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A Fundamental Study of the Co-Combustion of Coke and Charcoal during Iron Ore Sintering Jiapei Zhao, Chin Eng Loo, Jinliang Yuan, Fu Wang, Jiatang Wang, Houcheng Zhang, and He Miao Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.8b00939 • Publication Date (Web): 26 Jun 2018 Downloaded from http://pubs.acs.org on June 27, 2018

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Energy & Fuels

A Fundamental Study of the Co-Combustion of Coke and Charcoal during Iron Ore Sintering Jiapei Zhao1, Chin Eng Loo2, Jinliang Yuan1*, Fu Wang1, Jiatang Wang1, Houcheng Zhang3, He Miao1

1

Faculty of Maritime and Transportation, Ningbo University, Ningbo 315211, P. R. China 2

3

*

Iron Ore Utilisation Consultant, NSW 2280, Australia

Faculty of Science, Ningbo University, Ningbo 315211, P. R. China

Corresponding author: Prof. Jinliang Yuan

Email address: [email protected] Phone: +86 574 87609545 Postal address: Faculty of Maritime and Transportation, Ningbo University, Ningbo 315211, P. R. China

1 ACS Paragon Plus Environment

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Abstract: In a global carbon cycle, the net greenhouse gas (e.g., CO2) emissions can be significantly reduced if fossil fuels could be substituted with renewable and cleaner biomass-derived fuels. In the traditional iron ore sintering process, the complete replacement of coke - a coal derived fuel - with charcoal is not possible because the two fuels have very different properties and combustion behaviours, resulting in an unacceptable deterioration in sintering performance. Consequently, only low substitution ratios can be tolerated. However, research has indicated that this ratio can be increased through altering the combustion behaviour of charcoal. Most fuel particles in a sintering bed have an encapsulated layer of fine ore and flux particles. Through intentionally altering the properties of this adhering layer, combustion behaviour can be altered, leading to improved sintering performance. This work uses a newly-developed combustion model and a 2D sintering model to appropriately describe the combustion behaviour in sintering based on fuel properties and defines the optimum thickness and porosity of the adhering layer. In this study, the required properties of the adhering layer encapsulating charcoal particles - so as to match the combustion behaviour of coke particles - is determined theoretically. This study also shows that the conditions required for different fuels to have similar sintering performance are: (a) comparable ignition temperature and overall combustion rate, and (b) comparable rates of combustion at various temperatures. Matching the overall combustion rate alone does not necessarily result in comparable sintering performance. Meanwhile, the apparent density and water holding capacity of the substituting fuels should be close to the equivalent values for coke to ensure similar granulation performance and, subsequently, the properties of the bed prepared for sintering. Until these conditions are fully met, combustion efficiency, the properties of the formed flame front (e.g., width, temperature, and speed) and, consequently, sintering performance will deteriorate. In practice, fully matching all these conditions are difficult. The present work has given guidelines on which are the critical variables of the adhering fines layer that have to be considered when charcoal is introduced into sintering and also how the variable interact to determine flame front properties.

Keywords: charcoal and coke co-combustion; iron ore sintering; modelling; adhering layer; sintering performance 2 ACS Paragon Plus Environment

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Energy & Fuels

Nomenclature

Ap

Volumetric surface area of the porous media in sintering bed (m2/m3-bed)

ks

Thermal conductivity of solid phase (W/m K)

C

Oxygen mole concentration (kmol/m3)

kr0

Static effective radial conductivity (W/m K)

C0

Oxygen concentration at rc0 at first combustion stage (-)

krg

Gas effective radial conductivity (W/m K)

C∞

Oxygen mole concentration in bulk gas (kmol/m3)

krs

Solid effective radial conductivity (W/m K)

Cpg

Gas phase heat capacity (J/kg K)

kv

Reaction rate constant based on volume (m3/kmol s)

Cs

Instantaneous carbon mole concentration in coke (kmol/m3)

mc

Instantaneous carbon mass (kg)

Cs0

Initial carbon mole concentration in coke (kmol/m3)

mc0

Initial carbon mass (kg)

D

Sinter pot inner diameter (m)

pc

Packing parameter (-)

De1

Effective oxygen diffusion coefficient in reacting coke core (m2/s)

pc1

Loose packing parameter (-)

De2

Effective oxygen diffusion coefficient in coke ash layer (m2/s)

pc2

Dense packing parameter (-)

De3

Effective oxygen diffusion coefficient in adhering layer (m2/s)

qg

Volumetric heat source of the gas phase due to all reactions (W/m3)

Di

Diffusion coefficient of gas species i in gas phase (m2/s)

qg-w

Gas-to-wall heat flux at pot wall (W/m2)

Dk

Knudsen diffusion coefficient of oxygen in coke core (m2/s)

qs

Volumetric heat source of the solid phase due to reactions (W/m3)

DO2

Effective oxygen diffusion coefficient in bulk gas (m2/s)

qs-w

Solid-to-wall heat flux at pot wall (W/m2)

Eg

Total energy of the gas phase (J/kg)

r

Particle radius coordinate (m)

Es

Total energy of the solid phase (J/kg)

r3

Radius of granule (including adhering layer) (m)

G

Superficial gas mass flux (kg/m2 s)

rc

Instantaneous carbon radius (m)

Ji

Diffusion flux of species i (kg/m2 s)

rc0

Initial carbon radius (m)

Nad,i

Number density of adhering coke particles of size interval i in packed bed (1/m3)

rpore

Average radius of pores in coke core (m)

Nnu,i

Number density of nuclear coke particles of size interval i in packed bed (1/m3)

t

Time in coke combustion modelling system (s)

Nsize

Total number of coke size intervals (-)

t'

Time in CFD modelling system (s)


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Pg

Gas static pressure (Pa)

x0.5

Particle size giving a partition coefficient of 0.5 (m)

Pr

Prandtl number (-)

v

Gas phase physical velocity vector (m/s)

Rad,i

Reaction rate of adhering coke particles of size interval i (kmol/s)

Greek symbols

Rc

Reaction rate of single carbon particle (kmol/s)

∆t'

Length of time step in CFD modelling system (s)

Re

Reynolds number (-)

α

Conversion mode parameter in eq (23) (-)

Rnu,i

Reaction rate of nuclear coke particles of size interval i (kmol/s)

β

Conversion mode parameter in eq (24) (-)

Rc,all

Reaction rate of all nuclear and adhering coke particles for full size distribution per unit bed volume (kmol/m3 s)

χi

Partition coefficient of particle size interval i (-)

Ru

Universal gas constant (J/mol K)

δ

Adhering layer thickness of granulated fuel (m)

Sc

Schmidt number (-)

εal

Porosity of adhering layer (-)

Sh3

Equivalent Sherwood number (-)

εash

Porosity of ash layer (-)

Sm

Mass source term resulted due to all reactions (kg/m3-bed s)

εc0

Porosity of initial coke core (-)

Smc0

Specific surface area of coke core (m2/kg)

εal,cr

Critical adhering layer porosity for the combustion rate of coke and charcoal to be comparable

Sm,i

Mass source term of gas species i due to all reactions (kg/m3-bed s)

ε′

Emissivity of the solid phase (-)

Sm,j

Mass source term of solid species j due to all reactions (kg/m3-bed s)

φ0

Thiele modulus based on initial carbon particle radius rc0 (-)

Tair

Ambient air temperature around external pot wall (K)

φh

Thiele modulus based on instantaneous coke radius rc (-)

TE

Melting completion temperature (K)

ηh

Effectiveness factor defined in eq (29) (-)

Tg

Gas temperature (K)

ϕ

Dimensionless oxygen concentration (-)

TM

Melting starting temperature (K)

ϕ0

Dimensionless oxygen concentration at rc0 at first combustion stage (-)

Ts

Solid temperature (K)

ϕc

Dimensionless oxygen concentration at rc at second combustion stage (-)

Vs

Volume of carbon particle (m3)

ϕ2

Dimensionless oxygen concentration at rc0 at second combustion stage (-)

Wc

Carbon molecular weight (kg/kmol)

ϕ3

Dimensionless oxygen concentration at r3 (-)

WO2

Oxygen molecular weight (kg/kmol)

λk

ratio of solid to gas thermal conductivity (-)

X

Conversion ratio of carbon (-)

λW

Thermal conductivity of pot wall (W/m K)


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Energy & Fuels

Yi

Mass fraction of solid species j (-)

µg

Dynamic viscosity of the gas (kg/m s)

Yj

Mass fraction of gas species i (-)

ν

Stoichiometric coefficient for carbon and oxygen reaction (-)

dp

Granule diameter (m)

θ

Phase change factor (-)

hair

Convection heat transfer coefficient at external pot wall (W/m2 K)

ρc

Instantaneous carbon apparent density (kg/m3)

hi

Enthalpy of species i (J/kg)

ρc0

Initial carbon apparent density (kg/m3)

hg-w

Gas-to-wall heat transfer coefficients at inner pot wall (W/m2 K)

σ

Melt fraction (-)

hm

Mass transfer coefficient (m/s)

ωal

Tortuosity factor of adhering layer (-)

hrs

Solid radiation coefficient (W/m2 K)

ωash

Tortuosity factor of ash layer (-)

hrv

Void-to-void radiation coefficient (W/m2 K)

ωc0

Tortuosity factor of initial coke core (-)

hsg

Solid-to-gas heat-transfer coefficient (W/m2 K)

ξ

Dimensionless radius (r/rc0) (-)

hs-w

Solid-to-wall heat transfer coefficients at inner pot wall (W/m2 K)

ξal

Dimensionless thickness of adhering layer (δ/rc0) (-)

hw

Bed-to-wall heat transfer coefficient (W/m2 K)

ξ3

Dimensionless radius of granule (r3/rc0) (-)

kal

Adhering layer mass transfer coefficient (m/s)

ξc

Instantaneous dimensionless carbon radius at second combustion stage (rc/rc0) (-)

kash

Ash layer mass transfer coefficient (m/s)

ξc,cr

Critical instantaneous dimensionless carbon radius for the combustion rate of coke and charcoal to be comparable

kg

Thermal conductivity of gas phase (W/m K)

ψ

Dimensionless carbon concentration (-)

kgbl

Gas boundary layer mass transfer coefficient (m/s)

τ

Dimensionless time (-)

krg

Gas effective radial conductivity (W/m K)

τcr

Dimensionless time at the end of first combustion stage (-)

kpr

Mass transfer coefficient considering the combined effects of pore diffusion and surface reaction (m/s)

τ

Stress tensor (Pa)


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1. Introduction Iron ore sintering is widely used in the world to produce lumpy agglomerates for the ironmaking blast furnace 1. The sintering process involves granulating a mixture of predominantly iron ores, fluxes, and 4 wt.% coke breeze in a rotating drum. As the mixture cascades in the drum water is added and this causes fine particles (typically less than 0.25 mm) to adhere onto the surfaces of the larger particles. The granulated mix is then laid onto a strand to form a permeable, particulate bed. As the strand moves under an ignition hood, the bed surface is set alight to form a narrow flame front. Air is drawn down through the bed and this causes the flame front to move down the bed. In other words, the combustion of coke particles occurs sequentially, a layer at a time starting from the bed surface. Typically, a flame front is between 20 to 40 mm in thickness, broadening down the bed, and it provides around 80 % of the total heat generated in the sintering process 2. This means that the combustion behaviour of the fuel particles has a huge impact on the properties of the flame front and the strength of the product sinter. In recent years, this important area has been considered by many researchers 3-9. Sintering is widely considered as an energy intensive process and a major source of pollution. The combustion of conventional fossil fuels, e.g., coke breeze and anthracite, in sintering is a significant contributor towards the total CO2, SOx and NOx emissions from a steel plant 10. The drive to greatly reduce green house gas (GHG) emissions by the governments worldwide means that the replacement of fossil fuels by cleaner and renewable energy sources is an important consideration. In sintering, a potential option is solid biomass and biomass-derived fuels, as biomass is not only carbon neutral but has a calorific value close to that of solid fossil fuels. In addition, past sintering tests have shown that, when biomass or biomass char replaces coke breeze

8, 11, 12

, SOx and NOx emissions also reduces.

