Numerical Study on Pulverized Biochar Injection in Blast Furnace

Abstract

The possibility of injecting pulverized biochar instead of conventional pulverized coal in blast furnace ironmaking was investigated numerically. More detailed reactions including the water-related reactions were considered here. The combustion process from the tuyere to the raceway of a blast furnace was simulated. Oak char (volatile matter wt.% dry basis, VM = 27.11 wt.%-db) provided a lower temperature than Taiheiyo coal (VM = 44.60 wt.%-db). Increasing the O₂ concentration from 23 to 27 wt.% resulted in a higher combustibility of both solid fuels. However, the effect of increasing oxygen concentration was still insufficient for the Oak char at high injection rates because of its inadequate volatile content. Biochar properties become increasingly important as the injection rate increases. Compared with Oak char that provided a combustibility of 68% at an injection rate of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and hot blast of 27 wt.% O₂ concentration, Oak char 1 (VM = 32.09 wt.%-db) had a higher combustibility of 71%.

1. Introduction

An increase in the requirements for energy savings and reduced environmental impacts has led to investigations into innovative ironmaking technologies with the aim of reducing energy consumption and CO₂ emissions. CO₂ emissions in the ironmaking process account for approximately 70% of total CO₂ emissions in the steel industry.¹⁾ Blast furnaces will remain the predominant ironmaking equipment in the foreseeable future. In a blast furnace, preheated air and fuel (generally pulverized coal) are injected into the lower part of the furnace through a tuyere, forming a raceway in which the injected fuel and some of the coke descending from the top of the furnace are combusted and gasified. An analysis of the heat and mass transfer phenomena inside a blast furnace is important; however, it is almost impossible to measure all of the required information inside the blast furnace accurately.

Operation with pulverized coal injection (PCI) into a blast furnace tuyere is used to reduce the coke feed rate. The operating conditions, coal types, pulverized coal (PC) diameter and PC injection method influence the combustion behavior in the raceway zone.^2,3) The raceway zone is primarily responsible for the production of combustion gases in the blast furnace. When the PC feed rate increases, unburned char accumulated in the furnace causes a decrease in the permeability through the coke bed. A stable state of operation can therefore not be achieved. The combustion process in the blast furnace raceway is clearly complex, and detailed measurements are extremely difficult to obtain because of the existing high temperatures and pressures, presence of molten material, lack of accessibility and inevitable reduction in production. An alternative to direct measurements is the development of a mathematical model of PC injection to clarify the limit of the PCI rate.⁴⁾ Mathematical models for analyzing the effect of uncertain factors on the combustion characteristics of PC in blast furnace have been reported.^5,6,7) The simulation has emphasized also on the effect of the lance configuration⁷⁾ and the results have suggested that a coaxial double-lance configuration with cooling gas channel induces a stronger swirling flow than the lance configuration without cooling gas channel.

Energy system involving carbon-based vectors can achieve very low CO₂ intensity when it uses energy mix of carbon positive and carbon negative technologies. Interestingly, negative CO₂ emission can offset CO₂ emission generated by conventional fossil fuel-fired energy systems.⁸⁾ Energy from biomass is estimated as a promising candidate for carbon-neutral or even carbon negative energy systems with C-based energy vectors. Biomass is a solid fuel with high moisture and volatile content. Biomass also has a lower latent heat and density compared with coal. Biomass can generally be defined as a hydrocarbon fuel which consists mainly of carbon, hydrogen, oxygen and nitrogen. A low moisture content is the basis of high combustion quality; therefore, the biomass material should be stored out of the rain and aerated.⁹⁾ The carbonization of biomass enriches the carbon content and removes oxygen. The resulting biochar has an increased energy density,¹⁰⁾ although its characteristic differs for each biomass source and at each pyrolysis temperature.¹¹⁾ Biochar from biomass materials is a renewable energy carrier which is competitive on energetic basis and also as a viable option for atmospheric carbon sequestration.¹²⁾ However, the increased utilization of biochar is still limited in the iron and steel industry.^13,14,15)

This paper provides a numerical investigation into the combustion of pulverized biochar injected into a blast furnace for ironmaking. The purpose of this study was to investigate the potential of using pulverized biochar injection instead of conventional PCI in blast furnaces. Compared with our previous study,¹⁵⁾ we here use a more detailed reaction mechanism and focus on high injection rates. The temperature and combustible gas distribution profiles through the tuyere and raceway zone are explained. A comparison of calculated results between pulverized biochar injection and PCI is provided. The combustibility under various conditions is discussed. Overall, the findings provide information to be used when considering the implementation of pulverized biochar in the ironmaking blast furnace.

