Flame front progress in gas assisted iron ore sintering

Journal Pre-proofs Flame front progress in gas assisted iron ore sintering Nicos Tsioutsios, Christian Weiß, Johannes Rieger, Elmar Schuster, Bernhard Geier PII: DOI: Reference:

S1359-4311(19)35078-1 https://doi.org/10.1016/j.applthermaleng.2019.114554 ATE 114554

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Applied Thermal Engineering

Received Date: Revised Date: Accepted Date:

23 July 2019 11 September 2019 16 October 2019

Please cite this article as: N. Tsioutsios, C. Weiß, J. Rieger, E. Schuster, B. Geier, Flame front progress in gas assisted iron ore sintering, Applied Thermal Engineering (2019), doi: https://doi.org/10.1016/j.applthermaleng. 2019.114554

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Flame front progress in gas assisted iron ore sintering Nicos Tsioutsios1, Christian Weiß1*, Johannes Rieger2, Elmar Schuster3, Bernhard Geier3 1Chair

of Process Technology and Industrial Environmental Protection,

Montanuniversität Leoben, Austria 2K1-MET

GmbH, Linz, Austria

3 voestalpine

Stahl Donawitz GmbH, Leoben, Austria

*) Corresponding author: Tel.: +43 3842 402-5009; Email: [email protected] Montanuniversität Leoben, Department of Environmental and Energy Process Engineering, Chair of Process Technology and Industrial Environmental Protection Franz-Josef-Straße 18, Leoben 8700, Austria

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Abstract The present paper is focused on the control of heat release and on the propagation of the flame front in an iron ore sinter bed. Small scale sinter experiments are performed in cylindrical packings of the sinter mix which are ignited at the bed surface and vertically passed by the suction gas. The heat front propagation is found to be closely linked to the consumption of the fuel mix in the sinter bed. Based on this findings, a dual fuel approach is explored whereas front propagation of the solid fuel combustion (coke breeze) is supported by pulsed injection of a secondary gaseous fuel. A non-stationary, one dimensional model is established to analyse the heat wave travelling in the sinter bed. The model setup describes a local thermal non-equilibrium between the sinter bed being defined as porous medium and the permeating gas. The model is based on the balance equations for the mass and heat transports caused by the solid and gaseous fuel combustion. Model calculations support the experimental findings and demonstrate that satisfying sinter peak temperature levels and locally focused energy release are ensured by combined solid fuel burnout and high temperature gas combustion. During some labscale sintering tests, coke breeze content was reduced in the sinter feed mixture, and secondary fuel gas was supplied in pulsed mode to compensate the missing energy demand. The flame front speed as well as properties related to the sinter quality such as the tumble and shatter strength were determined, with the later index showing more sensitive results. The investigations demonstrate that pulsed methane injection can increase the flame front speed compared to the pure coke case with 5.1 wt.% coke. Furthermore, the pulsed gaseous fuel combustion can nearly compensate the strength loss for a moderately coke reduced rawmix with a coke content of 3.5 wt.% . Keywords: Iron ore sintering, gaseous fuel, flame front speed , sinter strength

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Highlights 

Control of flame front propagation in small-scale iron ore sintering studies



Partial substitution of the standard fuel coke breeze by a pulsed gas injection



Numerical modelling of the heat front travelling through the sinter bed



Lab-scale sintering tests with reduced coke breeze content in the sinter mixture

Graphical abstract

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Introduction According to current Worldsteel reports (www.worldsteel.org) the global crude steel production of 1.689 Mt in the year 2017 doubled compared to the one in year 2000 and steel represents the most important constructing material. In Europe, primary steelmaking is almost exclusively carried out in integrated steel mills [1]. Until now, the major share of the iron containing feed material for the pig iron production via the blast furnace route is processed in sinter plants. On the sinter strand, the fine ore particles are agglomerated into porous clinker, called sinter. The required heat for this step is generated by the combustion of fuels such as coke breeze (fine fraction of the metallurgical coke) being part of the feed mixture. Research activities focus on coke substitution strategies by alternative fuels for the standard sintering process due to feed material dependence considering coking coal as critical raw material [2] as well as due to environmental regulations. Beside nitrogen oxides NOX and sulfur oxides SOX, the reduction of fossil carbon dioxide CO2 in steel industry off gases is increasingly forced by European Union regulations. Considering opportunities for emission reduction during the sintering process offered by alternative fuels, Lu et al. [3] revealed that charcoal compared to coke breeze reduces off gas concentration of NOX and SO2 while increasing CO emissions. Charcoal and biomass based fuels are therefore considered as potential coke substitutes providing that efficiency and completeness of combustion is ensured. In an earlier contribution, Lovel et al. [4] performed sinter pot tests and found charcoal addition to fasten sintering progress by acceleration of flame front travelling. This fact was attributed to the higher reactivity of charcoal and the larger specific area compared to the coke breeze particles. These findings are in accordance with Cheng et al. [5,6], although, the later investigators argue that a too high charcoal addition might decreases

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sinter strength. Excessively high combustion rates of bio-based fuels generally promote their conversion already at moderate temperature level and turn out to contribute less to melt phase formation at high temperatures. Therefore, sinter quality weakens. Fan et al. [7] used straw based fuels as partial coke breeze substitute observing that sinter quality deteriorates with a too high content of these biomass fuels. Exemplary for an alternative process internal carbon source Lanzerstorfer et al. [8] investigated the utilization of blast furnace dust. The carbon replacement factor of blast furnace dust, which is close to 1.0 indicated the suitability for a recycling of steel mill internal carbon containing residuals as fuel source in the sinter plant. These findings additionally support Cheng et al. [6] viewing an equivalent carbon substitution approach being superior to the one purely based on equivalence of calorific heat. On the other hand considering gaseous fuel alternatives, the injection of natural gas investigated by Iwami et al. [9], it was shown in principle, that gas sintering can increase productivity and sinter quality. The positive contribution of the gaseous fuel was attributed to the formation of a wider and more homogeneous high temperature zone. Collectively, these studies highlight the need to define sintering conditions which maintain or even stimulate high temperature zone formation during coke substitution by alternative fuels. It is generally agreed in the iron ore sinter science community that agglomerate strength and grain bonding in the final sinter demands residence of the sintering mixture above a high temperature level of approximately 1100 °C to ensure intermediate melt phase formation [10]. The area of the material’s overheat above that temperature level versus time (cf. insert in Fig. 1) experienced during the ongoing sinter process is a general accepted measure to quantify the extend of sintering [11,12]. As already pointed out by Loo [10] all factors affecting the extend