Consequently, a number of reported experimental programs have considered how biomass or biomass char can be used in larger amounts in the sintering process 6, 9, 12-14. Unfortunately, the overall conclusion has been that replacing coke with untreated raw biomass, biomass char or charcoal in sintering results in a clear deterioration in sintering performance.


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Kawaguchi et al. 8 reported that the full replacement of coke with raw biomass decreased flame front temperature and the level and quality (e.g., flow and coalescencing behaviour) of the melts generated. The effect of this on sintering was not acceptable. Even with reduced biomass substitution ratios (mass percent of coke substituted by other fuels based on equivalent amount of fixed carbon or FC), the flame front temperature was adversely affected and sinter yield and strength decreased. At the same time, the tar generated in biomass devolatilization process caused unacceptable problems in the operation of the fan and de-dusting systems. As a result, the authors deemed that raw biomass is an unacceptable fuel and recommended high grade charcoal (with FC > 90 %) instead 8. Lower SOx, NOx and dust emissions were also observed in sintering tests using high grade charcoal. However, the fuel rate (weight FC consumed to produce a tonne of acceptable sinter, kg/t) for the charcoal tests were 85.2-109.2 kg/t compared to 50.5-53.9 kg/t for the coke tests. This means that when high grade charcoal is used to fully replace coke, fuel consumption rate has to be around 70 % higher. Unless the combustion behaviour of charcoal can be greatly improved, charcoal cannot be considered a viable total replacement for coke. Partial replacement is an option and this means that understanding the cocombustion behaviour of sintering beds containing a mixture of charcoal and coke is needed. Lovel et al. 7, 15 found in their sinter pot tests that when using high grade charcoal (FC of 94.3 %), around 50 % coke can be replaced without significant deterioration in sinter yield. For lower grade charcoal (FC of 72.7 %), sinter yield was significantly lowered. Lu et al.

16

also found that when the FC value of

charcoal was 70.9 %, the tolerable substitution ratio was lower than 50 %. All these results indicate that fuel reactivity has a large influence on sintering performance. It is well-established that the reactivity and thus combustion rate of charcoal are much higher than that of coke because of its higher porosity, specific surface area and intrinsic reactivity, and lower apparent density. These properties have resulted in lower combustion efficiency ηcomb (defined as the mole ratio of [CO2]/([CO2] + [CO]) of the exhaust gas i.e., exiting the bottom of the sintering bed ), lower flame front temperature, higher flame front speed and lower sinter yield and strength. Several experimental programs been consistent in confirming this influence of charcoal on the sintering process.


8, 15, 16

, have

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To increase the charcoal substitution ratio, various attempts have been made in the literature. Chen et al.

6

used iron ore fines to form a coating or adhering layer on the charcoal particles before they

were added into the granulation drum. They found that with the introduction of this additional pregranulation treatment step, substitution ratio can be increased and increases in sinter yield and strength were also obtained. In their sinter pot tests, Gan et al.

17

, using a similar additional step to

pre-granulate the charcoal particles with fines also obtained similar results. Compared to tests which did not involve this additional step, they also found that CO concentration in the waste gas decreased and sintering performance improved. In sintering, pre-granulation has an influence on the combustion behaviour of all types of fuel particles, not just charcoal. For example, Oba et al.

18

used a complex

mix of Fe2O3, Fe3O4, Fe2O3 and CaO, and also a binary mix of Fe3O4 and CaO to coat coke particles in their studies. Combustion rate determined using a thermo-balance electric combustor showed that coated coke particles burnt much slower compared to free coke particles. For a selected fuel type, combustion behaviour can also be altered by changing its properties. For example, Fan et al.

5, 19

found that through compression and then removing the volatiles through a pyrolysis process, it was possible to decrease both charcoal porosity and reactivity. In sintering, these changes resulted in decreased combustion rate, increased combustion efficiency and improved sintering performance. The studies reviewed in this paragraph show conclusively that the combustion rate of solid fuel particles can be effectively reduced by increasing the mass transfer resistance of oxygen to its surface or by decreasing its porosity and intrinsic reactivity. To date the effect of pre-granulating charcoal has only been considered experimentally. The aim of this study is to understand this area in greater detail, using a proven theoretical sintering model. A good model can provide information that would be costly to obtain experimentally and it can also allow the variables that influence the combustion process to be assessed individually and then collectively. Sintering is a very complex process and in most cases it is impossible to isolate the variables to study them one at a time. This fundamental study is aimed at obtaining important information on adhering layer, such as: (1) its thickness for charcoal to match that of coke at comparable combustion rates, (2) its optimum porosity for a defined layer thickness, (3) the optimum


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Energy & Fuels

thickness and porosity for charcoal of different particle sizes and reactivities, and (4) how catalytic effects of the layer influences charcoal combustion rate. Clearly, the accuracy of answers from such studies will depend on the robustness of the model. In the next section, the new solid fuel combustion sub-model

20

– which is a component in the 2D

sintering model – is described in some detail. Prior to the development of this combustion sub-model, coke was the only fuel considered in the modelling work. The model could not simulate highly porous fuel particles (e.g., charcoal), where some combustion will invariably occur inside the particles. The versatility of sintering model was also extended to simulate a cylindrical pot - in two dimensions across the pot diameter. In a recent study, the 2D model with the new coke combustion model has been validated using laboratory results 20. The study involving the new fuel combustion model – using fuel properties from the literature – shows trends that are fully in line with expectations from the literature. This indicates that the 2D model, incorporating the new fuel combustion model, adequately simulates a real sintering process. In this study, the model is used to study beds where coke and charcoal are combusting simultaneously.

2. Model development 2.1. Governing equations In this study, the newly-developed two-dimensional axisymmetrical transient sintering model is used to simulate sintering with charcoal and coke co-combusting. This takes place in a pilot-scale sinter pot, which is shown schematically in Figure 1. Most of the significant reactions and physical changes taking place during sintering are considered by the model. The ability to accurately validate experimental results suggests that this is the case

20

. The variations in energy, mass, species and

momentum in the axial and radial directions of the sinter pot are considered, but changes in the circumferential direction are neglected. Layers of the granulated bed are heated up as the flame front


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approaches. When a layer is in the front, localised melts form and the bed structure transforms as the solid-melt mixtures undergoes coalescence. The layer cools as the front departs. But before this happens a significant amount of heat from the front has been transferred to the layer below. At some point the amount of heat is sufficient to cause the combustion of the fuel particles in this lower layer. This sequential heating and cooling process is repeated down the bed, and manifests as a descending flame front. As expected factors like the flow rate of air through the bed – needed to sustain the combustion of the fuel particles – has a large influence on how quickly the flame front descends. The bulk density of the bed is also an important factor. Both these factors determine the changes in bed structure – which is an important variable because it influences the strength of the sinter product. As expected, the occurring structural changes are very complex. Like all models, a number of assmptions have to be made in this area and they have been discussed

21, 22

. As in the previous work

20, 23

, it is

assumed that only the heat and mass transfer processes are affect by the structural changes, which means that the bed can be considered as a reactive porous media. Only the gas and solid phases are considered in the model. The liquid water and generated melt are assumed static, and from the modelling point of view treated as solid phases. As sintering progresses the bed shrinks. The amount of shrinkage is dependent on a number of factors, e.g., ore type, basicity and bed temperature. The influence of bed shrinkage on sintering process is complex and has been less understood in the literature. As bed shrinkage is not large (usually less than 10 %) in our studies, its impact on the process is not considered. The combustion of the solid fuel particles, the heat and mass transfer processes, and complex reactions occurring in the porous media – such as limestone calcination and latent heats evolution during melt formation - are included in the model. The model is developed based on the porous media model in ANSYS Fluent 16.0, and the incorporation of user-defined functions (UDFs). The (default) superficial velocity formulation in Fluent is based on volumetric flow rate but, in sintering, this limitation greatly decreases the accuracy of model because of the important role of velocity values and gradients. Consequently, in this work the physical velocity formulation, where gas and solid phases have their individual temperatures, is adopted. The gas phase is assumed to be an incompressible ideal gas. As suggested by Komarov et al. 10 ACS Paragon Plus Environment

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Energy & Fuels

24

, gas flow in the sinter bed can be assumed laminar. The energy conservation equations for the gas

and solid phases are written as follows.

∂ (ερ g ) ∂t

(

)

+ ∇ ⋅ ερ g v = S m

(1)

where, ε is bed porosity, ρg is true gas density (kg/m3), v is gas phase physical velocity vector (m/s), Sm is the resulting mass source term attributable to various reactions (kg/m3-bed s). The conservation equation for solid mass is:

∂ [(1 − ε )ρ s ] = −Sm ∂t

(2)

where, ρs is the true density of the solid phase (kg/m3). The gas energy conservation equation can be written as 25:

∂ (ερ g E g ) ∂t

[

( )

]

  + ∇ ⋅ ε v (ρ g E g + Pg ) = ∇ ⋅ εk g ∇ Tg − ∑ (hi J i ) + ε τ v  i   ( ) + q g + hconv A p Ts − Tg

(3)

where, Eg is the total energy of the gas phase (J/kg); Pg is gas static pressure (Pa); kg is thermal conductivity of the gas phase (W/m K); hi and Ji are the enthalpy (J/kg) and diffusion flux of gas species i (kg/m2 s); τ is stress tensor (Pa); hconv is the heat transfer coefficient (W/m2 K) as expressed in previous work 23; Ap is the volumetric surface area of the porous media in the sintering bed (m2/m3bed); and qg is the volumetric heat source of the gas phase attributable to all reactions (kJ/m3 s). The last term on the right-hand side of eq (3) represents the source term because the gas and solid phases are not in equilibrium. The solid energy conservation equation can be expressed as 25:

∂ [(1 − ε )ρ s E s ] = ∇ ⋅ [(1 − ε )k s ∇ Ts ] + q s + hconv A p (Tg − Ts ) ∂t

(4)

where: ks is thermal conductivity of the solid phase (W/m K), qs is the volumetric heat source of the solid phase due to all reactions (kJ/m3 s). 11 ACS Paragon Plus Environment

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The gas momentum conservation equation is 25:

(

∂ ερ g v ∂t

) + ∇ ⋅ (ερ vv ) = −ε∇P g

g

( )

+ ∇ ⋅ ε τ + ερ g g − Fv

(5)

where, Fv is momentum loss due to the solid matrix consisting of the viscous and inertial resistance terms (kg/m2 s2). The Ergun equation

26

, applicable to flow for a wide range of Reynolds numbers

(defined based on the equivalent diameter of the particles in the packed bed), is used in this work. The species conservation equation for the gas phase can be written as:

∂ ( ερ g Yi ) ∂t

r + ∇ ⋅ ερ g vYi = ∇ ⋅ ( ε Di ∇ Yi ) + S m ,i

(

)

(6)

where, Yi is the mass fraction of species i, Di is the diffusion coefficient of species i (m2/s), Sm,i is the mass source term of gas species i due to all reactions (kg/m3-bed s). The species conservation equation for the solid phase can be written as:

[

∂ (1 − ε )ρ s Y j ∂t

]= S

(7)

m, j

where, Yj is the mass fraction of solid species j, Sm,j is the species soure term for solid species j due to all reactions.