2. Numerical Analysis

The numerical investigation focuses on the combustion behavior from the tuyere (where there is a single lance for fuel and blast injection for hot gas) to the raceway region of a blast furnace. The tuyere is used to implement a hot blast and inject the solid fuel from the lance (Fig. 1(a)). In the blast furnace, the raceway is surrounded by a packed bed of coke. An assumption regarding the raceway is the absence of any solid particles such as coke. Taiheiyo coal and two biochars (namely Oak char and Oak char 1) were used in this numerical simulation. Properties of the Taiheiyo coal,²⁾ Oak char¹⁶⁾ and Oak char 1¹⁴⁾ including the proximate and ultimate analyses for the solid fuels are summarized in Table 1.

Fig. 1.

(a) Illustration of the tuyere and raceway of a blast furnace (refer to [4]) and (b) geometry and computational domain.

Table 1. Solid fuel properties.

		Taiheiyo coal	Oak char	Oak char 1
Proximate [wt.%-db]	FC	39.80	55.60	62.85
	VM	44.60	27.10	32.09
	Ash	15.60	17.30	5.06
Ultimate [wt.%-daf]	C	77.50	78.11	78.56
	H	6.50	2.54	17.49
	O	14.80	18.75	3.54
	N	1.00	0.48	0.41
	S	0.20	0.12	0
HHV [MJ/kg]		26.40	23.05	28.16

db: dry basis; daf: dry ash free; HHV: higher heating value

2.1. Mathematical Model

Gas-particle flow plays a dominant role in multiphase flow in an ironmaking blast furnace. A comprehensive model based on continuum-(Eulerian) and discrete-(Lagrangian) types describing the hydrodynamics of gas-particle flow as a discrete particle model has been developed. In this model, the gas phase is treated with a Eulerian frame and described by the steady-state Reynolds-averaged Navier-Stokes equations and the k-ε turbulence model.¹⁷⁾ In discrete phase modeling, particles of known size distributions and properties are injected into the combustion chamber and tracked using a Lagrangian approach throughout the computational domain. Individual particle trajectories are tracked and solved for using Newton’s second law of motion. The conservation equations of this model have been described in detail in our previous study¹⁵⁾ and are summarized in Table 2. Here, this model is applied with the modification of some additional reactions as described below.

Table 2. Conservation equations.¹⁵⁾

Gas phase:
Mass	$\nabla .\left(\rho \overrightarrow{u}\right)=S_m$	(1)
Momentum	$\nabla .\left(\rho \overrightarrow{u}\overrightarrow{u}\right)=-\nabla p+\nabla .\left(\overline{\tau }\right)+\rho \overrightarrow{g}+\overrightarrow{F}$	(2)
Energy	$\nabla .\left(\overrightarrow{u}\left(\rho H+p\right)\right)=-\nabla .\left(\sum _j{h}_j{J}_j\right)+S_h$	(3)
Gas species i	$\nabla .\left(\rho \overrightarrow{u}{Y}_i\right)=-\nabla .\overrightarrow{J_i}+R_i+S_i$	(4)
Turbulent kinetic energy	$\frac{\partial }{\partial x_i}\left(\rho k{u}_i\right)=\frac{\partial }{\partial x_j}\left[\left(\mu +\frac{{\mu }_t}{{\sigma }_k}\right)\frac{\partial k}{\partial x_j}\right]+G_k-\rho \epsilon -Y_M+S_k$	(5)
Turbulent dissipation rate	$\frac{\partial }{\partial x_i}\left(\rho \epsilon u_i\right)=\frac{\partial }{\partial x_j}\left[\left(\mu +\frac{{\mu }_t}{{\sigma }_{\epsilon }}\right)\frac{\partial \epsilon }{\partial x_j}\right]+C_{1\epsilon }\frac{\epsilon }{k}{G}_k-C_{2\epsilon }\rho \frac{{\epsilon }^2}{k}+S_{\epsilon }$	(6)
Particle phase:
Mass	$\frac{d{m}_{\text{p}}}{dt}=\dot{m}$	(7)
Momentum	$\frac{d{u}_p}{dt}=F_D(u-u_p)+\frac{g({\rho }_p-\rho )}{{\rho }_p}+F$	(8)
Energy	$m_{\text{p}}{c}_{\text{p}}\frac{d{T}_{\text{p}}}{dt}=h_{i,conv}{A}_{\text{p}}\left(T_{\text{g},i}-T_{\text{p}}\right)+\frac{d{m}_{\text{p}}}{dt}{H}_{reac}+A_{\text{p}}{\epsilon }_{\text{p}}\sigma \left(T_{Rad}^4-T_{\text{p}}^{\text{4}}\right)$	(9)