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and shape of the temperature transient need to be carefully analyzed to maintain a high quality sinter process.

Fig. 1 Moving temperature transient in a sinter column (a), kinematic relation Lx = utp linking the 1D- representation of the heat wave to the temperature profile in the moving sinter bed (b).

Sintering starts by an ignition of the solid fuel in the top layer of the bed, usually with the help of a gas burner. Transient temperature profiles move through the bed due to the top-down directed air suction in form of a heat wave. The profiles result from the ignition zone’s heat flux into the deeper material layers interacting with the additional heat released from the fuel combustion and the heat demand of calcinations processes in the sinter mix. Starting with the early fundamental studies of iron-ore sintering (as summarized in [13] heat wave propagation is interpreted as two distinct moving fronts, called heat front and flame front, driven down the bed due to air suction. Tracking the moving fronts in mathematically elaborated form was established under the subject headings smolder combustion [14,15] or filtration combustion [16]. The former was defined as an oxygen-limited, flameless form of combustion with low propagation rates compared to flaming [14]. As it is the case in iron ore sintering the porous mineral

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matrix serves as carrier structure for the dispersed coke particles and additionally provides access of oxygen from the permeating gas stream to the fuel through its pore system. Furthermore, the high heat capacity enables the matrix material to act as a heat regenerator [17] feeding the heat to reaction sites further downstream from the current position of the combustion zone. For heat and flame front moving at comparable speed the temperature profile of the heat wave stays narrow with a high temperature maximum. This front matching is often termed “superadiabatic” [15,16] in the filtration combustion literature, referring to a reaction peak temperature exceeding the adiabatic combustion temperature which would be found for a given rate of conversion of the same mixture composition reacting uniformly under adiabatic conditions. Contrary, front mismatching is undesirable since it does not allow to maximize the peak temperature. Mismatching may occur in a reaction leading scenario, where increased oxygen supply in case of a high reactive fuel (e.g. use of charcoal instead of coke breeze) drives the flame front ahead of the heat front. Mismatching by reaction trailing however, may occur at low oxygen supply forcing the heat front to hurry ahead of the flame front. Therefore, front matching can be favoured with a medium oxygen supply adjusted to the fuel reactivity [10,13]. However, combustion of a solid fuel – independent of its reactivity – under practical conditions may stay incomplete as a fraction of the carbon is not fully converted into carbon dioxide. Some extend of carbon monoxide and excess oxygen remain in the off-gas reducing the theoretical amount of energy release in the high temperature region. In our study the above mentioned facts motivate to introduce a secondary gaseous fuel partially substituting the primary fuel to alter the conditions in the combustion zone. A research activity at JFE Steel demonstrated benefits offered by gaseous fuel injection in terms of sinter porosity optimisation [18]. In a practical

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examination of their approach in our laboratory it turns out that measures have to be taken to avoid the tendency of the combustion zones of primary fuel and gas to separate from each other due individual speeds of the associated reaction fronts. The current paper aims to further develop the dual fuel approach. It is explored how front propagation of the solid fuel combustion can be supported by pulsed injection of a secondary gaseous fuel. Satisfying sinter peak temperature levels and locally focused energy release are ensured by combined mechanisms of solid fuel preheating and high temperature gas combustion. Section 2 introduces a non-stationary, onedimensional model to support the analysis of the travelling heat wave in the sinter bed. The modelling was motivated by a set of experiments in a laboratory sinter test rig to demonstrate the dual fuel approach. Section 3 presents the experimental setup and findings. Section 4 provides a discussion and comparison with the theoretical results. A preliminary evaluation on the combined use of solid and gaseous fuel to effectively support sintering is drawn in the conclusions (section 5).

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2. Sinter model 2.1 LTE versus LTNE approach Due to the small solid-to-gas heat transfer coefficient as well as the high solid’s heat capacity a locale thermal non-equilibrium (LTNE) between porous solid and permeating gas [17] must be expected. The LTNE applies at least within the current extension of the heat wave in the sinter-bed. Therefore, as standard in advanced sintering models [19], a detailed examination of the heat and flame front dynamics call for a model setup including a combined solution of the energy conservation equations for the solid and gas temperature field (see section 2.2). Within a preliminary step however, valuable insight considering the travelling speed of the heat wave and its dependence on sintering operation parameters can be gained by examining a pseudo-homogeneous model view [16,20]. Within this local thermal equilibrium (LTE) assumption, implying infinitely fast solid-gas heat transfer, Cheng et al. [6] extracted a definition for the speed of the heat front (HTF)



v HTF  u g  g c g  g c g  (1  ) s c s



(1).

Characteristically, eq. (1) involves the linear dependence on the gas velocity ug in the solids pore space which was also proposed in former correlations for vHTF in [20-22]. Due to the dominance of the solid’s heat capacity compared to the one of the gas the heat front travelling speed vHTF turns out to reach much smaller values than ug. In addition, the HTF is considerably affected by the temperature dependence of the gas density g and heat capacity cg. Therefore, a change in the gas temperature profile and eventually gas composition will considerably alter the residence time of the heat wave in the sinter bed.