2.2. The physicochemical processes and sub-models The complex physicochemical processes in sintering process is defined by the ten reactions summarized in Table 1. More detailed information on these reactions can be obtained from previous publications

2, 20, 23

. Recently, a new solid fuel combustion model has been developed

20

to take into

account the effects of fuel porosity and the adhering layer on the combustion process. It is wellestablished that sinter mix composition affects the melting and solidification processes and this study also attempts to improve the accuracy of the sub-model that quantifies this behaviour. In moving from


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Energy & Fuels

a 1D to 2D sintering model, the heat transfer between the sinter pot wall and the gas and solid phases is also included. These changes are discussed in greater detail below.

2.2.1. Fuel combustion model A unique property of the solid fuel in sintering is that they are not free particles and can exist in different structural configurations. This is a direct result of the granulation process. The fuel particles range from around 0.01 mm to 10 mm and the larger particles are coated by a layer of fines, while the finer fuel particles end up in the adhering layer of some larger particles. Based on a fundamental understanding of the partitioning of particle sizes – either ending up as a nuclear particle or in the adhering layer – four configurations have been defined for fuel particles. However, they can be simplified down to a generalized configuration 20, which has a fuel nuclear particle encapsulated by an adhering layer of variable thickness, as shown in Figure 2(a). Through altering the thickness of the layer, the fuel particle can be changed to become a nuclear particle or an adhering particle. In this study, to fully quantify the fuel combustion process, three definitions are used. Firstly, a fuel particle refers simply to free coke or charcoal particle (an adhering layer of 0 mm) and is defined by its fixed carbon, ash and volatile matter (VM). Secondly, a granulated fuel particle is one with an adhering layer, regardless of the thickness of the layer. Finally, the term carbon particle is loosely used to refer to particles containing fixed carbon, regardless of whether the particle is coke or charcoal. Figure 2(b) and (c) show that the combustion of porous solid fuels occurs in two stages 20. In the first stage, the carbon particle size is essentially constant while its density decreases with conversion. At the end of this stage, there is no longer any carbon on the particle surface; the thin original carbon layer is now replaced by an ash layer. The second stage is considered to commence when the core of the carbon particle starts to shrink. Changes are observed in both radius and apparent density of the carbon particle, which does not happen in the first stage. The critical time to distinguish the two stages can be determined from the following equation:


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Cs r =rc 0 = Cs r =rc 0 − kvCs0C∞ϕ0∆t′ t =t ′+∆t ′

(8)

t =t ′

where: Cs r = rc 0

t = t ′ +∆ t ′

and Cs r = rc 0 (kmol/m3) are the concentration of carbon at outmost carbone particle t =t ′

surface at time t’+∆t’ and t’, respecitively; ∆t’ is the calculation time step (s); kv is the reaction rate constant based on volume (m3/kmol s); Cs0 is the initial carbon mole concentration in fuel (kmol/m3); C∞ is the oxygen mole concentration in bulk gas (kmol/m ); and k v Cs 0 C ∞ϕ 0 (kmol/m s) is the 3

3

combustion rate in the time interval ∆t’. The first stage is completed once the Cs r = rc 0

t = t ′ +∆t ′

= 0 , and it is

also the criteria for the determination of the critical time. In the above equation, ϕ0 is dimensionless oxygen concentration at rc0 at the first combustion stage, which can be given by 20:

 D  1 1  ϕ0 = 1 + e1 (φ0 coth φ0 − 1)1 − + 2   ξ3 ξ3 Sh3   De3

−1

(9)

where: De1 and De3 are the effective diffusion coefficients of oxygen through the reacting fuel core and the adhering layer (m2/s), respectively; rc0 is the initial fuel particle radius; φ0 is the Thiele modulus defined as φ0 = rc0(vkvCs0/De1)1/2; v is the stoichiometric coefficient for carbon and oxygen reaction, v equals 0.5 for C + 0.5O2 = CO; and Sh3 is the equivalent Sherwood number defined as Sh3 =hmrc0/De3 and hm is the mass transfer coefficient for the granulated fuel particle. The combustion rate of a fuel particle of the generalized configuration is Rc (kmol-C/s), which is applicable to both the first and second stages 20:

Rc =

4πrc0ξc De1C∞ [φ0ξc coth(φ0ξc ) − 1] ν   1 D D 1  1 + e1 [φ0ξc coth(φ0ξc ) − 1]1 − ξc + e 2 ξc 1 − + 2  De2 De3  ξ3 ξ3 Sh3  

(10)

Equation (10) can be rearranged into a form elucidating various combustion resistances as 20:

Rc =

4πrc20ξ c2C∞ ν 1 1 1 1 + + + k pr k ash k al k gbl

(11)


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Energy & Fuels

where,

1 rc 0ξ c = k pr De1 [φ0ξ c coth (φ0ξ c ) − 1]

(12)

1 r ξ = c 0 c (1 − ξ c ) kash De 2

(13)

1 1 r ξ2  = c 0 c 1 −  k al De 3  ξ3 

(14)

1 1 ξc2 = k gbl hm ξ32

(15)

where: 1/kpr, 1/kash, 1/kal, and 1/kgbl, are the internal pore diffusion and reaction resistance, the ash layer diffusion resistance, the adhering layer diffusion resistance and the gas boundary layer mass transfer resistance, respectively; ξc is the instantaneous carbon radius rc in the dimensionless form, ξc = rc/rc0; In the first combustion stage, ξc = 1, thus 1/kash is eliminated since there is no ash layer formed in this stage. De2 is the effective oxygen diffusion coefficient in the produced ash layer. The dimensionless radius of granulated fuel particle, ξ3, is defined as:

ξ 3 = 1 + ξ al

(16)

where, ξal is the dimensionless thickness of the adhering layer:

ξ al =

δ

(17)

rc 0

where, δ is the adhering layer thickness (m). The effective diffusion coefficients of oxygen through the reacting fuel core, ash product layer and the adhering layer, i.e., De1, De2 and De3 are given as:

De1 =

ε c20 1 1 + Dk DO2

(18)


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2 De2 = ε ash DO2

(19)

De3 = ε al2 DO2

(20)

where, D O 2 is the oxygen diffusion coefficient in the bulk gas, approximated by the molecular diffusion coefficient of oxygen in air (m2/s) 27: D O 2 = 1 . 78 × 10 5 (T g / 273 ) 1 .75

(21)

Dk is the Knudsen diffusion coefficient of oxygen in the fuel core (m2/s), and is given by 28-30:

8RuTg 2 Dk = rpore 3 πWO2

(22)

where, rpore is the average radius of the pores in the reacting fuel core (m), Ru is the universal gas constant (J/mol K), and Tg is gas temperature. By assuming that all the pores are cylindrical, rpore can be calculated from

30

i.e., rpore=2εc0/(Smc0 ρc0). Here, Smc0 (m2/kg) and ρc0 (kg/m3) are the initial

specific surface area and apparent density of carbon in the coke particle. The above rate equations require information on the variation of carbon particle size and density, and this information, for the two combustion stages, has been derived in a previous work

20

, and is

summarized as follows. For the combustion of a porous or non-porous spherical carbon particle, the variations of particle size and density follow the rules below 31: α

ρc  mc   = ρc0  mc0  rc  mc   = rc 0  mc 0 

(23)

β

(24)

where, rc, ρc and mc are the instantaneous carbon particle radius, apparent density and mass, respectively, while rc0, ρc0 and mc0 are their initial values. The relationship α + 3β = 1 holds for


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Energy & Fuels

spherical particles. The limiting values of α = 1 and α = 0 correspond to combustion in Regime I (kinetics-limited) and Regime III (diffusion-limited), respectively. For combustion in the first stage, carbon particle size is constant at rc0, while particle apparent density can be calculated from the reaction rate in eq (11) as:

rc = rc 0

(25)

dρc RW = − c c3 dt 4 3πrc

(26)

For the second combustion stage, the determination of α and β values needs more complex mathematical manipulations. It has been theoretically determined that α and β values can be related to the effectiveness factor ηh as 20:

α ≈ 0.4η h

(27)

β ≈ (1 − 0.4η h ) 3

(28)

And, ηh is given by:

ηh =

3

φξ

2 2 0 c

[φ0ξc coth (φ0ξc ) − 1]

(29)

The physical meaning of ηh is the ratio of actual carbon oxidation rate to the maximum carbon oxidation rate attainable when gas concentrations and temperatures in the particle interior are the same as those at the particle outer surface. Equation (27) has been found to be a good approximation of the exact solution for α 20. Based on the above results, the overall fuel combustion rate per unit bed volume, Rc,all (kmol-C/m3bed s) can be calculated from the following equation 20:

Rc, all =

N size

∑[χ N k

nu, k

Rnu,k + (1 − χk ) Nad , k Rad , k ]

(30)

i \ k =1


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where, Nsize is the total number of size intervals, χk is the partition coefficient of fuel particle size interval k, which is the mass fraction of particles of size interval k acting as nuclei due to granulation.

χk is calculated from a well validated granulation model given the granulation conditions

32, 33

. Nnu,k.

and Nad,k (1/m3) are the number density of the nuclear and adhering fuel particles for every unit sinter bed volume, respectively. Rad,k and Rnu,k are the combustion rates of nuclear and adhering fuel particles of size interval k, which can be calculated from eq (11) and relevant relationships as discussed above. The only difference between the calculations of Rad,k and Rnu,k is that in Rnu,k, the adhering layer thickness ξ3 is an non-zero value output from the granulation model, while in Rad,k, ξ3 is set to 0. Even though the amount of fuel particles used in sintering is only around 4 wt.% in the mix, such a detailed treatment of their combustion behaviours is justified and necessary on the basis that they provide approximately 80 % of the total heat required for the iron ore sintering process 2. On formulating the combustion model, it was then integrated into the 2D sintering model.