Note: S is source [units vary]; k and ε indicate turbulent kinetic energy [m²/s²] and turbulent dissipation rate [m²/s³]. G_k represents the generation of turbulence kinetic energy related to the mean velocity gradient and C_1ε, C_2ε, σ_k, σ_ε are the turbulent model constants.

The combustion process is composed of the following stages: inert heating, the devolatilization of solid fuel particles, and the gaseous combustion of volatiles, followed by the oxidation and gasification of char. Inert heating occurs until the particle temperature reaches the vaporization temperature. When the particle temperature reaches the vaporization temperature, devolatilization (R1) commences for a combusting particle. Significant devolatilization is initiated at approximately 600 K.¹⁸⁾ The devolatilization process releases volatiles (C_αH_βO_γN_δ) and char (C(s)).   

$$\text{Solid Fuel}\to {\mathrm{C}}_{\alpha }{\mathrm{H}}_{\beta }{\mathrm{O}}_{\gamma }{\mathrm{N}}_{\delta }+\mathrm{C}\left(\mathrm{s}\right)$$R1

Devolatilization indicates that volatiles are released according to the following expression:   

$$-\frac{d{m}_{\text{p}}}{dt}=A_0{f}_{\upsilon ,0}\left(1-f_{w,0}\right)m_{\text{p},0}$$(10)

where m_p, m_p,0, and A₀ represent the particle mass [kg], initial particle mass [kg] and rate constant [s^–1], respectively. f_v,0 and f_w,0 indicate the fraction of volatiles initially present in the particle and evaporating material, respectively.

The combustion of volatiles is represented by the gas reactions as follows:   

$${\text{C}}_{\alpha }{\text{H}}_{\beta }{\text{O}}_{\gamma }{\text{N}}_{\delta }+a{\text{O}}_{\text{2}}\to \mathit{\text{b}}\text{CO}+\mathit{\text{c}}{\text{H}}_{\text{2}}\text{O}+\mathit{\text{d}}{\text{N}}_{\text{2}}$$R2

  

$$\text{CO}+0.5{\text{O}}_{\text{2}}\to {\text{CO}}_{\text{2}}$$R3

  

$$\text{CO}+{\text{H}}_{\text{2}}\text{O}\leftrightarrow {\text{CO}}_{\text{2}}+{\text{H}}_{\text{2}}$$R4

  

$$\text{CO}+3{\text{H}}_{\text{2}}\leftrightarrow {\text{CH}}_{\text{4}}+{\text{H}}_{\text{2}}\text{O}$$R5

  

$${\text{CH}}_{\text{4}}+0.5{\text{O}}_{\text{2}}\to \text{CO}+2{\text{H}}_{\text{2}}$$R6

  

$${\text{H}}_{\text{2}}+0.5{\text{O}}_{\text{2}}\leftrightarrow {\text{H}}_{\text{2}}\text{O}$$R7

The stoichiometric coefficients of reaction R2 are summarized in Table 3. The formula of volatiles (C_αH_βO_γN_δ) is defined based on the solid fuel properties in the Table 1. Compared with our previous study¹⁵⁾ that only used the two overall reactions R2 and R3, more detailed reactions including the water-related reactions (R4, R5, R6, R7) are considered here.