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2.2 LTNE governing equations Focusing on the mass and energy transports in perpendicular direction towards the bed surface we apply the following one-dimensional mathematical model to describe the heat wave travelling through the bed theoretically. Eq. (2) and (3) formulate the balances of solid and the gas heat transport in a vertical 1D reference frame of the sinter bed considered as a homogeneous porous medium. All terms in eq. (2) and (3) are formulated per control volume (m-3) of the porous medium consisting of the granular solid (index s) and the interstitial gas phase (index g). Further explanation of the symbols used is given in the nomenclature section.

T     T  (1  ) s c s  s   (1  )   eff ,s s   (1  )a v h sg (Tg  Ts )  Q c  Q add z  z   t 

(2)

T   Tg   T         eff ,g g   u g  g c g  g   (1  )a v h sg (Tg  Ts )  Q ng  g c g  z  z   t   z 

(3)

As a simplification the porosity  is kept as a constant parameter throughout the bed, neglecting any change in the geometry due to bed shrinking and pore space reorganisation during ongoing sintering. The solid-gas heat transfer is proportional to the temperature difference (Tg – Ts) times the heat transfer coefficient hsg and the term (1-)av denoting the (outer) specific surface area of the solid per control volume of the bed. The factor ug in the convective heat transfer of the gas phase (second term on the rhs of eq. 3) denotes the superficial gas velocity. It should be noted that the inter-phase heat transfer will depend on the suction regime quantified by the gas velocity ug [m/s] within the bed’s pore space. The heat conduction terms (first term on the rhs of eqs. 2 and 3) deserve specific attention even though the pure material conductivities be of low significance. Effective properties eff,s and eff,g might gain importance since they depend on the structure of the solid, the radiative transport and flow conditions. In line with an effective thermal conductivity approach [23] eq. (4) introduces a volume-

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weighted average combining a temperature-dependent radiative plus a molecular diffusive conductivity for the solid. 3

 eff ,s   rad  (1  ) s  4d s Ts  (1  ) s

(4)

Considering the gas phase, dispersive contributions will be significant compared to molecular heat diffusion due to mixing within the interstitial gas at the pore scale and formation of tortuous flow channels. Taking orientation at Hsu and Cheng [24] we model heat dispersion in the gas by a Péclet number modification of the conductivity in the form eff,g = gPe(1-)/ with Pe = ugdp/g. Note the thermal diffusivity of the gas g = g/(cp)g effectively cancels a contribution of the molecular heat diffusion in the conduction term in eq. (3). The solid-gas heat transfer coefficient hsg in eqs. (2) and (3) is based on the well recognized Wakau-correlation for granular packings, which was reviewed recently for lower values of the complex (Re0.6Pr1/3) by Zanoni et al. (2017). Although our operation case studies in section 3 are characterized by rather small particle sizes with Renumbers Re < 250 we find the heat transfer for cool and hot bed conditions still laying well within the applicability range of the Wakao correlation. To keep the overall model layout compact considerable simplifications are introduced in the source terms of the eqs. (2) and (3), see the following sections 2.3.and 2.4. Besides the volumetric energy sources by the heat input of coke combustion Qc and natural gas combustion Qng some further latent heat fluxes (Tab.1) are considered which although comparable small in quantity are able to noticeably alter the shape of the temperature transient.

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2.3 Coke combustion The added coke in the considered rawmix is located preferentially in a fraction-of-amillimeter thin layer formed by buildup-agglomeration on the surface of the larger ore particles. In our model we view this coke content simplistically as a fraction of the total packing surface, which will be denoted further as the surface of the rawmix packing (rmp). If one neglects gas-film transport resistance of the oxygen supplied to the reacting surface only the reaction controlled term is maintained, which in intrinsic form acc. to Smoot & Smith [25] reads rc ,int  k 1 exp E a / RTs p O 2

(5)

Preliminary calculations however demonstrated that a realistic overall coke oxidation demands consideration of the declining intensity of the combustion rate during the progress of conversion towards full burnout. The rate declining effect of a Cranck type diffusion control on the apparent combustion rate is described by a conversion function f(x) acc. to Ginzling & Brounshtein [26] in the form

rc ,app  dx dt  rc ,int f ( x )

with

f ( x )  (3 / 2)[(1  x ) 1/ 3  1] 1

(6)

The coke conversion rate rc as applied in eq. (6) is defined as mass of coke oxidized per second and available coke surface area. Coke mass is balanced in terms of carbon content per bed volume, mc [mol/m³]. The corresponding mass balance equation with initial condition reads

m c   M c t

with

mc

t 0

m c ,0

(7)

 [mol C/(m³.s)] Converted carbon mass from coke per unit volume of the sinter bed M c  r A is relates to the coke combustion rate by M c c s ,coke , where a simplified geometrical estimate for the coke surface density within the sinter bed is provided in the form A s,coke  (1  ) A s,rmp  c . The Arrhenius term for the coke conversion rate rc depends on

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the solid temperature rc =f(Ts) and introduces a strong nonlinearity in the associated heat source in eq. (2). For the usual conversion temperatures exceeding 1000 K carbon monoxide (CO) will be the primary oxidation product, which will predominantly convert to carbon dioxide (CO2) due to the excess oxygen of the blast gas. The coke combustion heat input in eq. (2) involves the following Boolian condition

 H Qc  M c coke

if (m c  0)

(8)

where Hcoke denotes the enthalpy of coke combustion, which is in the order of 30240 kJ/kg for a pure dry coke acc. to Simmersbach [27].

