Transition Mechanism Between Combustion Regions in Swirling

Jan 3, 2017 - Nuclear Research Center of the Negev, Tel-Aviv 61070, Israel. ABSTRACT: Computational examination of pulverized coal combustion in an ...
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Transition Mechanism Between Combustion Regions in Swirling Entrained Flow Downer Reactors Tal Eluk,*,† Avi Levy,† Efim Korytnyi,‡ and Tali Bar-Kohany§,∥ †

Department of Mechanical Engineering, Ben-Gurion University of the Negev, Beer-Sheva 8410501, Israel Ernst David Bergmann Campus, Ben-Gurion University of the Negev, P.O. Box 653, Beer-Sheva 84105, Israel § School of Mechanical Engineering, Tel-Aviv University, Tel-Aviv 69978, Israel ∥ Nuclear Research Center of the Negev, Tel-Aviv 61070, Israel ‡

ABSTRACT: Computational examination of pulverized coal combustion in an entrained flow downer reactor exhibited the existence of two distinct combustion regions referred to as near burner combustion (NBC) and far region combustion (FRC) with lack of mediate state. Operating conditions and material properties were altered within a broad range in which the reactor’s behavior was studied. A transition mechanism between the two combustion structures was recognized and isolated to illustrate ranges of stable combustion. NBC was mainly affected by the hydrodynamics while FRC was mainly affected by the devolatilization kinetics. Non dimensional analysis of the gas solid behavior yielded Stokes values close to unity for the transition cutoff. A new methodology for the selection of a distribution representative particle diameter is presented and assumed to be applicable for various types of two phase exothermic reacting flows. The pulverized coal combustion inside the reactor was simulated using the Euler−Lagrange approach and validated against experimental results. Special attention was given to the backmix of hot gases combined with the inherent recirculation zone near the reactor’s entrance for their effects on the transition mechanism between combustion regions. Understanding these mechanisms will lead to better control over processes regarding phase continuous mixing vs short contact within entrained flow downer reactors.

1. INTRODUCTION The entrained flow downer reactor as a research and industrial facility has been gaining popularity for the past few decades. As this reactor becomes more common, the necessity to fully understand its behavior and to assess its functionality is crucial. The concept of the reactor is described by Zhu et al.1 as cocurrent aided by gravity downflow fluidized bed reactor, or simply “downer reactor”. Characteristics that have been attributed to downer reactors are the potential of short contact time between phases, low axial dispersion, and narrow residence time distribution, i.e., less back mixing of solids and products than the bubbling fluidized bed or the up flowing bed (riser).1,2,3,4 Due to these characteristics, downer reactors are chosen to perform tasks, such as catalytic cracking.4,5,6 They can also be found as a single unit or as part of a combined triple reactor cycle (riser, downer, bubble bed) for pyrolysis7 or gasification and combustion8,9 where the reactor is referred to as entrained flow gasifier\combustor. While some of the research on downer reactors is being done through performance evaluation of a full scale or a small scale industrial process, it can also be used exclusively as a test facility for the estimation of kinetic rates.9,10,11 A recent developing field is replacing traditional fuels with biofuel and biochar.12 These “new” fuel sources are widely diverse and would need to be characterized before implemented in industry. In these cases the same research methods for pulverized coal via downer reactor can be of use. It is evident that the supporting argument for the extensive use of downer reactors throughout a wide range of applications is its similarity to a plug flow reactor (PFR) behavior,1,3 i.e., a tubular vessel with no axial mixing and full radial mixing. Nonetheless, for some applications, namely © XXXX American Chemical Society

combustion and exothermic gasification, recirculation of hot gases can be vital for the maintenance of a steady process; thereby downer reactors can still prove themselves as useful apparatuses. Clearly, it is important to choose a suitable design and operating conditions in order to meet the industrial and research demands. When discussing nonpremixed combustion, the near burner region is of high importance since it is the place where fuel and oxidizer stream are injected and mixed. Usually, their contact at the near burner region regarding mass flow, velocity component, and axisymmetric design should promote back mixing of hot gases for flame stability.13 Swirl dominated flow has long been recognized as a method for maintaining combustion stability,14 as swirl creates a low pressure core resulting in a recirculation zone. Recent works on swirl dominated combustion of pulverized coal with double15,16 and triple17,18 inlet lines captured the hydrodynamics of the hot gas recirculation zone. Depending on the ratio between axial momentum to tangential momentum, i.e., swirl number, at the inlet, the flow structure can take a few known forms. For low swirl numbers, it is possible for the central axial jet to penetrate the toroidal recirculation zone while for high swirl numbers the central jet is absorbed by the recirculation zone. In this case a much larger back flowing core will be evident. As mentioned earlier, most processes require enhanced contact between solids and gaseous phase, while mixing through backflow and recirculation would only be desirable in some cases. Though Received: September 3, 2016 Revised: December 4, 2016 Published: January 3, 2017 A