2.2.2. Melting and Solidification model

Other than carbon particles, the other important components of a sinter mix are iron ore (e.g., hematite and magnetite) and fluxes (e.g., limestone, dolomite and serpentine). When in close proximity, these materials react at low temperatures to form various low melting-point eutectic systems, e.g., SiO2-FeO, Fe3O4-2FeO-SiO2, MnO-Mn2O3-SiO2. As temperatures increase, these eutectic mixtures melt and the volume of melt continues to increase as ore and flux particles are assimilated, some fully and others partially. In regions of the flame front exceeding about 1373 K, three phases i.e., solid, liquid and gas, co-exist. With the departure of the flame front and falling temperatures, the melt solidifies to become the bonding phase of the sinter. Local melts have varying compositions and solidify to form different types of bonding phases. The cooling rate also has an influence on the type and structure of these bonding phases. The complexity of the melting and solidification process means


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Energy & Fuels

simplifications have to be made in the mathematical modelling of this sub-process. A simplified thermodynamic-based model of the melting and solidification process was developed in previous work

23

. It assumed that for a given solid mixture composition, the

melt fraction σ (mass fraction of liquid melt in the solid-liquid-gas mixture) is a function of solid temperature Ts, the phase change factor θ, the temperature at which melting commences

TM, and the temperature at the completion of melting TE. Below TM no melt forms in the system, i.e., σ = 0. Above TE all solids are fully melted, i.e., σ = 1. Solidification and melting are inverse processes, except that they have different latent heat values. The melt fraction σ can therefore be expressed by 23:

  1,   T − T σ =  s M  TE − TM   0, 

Ts > TE θ

  , 

TM ≤ Ts ≤ TE

(31)

Ts < TM

In the above equation, θ, TM, and TE are strongly dependent on mix composition and sinter basicity (defined as lime to silica ratio, and is one of the most important variables in sintering). These dependences are very complex and have not been considered in all our previous studies. In this study a thermochemical software, FactSage 6.4, is used to model the melting and solidification behavior of the sinter mix. How exactly this is done is best demonstrated by an example. Table 2 shows the composition of the chosen high basicity sinter mix. Figure 3 shows the dependence of melt fraction on temperature, as predicted by FactSage. Using the melting and solidification model, expressed by eq (31), it is possible to obtain a curve that best-fit the FactSage results. This curve is also shown in Figure 3 and its θ, TM and TE values are 1.34, 1403 K and 1643 K, respectively. In this work, the melting and


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solidification model is improved through using this technique to take into account differences in sinter mix composition. Clearly this will improve the predictions of the 2D sintering model.

2.2.3. Heat transfer between the bed and wall In the 2D sintering model, heat transfer between the transforming particulate bed and pot (steel) wall needs to be taken into account, as there will be some inevitable heat loss from the the pot wall to the surrounding air. It is assumed that heat flow from the bed to the wall is only through conduction, convection and radiation, and heat flow from the outer wall to the surrounding air is through convection. Heat losses from the gas and solid phases, qg-w and qs-w , are given by 20:

qg − w =

qs−w =

Tair − Tg

(32)

1 / hg − w + δ w / λw + 1 / hair

1 / hs − w

Tair − Ts + δ w / λ w + 1 / hair

(33)

where, Tg, Ts and Tair are the temperatures of the gas and solid phases adjacent to the inner pot wall, and ambient air around the pot; hg-w and hs-w are the heat transfer coefficients at the gas-to-wall and solid-to-wall interfaces, respectively; hair is the heat transfer coefficient between the outside pot wall and the atmosphere; δw and λw are the thickness and thermal conductivity of the pot wall, respectively. The correlations for the heat transfer coefficients of hg-w and hs-w are described below. The gas-to-wall heat transfer coefficient, hg-w, can be expressed as 34, 35:

hg −w = hwkrg /(krg + krs )

(34)

While the solid-to-wall heat transfer coefficient, hs-w, is given by 34, 35:

hs −w = hwkrs /(krg + krs )

(35)

where, the bed-to-wall heat transfer coefficient, hw, is written as 34, 35:


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Energy & Fuels

hw = 2.44kr0D−4 3 + 0.033kg Pr Re d p−1

(36)

The static effective radial conductivity, kr0, is 34, 35:

 d h k r 0 = k g ε 1 + p rv  kg 

 k g (1 − ε ) + −1     1 hrs d p  2 + + p k g  3kλ  c

(37)

And the gas effective radial conductivity, krg, is expressed as 34, 35: 2   d h    d p    p rv   krg = k g ε 1 + + 0.14 Pr Re/ 1 + 46     kg    D     

(38)

The solid effective radial conductivity, krs, is 34, 35:

 1 h d −1 2   krs = kg (1 − ε ) /  + rs p  + ks  3λk   pc 

(39)

The solid conductivity, ks, is given by:

k s = 3 .14 × 10 −3 Ts + 1 .797

(40)

The void-to-void radiation coefficient, hrv, is written as 34, 35:

 ε  1 − ε ′  hrv = 2.27 ×10−7 Tg3 / 1 +    2(1 − ε )  ε ′ 

(41)

Solid radiation coefficient, hrs, is described as 34, 35:

 ε′  3 hrs = 2.27 × 10− 7  Ts  2 −ε 

(42)

Packing parameter, pc, is 34, 35:

pc = pc 2 + ( pc1 − pc 2 )

ε − 0.260

(43)

0.476 − 0.260

where, pc1 and pc2 are given by 34, 35:


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2

 λ −1  0.3525  k  2  λk  − pc1 = 0.4569 ( λk − 1) 3λk ln λk − 0.5431( λk − 1)  − λ

(44)

k

2

 λ −1  0.07217  k  λk  2  − pc 2 = 0.07498 ( λk − 1) 3λk ln  λk − 0.9250 ( λk − 1)  − λ

(45)

k

The solid-to-gas heat-transfer coefficient, hsg, is described as 34, 35:

hsg =

2.06Cp g G

ε

Re −0.575 Pr − 2 3

(46)

In the above equations, Reynolds, Prandtl and Schmidt numbers are defined as, Re = d p G / µ g ,

Pr = Cp g µ g / k g , Sc = µ g /( ρ g Dm ) . λ k = k s / k g . G is the superficial gas mass flux (kg/m2 s). D is the diameter of the sinter pot. Dm is the diffusion coefficient. ε ‘ is the emissivity of the solid phase.

2.2.4. Other sub-models All the sub-models in the 2D sintering model are listed in Table 1. Those not discussed so far are: (1) CO oxidation, (2) limestone (CaCO3) calcination, (3) dolomite (CaMg(CO3)2) calcination, (4) serpentine (Mg3Si2O5(OH)4) dehydroxylation, (5) goethite (α-FeOOH) dehydration, (6) magnetite (Fe3O4) oxidation, (7) moisture drying and (8) condensation. The interested reader can refer to previous work

2, 20, 23

for detailed information on these reactions e.g., the mechanisms and associated

physicochemical processes.

2.3. Boundary and initial conditions


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Energy & Fuels

The governing equations described in Section 2.1 are subject to important boundary and initial conditions. These are briefly summarised here. The boundary conditions have to reflect both the ignition and sintering periods. For the conservation equation of gas energy, the temperature of the ignition hood was used as the initial gas temperature and ambient temperature was assumed thereafter. Throughout the two periods fully developed flow at the pot outlet is assumed. The wall of the sinter pot is a boundary and heat exchange in this area is an important consideration. Eq (32) is assumed to fully describe the heat exchange between the gas and the pot wall. For the conservation equation of solid energy, during ignition, radiative and convective heat is transfer between the ignition hood and the top bed surface. While during sintering, the top surface exchanges heat with the ambient air. At the bottom surface, the wind box temperature determines the level of radiative and convective heat exchange. Heat transfer between the solid and pot wall is quantified by Eq (33) during the whole sintering period. Just as temperatures have to be defined to satisfy the gas energy conservation equation, pressures have to be defined for the gas momentum conservation equation. As with temperature, pressures at the top and bottom of the bed are defined by the sintering conditions used i.e., the inlet and outlet gas pressures. The flow at the bottom of the bed is assumed to be fully developed. Zero shear stress condition is assumed at the pot wall. The inlet gas species during and after ignition is defined by the composition of the gases in the ignition hood and of the ambient air. The absence of shear stress at the pot wall also means that a zero flux condition exists. There are no boundary conditions for the solid species conservation equation, Eq (7), since solid species are assumed to only vary with time. The properties of the original sinter bed defines all the initial conditions for gas and solid temperatures, compositions and momentum.


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2.4. Solution methodology The 2D axisymmetrical sinter bed was discretized into 33 cells in the radial direction and 120 cells in the axial direction. Denser meshes were generated in the zone close to the pot wall. To discretize the spatial terms, the second-order upwind scheme was employed. The first-order implicit scheme was adopted to desritize the transient terms. The governing equations are non-linear and were solved iteratively. It was found that CO oxidation reaction is more difficult to converge compared to other variables. For each time step, the solutions were considered to be converged when all variables met their own convergence criteria.

3. Experimental Results from carefully controlled tests on a pilot-scale sinter pot were used to validate the 2D sintering model. All the experimental results have been discussed in earlier publications 2, 20, 36, 37. The dimension of the cylindrical sinter pot is 600 mm in height and 330 mm in diameter, as shown in Figure 1. The inner pot wall was made of stainless steel. To reduce heat lossed from the pot wall, the wall was covered by a thick layer of insulation materials. However, a small amount of heat loss is still expected. The purpose of this paragraph is to provide an understanding of how sinter tests are carried out in a laboratory and what experimental results are used to validate the sintering model. Typically, a sinter test starts with the granulation of a defined sinter mix blend, comprised of iron ore particles (~65 wt. %, a blend of hematite, magnetite, etc.), returned sinter fines (~20 wt. %), coke particles (~4.0 wt. %), and fluxes particles (~11 wt. %). These are mixed in their as-received state in a rotating granulation drum for a couple of minutes before water - around 8 wt. % (based on total dry blend mass) - is sprayed onto the cascading mix surface. The dispersed water appear as little lenses adhering fine particles onto the surface of the larger particles. Some water is absorbed by the porous particles and it is important to make allowance for this intra-particle water otherwise granulation efficiency will be adversely affected. After a standard granulation period, the entire mix was carefully removed 24 ACS Paragon Plus Environment

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from the drum and charged into the sinter pot to form a packed bed, as illustrated in Figure 1. The bed is then ignited at the top surface for 1.5 minutes by a natural gas ignitor to establish a thin flame front. The suction in the wind box draws the air and combustion gas through the bed. As a result, the flame front is slowly drawn down the bed (~0.5 mm/s). The end of sintering is when the flame front has reached to the bottom of the bed, and the time required to sinter the entire bed, or sintering time, is dependent on factors such as the permeability and bulk density of the bed, and the suction across the bed. Fairly typical values are 25 to 35 minutes. All tests aimed at obtaining results for use in model validation have thermocouples inserted at 100, 300 and 500 mm down the bed to record the temperature of flame front. There is also a thermocouple located in the wind box to record the waste gas temperature, and the maximum recorded temperature occurs at the bed burnt-through point. Results from the bed thermocouples are termed temperature-time profiles. From the profiles, three important parameters can be derived