2.4 Homogeneous gas reaction and further source terms Oxidation kinetics of carbon monoxide (CO) as used in advanced sinter models (among others [19,28]) indicate a high reaction rate of CO oxidation in dependence of oxygen as well as water vapour concentration and gas temperature. Exemplary the kinetic equation suggested in [19] as originally developed by Howard [29] is taken as a reference: 0.5

rCO  3.25.10 7.c CO .c O 2 .c H2O

0 .5

exp 15098 / Tg  [mol / s]

(9)

In line with recent experimental findings [30] for oxidation of lean CO containing mixtures under moist conditions, eq. (9) indicates, in case of sufficient oxygen supply, a fast CO oxidation leading to complete burnout within reaction times of some hundred milliseconds for temperature levels exceeding 1000 K and nearly vanishing conversion for cooler gas with temperature below 925 K. Due to the fast homogenous oxidation compared to the moderate kinetics of the coke combustion for our model in case of limited oxygen supply we assume preferential conversion of gaseous combustibles in the reaction zone. As secondary gaseous fuel methane (CH4, Hcombut = -890 kJ/mol at standard cond.) is intermittently added to the suction gas at the sinter bed surface. We apply a

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simplified 1-step kinetic model of Gosiewski et al. [31] for the overall reaction CH4 + 2O2  CO2 + 2H2O, which was originally developed for combustion of lean methan/air mixtures under laminar pore flow conditions and excess of oxygen in a monolithic ceramic reactor, composed of parallel channels with 2 mm channel width.

rCH4  6.86.10 6.c CH4 . exp 130622 / Tg  [mol / s]

(10)

Further source or sink terms of latent heat as considered in the solid-energy and species mass equations are summarized in Tab. 1. The thermal decomposition of ferrous carbonate is known to show moderately high kinetics not below a temperature of 800 °C (FeCO3(s)  FeO(s) + CO2(g), eventually followed by oxidation of wüstite to magnetite and hematite [32]). Although the considered ore mixture in our experiments is rich in siderite (approx. 40 wt.% on dry ore basis), a weak exothermic contribution of the oxidative siderite decomposition [32] was considered of low significance during the raw-mix heat-up compared to the dominating endothermic processes of moisture evaporation, calcinations of limestone or dolomite and melt phase formation, which sum up to an order of 100 kW/m3 of heat sink in the sintermix.

Tab. 1 Latent heat contributions of non-fuel related reactions and phase transitions Rate model

Ref. Latent heat

Source / Sink-terms in

Drying of solids Condensation of moisture

[19]  2.350 kJ/mol H2O

Energy eqs. and species mass balances

Melting of sinter-mix Solidification of sinter-mix

[19] -0.254 MJ/kg +0.117 MJ/kg rawmix

Solid energy eq. Solid energy eq.

Calcinations of limestone [33] -0.879 MJ/kg rawmix above 800 °C

Solid energy eq. and species mass balances

2.5 Ignition boundary conditions To solve the solid- and gas energy equations (2) and (3) continuity of temperature and heat fluxes has to be prescribed as inlet boundary conditions at the top surface of the

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granular bed for both media. In preliminary calculations it was detected that various choices of the initial and boundary conditions strongly effect the calculated heat wave dynamics. To obtain a fair agreement with the temperature response of the thermosensor located nearest to the bed surface (see experimental section) the following procedure was adopted. As an initial temperature pulse LTE conditions were prescribed during the ignition period with Ts and Tg maintained at the adiabatic flame temperature T0 of the ignition burner within a surface layer of L=9 mm. After ignition end the Tg(z=0) at the bed surface switches back to ambient conditions whereas Ts(z=0) is defined in time-dependent form as Ts (t )  T0  293  exp t  th   293

for (z  0)

(11).

The post-ignition transient of the bed surface temperature acc. to eq. (11) follows from a heat balance of the ignition zone of thickness L, with the thermal relaxation time th = sLcs/hsg,0

(12),

which can be estimated by fitting eq. (11) to the monitored top-surface temperature decrease as recorded by thermographic imaging during the experiments. The solid-gas heat transfer coefficient hsg,0 in eq. (12) denotes the conditions of air suction at the bed surface (marked by the index 0).

2.6 Model scope The model in its present form comprises eqs. (212) and is limited to a one-dimensional description of axial gas and heat transport in a porous sinter bed. It is intended to be applied to the heat wave dynamics in a lab-scale sinter column experiments as introduced in the following section.

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3. Experimental investigations 3.1 Construction and operation of the SASITE facility Batch experiments in sinter pots provide a common methodology for iron ore sintering on a lab-scale basis. Typically, sinter pots have diameters in the range of 30 - 60 cm and a material packing height of 40 - 60 cm comparable to the bed layer thickness of industrial sinter belts; sinter pots of this dimension produce 30 - 300 kg of sinter per test run. A batch operation on this scale involves time-consuming steps of material preparation for the green mix, post-processing of the sintered material and data evaluation cycles. Therefore, an important aim of the current research work was to minimize the effort for material pre- and post-processing by miniaturization of the labscale sinter test. A small sinter test facility (SASITE) was developed based on an inner diameter of the suction tube section of only 70 mm in close analogy to the dimensioning of the sinter experiments described by Ooi et al. [34]. A schematic view of the SASITE experimental setup is provided by Fig. 2. The cylindrical sinter packing of 70 mm diameter and a bed height of 40 cm is enclosed in a refractory ceramic tube, equipped with a hearth layer on a perforated carrier disk. The setup further consists of a natural gas / oxygen fueled ignition burner, a secondary gas injection, an infrared camera inspection of the ignited bed surface, an off-gas suction system with integrated gas analysis as well as control and data acquisition related instrumentation. During the development phase, prior experiments were run in a quartz glass tube as transparent containment for the sinter packing providing direct visibility of the flame front dynamics. Further analysis however revealed the necessity to restrict radiative heat losses and to include a direct monitoring of the developing temperature profile by a set of up to eight equally spaced thermocouples (type K or alternatively N) along the axial extension of the bed. This specifications were found easier to be met by a cylindrical

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refractory enclosure (alumina) including precision drilled holes for the hose mounting of the thermocouples and optional pressure transducers.