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels the field of swirl combustors has been studied for decades, modern industrial demands incorporated with development of computational and experimental research methods leads to ongoing studies regarding swirl dominated combustion. Recent work19 demonstrated changes in flame-flow dynamics and flame structure for premixed combustion for different equivalence ratios. The present study aims to investigate transition of flame-flow strictures in the case of combustion of pulverized coal. The behavior of the swirl-dominated downer reactor can be roughly separated into two sections, the mixing zones at the entrance and the plug flow zone that follows. The CFD study is expected to give us insight about the extent of each section and help us design and operate the reactor.

Where F is the fuel stream built from the elements specified in Table 1 and A is the air stream.

Table 1. Melawan Coal Properties property proximate analysis (DCB)a (wt%) total moisture residual moisture ash volatile matter fixed C gross CV MJ/kg, wet net CV MJ/kg, wet Ultimate analysis (DAF)b (wt%) carbon, C hydrogen, H oxygen, O nitrogen, N sulfur, S

2. METHODS At the present study, a comprehensive model for pulverized coal combustion in an entrained flow downer reactor was built based on integration of multiple models formerly developed via commercial CFD code as commonly used for the study of multiphase reactive flows. The model is used for computational investigation of the twophase reactive flow inside the reactor, and as such, had to be primarily validated against experimental results. A brief review of the main models used in this work is presented here, followed by the geometry and operating conditions in a 50 kW test furnace10,20 at the Laboratory for Clean Combustion of Ben-Gurion University, whom the validation experiments where carried by. A comprehensive description of the various models can be found in the literature and in commercial CFD code’s documentation.21 The dilute solid phase was modeled using the Lagrangian approach applied by CFD-DPM coupling of heat, mass, and momentum. Turbulent effect on particle trajectories was taken into account using Discrete Random Walk model, modifying the slip velocity on the particle with an additional fluctuating component. The highest Biot number for coal particles reached 0.25, while 99.7% of the coal mass remained below 0.1, thus lumped heat assumption is acceptable The void fraction was above 99.9% and interparticle collisions were neglected, taking into account that certain deviation from optimal particle trajectories might occur even for highly diluted flows as suggested in related studies.22 Turbulent flow with strong swirl and a Reynolds number spanning on the lower range of the turbulent spectrum (3000 < Re < 5000) was modeled using the two equation RNG k-ε model. The ability of RANS-based turbulent models to predict the behavior of pulverized coal combustion is an ongoing discussion that evolves with the enhancement of computational capabilities and studies in the field, mainly with comparison to LES. Though LES models would yield more accurate microstructures; thereby presumably better quantitative results, it would also increase the computing time drastically. Moreover, recent reviews23,24 state that RANS models can yield acceptable accuracy for a wide range of industrial flows, as can be seen at the present study. Thermal radiation was taken into account using the discrete ordinates (DO) model that is suitable for particle radiation interaction. The milling output of the pulverized coal agrees well with the Rosin-Rammler distribution:

a

DCB: dry coal base. bDAF: dry ash free.

m v (t ) = (mp,0 − ma )

∫0

t

t

(α1R1 + α2R 2)e− ∫0 (R1+ R2)dt dt

(3)

Where mv is the volatile yield from a particle with initial mass mp,0 and ash content of ma. αi are yield factors and Ri kinetic rates with Ai and Ei as the rate constant and activation energy respectively, i = 1,2 for the first kinetic rate and second kinetic rate:

R i = Ai e−(Ei / RT̅ )

(4)