Table 3. Stoichiometric coefficients of volatile combustion (R2).

Reaction coefficient	Taiheiyo coal	Oak char	Oak char 1
α	1.44	0.83	0.91
β	3.60	2.30	3.11
γ	0.52	1.07	0.97
δ	0.0399	0.0316	0.0259
a	1.36	0.45	0.75
b	1.44	0.83	0.91
c	1.80	1.15	1.55
d	0.0199	0.0158	0.0129

The finite-rate/eddy-dissipation model for the gas reaction mechanisms in turbulent flow was employed where the reaction rates of the Arrhenius model, R_g,Arr, and the Eddy Break-Up turbulence chemistry interaction model, R_g,EBU,R for reactants and R_g,EBU,P for products,¹⁹⁾ were calculated. The net reaction rate is taken as the minimum of these three rates. R_g,Arr, R_g,EBU,R and R_g,EBU,P are expressed as follows:   

$$R_{\text{g,Arr}}=\left({\nu }_P-{\nu }_R\right)M_w\left(k_{r,f}\prod _R{C}_R^{{\alpha }_R}-k_{r,b}\prod _P{C}_P^{{\alpha }_P}\right)$$(11)

  

$$R_{\text{g,EBU,}R}={\nu }_R{M}_{w}A\rho \frac{\epsilon }{k}\underset{R}{min}\left(\frac{Y_R}{{\nu }_R{M}_{w,R}}\right)$$(12)

  

$$R_{\text{g,EBU,}P}={\nu }_P{M}_{w}AB\rho \frac{\epsilon }{k}\frac{\sum _P{Y}_P}{\sum _P{\nu }_P{M}_{w,P}}$$(13)

ν, M_w, Y, C, are stoichiometric coefficient, molecular weight, mass fraction and molar concentration [kmol/m³] of the corresponding species, respectively. α is rate exponent for reactants and products. Subscript R and P indicate reactant and product of reactions. Subscript f and b are forward and backward reactions. ρ represents the density [kg/m³]. A and B are the empirical parameters. The Arrhenius kinetic rate of reaction is defined as:   

$$k_r=A_r{T}^{{\beta }_r}{e}^{\stackrel{-E_r}{}/\underset{RT}{}}$$(14)

where subscript f and b represent forward and backward reaction. R is the universal gas constant. A_r, β_r and E_r are the pre-exponential factor, temperature exponent and activation energy for reaction, respectively.

For heterogeneous surface reactions, the following reactions are considered during the combustion process.   

$$\text{C(s)}+0.5{\text{O}}_{\text{2}}\to \text{CO}$$R8

  

$$\text{C(s)}+{\text{CO}}_{\text{2}}\to \text{2CO}$$R9

  

$$\text{C(s)}+{\text{H}}_{\text{2}}\text{O}\to \text{CO}+{\text{H}}_{\text{2}}$$R10

The chemical reaction rate is expressed as follows:^20,21)   

$$R_{j,r}=A_{\text{p}}{\eta }_r{Y}_{j}p\frac{k_r{D}_{0,r}}{D_{0,r}+k_r}$$(15)

A_p and η_r indicate the particle surface area [m²] and effectiveness factor, respectively. Y_j is the mass fraction of species j. p represents the partial pressure [Pa]. The diffusion rate coefficient, D_0,r, for reaction r is computed as:   

$$D_{0,r}=C_{1,r}\frac{{\left[\left(T_{\text{p}}+T_{\text{g}}\right)/2\right]}^{0.75}}{d_{\text{p}}}$$(16)

where C_j,r is the molar concentration of species j in reaction r. d_p represents particle diameter [m]. The kinetic parameters to determine k_r using the Arrhenius expression are summarized in Table 4.²²⁾

Table 4. Kinetic reaction parameters.