Fig. 2 Lab-scale sinter test (SASITE) with secondary fuel gas injection; flame monitoring and automation of the ignition burner not shown.

The granulated raw mix for the sinter experiments was directly obtained from the industrial project partner. The standard mix consisted of granules with a means size of d50 = 3.2 mm and a moisture content of typically 3.4 wt.%. More details on the size and composition characteristics of the raw mix are summarized in Table 2. Special care is taken to maintain the nominal moisture content and homogeneous moisture distribution of the native raw mix from the sinter plant in the SASITE sample material. Thermal

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gradients during storage have to be avoided to ensure a stable granule size distribution without moisture and/or particle size segregation in the packed material. Charging of the sinter bed tube starts with a heard layer of sintered material with particle sizes up to 10 mm. The thermocouples are protected by corundum wrappings with similar thermal capacity as the sinter material and have to be carefully embedded in the granular packing which is charged layer by layer to ensure homogeneous compaction. Prior to the sinter experiment packing density and cold flow permeability (in each thermocouple depth) are checked to confirm the generated SASITE packing structure to show comparable characteristics to conventional sinter pot and industrial sinter belt conditions. The range of packing related parameters obtained for the green bed prior to the experiments are summarized in Table 2. Tab. 2 Size and composition characteristics of the standard sinter raw mix Raw material composition: Iron ore, wt. % Return fines, wt.% Lime and dolomite, wt.% Other additives Coke, wt.% Initial moisture (based on 100% dry mix) Size distribution: Mean granule size, d50 RRS*) skewness, n [-] RRS*) characteristic diameter, dp’ Green bed packing characteristics: Initial packing height Bulk density Bed porosity Specific surface area Pressure drop

56 18 4 17 5 3.4  0.4 wt.% 3.2 mm 1.4 4.0 mm 400 mm 1950  32 kg/m³ 42.5  0.9 vol.% 2500  300 m²/m³ 135  15 mbar/m

*) In the RRS size distribution model the cumulative mass fraction R(d) of particles larger than a given grain size d is described as R(d) = exp[-(d/dp’)n] with the characteristic grain size parameter dp’ and the skewness parameter n.

Measured from the electrical spark ignition of the burner the ignition period lasts for 90 seconds. After ignition, the air blast drawn through the bed immediately starts cooling

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the top layer which could be well observed by an infrared thermo-camera, which was positioned to monitor the bed surface temperature. As soon as the downward directed heat flux rises the raw material’s temperature ahead of the flame front up to coke ignition the combustion zone shifts downward. When the flame front has reached the bottom of the packing and the temperature at all thermocouple positions have dropped below 100 °C a additional permeability test for the sinter product is conducted. At the end of the experiment, the vacuum pump is switched off and the sinter material is dismounted, photographed and its strength measured by disintegration tests according to the corresponding norms for the shatter (JIS-M 8711) and the tumble (ISO 4696-1 ) indices.

3.2 Case studies A test case study was performed to investigate the substitution of coke breeze by methane as a secondary gaseous fuel. As already known from previous tests the total amount as well as the temporal variation of the fuel gas feed need proper adjustment to counterbalance the combustion heat loss due the reduced amount of coke effectively. The test design aims to develop a gas injection scheme capable to maintain the product sinter strength at an acceptable level for the coke reduced sinter cases. In previous investigations of gas assisted sintering, the burning gas feed was kept constant [18, 35, 36], or was varied stepwise [37] to support the heat pattern of the flame front. In our investigation, we inject the fuel gas in form of cyclic pulses with constant frequency. For our suction regime and a maximum methane flux of approx. 2 Nm³/h preliminary experiments revealed an injection time interval exceeding 60 s may eventually lead to complete consumption of oxygen in the suction gas, which was supplied at the bed surface with a superficial gas velocity of 1.2 m/s. Therefore, pulse duration was limited

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to 30 s followed by an off-period kept constant at 60 s. Fuel gas flux and number of gas pulses applied in the experiments are specified in Table 3. The coke breeze level for the experiments with reduced amount of coke was defined at 3.5 and 2.5 wt.%, the later corresponding to a saving of roughly half of coke compared to a standard sinter rawmix. For a coke content of 3.5 wt.% in a SASITE-experiment, heat balancing reveals 1.5 MJ of calorific value loss, which has to be compensated with the secondary fuel gas. At a fixed methane gas flow of 0.43 moles per minute, a total gas injection time of 4 minutes can deliver this missing calorific value. In this case, within the injection time intervals methane concentration in the suction gas reaches 3.2 vol.%. The fuel gas pulsing scheme is shown in the schematic graphs of Fig. 3. The base case (Exp01) and one of the reduced coke experiments (Exp05) are carried out without any gas intervention. The gas supply nozzle for the tests with added secondary fuel gas is put in place shortly after the ignition of the material. The gas supply is positioned central approx. 3 cm above the bed surface and the rising edge of the first gas pulse is initiated 5 min after the end of ignition. Roughly within this time, the flame front reaches the TC 3 thermocouple which is located in a sinter bed depth of 13 cm. Tab. 3 Sinter test conditions of the standard coke and reduced coke mixtures; methane injection scheme Experiment nr.

Exp01 base case Exp02 reduced coke content Exp03 reduced coke content Exp04 reduced coke content Exp05 reduced coke content Exp06 reduced coke content

Coke DCGF content *) **) [wt.%]

[%]

5.0 3.5 3.5 3.5 2.5 2.5

-100 120 140 0 138

Methane injection Number of gas pulses -8 8 8 -8

Pulse duration [s] -30 30 30 -30

Gas flux [Nm³/h] -0.59 0.98 1.37 -1.72

*) specified on dry raw-mix basis **) energy equivalent of coke reduction – degree of compensation by gaseous fuel (DCGF)

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Fig. 3 On/off-pulsing scheme for the fuel gas injection during sintering time interval.