2, 23

and they are: the maximum flame front temperature (FFT),

the residence time above 1373 K (RT), and the area enclosed (EA) between the 1373 K line and the temperature-time profile. Reported FFT, RT and EA results are averaged values for the three bed locations. Flame front speed (FFS) is the other important parameter in sintering, which is the average speed of the flame front as it descends in the bed. FFS is the ratio of bed height to sintering time (ST). Indicators of sintering performance are these six parameters: FFT, RT, EA, FFS, ST and ηcomb (combustion efficiency). The sintering model generates theoretical temperature profiles from which five of these parameters are obtained. Combustion efficiency in pot tests is determined from waste gas analysis; and theoretically from the combustion sub-model. Good comparisons in all these areas indicate that a high degree of validation is achieved. Regardless of the fuel type used – in this case coke or charcoal – these parameters are the indicators of sintering performance. On a commercial sinter machine, the measures of performance are: (a) solid fuel rate, which is the kg fuel required to produce a tonne of blast furnace sinter, (b) productivity, which is the kg of blast furnace sinter produced per unit area on the sinter machine, and (c) the quality of the sinter product, of which tumble strength is a critical measure. Measured temperature profile is an expression of the heat treatment process available at a particular location in the bed. For the same total heat available, a


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narrow and high temperature profile is preferable to a broader and flatter profile because the utilisation of heat is less efficient in the latter case. For high productivity operations the flame front has to descend at a higher speed and this narrows the temperature profile, which usually leads to reduced sinter strength. Past studies indicate that EA, which depends on all the other five parameters, has a large influence on sinter strength – which is not surprising because it is a measure of the total heat available at melt formation temperatures. It has also been shown that another sinter quality parameter – low-temperature reduction degradation index – is dependent on the cooling rate experienced by the sinter, which is the downward sloping section of the profile after the maximum temperature value. Because of the important influence of temperature profile on the sintering process, its use as a validation criterion is not only justified but appropriate.

4. Results and discussion 4.1. Model validation Figure 4 shows an example of how well the predicted and measured temperature-time profiles compare. Predicted maximum flame front temperatures at the three bed locations agree well with the experimental results. The predicted time at which the flame front arrives and departs from each of the the three locations are close to the experimental results. This indicates that flame front speed predictions are good. As coke combustion is the dominant heat source in the sintering bed, the ability to accurately predict flame front position and temperature implies the developed combustion model is reasonable. At the same time, a good match of the heating and cooling sections of the temperaturetime profiles also means that the melting and solidification and heat transfer sub-models are realistic. It is to be noted that the measured maximum flame front temperature at 100 mm down the bed is higher than the predicted value. A reason for this could be the influence of the natural gas ignition flame on the tip of this top thermocouple (especially true with bed shrinkage) or because the heat input into the bed by the flame is underestimated in the model. It is also noted that the measured maximum flame front temperature down the bed did not show an increasing trend, with the 26 ACS Paragon Plus Environment

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temperature at 500 mm slightly lower than that at 300 mm. Such behaviour is probably caused by local differences in the way the melt has coalesced and reshaped, and the amount of gases generated there

22, 38, 39

. Such variations in melt behaviour coupled with bed compression (due to gravity) can

have a significant influence on bed structure and, consequently, the local heat and mass transfer processes occurring at high temperatures. Ohno et al.

4

suggested that the formation of melt will

decrease the diffusion rate of oxygen in the coke combustion process. At high temperatures, if combustion rate is lowered, FFT is expected to decrease.

4.2. Sintering performance with coke and untreated charcoal In this study, the combustion of different fuels - coke, untreated charcoal and treated (pregranulated) charcoal – is simulated using the new combustion sub-model and the effect of this on the sintering performances is determined using the 2D sintering model. Eight cases are designed and they are shown in Table 3. The properties of the fuels (except size distribution), green bed and sintering conditions are given in Table 4. Normal size distribution of coke and charcoal used in this work are given in Table 5. The properties of coke and charcoal are from the paper by Murakami et al. 13, and are Coke 3 and Charcoal 3 in their paper. Compared to coke, charcoal has a higher porosity, higher specific surface area, lower apparent density, higher frequency factor and lower activation energy. Ignition temperature of charcoal is also believed to be lower than that of coke. These properties are expected to increase the combustion rate of charcoal and alter sintering performance. Figure 5 shows the temperature-time profiles at 100, 300 and 500 mm down the bed for sintering with coke (Case 1) and charcoal (Case 2), respectively. Table 3 shows the results of FFS, ηcomb, FFT, RT and EA for sintering with coke and charcoal. It is seen that the FFS for charcoal is around 50 % faster than that of coke, but combustion efficiency ηcomb is 30 % lower than that of coke. FFT for charcoal is 239 K lower than the equivalent value for coke. It is seen in Figure 5 that the maximum temperature reached for coke is 1650 K, which is much higher than that of charcoal at 1350 K. This temperature is too low and incapable of sustaining the endothermic sintering reactions. It is generally agreed that significant


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sintering reactions cannot take place below 1373 K 23, 40. The reasons for lower FFT for sintering with charcoal include 20: (1) insufficient heat generation due to the low ηcomb value for charcoal combustion, (2) low heat utilization efficiency because of the mismatch between FFS and HTFS (heat transfer front speed), (3) low igntition temperature and high combustion rate at low temperatures which broadens the distribution of heat and the temperature profile. In the following discussions, attempts will be made to match charcoal and coke combustion rates; if this can be achieved then the use of charcoal in iron ore sintering is certainly a viable option.

4.3. Sintering performance with coke and pre-granulated charcoal 4.3.1. Effect of adhering layer properties One objective of the present work is to define the structural configuration of charcoal so that its combustion rate approaches that of coke. Equation (11) shows that fuel combustion rate varies with fuel physical, chemical and configuration properties. In this work, charcoal and coke are assumed to be of the same size distribution. Thus, in order to make the combustion rates of the two fuels comparable, the only adjustable parameters are ξ3 and De3, i.e., the thickness of the adhering layer and the effective diffustion coefficient of oxygen through the adhering layer – which, according to eq (20), changes when the porosity of the adhering layer, εal, is altered. Using the model, the combustion rates of coke and pre-granulated charcoal with the adhering layer of varying thickness and porosity were determined. Figure 6 shows the combustion rates of single coke and charcoal particles of the same size, with no adhering layer, at different temperatures (1073, 1373 and 1673 K). The combustion rates are plotted against carbon conversion ratio X (cumulative mass fraction of carbon consumed during combustion). Typical sintering conditions were used and the initial coke and charcoal particle size dp = 1 mm, physical velocity of bulk gas ug = 3.87 m/s, oxygen molar fraction in bulk gas NO2 = 0.13. Results show clearly that the combustion rate of charcoal is much higher than that of coke. At the temperature of 1673 K, charcoal combustion rate is initially 3.1


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times that of coke, then increasing to 13.9 times that of coke near burnout. At the the lower temperature of 1073 K, the corresponding values are much larger - 11.1 times at the start and increasing to 74.2 times close to burnout. It is seen that the difference between charcoal and coke combustion rates is larger at lower temperatures. This is due to the fact that at lower temperatures, combustion tends to be controlled by intrinsic reaction kinetics, and as has been discussed above that the intrinsic reaction rate of charcoal is expected to be much higher than coke due to its physicial and chemical properties shown in Table 4. In this study, it is assumed that the adhering layers on the charcoal particles are placed there using a pre-granulation process. The pre-granuated charcoal together with all the other sinter mix components are then granulated under standard conditions and the properties of the granulated mix is given by the granulation sub-model. For the pre-granulation stage, it is assumed that charcoal particles of certain size or size range is granulated with alumina powder (Al2O3, -60 µm) and sufficient water to yield composite fuel particles of a larger uniform size, e.g., 2 mm. Binder materials, e.g., bentonite, can be added in pre-granulation to enhance granulation and yield the desired structure parameters of the adhering layer. Alumina powder is assumed to be inert in the combustion conditions and, therefore, it can be assumed that the layer only acts as a physical barrier of gas diffusion during combustion. As alumina powder is very fine, it is assumed that the adhering layer is strong and rigid, and the produced charcoal granules have the generalized configuration as shown in Figure 2. To illustrate the effect of adhering layer thickness on combustion rate, Figure 7 shows the combustion rate of single charcoal particles with and without an adhering under the same sintering conditions – which are also shown in Figure 6. Results show that the existence of an adhering layer greatly decreases combustion rate. Essentially, oxygen diffusion resistance increases because of the presence of an adhering layer. Figure 8 shows the effect of adhering layer porosity on combustion rate. Under sintering conditions, similar to those used to obtain results in Figure 6 and Figure 7, Figure 8 shows that as the adhering layer porosity for pre-granulate charcoal is decreaed from 0.315 to 0.215, combustion rate decreased significantly. The deceases are large because, as given in Eq (20), the effective diffusion coefficient is proportional to the square of adhering layer porosity. 29 ACS Paragon Plus Environment

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4.3.2. Equalizing the combustion rate of charcoal and coke Results from this study can have practical implications if it can provide guidelines on the properties of the pre-granulated charcoal product and, in particular, a definition of the adhering layer structure parameters (i.e., ξal and εal). This information can be obtained through using different permutations and combinations of ξal and εal until optimum conditions are reached. However, this tedious approach is avoided in this work. Alternatively, these optimum parameters are determined by adopting the following theoretical analysis. Firstly, the combustion resistance of 1 mm coke and charcoal particles with no adhering layer at typical sintering conditions shows that 1/kpr is the dominant factor controlling coke combustion rate, and this resistance is much larger than that for charcoal. The resistance for charcoal combustion, 1/kpr, is of very similar magnitude to the other resistances, i.e., 1/kash, and 1/kgbl for charcoal. To equalize the combustion rate of coke and charcoal particles of the same size, the charcoal particle needs to have a thicker adhering layer to decrease its combustion rate. Coke combustion rate is dominated by 1/kpr, and its value is much larger than the values of 1/kash and 1/kgbl for coke. For both coke and charcoal of the same size, 1/kash of both fuels have quite similar magnitudes. This is also the case for 1/kgbl. Therefore, to equalize the combustion rate of coke and charcoal, it is reasonable to equalize the 1/kpr for coke (more strictly termed 1/kpr_coke) to 1/kal for charcoal (termed 1/kal_char) for the same size:

1 k pr _ coke

=

1

(47)

k al _ char

as,

rc 0ξ c ,cr

De1 [φ0ξ c,cr coth(φ0ξ c ,cr ) − 1]

=

rc 0ξ c2,cr  1 1 −  De3  ξ 3 

(48)

where, ξc,cr is the instantaneous dimensionless carbon radius at the critical adhering layer porosity εal,cr.