The released volatiles represent the fuel stream being used to calculate f and f ′ used by the mixture fraction model. Rate constants and activation energies were taken from existing data of the specific coal type.10 Char surface reactions were taken into account using a kinetic\diffusion limited model based on the work of Baum and Street.28 Surface reaction products were also represented by the mixture fraction model and considered as source term for the conserved scalar f transport equation. The complete model of the reactor was solved at steady state using commercial CFD code by Ansys-Fluent ver.16.2. Experimental results used for validation were performed on a 50 kW cylindrical test furnace seen in Figure 1 and thoroughly described by Spitz et al.20 The furnace is a 4.8 m long entrained flow downer reactor of 0.2 m diameter. Primary air containing the pulverized coal enters at low temperature and low velocity normal to the primary inlets while secondary air is preheated and injected at high axial and tangential velocities generating a strong swirl. Proximate and ultimate analysis of Indonesian sub-bituminous Melawan coal followed by operating conditions are specified in Table 1 and Table 2, respectively. Measurements of temperature and various flue gases were taken from opening locations along the reactor. Calculations were performed to simulate the first 3 m of the reactor. Outlet boundary conditions were set accordingly to match non developed swirl flow. Mesh independency was verified against a detailed mesh consisting of 1000 K cells, so the final mesh consists of 286 K hexahedron cells and has a maximum local error of 1.7% in gauge pressure, 0.17% in temperature, and 1.3% in axial velocity component sampled at the reactor’s axis and along a cylinder of half

(1)

Where Yd is the mass fraction of particles with a diameter greater than dp, d̅p is the mean diameter and n is the spread parameter. Particle size distribution was fitted to measurements with a mean diameter and spread parameter of 90.23 μm and 1.1215. According to the particle size distribution, residence time distribution and feeding mass flow rate, it was evaluated that the reactor contains approximately 109 particles at each given moment. The coal particles were represented by approximately 6K parcels that entered the domain by surface injection normal to the primary air inlets. Combustion was simplified to a mixing problem using the approach of mixture fraction ( f) as a conserved scalar which represents the fuel stream to air stream ratio:

F /A = f /(1 − f )

74.59 5.35 18.42 1.42 0.22

Assuming chemical equilibrium within the gas phase, one can account for temperature, density, and relevant species concentration with low computational cost. Active species that were considered are C, CO, CO2, CH4, C2H6, H, H2, H2O, HO2, N, N2, O, O2, OH. Discrepancies from chemical equilibrium are taken into account using the mixture fraction variance f ′ which is affected by the turbulent regime. The devolatilization mechanism was based on the work of Kobayashi et al.25 with two competing rates:

n

Yd = e−(dp/ dp̅ )

22.3 15.5 3.63 39.8 41.07 22.41 21.01

(2) B

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Figure 1. Furnace general view and sampling locations (left), cross section of burner (right) based off the work of Spitz et al.20 reactor and qualitative conclusions can be deduced from an extended computational study.

Table 2. Operating Conditions for 50 kW Test Furnace parameter primary air flow kg/h primary air temperature K secondary air flow kg/h secondary air temperature K total coal flow kg/h

11.255 340 33.54 575 5

3. RESULTS AND DISCUSSION 3.1. Transition between Combustion Regions. A thorough computational examination of the reactor exhibited the existence of two distinct combustion regions. The most obvious difference between them is the location of the combustion front or the location of intense exothermic reactions. One combustion region can be described as near burner combustion (NBC) while the second will be referred to as far region combustion (FRC) with combustion front taking place beyond 0.5 m from the entrance. It is clear from Section 2 that the experiment samples for the validation case is in NBC form. Evidence of FRC can be found in the literature13 specifically for pulverized coal with recirculation near the burner. Since for gasification and combustion FRC is not necessarily the desired case, it might have gone unmentioned while conducting an experiment. As it can be seen in Figure 3, FRC (a and b) and NBC (c and d) are easily detected by appearance alone regarding the gas temperature and the size of the backflow core confined by the zero axial velocity iso-surface, nevertheless, quantitative differences are introduced further on. Although notable changes in the volume and length of the backflow core are evident, it can be seen from the velocity vectors that the swirl induced flow structure does not change its characters during the transition. In both cases the central jets are overcome by the swirling jets and end up being absorbed by the recirculation zone, as identified with high swirl number flows. The unchanged swirl type structure can be attributed to the constant inlet conditions regarding the ratio between axial and tangential momentum. Transition between the two combustion regions can be obtained by altering various parameters, both from the fuel material properties point of view and the operating conditions. While both combustion structures exist at a broad range of operating conditions and fuel properties, the interesting thing to notice is that the transition is not a smooth and gradual one but rather abrupt. The difference between Figure 3a,b and c,d is a minor 2.1% change in devolatilization activation energy, which primarily is not an easily predicted parameter and this kind of change is well within deviation.26 It is more accurate to describe it as a “jump” between two states. As can be seen in Figure 4, when

radius distance. Since both gas and temperature measurements are suction based, it is more accurate to say that each sample represents a small volume inside the furnace. Therefore, experimental measurements taken at the furnace’s axis are compared to simulated results presented by an area weighted averaging of a 1 cm diameter cylinder around the axis. Experimental and simulated results are presented in Figure 2 showing good match of temperature and CO2 mole fraction. Relying on the validated results, the complete model can be further used for investigating the behavior of such entrained flow downer

