overall rate of char combustion in the bed is determined by combination of the rate for film mass transfer and the rate for chemical reactions. It is expressed as:

filmbedickbedoverallbed RRR ,int,,

111+= (39)

This section has addressed these topics:

Diffusion of reacting gases to char bed The mass transfer to the char bed from the gas flow above the bed is given by:

iii CKm =*

(40)

Where im*

are the mass transfer rates of the gas component to the bed (mol/m2-s); iK , the corrected mass transfer coefficients (m/s); iC the local concentrations of gas copmonemt at the cell adjacent to the char bed (mol/m3). Brown et al. [31] investigated that the mass transfer coefficient for oxygen and found that it increased nearly linearly with increasing gas velocity, The correlation was obtained for velocities 3-15 m/s, oxygen contents 7-21%, and furnace temperature 750-800oC. Their data correlated as:

9107822

.', .

,VK bedO = (41)

where V is the local gas velocity magnitude (m/s). This correlation has been used in present char bed model with a correction for gas film temperature. The mass transfer coefficients for other reacting gases are assumed to be directly proportional to the ratio of their diffusivity with that of oxygen. Thus, it gives:

',

.

,

)(.bedO

bedgbedO K

TTK

20

107350

2

+= (42)

17

bedOO

CObedCO K

DD

K ,, 2

2

2

2= (43)

bedOO

CObedCO K

DD

K ,, 2

2

2

2= (44)

where the diffusivities of oxygen, carbon dioxide, and water vapour are taken as 1.8E-5, 1.44E-5, and 2.38E-5 respectively; gT the local gas temperature at the cell adjacent to the char bed.

Chemical Kinetics in the Char Bed In event of the low bed temperatures, the reaction rate of O2, CO2 and H2O vapour with the char carbon can become slow enough so that chemical kinetics must limits the overall rate. The most important reactions in this regard are The kinetic rate for the chemical reactions is calculated according the following principles. The rate of reaction with oxygen is given by the Cameron and Grace’s sulphate-sulphide kinetics [32]. The rate expression for the SOC/Na 42 reaction taken the following form:

[ ] [ ][ ][ ] [ ]bedbedCS TCSO

SONaMR /exp., 1469600110

52404

42 −

+= (45)

M = Multiplier for Cameron-Grace reduction rate taken as 50

[ ]bedNa ,2 = 0460.

,bedNabed Xρ Sodium concentration, moles Na2/bed volume (mol/m3)

[ ]4SO = bedNa

bedSO

XX

,

,. 4240 Sulphate concentration, mol SO4/mol Na2

[ ]C = bedNa

bedC

XX

,

,.921 Carbon concentration, mol C/mole Na2

Fredrick and Hupa discussed that the above reduction rate from Cameron and Grace underestimates the rate of sulphate reduction and it needs to be increased by a factor of 50. The rate of reaction of carbon directly with oxygen 2OC / uses the Wendt’s modification of Smith’s correlation of carbon combustion data. The direct carbon oxidation rate was reduced by a factor to account for the fraction of the carbon surface area available for contact with the reacting gases according Sumnicht. The rate equation for direct C/O2 reaction is expressed as:

18

)/exp(int bedOgsumnCO TYPAPR 170863166722

−= (46)

sumnP = Sumnicht factor, for carbon surface area available for direct oxidation.

intA = Internal carbon surface area of bed (m2).

gP = The local pressure of gas mixture at the cell adjacent to the char bed (atm).

2OY = The local oxygen mole fraction at the cell adjacent to the char bed. Sumnicht factor would have a value of 0.7 for a char particle at the end of devolatilization, and zero when char combustion is complete. In this work, we used a value of 0.6, which corresponds to 50% completion of sulphur reduction. The internal carbon surface area of bed intA is given by:

bedCbedbedsp XVAA ,int ρ= (47)

spA = Internal specific surface area of char, taken as 11000 (m2 /kg)

bedV = Char bed volume per unit surface area (m3/m2), bedbed HV = The rate of reaction with CO2 and water vapour were taken from Li and van Heiningen [33,34]. The rate equations for gasification reaction C/CO2 and C/H2O have the following forms:

[ ] )exp(..

" bedCOCO

CObedCCO T

EYY

YCR 22500623

432

2−

+= (48)

[ ] )exp(.. bedHOH

OHbedOCH T

EYY

YCR 253009562

4122

2

2−

+= (49)

[ ]bedC = Mole concentration of carbon in the bed, (mol C/m3)

2COY , OHY2

, COY and2HY = Mole fractions of carbon dioxide, water vapour, carbon

monoxide, and hydrogen at the cell adjacent to the char bed. [ ]bedC is calculate by:

[ ] 0120./.bedCbedbed XC ρ= (50)

Overall rate of Char Bed reactions In this work, equation 51 is applied separately to each of the reacting gases O2, CO2, and water vapour because the chemical reaction rates differ by a factor of 26/1/14 at typical bed temperatures and above-be gas compositions, and the mass transfer rates by