Tab. 4 Thermo-physical and chemical-kinetic parameters applied in the simulations Parameters: Specific heat of dry suction gas, J/(kg.K)

910+0.303Tg-6.7e-5.Tg2

Ref. [39]

Specific heat of solid, J/(kg.K)

672+0.386Ts-1.08e-7/Ts2

[39]

Calorific value of dry coke

30240 kJ/kg

[27]

Melting start temperature (begin of melt formation)

1300 K

[19]

Temperature of total melting (fusion)

1600 K

[19]

Bed surface temperature at the end of ignition

1800 K

exp. 1.1(Re0.6Pr1/3)

Correlation for the solid-gas heat transfer coefficient

Nu = 2 +

Coke combustion; Arrhenius term in eq. (5) at pO2=21 kPa

rc=106exp(18000/T) mol/(m³.s)

[17] [25]

21

4. Results and discussion 4.1 Model validation for the experimental base case Experimentally, during sinter progress a typical decrease of the bed height was observed reaching 2.5 – 3.0 vol.% in total. The bed shrinkage at least partly compensates the increase of the sintered bed‘s porosity, which mainly follows the evaporation of moisture and the carbon burnout. Making use of the density of the product sinter material (derived by helium pycnometry as Sinter = 3.7 kg/m³) the porosity in the sintered bed was estimated to be 47.6 vol.%. A value exceeding the green bed porosity of 42.5 ± 0.5 vol.% given in Tab. 2 by roughly 5 percentage points. Theoretically, for suction under fixed pressure drop increasing permeability will rise the gas flux within a porous domain and eventually induce a progressive growth of the heat wave speed as indicated by the proportionality of ug and vHTF in eq. (1). However, as will be analyzed in more detail in section 4.2, we always find a constant heat wave speed for the wave travelling downward in the bed. The reason is, the total flow resistance which is always dominated by the structural transformation in the flame front as already pointed out by Loo and Leaney [38]. From this fact, we conclude that for suction under fixed pressure drop the change in porosity above and below the heat wave does not alter the volumetric gas flux significantly for conditions sufficiently far from the burn-through point.. Therefore, as a simplification in all calculations, it was decided to keep porosity at its green-bed level and constant over bed height. Models of thermo-physical and chemical kinetic parameters maintained unchanged in all simulations are summarized in Table 4. Additionally, the moisture dependence of the gas density and heat capacity as a function of temperature was introduced via a mass-weighted mixture formulation based on dry gas and pure steam properties [40]. Experimental conditions of the base case (Exp01) as specified in Table 2 and Table 3

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were used to investigate the quality of the numerical model prediction by a comparison of the calculated temperature transients with the temperature signals obtained from the thermocouples positioned within the bed. The thermal initial and boundary conditions for the simulation were extracted from the experiment as follows. Solid and gas temperature at the bed surface were prescribed as constant during the ignition period of 90 seconds followed by an exponential decrease of the solid temperature at the begin of the suction period. Area averaged mean temperatures were applied for this timedependent boundary condition, which have been evaluated from the thermographic imaging as shown in Fig. 4. A fit of the bed surface temperature transient by the decay function (11) provides a thermal relaxation time th, which according to the definition (12) can be used to estimate the heat transfer coefficient specifically at the bed surface at the end of ignition. A range of 175 < hsg,0 < 200 Wm-2K-1 is evaluated from the fit, in fair agreement to the value provided by the Wakao-correlation in Tab. 4 which for end of ignition conditions at the bed surface evaluates to hsg,0 = 228 Wm-2K-1. The time interval of our simulations starts at the end of the ignition and reaches up to the burn-through point. Calculated and measured temperature profiles for the base case Exp01 are compared in Fig. 5. The simulation is able to reproduce the arrival time of the heat wave by the peak temperature at the sensor positions quite satisfactorily. The nearby exponentially decaying shape of the temperature decline behind the heat wave can be described satisfactorily, due to the diffusion limited down-weighting of the intrinsic coke combustion rate as introduced in eq. (6). The increasing importance of diffusion limitation for growing coke conversion may be explained by the formation of an ash layer or a melt film. Furthermore, the duration of the travelling heat wave in Fig. 5 is stated realistically, as indicating the time interval a given sensor position is occupied by the high temperature zone. Some deficiencies exist in the exact height of

23

the simulated temperature maximum – which is not fully reached for the first thermocouple position – as well as with respect to the onset characteristics of the temperature rise. It should be considered however, that the heat front dynamics additionally depend on the specific shape of the dry-out zone in the upper part of the bed at the end of the ignition step.

Fig. 4 Post ignition temperature decline at the sinter bed surface; for the simulation runs the bed surface temperature transient is derived from the area averaged mean temperature, which was extracted from the thermographic imaging. The fitted exponential decay “T_solid” approximates the time-dependent boundary condition for the solid’s surface temperature in the model calculations, acc. to eq. (9). The fit’s initial decline is used as linear approximation “T_gas” for the inlet gas temperature.

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Fig. 5 Calculated solid and gas temperatures compared to measured bed temperature profiles at thermocouple positions of 9, 18 and 27 cm bed depth; lab-sinter test Exp01.