Figure 2. Experimental (open grey diamond) and simulated results at furnace axis of temperature (up) and CO2 mole fraction (down). C

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Figure 3. Temperature contours with velocity vectors and iso-surface of zero axial velocity at furnace cross section for minor changes in devolatilization activation energy: (a) and (b) FRC with E = 9.6 × 107 J/kmol. (c) and (d) NBC with E = 9.4 × 107 J/kmol. Inner part of the zero axial velocity iso-surface represents the backflow core.

the nature of the developing swirl dominated flow at that region. This region is characterized by axial volumetric flow located mostly within close proximity of the wall (2/3 of the flow rate is within the 1/3 outer radius of the tube) together with high tangential velocity causing low pressure backflow core. A significant increase in positive axial velocity is evident at the location in which combustion takes place while tangential velocity remains relatively unchanged, thus decreasing the radial pressure gradient and ending the recirculation zone. As the stagnation point, followed by the combustion front, regresses toward the entrance of the reactor, the stagnation point starts being dominated by the inlet (burner) hydrodynamics and pierces the combustion front. When sufficient hot gases are recirculated so fresh coal devolatilization occurs inside the backflow core, the transition to NBC takes place. This observation is vital since it implies that each of the two combustion structures is governed by a different mechanism. In the case of NBC the burner design, the hydrodynamics and the mixing control the combustion location while in the case of FRC the kinetics control the combustion location. In this study where volumetric reactions are based on chemical equilibrium, the kinetics refers to the devolatilization rate. This notion is demonstrated in Figure 4 where the cases of FRC are visibly dominated by the devolatilization rate, while the same parameter has negligible effect in the cases of NBC. Understanding these two states is crucial regarding the ability to control them. This can have profound effect on designing and operating reactors for the variety of processes described earlier. 3.2. Parametric Investigation. Assuming swirl combustion behavior is dominated by the recirculation and backmix of hot gases combined with its interaction with fresh incoming pulverized coal, the parameters chosen for the study are (a)

Figure 4. Temperature at furnace axis of different devolatilization activation energy exemplifying the “jump” between FRC at higher activation energies and NBC at lower activation energies.

activation energy for the devolatilization rate (E1 is gradually decreased from an initial value of 1 × 108 J/kmol, the continuous response occurs only up to the point of combustion region change from FRC to NBC with a lack of mediate state. In order to properly elaborate the discussion, two entities are defined: the backflow core and the stagnation point. The backflow core is defined as an enclosed volume containing negative axial velocity next to the reactor’s inlet and the stagnation point is defined as the farthest location along the reactor’s axis with negative axial velocity, assuming axisymmetric behavior. Both can be observed in Figure 3. It was evident that in most cases of FRC the stagnation point is located in the vicinity of the combustion front, and only a small amount of hot gases are recirculated (backflow core consist of more than 90% fresh inlet gases). The relationship between the stagnation point and the combustion front can be drawn from D

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels particle injection velocity, (b) particle mean diameter, and (c) secondary air inlet temperature. Usually, one would have control over some of those parameters and could choose which is preferably altered to achieve a steady process. Clearly, many more parameters can control the combustion structure from the operating conditions and material properties point of view, but mostly they will be predetermined, not easily changed, or totally unchangeable. A good example for a predetermined property is the set of parameters affecting the devolatilization kinetic rate discussed in Section 3.1. In order to investigate the effect of different parameters on the transition between combustion structures, a baseline case was created suitable for FRC at a state not far from the transition to NBC. The operating conditions and material properties for the baseline case were deliberately chosen not to promote NBC so that when a specific parameter is altered it will need to solely compensate for other NBC discouraging parameters to achieve transition. Table 3 contains relevant Table 3. Operating Conditions for Baseline Case parameter secondary air temperature K particle injection velocity m/s particle mean diameter μm particle Spread devolatilization A1 1/s devolatilization E1 J/kmol wall temperature K