19

0.4871/0.51. The overall rates of char bed reactions are calculated using the following forms:

ii

i

Rm

R1

2

11

+=

*

(51)

iR is the rate and i is the overall oxidation , CO2 gasification and H2O gasification The

total consumption rate of carbon in the bed is then given by:

)(. ,,,, overallOCHoverallCCOoverallCoxoverallC RRRR22

0120 ++= (52)

Mass Exchange between Char Bed and Over-Bed Gases

Condition 1: overallCbed

Ccharinbed R

AM

C ,, ≥=

Where bedA is the surface area of a particular char bed region and inbedC the incoming flux of carbon to the char bed region. Condition 1 indicates that the incoming flux of carbon is grater or equal to the total consumption rate of carbon in the bed. In this case, the bed char may grow or maintain a quasi-stable shape. Thus, the interface exchange terms can be determined using the following relations. The amount of carbon consumed by the sulphate reduction is calculated by:

bed

inoutScharoverallCS A

RRMR

03204

.

)(,,

−= (53)

The amount of carbon consumed by direct C/O2 reaction is then given by:

overallCSoverallCoxoverallCO RRR ,,, −=2

(54)

The exchange flux of water vapour due to the remaining water in the droplets falling on the char bed region is determined through:

bedOHchar AMm Water /,

'

2= (55)

In a similar way, the exchange flux of volatiles is given by:

20

bedVMcharVM AMm /,

.

= (56)

As described previously, the volatiles are composed of three components: CH4, CO and H2. Their proportions are determined during the in-flight combustion stages. Thus, the exchange fluxes of CH4, CO and H2 due to remaining volatiles in the droplet/particles are calculated by:

VMCHVMCH mFm.

,

.

¤=4 (57)

VMCOVMCO mFm.

,

.

= (58)

VMHVMH mFm.

,

.

22 = (59)

The exchange fluxes of CO, H2, O2, CO2, and H2O due to chemical reactions in the bed are calculated by the following equations:

)(. ,,,,,

.

overallOCHoverallCCOoverallCOoverallCSbedCO RRRRm222

20280 +++= (60)

overallOCHbedH Rm ,,

.

.22 0020= (61)

overallCoxbedO Rm ,,

.

.03202 −= (62)

overallCCObedCO Rm ,,

.

.22 0440−= (63)

overallOCHbedOH Rm ,,

.

.22 0180−= (64)

Where the minus sign means the flux from the above-bed gases to the bed. The source/sink terms of gas species in a cell adjacent to char bed surface can then be determined using the relations as below:

cellcellVMCHcellCH VAms /,

.

, 44= (65)

cellcellbedCOVMCOcellCO VAmms /)( ,

.

,

.

, +=4

(66)

cellcellbedHVMHcellH VAmms /)( ,

.

,

.

, 2242+= (67)

cellcellVMcellO VAms /,

.

, 220= (68)

21

cellcellVMCcellCO VAms /,

.

, 220= (69)

cellcellbedOHwatercellOH VAmms /)( ,

..

, 22+= (70)

Where cellA and cellV are the interface area and volume of the cell in question.

Condition 2: overallCbed

Ccharinbed R

AM

C ,, <=

In this case, the char bed will shrink. To simplify the bed calculations, we assume that the total consumption rate of carbon in the bed is equal to the incoming rate of carbon to the bed. In addition, the fractions of carbon which is consumed via different chemical reaction paths are assumed to be the same as in the condition 1. The modified interface exchange terms for the condition 2 are given as the following:

overallCinbedoverallOCHbedH RCRm ,,,

.

/.22 0020= (71)

overallCinbedoverallCoxbedO RCRm ,,,

.

/.03202 −= (72)

overallCinbedoverallCCObedCO RCRm ,,,

.

/.22 0440−= (73)

overallCinbedoverallOCHbedOH RCRm ,,,

.

/.22 0180−= (74)

The other interface exchange terms remain the same as in the condition 1.