The exact moisture profile at the beginning of the suction period may not be fully identical for all experiments. Therefore, minor

deviations of the calculations will

always exist, due to the simplified prescription of a complete dry-out of 4.5 cm plus 4 cm ramp up to the initial bed moisture. Reasonable reproducibility is given for repeated experiments as demonstrated in Fig. 6. Two realizations are shown for the base case as well as the reduced coke experiment and three realizations for the experiment with burning gas injection. For a better orientation, the burn-through point is marked by arrows at the time axis in Figs. 6 at the off-gas peak temperature for each case. Reproducibility is better for the pure coke cases compared to the case with additional gas injection. Generally, the signals at thermocouple positions 9, 18 and 27 cm show better consistency compared to the temperature readout near the lower end of the bed

25

and in the off-gas. Therefore, the averaged temperature responses from the central bed at positions 9, 18 and 27 cm where taken as a basis for comparisons with the theoretical calculations (cf. Figs. 5 and 7).

Fig. 6 Experimental variation of thermocouple temperatures for repeated experiments; Exp01 – base case (a), Exp05 – reduced coke without gas (b), Exp06 – reduced coke with gas (c). Dash-dotted lines indicate averaged profiles.

As in the simulation study of Zhang et al. [39] based on a comparable implementations of the LTNE approach, we find the gas temperature response running ahead of the solid heat wave. The calculated case in Fig. 5 shows a temporal offset between gas and solid temperature rise of up to 40 s within the temperature regime above 600 K. As the conduction within the solid grains is the slowest heat transfer process involved, convective heat transport in the low temperature region as well as radiative heat

26

transport between neighbouring grain surfaces in the high temperature region will distinctly enhance the heat transport through the packing. Although the exact positioning of the thermocouple’s tip within the granular packing is not exactly defined, assuming the tip temperature to be preferentially driven by the most dominating mechanism of heat transport, one can expect the measured temperature to be close to the LTNE-derived Ts at least in the radiation dominated high temperature region.

Fig. 7 Calculated solid and gas temperatures compared to measured bed temperature profiles at thermocouple positions of 9, 18 and 27 cm bed depth; lab-sinter tests Exp05 and Exp06 with reduced coke (a) without and (b) with additional fuel gas injection.

4.2 Reduced coke cases and effects of fuel gas injection The diagrams in Fig. 7 show measured and calculated temperature profiles for the reduced coke cases (a) without and (b) with additional fuel gas. The peak temperature decline of the heat wave – even in the extreme case of a coke content reduced down to 2.5 wt.% – could obviously be counterbalanced with the pulsed gas injection. Test runs with up to four thermoelements where performed to document the thermal evolution of the base case Exp01 and compare it to the gas injection case Exp06 in the temperatureisolines plot Fig. 8. The travel speed of the heat wave on its way down the bed can be estimated from the inclination of temperature isolines. Additionally, the heat wave

27

speeds can be summarized in form of peak temperature arrival times at the defined sensor positions of the experimental and simulated runs as shown in the insert of the diagram in Fig. 8.

Fig. 8 Temperature isolines for the coke-only base case Exp01 and the reduced-cokewith-gas test case Exp06 (two consecutive realizations). Experiments are artificially shifted per 100 seconds on the time axis for better visual representation. Figure insert shows comparison of experimental and simulated temperature peak times at selected sensor positions.

With 37 mm/min, the heat wave in the gas assisted sintering case is significantly faster than the one in the pure coke sintering cases Exp01 and Exp05 which travel with 26 and 27 mm/min respectively. According to equ. (1) at constant suction conditions any

28

increase of the gas volumetric heat capacity gcg will shift the heat front speed vHTF to higher values. Particularly, water vapor turns out to vigorously enhance the vHTF, if one considers the contribution of the vapor’s heat capacity especially in the high temperature range within the heat front. Two moles of vapour are produced per mole methane burned. For this reason, combustion generated moisture present in the high temperature zone is able to more actively alter the vHTF, compared to the vapour from the moisture evaporation, which predominantly is moving ahead of the heat front. Despite of the excess oxygen typically being present in the sinter off-gas it turns out from the experiments that fuel carbon does not burn completely to carbon dioxide. Some small fractions of carbon monoxide are always detectable in the sinter off-gas. In all experiments CO concentration increases with falling residual oxygen content, cf. Fig. 9. A roughly linear correlation is found for experiments with different secondary fuel compensation rates DCGF = 100 % and alternatively 120 %. (Results for DCGF = 140 % are similar to the 120 % - case and therefore are not shown in Fig.9. In detailed view, the state point of the CO- versus O2-concentration follows a loop-shaped trajectory with the highest CO-contents found during the begin of the off-gas temperature-rise, even before the flame front is approaching the burn-through point. In case of higher secondary fuel input flux there is a tendency to shift the off-gas COconcentration to higher peak values since carbon burnout becomes increasingly incomplete for lower residual oxygen concentrations.

29

Fig. 9 Scatter plot of CO versus residual O2 concentration in the off-gas during tests with 3.5 wt.% coke (Exp02 and Exp03 of Tab. 3). The trajectory of the state point for each experiment generally follows a loop-shape as shown in the figure insert.

4.3 Quantification of sinter strength parameters Sinter strength of the SASITE samples was measured by disintegration tests. Thereby, the sinter’s resistivity against abrasive wear and forced breakage of the primary sinter lumps was quantified in form of scatter and tumble indices. After finish of the sinter experiment, the sinter column was cooled by forced convection under defined air suction. The cooling procedure was kept constant to maintain comparability between all experiments. Dismantling of the produced sinter from the test

30

rig and removal of the corundum wrappings of the thermocouples from the sinter was performed immediately after cooling directly followed by the disintegration tests. It was found important to avoid any bias of the strength results by potential aging effects from storage of the sinter samples under atmospheric contact. Fig. 10 summarizes strength results for a selected set of sinter samples produced under varied operating conditions in the SASITE facility. In the tumble test, the dominant sliding motion of the bulk material causes wear due to abrasion. Alternatively, a defined series of drops in the shatter test breaks the sinter parts into smaller fragments. Despite the fact that dynamic stress and wear conditions are different in both tests, tumble and shatter results roughly correlate. For the reduced coke sinter tests, strength is significantly improved when increasing amounts of secondary fuel gas are applied. The gain of strength for the coke reduced rawmix with 3.5 wt.% coke is high enough to reach an acceptable level of sinter strength in comparison to the base case with standard rawmix (coke content of 5.1 wt.%). In case of the coke reduced rawmix with 2.5 wt.% coke, the strength loss in relation to the base case can not be fully compensated even if the gaseous fuel provides 1.4 times the missing energy equivalent of the coke reduction.