450 15 240 2 1.66 × 106 8.4 × 107 1100

parameters for the baseline case where A1 and E1 are the preexponent factor and activation energy of the first devolatilization rate, respectively. Those were altered from the validation case to match the above description. Wall temperature was set to match the wall temperature for which ignition was obtained for the experiments that were carried out in the reactor. During the experiments, the walls are heated to 1100 K with diesel combustion before the pulverized coal is introduced to the reactor. Thus, when the coal is subsequently injected, the heat flux from the wall acts as an ignition mechanism. Parameters not appearing in Table 3 can be assumed to have the same value as for the validation case. It is important to notice that in the present work only the transition from FRC to NBC was investigated. Since NBC is assumed to be a robust structure, it is prone to exemplify hysteresis and results can only be regarded to one way transition. From a numeric point of view, this problem is initial condition dependent, thus all simulations were initialized from the baseline case described. 3.3. Results. In order to focus on the recirculation zone characteristics, the average temperature of the backflow core was calculated in addition to the location of the stagnation point at its end. Figure 5 shows the results of the parametric investigation for the transition from FRC to NBC. First of all, the phenomenon described in Section 3.1 is evident quantitatively. Also, the abrupt nature of the transition is observed. It is clear that all three parameters were able to generate transition between combustion structures at some point within the investigated range, a range which was supposed to resemble practical values. When the transition occurs, we see an instant regression in the stagnation point of 10−15 cm and a rise of 500−600 K in the average backflow temperature, which means the backflow core is shorter and full of hot gases, compatible with the NBC state. It can be deduced

Figure 5. Average backflow temperature and stagnation point response to (a) particle injection velocity, (b) particle mean diameter, and (c) secondary air temperature. Δ Represents the baseline case detailed in Table 3.

that during the transition, the flame front is being relocated to the reactor’s entrance from a previous position of more than 0.5 m downstream. 3.3.1. Transition Based on Particle Mean Diameter and Injection Velocity. Understanding how particle mean diameter and injection velocity influence the combustion structure (Figure 5a,b) can come from examining the particle-gas behavior. Particles are injected from primary air inlets with a certain velocity and immediately face counter flow from the recirculation zone. Here an examination via Stokes number is suggested. The Stokes number represents the ratio of the characteristic time of the coal particles to the characteristic time of the flow that is evaluated by tf = u/D. Stk = τ ·u/D

(5)

Here u is the free stream velocity which in this case was taken as the injection velocity of the particle, D is the characteristic length which is the reactor’s diameter, and τ is the particle’s relaxation time calculated with Schiller and Naumann’s drag coefficient: τ = ρp d p2/18μ(1 + 0.15Re 0.687 ) p E

(6) DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Here ṁ p,0 is the coal feed rate at the entrance of the reactor. The relevant Yd is returned to eq 1 for the diameter in Stk particle relaxation time. 3.3.2. Transition Based on Inlet Temperature. Returning to the effect of inlet temperature demonstrated in Figure 5c; retaining the same material properties and considering hydrodynamics do not change drastically, the particle trajectories and residence time do not diverge much either. Hence, the remaining factor determining the combustion front is the distance needed for the gas and particles near the wall to reach a critical ignition temperature, therefore directly related to the inlet temperature. Here also the amount of combustible mass needed to generate transition changes at each case since higher inlet temperature contributes to the energy balance. It should be noted that aside from the transition itself, only mild changes are evident in average backflow temperature and stagnation point for the examined range as compared to the other studied parameters. 3.3.3. Effects on Combustion Efficiency. In order to look at the reactor efficiency, the fraction of unburned fuel along the reactor is illustrated in Figure 7 for various cases regarding

Velocity for Rep was again taken as the injection velocity of the particle. Smaller particles with lower injection velocities, namely low Stk number, are more likely to be suspended in the recirculation zone, then follow the stream line with tangential acceleration bringing the unburned portions closer to the wall by centrifugal forces, thus being heated either by the recirculating zone or the wall. Larger particles with higher injection velocities, namely high Stk number, are more likely to be shot through the recirculating zone, and then gain enough tangential velocity to meet the wall in a deeper part of the reactor, thus responding in a lesser degree to both wall and recirculation heating mechanisms. A comparison of the effect of particle diameter and injection velocity on the transition between the NBC and FRC structures, as a function of Stk number and normalized, averaged, backflow temperature can be seen in Figure 6 where both parameters achieved transition