Reaction	Ar	βr	Er [J/kmol]
R2	2.1 × 1011	0	2.03 × 108
R3	2.2 × 1012	0	1.67 × 108
R4f	2.75 × 1011	0	8.38 × 108
R4b	2.65 × 10–2	0	3.96 × 103
R5f	5.12 × 1014	0	2.73 × 104
R5b	4.4 × 1011	0	1.26 × 108
R6	3 × 108	–1	1.68 × 108
R7	6.8 × 1015	0	1.26 × 108
R8	1.36 × 106	0.68	1.30 × 108
R9	6.78 × 104	0.73	1.63 × 108
R10	8.55 × 104	0.84	1.40 × 108

f: forward reaction; b: backward reaction

The change in particle temperature is determined using an energy balance for particles (Eq. (9)) governed by convective heat transfer and latent heat transfer associated with mass transfer and radiative heat transfer.^23,24) c_p, H_reac and T in the Eq. (9) represent heat capacity [J/kg-K], reaction heat [J/kg] and temperature [K], respectively. Subscripts p, g and Rad, respectively, indicate particle, gas phase and radiation term. ε_p and σ are the particle emissivity and Stefan-Boltzmann constant, respectively. P-1 radiation model is used and heat source due to particle radiation expressed as follows:   

$$-\nabla .q_r=-4\pi \left(a{n}^2\frac{\sigma T^4}{\pi }\right)+\left(a+a_p\right)G$$(17)

where E_p is the equivalent emission of the particles, a indicates the equivalent absorption coefficient, and n represents the refractive index of the medium. T_Rad is calculated using the following expression:   

$$T_{Rad}={\left(\frac{G}{4\sigma }\right)}^{\stackrel{1}{}/\underset{4}{}}$$(18)

G is the incident radiation [W/m²] expressed as follows:   

$$G={\int }_{\mathrm{\Omega }=4\pi }{I}_{}d\mathrm{\Omega }$$(19)

where I is the radiation intensity and Ω is the solid angle.

h_i,conv (Eq. (9)) is associated with the Nusselt number, which is a function of particle Reynolds number and gas Prandtl number^25,26) as follows:   

$$N{u}_i=\frac{h_{i,conv}{d}_{\text{p}i}}{k_{\alpha }}=2.0+0.6{Re}_d^{\stackrel{\text{1}}{}/\underset{\text{2}}{}}{Pr}^{\stackrel{1}{}/\underset{3}{}}$$(20)

Pr is the Prandtl number of the continuous phase (= c_pμ/kα) and Re_d is the relative Reynolds number based on the particle diameter and relative velocity as follows:   

$${Re}_d=\frac{{\rho }_{\text{g}}{d}_{\text{p}}\left|u_{\text{p}}-u_{\text{g}}\right|}{\mu }$$(21)

where k_α is the thermal conductivity [W/m-K]. μ and u indicate the molecular viscosity [Pa-s] and velocity [m/s], respectively.

2.2. Solution Procedure and Calculation Conditions

An iterative solution procedure is used in the overall gas-particle coupling for solving the governing equations described above.²⁷⁾ As the particle trajectory is computed, the two-way coupling incorporates the effect of the discrete phase trajectories on the continuum and the continuous phase always impacts the discrete phase. The iterative cycle is repeated until overall convergence is reached for both phases. The discretization of the gas phase governing equations is based on the finite volume method employing a staggered grid and solved by the SIMPLE algorithm²⁸⁾ for pressure-velocity coupling.

2.3. Calculated Conditions

A 2D simulation was used for the tuyere (consisting of lance and hot blast) and raceway regions of the blast furnace. Figure 1(b) shows the geometry for the computational domain. A fine mesh was developed with 76646 cells. The numerical investigation focused on simulating the tuyere (where there are single lances for injecting solid fuel and a blast injection for hot gas) and the raceway region of the blast furnace. The lance had an inner diameter of 14 mm. The tuyere had an inner diameter of 180 mm at the inlet side and 140 mm at its end (Fig. 1(b)). The wall along the tuyere is maintained at adiabatic conditions of zero heat flux. The raceway is modeled as a tube 0.5 m in diameter and 1.5 m in length. In the blast furnace, the raceway is surrounded by a packed bed of coke and is assumed to be at 2073 K.²⁾ A particle diameter of 70 μm has been used generally. The computational conditions in Table 5 were selected for investigating the effect of injection rate and O₂ concentration on combustibility.