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When the combustion rate of coke and charcoal are comparable, the value of De3 can be expressed as:

 1 De3 = De1ξ c, cr 1 − [φ0ξc , cr coth (φ0ξ c, cr ) − 1]  ξ3 

(49)

Substituting the expressions of ξ3 (eq (16)) and De3 (eq (20)) into eq (49) yields,

ε al ,cr

 D  1 = ξ c ,cr e1  1 − ξ al D 1 + O32  

   ϕ0ξ c ,cr coth (ϕ 0ξ c ,cr ) − 1   

−1

(50)

Equation (50) shows that the critical porosity of the adhering layer - for the same coke and charcoal combustion rate - is dependent on the thickness of the adhering layer around the charcoal particle ξal, the reaction extent (reflected by ξc,cr), the pore structure of coke core (reflected by De1), as well as the reactivity of coke (reflected by φ0). This basically indicates that to reduce the combustion rate of charcoal particles to that of coke, a thicker adhering layer of lower porosity is required. Figure 9 shows the variation of εal,cr with ξc,cr for 1 mm coke and charcoal particles at various temperatures (i.e., 1073, 1373 and 1673 K). The thickness of the adhering layer around the charcoal particle ξal is fixed at 0.4. Results indicate that the relationship between εal,cr and ξc,cr is nearly linear. At 1373 K, in order to match the initial combustion rate of granulated charcoal (when ξc,cr = 1) to that of coke, the critical porosity of the adhering layer convering charcoal, εal,cr, needs to be 0.24. As ξc,cr reduces to 0.5, the porosity of the adhering layer has to be reduced to 0.12. It is also seen that at lower temperatures,

εal,cr needs to be lowered in order to reduce the combustion rate of pre-granulated charcoal to that of coke. This is to be expected because at lower temperatures, combustion is controlled more by 1/kpr, i.e., pore diffusion and surface reaction, and 1/kpr for coke combustion is much smaller than that of charcoal. Therefore, a lower adhering layer porosity is required in order to reduce the combustion rate of charcoal. The combustion rate of coke and pre-granulated charcoal particles changes with carbon conversion. For different ξc,cr values, i.e., 1, 0.5 and 0.1, Figure 10 presents the combustion rate results at 1373 K


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for dp = 1 mm and ξal = 0.4. Figure 10(a) shows the variation in the combustion rate of coke and pregranulated charcoal with ξc assuming the three ξc,cr values. At ξc,cr = 1, the combustion rate of pregranulated charcoal is slightly slower than that of coke at ξc = 1. Similarly, at ξc,cr = 0.5 and 0.1, the combustion rate of pre-granulated charcoal is slightly slower than that of coke at the same ξc values. These results not only provide greater insights into the meaning of ξc,cr but also shows the value of the theory developed in this work. Clearly, ξal, ξc,cr and εal,cr, are the critical parameters that control the combustion rate of pre-granulated charcoal, and sintering performance can only be improved through changing these values. Furthermore, these three parameters for pre-granulated charcoal – to have comparable combustion performance as free coke particles - can be defined. Nonetheless, it should be noted that this analysis assumes that the combustion of free coke in iron ore sintering are controlled by pore diffusion and surface reactions in the reacting carbon core, and the overall combustion rate of free coke particles is much slower that of free charcoal particles of the same particle sizes. If coke combustion is controlled by mass transfer through the ash layer or the gas boundary layer, the results above would not be applicable. Generally, the ash value for coke is not large (10 to 15 wt. %), thus the formed ash layer will be quite porous and may only create a small hindrance to gas diffusion through the ash layer. To obtain a great understanding into the importance of carbon conversion, Figure 10(b) shows the variation of the combustion rates of coke and pre-granulated charcoal for the three ξc,cr values. It is seen that when mass transfer through the adhering layer dominates, the combustion rate of charcoal is nearly constant until the particle is almost fully consumed. It is to be noted that when any two of the three parameters, i.e., ξal, ξc,cr and εal,cr , are defined, the third can be calculated using the above methodology. In iron ore sintering, the permeability of the prepared granulated bed on the strand is a crucial parameter. Larger particles form more permeable bed and are, therefore, preferable over finer particles and this is also true for fuel particles. Based on this consideration, the following pre-granulated charcoal structure configurations are proposed. The size of charcoal particles is classified into four size groups: Group 1: -0.1 mm (smaller than 0.1 mm)


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Group 2: +0.1 -1 mm (larger than 0.1 mm and smaller than 1 mm) Group 3: +1 -3 mm Group 4: +3 -10 mm In the four size groups, Group 1 refers to ultra-fine fuel particles. This type of particles are usually embedded into the adhering layer of the granules, and they have a high combustion rate. Therefore, it is beneficial that Group 1 charcoal particles are converted into partcles of the P-type configuration (pellet of well-mixed fine charcoal and fine alumina) during pre-granulation prior to their entering the normal granulation process. The aim would be to obtain uniform 2 mm particles at the completion of pre-granulation step. From the results of this study the porosity of the pre-granulated fuel should be around 0.05 to reduce the combustion rate to optimum levels. Group 2 charcoal particles will gravitate towards the S-type particle configuration (single coarse coke particle coated with fine alumina), with an adhering layer. For these particles, the target granulated size should be 3 mm, and the adhering layer porosity 0.15. For Group 3 charcoal particles, the corresponding values should be 4 mm in size and 0.35 in porosity. For Group 4 charcoal particles, optimum values are 10 mm and 0.5. In deriving the porosity and thickness of the adhering layers, the value of ξc,cr used is 0.75 regardless of the particle groups. The robustness of these proposed values will be assessed in the next section when charcoal particles covering the full size distribution are considered.

4.3.3. Combustion of pre-granulated charcoal of full size distribution In commercial sintering operations, fuel particles of a wide size distribution are used and they are added to the mix on a weight basis. Charcoal has a lower apparent density and higher porosity than coke and, if both materials have the same size distribution, there will be some differences in the number of particles and, therefore, surface area available for reactions. The larger number of charcoal particles will result in increased overall reaction rate. Simulations were carried out to study the combustion of pre-granulated charcoal and free coke particles at different temperatures (1073, 1373


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and 1673 K). The combustion of each charcoal size fraction is then evaluated and combined based on the size distribution to obtain an overall combustion rate. It is assumed that both coke and charcoal have the same fixed carbon mass and size distribution. The structural parameters of the adhering layer on charcoal particles are similar to those used in the previous section. Figure 11 shows the results of carbon conversion with time for the two fuels. As observed earlier, the existence of the adhering layer effectively suppressed the combustion rate of charcoal. Without an adhering layer, the combustion rate of charcoal is much higher that of coke. With an adhering layer, the combustion rate of charcoal can be lowered to approach that of coke at the temperatures of 1373 and 1673 K. However, at the lower temperature of 1073 K, the retardation role of the adhering layer is not as effective. Similar results are observed in Figure 9. At this low temperature, to match the combustion rate of charcoal with that of coke, the adhering layer thickness of charcoal must be further increased and the porosity decreased. The results in this section show that the parameters of the adhering layer can effectively decrease charcoal combustion rate especially at the temperatures higher than 1373 K. The effectiveness of the adhering layer in decreasing charcoal combustion rate will be further validated against pilot-scale sintering results. This is the topic of the next section.

4.3.4. Combustion of pre-granulated charcoal in iron ore sintering The structural properties of pre-charcoal required to obtain comparable combustion rate to coke has been defined. The sintering performance of this charcoal is assessed under the conditions as given in Section 4.2. Figure 12 shows the temperature profiles at 500 mm down the bed for sintering with coke (Case 1) and pre-granulated charcoal, at the charcoal substitution ratio of 30, 40, 50, and 70 % (Case 3 – Case 6). The sintering performance parameters are shown in Table 3. In this work, the substitution is not based on equivalent energy but on fixed carbon, which is a more meaningful approach - as suggested by Chen et al. 6. The substitution ratio of 30 % means 30 % of the fixed carbon of coke is replaced by the same mass of fixed carbon from charcoal. It is seen from Figure 12 and Table 3 that the FFS for sintering with pre-granulated charcoal matches very well with that of


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coke for various substitution ratios. This indicates that the overall combustion rate of charcoal is close to that of coke, which suggests the assumed properties of the adhering layer on the pre-granulated charcoal is reasonable. As charcoal substitution ratio is increased from 30 % to 70 %, FFT is decreased from 1510 K to 1463 K, ηcomb is decreased from 88.3 % to 78.4 %, RT decreased from 2.66 min to 2.16 min, and EA decreased from 414.1 min K to 265.7 min K. The changes in these values indicate that sintering performance has decreased significantly with increasing substitution ratio. At 30 % substitution ratio, the EA value is around 80 % of that for the coke only test. This could mean that a substitution ratio 30 % is acceptable when charcoal pre-granulation is introduced. If charcoal is not pre-granulated, results of Case 8 indicate that for the same decline in EA value, the substitution ratio is around 10 % to 15 %. Essentially, with pre-granulation the amount of charcoal added to a sinter mix could be doubled. It is interesting to note from the temperature profiles in Figure 12 that temperature increases are slower above 1200 K when pre-granulated charcoal is used. This is believed to be related to the combustion characteristics of the pre-granulated charcoal. As shown in Figure 10, above the critical

ξc,cr value, the combustion rate of pre-granulated charcoal is lower than that of coke – resulting in a lower rate of heat generation at higher temperatures and, therefore, reduced temperature rise rate. At temperatures below 1200 K, the combustion rate of charcoal is faster. In iron ore sintering, higher combustion rate at low temperatures is not desirable, as it causes the temperature profile to broaden. The resulting lower temperatures mean that the generated CO from carbon oxidation is not converted to CO2 as efficiently, and combustion efficiency decreases. This is verified by the tests where the fine charcoal particles in the fuel have been removed (Case 7). Figure 12 also shows the temperature profile for pre-granulated charcoal that have the minus 0.5 mm fraction removed and substituted with the same amount of +1 -2 mm particles. For a substitution ratio of 40 %, results in Figure 12 and Table 3 show that the FFT and ηcomb for Case 7 is much higher than that for pre-granulated charcoal containing fine particles (Case 3 to Case 6) and is even higher than that for Case 1 – where coke is the fuel source. However, because the combustion rate of fine charcoal particles is higher that of coke at low temperatures, this results in reduced combustion efficiency. 35 ACS Paragon Plus Environment

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Removing the fine particles can greatly increase combustion efficiency and increase heat generation at higher temperatures, which then raises the FFT value. This indicates that the time at which combustion heat is released is also an important factor influencing FFT. Releasing the combustion heat at high temperatures, e.g. 1200 K, is the preferred option because releasing too much heat at lower temperatures inevitably results in decreased maximum flame front temperature and lower combustion efficiency.