Figure 6. Average backflow temperature normalized by the stoichiometric adiabatic flame temperature (Tst = 2226.1 K), with different Stk numbers changed by the particle’s injection velocity and the particle representative diameter.

approximately at same Stk values (Stk ∼1). Since the pulverized coal was said to have a specific distribution illustrated in eq 1, some attention needs to be dedicated for choosing the single diameter that represent the relaxation time for a given case. The mean diameter of the distribution might be an easy choice but it lacks physical relation to the described phenomenon. The following methodology is proposed for diameter selection. The chosen diameter should be one that has the ability to account for a transition between the combustion structures. It is fair to say that the progression of particle incineration is from small to large. Hence, the diameter that should be considered is the one containing enough combustible matter to reach transition temperature (Ttr) inside an adiabatic well-mixed control volume, representing the backflow core. To simplify the discussion, Ttr will be considered as an empirical constant drawn from Figure 5 (Ttr = 1000K). From the mixture fraction model, Ttr = T(f tr) is predictable and dictated by chemical equilibrium inside the control volume by the ratio of air stream Ȧ and fuel stream Ḟ tr, f tr = f(Ḟ tr, Ȧ ) as in eq 2. Here Ȧ can be taken as the total air mass flow rate at the inlets and Ḟ tr as the mass flow rate of the combustible portion of the coal up to the particle diameter which will represent the case. Thus, we can write the relation between Ḟtr and Yd as Ftṙ = (mp,0 − ma ) ·(1 − Yd) ·ṁ p,0 /mp,0

Figure 7. Fraction of unburned fuel along the reactor for selected cases of (a) different particle injection velocity, (b) particle mean diameter, and (c) secondary air temperature.

(7) F

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels different values of particle injection velocity, particle mean diameter and secondary air temperature. In Figure 7, the effect of the two combustion structures is evident and bears similarity to an analysis for combustion in PFR vs combustion in continuous stirred tank reactor (CSTR) and PFR in series found in literature27 and illustrated in Figure 8. An ideal CSTR

Figure 8. Fraction of unburned fuel in ideal reactor models for Anthracite combustion. (a) PFR model and (b) combined CSTR-PFR in series model. Based on the work of Beer and Lee.27

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is based on the assumption of complete mixing and can correlate to the mixing within the recirculation zone at entrance of the reactor. Considering this analogy, the FRC resembles PFR behavior and the NBC resembles the behavior of CSTR and PFR in series. This similarity emphasizes the nature of the recirculation zone in FRC which is in fact, nonreactive mixing. Such feature is valuable recalling that the entrained flow downer reactor is used, in some processes thanks to its ability to achieve near PFR behavior from an initial mixed state. It is interesting to notice that the efficiency at the outlet zone of the reactor shows almost no response to high inlet temperatures (in Figure 7c) in comparison to the effect of particle injection velocity and particle mean diameter (seen in Figure 7a,b) on the fraction of unburned fuel. Progress in efficiency at the latter part of the reactor can be majorly related to the diffusion rate dominating the resistance on char combustion for the larger particles (namely proportional to 28 d−1 Since p ) as suggested and modeled by Baum and Street. these large particles account for the remaining unburned fuel at the end of the reactor, manners for compensating on limitations of diffusion rate can be in the form of longer residence time or smaller particles (as seen in Figure 7a,b). Recent work29 on biochar gasification showed similar dependency of particle mean diameter on carbon conversion. Increasing the air inlet temperature hardly effects residence time or diffusion limitation at the end of the reactor.

• The dominating mechanism for the transition from FRC to NBC seems to be the amount of hot recirculating gases, and the trajectories of fresh particles at the recirculating zone and near the wall. • Cutoff values of Stk number close to unity can define the transition from FRC to NBC for parameters affecting the gas−solid behavior. Methodology for the selection of a distribution representative particle diameter was presented. It is assumed that the suggested methodology could be applied on other types of two phase tubular exothermic reacting flows. • Differentiation between NBC governed by inlet hydrodynamics and FRC governed by devolatilization kinetics was made. • Limitation on the fraction of unburned fuel at the latter part of the reactor is mainly dominated by diffusion rate for larger particles. • Findings regarding instrument utilization and transition between combustion regions are expected to be applicable for a broad range of processes involving two phase swirling reacting flows. • Controlling these two combustion regions can be used as a tool to control the level of complete mix vs short contact when operating a downer reactor.