Table 5. Computational conditions.

Hot Blast:	Inlet velocity	= 188 [m/s]	O2 composition	= 23, 27 [wt.%]
	Temperature	= 1450 [K]
Solid Fuel:	Feed rate	= 25 – 200 [(kg solid fuel)/(1000 Nm3 feed gas)] *
	Temperature	= 300 [K]
	Particle diameter	= 70 [μm]
Carrier:	Inlet velocity	= 1.13 [m/s]	O2 composition	= 23 [wt.%]
	Temperature	= 300 [K]

*  Feed gas = Hot blast + Carrier

3. Results and Discussion

A comparison between the results from the model described above and experimental data²⁹⁾ using a drop tube reactor is shown in Fig. 2. The electrically-heated drop tube furnace had a length of 2.5 m and an internal diameter of 200 mm. Bituminous coal was injected at a rate of 1 kg/h with carrying air of 1.5 Nm³/h at 473 K and total gasifying air of 8 Nm³/h at 523–623 K. The wall of the reactor made from ceramic was maintained at 1523 K. A level of confidence in the predicted O₂ and CO₂ was established through comparison with experimental results. It is expected that the simulation method is capable of providing realistic predictions.

Fig. 2.

Comparison between experimental²⁹⁾ and calculated results.

The Oak char provided a lower temperature distribution than the Taiheiyo coal at an injection rate of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and hot blast of 23 wt.% O₂ (Fig. 3(a)). The field temperature of the Oak char was lower than that of the Taiheiyo coal. This result is expected because the combustion heat of Oak char is lower than that of Taiheiyo coal. The temperature distribution influences the gas composition (Fig. 3(b)). As shown in Fig. 3(b), the injection of Oak char with low volatile content causes a lower overall burnout. The volatile content results in the first ignition in the tuyere. Volatile contents are released from solid fuels during devolatilization. Solid fuels with high volatile contents are normally injected because of their generally superior combustion performance resulting from better C(s) reactivity and hence burnout. Because of volatile oxidation (R2), the reduced release of volatile content by the Oak char compared with that of the Taiheiyo coal results in the Oak char having a higher oxygen composition than the Taiheiyo coal (Fig. 3(b)).

Fig. 3.

(a) Temperature profile and (b) gas compositions at 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and 23 wt.% O₂. (Online version in color.)

The released volatiles burn rapidly with oxygen in the hot blast thereby increasing the gas and particle temperatures. Both the average gas and particle temperatures of Oak char were lower than the Taiheiyo coal at an injection rate of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and a hot blast of 23 wt.% O₂ (Fig. 4(a)). A decrease in average particle temperature occurred at an axial distance approximately 0.2 m into the tuyere because of the low carrier gas temperature. The temperature history of the particles influences the local concentrations of gas species (Fig. 4(b)). Compared with the Taiheiyo coal, the Oak char cannot yield a higher volatile combustion (R2) because it contains an inadequate volatile content. As a result, the Oak char yielded a higher oxygen concentration than the Taiheiyo coal. Compositions of CO and CO₂ are influenced primarily by reaction R3. The related water reactions, mainly R7, favor water production.

Fig. 4.

Comparison of average (a) temperature and (b) gas compositions for Taiheiyo coal and Oak char at 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and 23 wt.% O₂ in the axial direction.

Figures 5(a) and 5(b) show the effect of O₂ concentration on average temperature and gas compositions along the axis at an injection rate of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] for the Taiheiyo coal and Oak char, respectively. For both samples, using a hot blast with 27 wt.% O₂ increases the temperature and influences gas concentrations. This occurs because O₂ enrichment improves the combustion efficiency. This additional O₂ portion from the hot blast affects the temperature history of the particle and can influence the local concentrations of gaseous species. An increase in O₂ concentration increases the temperature and CO₂ content because of the exothermic reactions C(s) + 0.5O₂ → CO (R8) and CO + 0.5O₂ → CO₂ (R3). Furthermore, oxygen in the raceway is consumed rapidly as the pulverized solid fuel is injected. The maximum CO₂ concentration corresponds to that of highest temperature. Due to the exothermic reaction of C(s) + 0.5O₂ → CO (R8), the higher temperature at the end part of raceway occurs (see also Fig. 3(a)).