5. Conclusions In iron ore sintering, the ability to replace coke with renewable biomass fuels will benefit the environment. However, results from studies in this area indicate that a complete replacement is not possible because the properties of the biomass/charcoal and coke are so different. To obtain a greater understanding of acceptable substitution ratios when using charcoal, the co-combustion of charcoal and coke in sintering is studied theoretically in this work. A mathematical combustion model capable of taking into consideration the structural configuration of fuel particles in sintering was developed. Basically in sintering, fuel particles have a fine adhering layer on their surface. When a fuel particle has an adhering layer, its combustion rate decreases. This means that intentionally forming an adhering layer of specified properties on charcoal particles is a technology that can be used to make the combustion behaviour of charcoal approach that of coke. Theoretical analyses were carried out in this area, using single fuel particles as well as particles in a typical full size distribution, with the aim of defining the required properties of the adhering fines layer. Theoretical analysis indicate that it is possible to match the overall combustion rate of charcoal to that of coke if the adhering layer configuration parameters, ξal, ξc,cr and εal,cr, are appropriately determined. Results from the combustion model were then integrated into a 2D sintering model. This model has been calibrated to simulate an experimental pilot-scale sintering process. The combustion and


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sintering performances for different fuels, i.e., coke, untreated charcoal and treated (pre-granulated) charcoal were studied in detail and compared with available data in the literature. Results indicate that fuel particle size, apparent density, porosity and intrinsic reactivity are important factors controlling combustion behaviour, and the values of these parameters vary significantly for the different fuels. Results show that the combustion rate of pre-granulated charcoal particles - which have similar overall combustion rates as that of coke - was fast at low temperatures (e.g., 973 – 1373 K), but slow at higher temperatures (e.g., 1373 – 1673 K). Such combustion characteristics resulted in a FFS value similar to that for coke tests. However, sintering performance deteriorated because of lower overall combustion efficiency, lower flame front temperature, and shorter residence time at high temperatures. This study also shows that sintering performance with charcoal in the mix can be improved through using only coarser charcoal particles. Fine charcoal particles (e.g., -0.5 mm) have a very detrimental influence on combustion efficiency and sintering performance. Removing these fine charcoal particles greatly improved sintering performance. Commercial coke breeze usually has significant -1 mm particles. If charcoal has a comparable size distribution as coke, then substitution ratio can hardly exceed 50 % even if the charcoal is pre-granulated to obtain optimum adhering fines layer properties. Although previous studies have highlighted the significance of matching the overall combustion rate of charcoal to that of coke, results of this work show that achieving this does not necessarily mean that sintering performance is able to match that obtained for coke. The conditions necessary for different fuels to have comparable sintering performance are: (a) comparable ignition temperature and overall combustion rate, and (b) comparable rates of combustion at various temperatures. These conditions determine combustion efficiency, flame front temperature, flame front speed and, therefore, the sintering performance. The apparent density and water holding capacity of the substituting fuels should also be close to the equivalent values for coke to ensure similar granulation and green bed conditions.


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Acknowledgements The authors acknowledge the financial support from the National Natural Science Foundation of China (Nos. 51706113 and 51706112), and the K.C. Wong Magna Fund in Ningbo University.

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(11) Ooi, T. C.; Aries, E.; Ewan, B. C.; Thompson, D.; Anderson, D. R.; Fisher, R.; Fray, T.; Tognarelli, D. The study of sunflower seed husks as a fuel in the iron ore sintering process, Miner. Eng. 2008, 21 (2), 167-177. (12) Ooi, T. C.; Thompson, D.; Anderson, D. R.; Fisher, R.; Fray, T.; Zandi, M. The effect of charcoal combustion on iron-ore sintering performance and emission of persistent organic pollutants, Combust. Flame 2011, 158 (5), 979-987. (13) Murakami, K.; Sugawara, K.; Kawaguchi, T. Analysis of combustion rate of various carbon materials for iron ore sintering process, ISIJ Int. 2013, 53 (9), 1580-1587. (14) Zandi, M.; Martinez-Pacheco, M.; Fray, T. A. Biomass for iron ore sintering, Miner. Eng. 2010, 23 (14), 1139-1145. (15) Lovel, R. R.; Vining, K. R.; Dell 'Amico, M. The influence of fuel reactivity on iron ore sintering, ISIJ Int. 2009, 49 (2), 195-202. (16) Lu, L.; Adam, M.; Kilburn, M.; Hapugoda, S.; Somerville, M.; Jahanshahi, S.; Mathieson, J. G. Substitution of charcoal for coke breeze in iron ore sintering, ISIJ Int. 2013, 53 (9), 1607-1616. (17) Gan, M. Fundamental research on iron ore sintering with biomass energy. Thesis. Changsha, Hunan, 2012, 67-72. (18) Oba, Y.; Konishi, H.; Ono, H.; Kawabata, H.; Takeuchi, E. Combustion behavior of coated cokes with fine Fe3O4 and CaO, Tetsu-to-Hagane 2017, 103 (6), 299-304. (19) Fan, X.; Ji, Z.; Gan, M.; Chen, X.; Yin, L.; Jiang, T. Characteristics of prepared coke–biochar composite and its influence on reduction of no x emission in iron ore sintering, ISIJ Int. 2015, 55 (3), 521-527. (20) Zhao, J.; Loo, C. E.; Zhou, H.; Yuan, J.; Li, X.; Zhu, Y.; Yang, G. Modelling and analysis of the combustion behaviour of granulated fuel particles in iron ore sintering, Combust. Flame 2018, 189 257-274. (21) Kasai, E.; Batcaihan, B.; Omori, Y.; Sakamoto, N.; Kumasaka, A. Permeation characteristics and void structure of iron ore sinter cake, ISIJ Int. 1991, 31 (11), 1286-1291. (22) Kasai, E.; Rankin, W. J.; Lovel, R. R.; Omori, Y. An analysis of the structure of iron ore sinter cake, ISIJ Int. 1989, 29 (8), 635-641. 39 ACS Paragon Plus Environment

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(23) Zhou, H.; Zhao, J. P.; Loo, C. E.; Ellis, B. G.; Cen, K. F. Numerical modeling of the iron ore sintering process, ISIJ Int. 2012, 52 (9), 1550–1558. (24) Komarov, S. V.; Shibata, H.; Hayashi, N.; Kasai, E. Numerical and experimental investigation on heat propagation through composite sinter bed with non-uniform voidage: part I: Mathematical model and its experimental verification, Journal of Iron and Steel Research, International 2010, 17 (10), 1-7. (25) ANSYS Incorporation. ANSYS Fluent 16.0 User's Guide, 2015. (26) Ergun, S. Fluid Flow Through Packed Columns, Chem. Eng. Prog. 1952, 48 (2), 89-94. (27) Gray, D. E., American institute of physics handbook, McGraw-Hill, 1982. (28) Wang, X.; Zeng, X.; Yang, H.; Zhao, D. General modeling and numerical simulation of the burning characteristics of porous chars, Combust. Flame 2012, 159 (7), 2457-2467. (29) Krishna, R. Problems and pitfalls in the use of the fick formulation for intraparticle diffusion, Chem. Eng. Sci. 1993, 48 (5), 845-861. (30) Petersen, E. E., Chemical reaction analysis, Prentice Hall, 1965. (31) Haugen, N. E. L.; Tilghman, M. B.; Mitchell, R. E. The conversion mode of a porous carbon particle during oxidation and gasification, Combust. Flame 2014, 161 (2), 612-619. (32) Litster, J.; Waters, A. Influence of the material properties of iron ore sinter feed on granulation effectiveness, Powder Technol. 1988, 55 (2), 141-151. (33) Litster, J.; Waters, A.; Nicol, S. A model for predicting the size distribution of product from a granulating drum, Trans. ISIJ 1986, 26 (12), 1036-1044. (34) Hobbs, M. L.; Radulovic, P. T.; Smoot, L. D. Combustion and gasification of coals in fixed-beds, Prog. Energy Combust. Sci. 1993, 19 (6), 505-586. (35) Hobbs, M. L.; Radulovic, P. T.; Smoot, L. D. Modeling fixed-bed coal gasifiers, AIChE J. 1992, 38 (5), 681-702. (36) Zhao, J. Numerical modelling of the iron ore sintering process and its experimental validation. PhD Thesis. Hangzhou, 2012. (37) Zhao, J.; Loo, C. E. Dependence of flame front speed on iron ore sintering conditions, Miner. Process. Extr. Metall. (TIMM C) 2016, 125 (3), 165-171.


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Energy & Fuels

(38) Loo, C. E.; Heikkinen, J. Structural Transformation of Beds during Iron Ore Sintering, ISIJ Int. 2012, 52 (12), 2158-2167. (39) Inazumi, T.; Fujimoto, M.; Sato, S.; Sato, K. Effect of Sinter-cake Load Reduction by Magnetic Force on Iron Ore Sintering, ISIJ Int. 1995, 35 (4), 372-379. (40) Zhou, H.; Zhao, J. P.; Loo, C. E.; Ellis, B. G.; Cen, K. F. Model predictions of important bed and gas properties during iron ore sintering, ISIJ Int. 2012, 52 (12), 2168–2176.


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Table captions: Table 1. Reactions considered in the present model Table 2. A sinter mix chemistry with high basicity Table 3. Sintering performance for different fuel properties Table 4. Fuel and green bed properties, and sintering conditions used for simulation Table 5. Normal size distribution of coke and charcoal used in this work


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Energy & Fuels

Table 1. Reactions considered in the present model No.

Rate equation

1

Reaction Coke oxidation

Formula 2C + O2 → 2CO

2

CO oxidation

2CO + O2 → 2CO2

23

3

Limestone calcination

CaCO3 → CaO + CO2

23

4

Dolomite calcination

CaMg(CO3)2 → CaO + MgO + 2CO2

23

5

Serpentine dehydroxylation

2Mg3Si2O5(OH)4 → 3Mg2SiO4 + SiO2 + 4H2O(g)

2

6

Goethite dehydration

2α-FeOOH → α-Fe2O3 + H2O(g)

2

7

Magnetite oxidation

4Fe3O4 + O2 → 6Fe2O3

23

8

Drying/condensation*

H2O(l) ↔ H2O(g)

23

9

Melting+

Solid mixture → Liquid melt

Present work

10

Solidification+

Liquid melt → Precipitated mineral and phases

Present work

20

+ New reactions included in the present model * From a modeling perspective, drying and moisture condensation are considered as reactions


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Table 2. A sinter mix chemistry with high basicity Sinter mix Mass fraction chemistry (wt.%) Fe2O3 81.40 CaO 10.74 SiO2 4.88 Al2O3 1.48 MgO 1.50


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Energy & Fuels

Table 3. Sintering performance for different fuel properties No.