22

References: 1 Star-CD manual “User Guide” 1999 2 Whitty, K., “Pyrolysis and Gasification Behaviour of Black Liquor under Pressurized Conditions,” PhD thesis, Report 97-3 Åbo Akademi 1997 3 Hupa, M., Solin, P., Hyöty, P., “Combustion Behaviour of Black Liquor Droplets“ Journal of Pulp and Paper Science: vol. 13 no. 2 March 1987 4 Frederick, W. J., Noopila, T., Hupa, M., “Swelling of Pulping Liquor Droplets During Combustion,“ Journal of Pulp and Paper Science Sep. 1991 5 Clay, D. T., Lien S. J., Grace T.,M. Macek A., Semerjian, H. G., Amin, N., “ Fundamental Studies of Black Liquor Combustion“ Report no. 2- Phase I, DE88005756 6 Frederick, W. J., Hupa M., “ Combustion Properties of Black Liquors,“ US DOE Report DOE/CE/40936-T1 1993 7 Feuerstein, D.L., Thomas, J.F., Brink, DL., “Malodorous Product From the Combustion of Black Liquor. I. Pyrolysis and Combustion Aspect.“ Tappi Journal 50(6): 258-262, 1967 8 Brink, DL., Thomas,J.F., Feuerstein, D.L., “Malodorous Product From the Combustion of Black Liquor. II. Analatical Aspect.“ Tappi Journal 50(6): 276-285, 1967 9 Frederick W. J. Hupa M., “ Combustion Properties of Kraft Black Liquors“ Åbo Akademi Report. 93-3, 1993 10 Makeough, P., Arpiainen, V., Venelampi, E., Alen, R., “ Rapid Pyrolysis of Black Liquor . Part I. Release of carbon .“ Paper ja Puu-Paper and Timber 76(10):650-656 1994 11 Grace T. M., Cameron J. H., and Clay D. T.,“ Role of The Sulfate-Sulfide Cycle in Char Burning: Experimental Results and Implications“Tappi Journal 108-113 Oct 1986 12 Milanova E., and Kubes G. J. “ The Combustion of Kraft Liquor Chars“ Pulp and Paper Research Institute of Canada 13 : T.M. Grace. Cameron J.H and Clay D.T. “Role of the sulphate-sulphide cycle in char burning: experimental results and implications” Tappi Journal, Oct 1986

14 Baklien, A., “The Effect of Extraction on Black Liquor from Eucalypt Pulping,“ APPITA 14(1) 1960 15 Hupa ,M. Backman R., Hyöty P., “Investigation of Firside Deposition and Corrosion in Sulphate and Sodium Sulphite Recovery Boilers“ Black liquor recovery boiler seminar D1-1 D1-11 1982

23

16 Miller, P. T., Clay, D. T., Prsentation at AIChE Mtg. Seattle WA, Aug. 26-28, 1985. 17 Miller, P.T. Clay , D.T.,Lonsky F. W. “ The Influence of Composition on The Swelling of Kraft Black Liquor During Pyrolysis“ Eng. Conf. 1986 225-233 18 Hupa ,M. Solin, P.Hyöty, P.,“ Combustion Behaviour of Black Liquor Droplets“ Journal of Pulp and Paper Science: vol. 13 no. 2 March 1987 19 Söderhjelm L. , Hupa M., Noopila T., “Combustibility of Black Liquors with Different Reological and Chemical Properties“ Journal. Pulp and paper SC. 15(4): J117-121 July 1989 20 Frederick W. J., Noopila, T. Hupa M., “ Combustion Behaviour of Black Liquor at High Solids Firing, “ Tappi Journal, 74(12):163-170, 1991. 21 Frederick W. J. Noopila, T. Hupa,M., “Swelling of Pulping Liquor Droplets During Combustion“ Journal of Pulp and Paper Sci. Sep. 1991 22 Noopilar T. , Alen R. , Hupa M. “ Combustion Properties of Laboratory-Made Black Liquors“ Journal of Pulp and Paper Sci. Vol. 17 no. 4 July 1991 23 Frederick W. J. Hupa M,“ The Effects of Temperature and Gas Composition On Swelling of Black Liquor Droplets During Devolatilization“, Journal of Pulp and Paper Sci., 20, 10 pp j274-j280, 1994 24 Walker, P. L., JR .and Hippo, E. J., “Factors Affecting Reactivity of Coal Chars “, Am. Cem. Soc. Div. Fuel Chemistry Preprints 20 (3), 45, 1954 25 Brink, A., “Evolution of Particle Composition and Gas Phase Chemistry in Recovery Boilers”, Internal report, Combustion Chemistry Research Group, Åbo Akademi University, Finland, 1995. 26 Lockwood, F.C. and N.G. Shah, “A New Radiation Solution Method for Incorporation in General Combustion Prediction Procedures”, 18th International Symposium on Combustion, The Combustion Institute, p.1405, 1981. 27 Tao, L., “Mathematical Modelling of Kraft Recovery Boiler,” NUTEK/IEA project (nr.92-03379) Recovery Furnace modelling, report no. 3 June 1997 28 Goreg, K., Cameron, J., “Kinetic study of Kraft Char gasification with Carbon Dioxide,” AIChE Mtg Boston Aug 24-27, 1986 29 Grace, T.M., Cameron, J.H., and Clay, D.T., Tappi J. 69(10): 108, 1986 30 Green R. P., Hough G., “Chemical Recovery in Alkaline Pulping Processes,” ISBN 0-89852-2555-2, Tappi Press B-046 1992. 31 Brown, C. A., Grace, T. M., Lien, S.J., Clay, D. T. “ Char Bed Burning Rates- Experimental Results,” International Chemical Recovery Conference, Proceedings

24

Ottawa, April 3-5, 1989 32 Cameron, J. H., Garce T. M., Tappi Journal 65 (7): 84-87, 1982