Fig. 5.

Effect of O₂ concentration at 200 [(kg solid fuel)/(1000 Nm³ feed gas)] on average temperature and gas compositions for (a) Taiheiyo coal and (b) Oak char. (Online version in color.)

As shown in Figs. 5(a) and 5(b), the effect of oxygen enrichment for the Oak char was smaller than that for the Taiheiyo coal. An increase in oxygen concentration in the hot blast promotes volatile combustion (R2). This causes an increased concentration of combustible gas CO that results in an increased temperature, as reported in our previous study.^30,31) However, the effect of oxygen concentration decreases at an axial distance of 1.8 m for the Taiheiyo coal (Fig. 5(a)) at the 27 wt.% O₂ concentration. In this position, the average temperature decreased and became lower than that at 23 wt.% O₂ concentration. The decrease in average temperature and increase in oxygen concentration result from the completion of volatile oxidation (R2). A decrease in CO₂ concentration occurs at this position because there is insufficient CO for CO oxidation to proceed (R3). As for the Taiheiyo coal, the average temperature at 27 wt.% O₂ concentration decreased and became lower than that at 23 wt.% O₂ concentration at almost 2 m for the Oak char (Fig. 5(b)). For both the Taiheiyo coal and the Oak char, the effect of O₂ enrichment is limited at high injection rates of 200 [(kg solid fuel)/(1000 Nm³ feed gas)]. A limitation of O₂ enrichment is more apparent for the Oak char compared with the Taiheiyo coal. The Oak char cannot sustain further volatile combustion at high O₂ concentration because it contains insufficient volatile matter. The volatile content therefore becomes increasingly important at high injection rates.

Figures 6(a) and 6(b) show the effect of O₂ concentration on the reaction rate of volatile combustion R2 at an injection rate of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] for the Taiheiyo coal and Oak char, respectively. These figures provide added information regarding the limitation of oxygen enrichment and illustrate why the average temperature at 23 wt.% O₂ was higher than that at 27 wt.% O₂ at the end of the raceway zone. An increase in temperature occurs late at 23 wt.% O₂ because the exothermic C(s) oxidations are delayed by the lag in volatile oxidation. It also causes the particle temperature of Oak char lower than that of Taiheiyo coal (see Fig. 4(a)). Compared with the Taiheiyo coal in Fig. 6(a), the Oak char (Fig. 6(b)) has a low volatile oxidation rate and a delay in volatile oxidation for both the 23 and 27 wt.% O₂ concentrations.

Fig. 6.

Effect of O₂ concentration at 200 [(kg solid fuel)/(1000 Nm³ feed gas)] on reaction rate of volatile combustion R2 for (a) Taiheiyo coal and (b) Oak char. (Online version in color.)

Figure 7 shows the variation in combustibility with injection rates under various conditions. The combustibility was determined using the C(s) mass fraction loss (m_out) at the tuyere exit and the C(s) mass fraction from the original solid fuel (m_in) at the entrance of the lance:   

$$\eta =\frac{m_{\text{in}}-m_{\text{out}}}{m_{\text{in}}}\times 100\%$$(22)

Figure 7(a) presents a variation in combustibility with hot blast at 23 and 27 wt.% O₂ for the Oak char and Taiheiyo coal. Oak char with a lower volatile content tended to provide a lower combustibility. At 27 wt.% O₂, the combustibility for both Oak char and Taiheiyo increased because the consumption oxygen increases for reaction with unburnt C(s). While Oak char appears to have little significant impact on blast furnace operation at a low injection rate, biochar properties become increasingly important at high injection rates. With higher volatile content, a higher oxygen concentration in the hot blast provides an increased combustibility at a high injection rate.

Fig. 7.

Combustibility with injection rates under various conditions: (a) using the model, (b) comparison between models for Oak char with and without water reactions, (c) comparison of Oak char for models with and without radiation and (d) comparison between Oak char and Oak char 1 using the model. (Online version in color.)