31

Fig. 10 Tumble and shatter test indices for SASITE product sinter samples; Exp01 – Exp06 according to Tab. 3

Conclusions Potential benefits of gas assisted iron ore sintering was investigated by small scale batch sinter experiments accompanied by theoretical calculations based on a onedimensional heat front propagation model. The combustion heat input diminished by reduced amounts of coke breeze in the prepared rawmix was counterbalanced by additional methane injection as a fuel substitute. Contrary to most prior investigations of gas sintering, the secondary fuel was added to the suction gas in a pulsed injection mode, forming series of on/off-cycles with a cycle periods of 90 s which during the “on”-phase of 30 s provided a constant molar flux of methane into the suction gas.

32

Effects of fuel gas injection on peak temperatures, heat wave speed, off-gas parameters and sinter strength were studied. Main findings obtained in the study are as follows: (1) The pulsed fuel gas injection mode prevents decoupling of the solid and gaseous fuel’s flame front and ensures a localized consumption of coke and methane in a compact combustion zone. In the dual fuel operation satisfactory peak temperature levels can be maintained within the heat wave even if coke content in the raw-mix is reduced down to one half of the original value. (2) In accordance with model based considerations, at constant suction conditions, methane combustion in the flame front changes the heat transport capacity of the suction gas and causes an increase of the propagation speed of the heat wave compared to the pure coke combustion case. (3) Incomplete carbon burnout during sinter progress produces a volume percentage fraction of carbon monoxide. The CO emission follows the residual oxygen content in the off-gas with mainly linear dependence and shows a minor increase with the feeded amount of gaseous fuel. (4) For coke reduced rawmixes, both, tumble and scatter indices could be significantly increased by the additional fuel gas supply. This findings are in agreement with recent results on gas assisted sintering of other investigators, e.g. [37]. However, for our fuel gas pulse scheme defining subsequent pulses at constant methane concentration, a certain over-compensation of the missing energy equivalent due to the coke reduction is needed to fulfil sinter strength requirements. Compared to the base case with 5.1 wt.% coke, the lower acceptance limit for the product sinter strength could be reached when the added gaseous fuel at least supplied 1.2 times the missing heating energy equivalent of the coke reduction. Contrary, for a rawmix with only 2.5 wt.% coke, the strength limit was missed, even with the fuel gas adding 1.4 times the missing energy equivalent

33

of coke reduction. Since the net heat input may not be equal in the upper and the lower bed, further optimization of the pulse scheme with changing pulse concentration along sinter progress will be needed. (5) If the proposed intermittent gas injection scheme is realized at an industrial sinter belt, the pulsed injection mode developed in the batch sinter experiment will transform to strip-shaped domains at constant distance along the sinter belt, where at defined positions, a continuous supply of fuel gas is feeded to the suction stream at the bed surface.

Acknowledgment The authors gratefully acknowledge the funding support of K1-MET GmbH, metallurgical competence center. The research programme of the K1-MET competence center is supported by COMET (Competence Center for Excellent Technologies), the Austrian programme for competence centers. COMET is funded by the Federal Ministry for Transport, Innovation and Technology, the Federal Ministry for Digital and Economic Affairs, the provinces of Upper Austria, Tyrol and Styria as well as the Styrian Business Promotion Agency (SFG). Beside public funding, the project was partially financed by the industrial partners voestalpine Stahl Donawitz, voestalpine Stahl and Primetals Technologies Austria.

34

Nomenclature av

volume-specific surface area, m² m-³

As

granular surface area, m²

c

mass specific heat capacity, J kg-1 K-1

dp

equivalent diameter of solid material, m

D

mass diffusion coefficient, m² s

hsg

solid-gas heat transfer coefficient, W m-2 K-1

m

particle mass

M

molecular weight, kg mol-1

Nu

Nusselt number = hsg dp/g, [-]

Pr

Prandtl number = µgcg/g, [-]

rc

coke conversion rate per m² reactive surface, kg m-²s-1

rCO

rate of homogeneous oxidation of CO based on gas volume, mol m-3s-1

Re

Reynolds number = g ug dp/µg, [-]

RRS

Rosin-Rammler-Sperling size distribution function, cf. Tab. 2

t

time, s

T

temperature, K

ug

gas velocity within porous medium, m/s

v

front speed, m s-1

x

coke conversion at time t defined as (minit – mt)/(minit – mfinal), [-]

z

vertical coordinate direction in the sinter bed, m

Q

heat release by reaction, J kg-1

Hevap enthalpy of evaporation/condensation, J kg-1

35

Greeks 

thermal conductivity, m² s-1



sinter bed porosity, -

c

volume fraction of coke in the raw-mix, [-]

eff

effective thermal conductivity, Wm-1K-1

µ

dynamic viscosity, Pa.s



mass density, kg m-3

th

thermal relaxation time, s

Subscripts and superscripts 0

ignition condition at bed surface

add

additional

c

coke

g

gas phase

ng

natural gas

rmp

rawmix packing in the sinterbed

s

solid phase

36

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Highlights • Control of flame front propagation in small-scale iron ore sintering studies • Partial substitution of the standard fuel coke breeze by a pulsed gas injection • Numerical modelling of the heat front travelling through the sinter bed • Lab-scale sintering tests with reduced coke breeze content in the sinter mixture

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41