Fuel type

Fuel size

FFS

ηcomb

FFT

RT

EA

Case 1

100% coke

Normal size distribution*

28.6

94.8

1551

3.20

527.0

Case 2

100% untreated charcoal

Normal size distribution

37.4

67.2

1312

0.0

0.0

Case 3

30% pre-granulated charcoal+70% coke


27. 8

88.3

1510

2.66

414.1

Case 4



27.6

85.9

1498

2.53

377.6

Case 5



27.4

83.4

1486

2.42

340.7

Case 6



27.3

78.4

1463

2.16

265.7

Case 7

40% pre-granulated charcoal (without 0.5 mm size)+60% coke 30% untreated charcoal+70% coke

Normal size distribution for coke and charcoal, but 0.5 mm size removed for charcoal

26.0

92.8

1572

2.79

561.5


29.3

83.3

1471

2.47

287.8

Case 8

* Normal size distributions for coke and charcoal are given in Table 5.


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

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Table 4. Fuel and green bed properties, and sintering conditions used for simulation Parameter Coke Charcoal Ash content of fuel (wt. %) 11.6 3.6 Fixed carbon content of fuel (wt. %) 87.4 97.7 Fuel porosity (-)

0.42

3

Fuel apparent density (kg/m ) 1160.0 2

Fuel specific surface area (m /g) Frequency factor of carbon oxidation (1/s) Activation energy of carbon oxidation (kJ/mol)

20 10

10.1

0.55 823.5 40 1011.5

150

140

Sinter mix moisture (wt. % wmb ) Ore A in sinter mix (wt. % dmb**) Ore B in sinter mix (wt. % dmb) Limestone (wt. % dmb) Serpentine (wt. % dmb) Dolomite (wt. % dmb) Solid fuel (wt. % dmb) Return fines (wt. % dmb) Granule Sauter mean diameter (SMD) (mm)

7.5 30.79 30.79 10.62 1.82 2.12 3.94 19.93 1.59

9 30.79 30.79 10.62 1.82 2.12 3.94 19.93 1.78

Green bed voidage (-) Green bed bulk density (kg/m3)

0.35 1847

0.35 1747

*

Note: wmb is wet mix basis, and dmb is dry mix basis.


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Energy & Fuels

Table 5. Normal size distribution of coke and charcoal used in this work Mean fuel size (mm) 9.51 7.10 4.45 1.93 0.84 0.42 0.21 0.126 0.089 0.0315 Mass percent (%) 0.06 0.44 12.19 28.19 13.87 12.22 11.79 4.91 3.8 12.53


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Figure captions: Figure 1. Schematic diagram of the sinter pot rig used for model validation Figure 2. Schematic of the generalized fuel configuration in iron ore sintering and the two stages identified in the combustion of porous solid fuels Figure 3. Variation of melt fraction σ with temperature Ts for a high basicity sinter mix predicted by the thermochemical modelling with FactSage (represented by the circles). The thick solid line indicate the results fitted to the melting and solidification model according to eq (31). Fitted results are θ = 1.34, TM = 1403 K and TE = 1643 K. Figure 4. Comparison of the measured and predicted temperature profiles of a typical sinter pot test Figure 5. Temperature-time profiles predicted at 100, 300 and 500 mm down the bed for sintering with coke and charcoal, respectively. Figure 6. Combustion rates of single coke and non-granulated charcoal particles without adhering layer (i.e., ξal = 0) at the temperatures of 1073, 1373 and 1673 K. Other conditions: initial coke/charcoal particle size dp = 1 mm, physical velocity of gas ug = 3.87 m/s, oxygen molar fraction in bulk gas NO2 = 0.13. Figure 7. Combustion rates of single untreated charcoal (ξal = 0) and pre-granualted charcoal (ξal = 0.4) particles at the temperatures of 1073, 1373 and 1673 K. Other conditions: dp = 1 mm, ug = 3.87 m/s, NO2 = 0.13 for both fuels. εal is 0.315 for pre-granualted charcoal. Figure 8. Combustion rates of single untreated charcoal (ξal = 0) and pre-granualted charcoal (ξal = 0.4) particles at the temperatures of 1073, 1373 and 1673 K. Other conditions: dp = 1 mm, ug = 3.87 m/s, NO2 = 0.13 for both fuels. εal is 0.215 for pre-granualted charcoal. Figure 9. Variation of εal,cr with ξc,cr for the combustion of pre-granulated charcoal particles aiming to obtain comparable combustion rates with coke. dp = 1 mm, ξal = 0.4, Ts =1073, 1373 and 1673 K, respectively. The horizontal and vertical lines locate the points discussed in the text. Figure 10. Variation of the combustion rates of coke and pre-granulated charcoal particles for different ξc,cr, i.e., 1, 0.5 and 0.1. dp = 1 mm, ξal = 0.4, Ts = 1373 K. (a) combustion rate against ξc. (b) combustion rate against X. Figure 11. Conversion version time relationship for coke and pre-granulated charcoal particles of full size distribution at the temperatures of (a) 1073, (b) 1373, and (c) 1673 K. Coke and charcoal particles have the same mass of fixed carbon and same size distribution. Figure 12. Temperature-time profiles at 500 mm down the bed for sintering with coke (Case 1) and 30, 40, 50 and 70 % pre-granulated charcoal (Case 3 – Case 6), and with 30 % pre-granulated charcoal without -0.5 mm size fractions (Case 7), respectively.


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Energy & Fuels

Figures

Figure 1. Schematic diagram of the sinter pot rig used for model validation


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Figure 2. Schematic of the generalized fuel configuration in iron ore sintering and the two stages identified in the combustion of porous solid fuels


Page 50 of 60

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Energy & Fuels

Figure 3. Variation of melt fraction σ with temperature Ts for a high basicity sinter mix predicted by the thermochemical modelling with FactSage (represented by the circles). The thick solid line indicate the results fitted to the melting and solidification model according to eq (31). Fitted results are θ = 1.34, TM = 1403 K and TE = 1643 K.


Energy & Fuels

2000 Experimental Model

1800 1600 Temperature (K)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 52 of 60

300 mm

500 mm

100 mm

1400 1200 1000 800 600 400 200 0

10

20

30

40

Time (min)

Figure 4. Comparison of the measured and predicted temperature profiles of a typical sinter pot test


Page 53 of 60

Coke at 100 mm Coke at 300 mm Coke at 500 mm Charcoal at 100 mm Charcoal at 300 mm Charcoal at 500 mm

2000 1800 1600 Temperature (K)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

1373 K

1400 1200 1000 800 600 400 200 0

10

20

30

40

Time (min)

Figure 5. Temperature-time profiles predicted at 100, 300 and 500 mm down the bed for sintering with coke and charcoal, respectively.


Energy & Fuels

10 Reaction rate x 1e9 (kmol/s)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 54 of 60

Coke at 1073 K Coke at 1373 K Coke at 1673 K Charcoal at 1073 K Charcoal at 1373 K Charcoal at 1673 K

8

6

4

2

0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

Conversion X (-)

Figure 6. Combustion rates of single coke and non-granulated charcoal particles without adhering layer (i.e., ξal = 0) at the temperatures of 1073, 1373 and 1673 K. Other conditions: initial coke/charcoal particle size dp = 1 mm, physical velocity of gas ug = 3.87 m/s, oxygen molar fraction in bulk gas NO2 = 0.13.


Page 55 of 60

10 1073 K, ξ al = 0

Reaction rate x 1e9 (kmol/s)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

1373 K, ξ al = 0

8

1673 K, ξ al = 0 1073 K ξ al = 0.4 1373 K ξ al = 0.4

6

1673 K ξ al = 0.4

4

2

0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

Conversion X (-)

Figure 7. Combustion rates of single untreated charcoal (ξal = 0) and pre-granualted charcoal (ξal = 0.4) particles at the temperatures of 1073, 1373 and 1673 K. Other conditions: dp = 1 mm, ug = 3.87 m/s, NO2 = 0.13 for both fuels. εal is 0.315 for pre-granualted charcoal.


Energy & Fuels

10 1073 K, ξ al = 0

Reaction rate x 1e9 (kmol/s)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 56 of 60

1373 K, ξ al = 0

8

1673 K, ξ al = 0 1073 K, ξ al = 0.4 1373 K, ξ al = 0.4

6

1673 K, ξ al = 0.4

4

2

0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

Conversion ratio X (-)

Figure 8. Combustion rates of single untreated charcoal (ξal = 0) and pre-granualted charcoal (ξal = 0.4) particles at the temperatures of 1073, 1373 and 1673 K. Other conditions: dp = 1 mm, ug = 3.87 m/s, NO2 = 0.13 for both fuels. εal is 0.215 for pre-granualted charcoal.


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Energy & Fuels

Figure 9. Variation of εal,cr with ξc,cr for the combustion of pre-granulated charcoal particles aiming to obtain comparable combustion rates with coke. dp = 1 mm, ξal = 0.4, Ts =1073, 1373 and 1673 K, respectively. The horizontal and vertical lines locate the points discussed in the text.


Energy & Fuels

(a) Combustion rate x 1e9 (kmol/s)

1.4 Charcoal, ξ c,cr = 1.0

1.2

Charcoal, ξ c,cr = 0.5 Charcoal, ξ c,cr = 0.1

1.0

Coke

0.8 0.6 0.4 0.2 0.0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

ξ c (−)

(b) 1.4 Combustion rate x 1e9 (kmol/s)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 58 of 60

Charcoal, ξ c,cr = 1.0

1.2

Charcoal, ξ c,cr = 0.5 Charcoal, ξ c,cr = 0.1

1.0

Coke

0.8 0.6 0.4 0.2 0.0 0.0

0.2

0.4

0.6

0.8

1.0

1.2

X (-)

Figure 10. Variation of the combustion rates of coke and pre-granulated charcoal particles for different ξc,cr, i.e., 1, 0.5 and 0.1. dp = 1 mm, ξal = 0.4, Ts = 1373 K. (a) combustion rate against ξc. (b) combustion rate against X.


Page 59 of 60

(a)

Carbon converstion ratio (-)

100

(a)

80

60

40

20

Coke Pre-granulated charcoal Untreated charcoal

0 1

10

100

1000

10000

T ime (s)

(b) 100 Carbon converstion ratio (-)

(b) 80

60

40

20


0 1

10

100

1000

10000

T ime (s)

(c) 100

(c) Carbon converstion ratio (-)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

80

60

40

20


0 1

10

100

1000

10000

T ime (s)

Figure 11. Conversion version time relationship for coke and pre-granulated charcoal particles of full size distribution at the temperatures of (a) 1073, (b) 1373, and (c) 1673 K. Coke and charcoal particles have the same mass of fixed carbon and same size distribution.


Energy & Fuels

2000 Coke 30 % charcoal 40 % charcoal 50 % charcoal 70 % charcoal 40 % charcoal without fines

1800 1600 Temperature (K)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 60 of 60

1373 1400

1200 1000 800 600 400 200 15

20

25

30

35

Time (min)

Figure 12. Temperature-time profiles at 500 mm down the bed for sintering with coke (Case 1) and 30, 40, 50 and 70 % pre-granulated charcoal (Case 3 – Case 6), and with 30 % pre-granulated charcoal without -0.5 mm size fractions (Case 7), respectively.


A Fundamental Study of the Cocombustion of Coke ... - ACS Publications

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