Figure 7(b) compares the combustibility between the model with and without the water reactions for Oak char at 200 [(kg solid fuel)/(1000 Nm³ feed gas)]. The reactions R4, R5, R6, R7 were neglected in the no-water reaction model. Compared with the model without water reactions, the combustibility of the model including water reactions offered a better result for both the 23 and 27 wt.% O₂. The model including the additional water related reactions provided a higher gas temperature distribution in the raceway than that without because of the significant effect of the exothermic reactions from the oxidation reactions of R6 and R7.

Figure 7(c) explains the combustibility between the no-radiation and radiation models for Oak char at 200 [(kg solid fuel)/(1000 Nm³ feed gas)]. The third term on the right hand side of Eq. (9) was not considered in the no-radiation model. Consequently, the no-radiation model did not include the radiation heat flux (Eq. (17)) in the heat sources of the energy equations. The model including radiative heat transfer showed a significant effect on combustibility for both the hot blasts of 23 and 27 wt.% O₂. This is because the energy equation accounts for radiative heat transfer and increases the particle temperature history thereby increasing the gas temperature distribution. The radiative heat transfer therefore becomes important to account for the prediction.

Figure 7(d) compares the combustibility between Oak char (VM = 27.11 wt.%-db) presented on the red colored line and Oak char 1 (VM = 32.09 wt.%-db) on the black line. Oak char 1 with higher volatile content and calorific heating value (see Table 1) provided a higher combustibility. At high injection rates of 200 [(kg solid fuel)/(1000 Nm³ feed gas)] and a hot blast of 27 wt.% O₂ concentration, Oak char 1 offered a higher combustibility of 71% compared with Oak char that provided the combustibility of 68%. Biochar properties become increasingly important at high injection rates (also see Fig. 7(a)).

4. Conclusions

The possibility of pulverized biochar injection in blast furnace ironmaking was numerically investigated. In the present model, heat and mass transfer with more detailed reactions including the water-related reactions for the volatile matter as well as the devolatilization of the solid fuel particle were taken into account. Pulverized Oak char was especially studied as a candidate of the solid fuel to the blast furnace. Its combustion behavior in the tuyere and raceway of the blast furnace was studied and compared with conventional pulverized Taiheiyo coal injection. The major results and findings were listed below:

(1) The present model reasonably agreed with the experimental data from the literature.

(2) Temperatures of gas and particle in the tuyere and raceway with Oak char were lower than those with Taiheiyo coal by several hundred degrees. Lower concentration of CO₂ and higher concentrations of O₂ and CO were obtained for the case of Oak char compared with the case of Taiheiyo coal. These results were attributed to the lower content of volatile matter of Oak char than that of Taiheiyo coal. The content of volatile matter played an important role for the gas temperatures and the gas concentrations in the tuyere and raceway.

(3) The effects of increases in O₂ concentration on the temperatures and the gas concentrations were limited for the case of Oak char compared with Taiheiyo coal. Oak char could not sustain further volatile combustion at high O₂ concentration because of its insufficient content of volatile matter. The content of volatile matter therefore became significant at high injection rates.

(4) According to the calculation results using three kinds of solid fuel, the combustibility of the solid fuel increased with its content of volatile matter.

(5) The radiation heat transfer strongly contributed to the temperature distributions in the tuyere and raceway and the combustibility of the solid fuel. Therefore, not only the properties of the radiation heat transfer such as the coefficient of adsorption but also concentrations of CO₂ and H₂O were important for the prediction.

(6) The reactions concerning water significantly influenced the temperature distribution in the tuyere and raceway and the combustibility of the solid fuel because of the exothermic oxidation reactions as well as the radiation heat transfer.

These results and findings can contribute to an understanding of the pulverized biochar injection with the aim of achieving low emission blast furnace ironmaking.

Acknowledgments

This research work was partially supported by the Japan Society for the Promotion of Science (JSPS) Scientific Research (A), 2010–2011, Research Number 22241020. The authors also gratefully acknowledge a grant from the Global Center of Excellence in Novel Carbon Resource Sciences, Kyushu University.

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