AUTHOR INFORMATION

Corresponding Author

*Telephone: +972 509449962; E-mail: [email protected]. ORCID

Tal Eluk: 0000-0002-4791-9343 Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors wish to acknowledge the support of CHE-IAEC research grant by the Pazy foundation and to thank Mr. Stas Zinchik for his assistance at the early stages of this work.

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REFERENCES

(1) Zhu, J.; Yu, Z.; Jin, Y.; Grace, J.; Issangya, A. Cocurrent downflow circulating fluidized bed (downer) reactors-a state of the art review. Can. J. Chem. Eng. 1995, 73, 662−677. (2) Kim, Y.; Lee, S.; Kim, S. Coal gasification characteristics in a downer reactor. Fuel 2001, 80, 1915−1922. (3) Cheng, Y.; Wu, C.; Zhu, J.; Wei, F.; Jin, Y. Downer reactor: From fundamental study to industrial application. Powder Technol. 2008, 183, 364−384. (4) Wu, C.; Cheng, Y.; Ding, Y.; Jin, T. CFD-DEM Simulation of Gas-Solid Reacting Flows in Fluid Catalytic Cracking (FCC) Process. Chem. Eng. Sci. 2010, 65, 542−549. (5) Talman, J.; Geier, R.; Reh, L. Development of a downer reactor for fluid catalytic cracking. Chem. Eng. Sci. 1999, 54, 2123−2130. (6) Deng, R.; Wei, F.; Jin, Y.; Zhang, Q.; Jin, Y. Experimental Study of the Deep Catalytic Cracking Process in a Downer Reactor. Ind. Eng. Chem. Res. 2002, 41, 6015−6019. (7) Dong, P.; Wang, Z.; Li, Z.; Li, S.; Lin, W.; Song, W. Experimental Study on Pyrolysis Behaviors of Coal in a Countercurrent Downer Reactor. Energy Fuels 2012, 26, 5193−5198. (8) Cheng, Y.; Wang, C. Numerical Study on Coal Gasification in the Downer Reactor of a Triple-Bed Combined Circulating Fluidized Bed. Ind. Eng. Chem. Res. 2014, 53, 6624−6635. (9) Matsumoto, K.; Takeno, K.; Ichinose, T.; Ogi, T.; Nakanishi, M. Gasification reaction kinetics on biomass char obtained as a by-product

4. CONCLUSIONS The computational study exhibited the existence of two distinct combustion regions referred to as NBC and FRC for pulverized coal combustion in a downer reactor. • It was evident that recirculation can play a dominant part in the reactor’s behavior. • The transition demonstrated response to operating conditions, such as particle injection velocity, inlet air temperature, and particle mean diameter, that are, in most cases, under the control of the operator. G

DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX

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Energy & Fuels of gasification in an entrained-flow gasifier with steam and oxygen at 900−1000 C. Fuel 2009, 88, 519−527. (10) Korytnyi, E.; Saveliev, R.; Perelman, M.; Chudnovsky, B.; BarZiv, E. Computational fluid dynamic simulations of coal-fired utility boilers: An engineering tool. Fuel 2009, 88, 9−18. (11) Bar-Ziv, E.; Saveliev, R.; Korytnyi, E.; Perelman, M.; Chudnovsky, B.; Talanker, A. Evaluation of performance of AngloMafube bituminous South African coal in 550 MW opposite-wall and 575 MW tangential-fired utility boilers. Fuel Process. Technol. 2014, 123, 92−106. (12) Meyer, S.; Glaser, B.; Quicker, P. Technical, Economical, and Climate-Related Aspects of Biochar Production Technologies: A Literature Review. Environ. Sci. Technol. 2011, 45, 9473−9483. (13) Beer, J.; Chomiak, J.; Smoot, L. Fluid Dynamics of Coal Combustion: A Review. Prog. Energy Combust. Sci. 1984, 10, 177−208. (14) Chigier, N.; Beer, J. Velocity and static pressure distributions in swirling air jets issuing from annular and divergent nozzles. J. Basic Eng. 1964, 86, 788−79. (15) Kumar, M.; Ghoniem, A. Multiphysics Simulations of Entrained Flow Gasification. Part I: Validating the Nonreacting Flow Solver and the Particle Turbulent Dispersion Model. Energy Fuels 2012, 26, 451− 463. (16) Yin, C.; Kaer, S.; Rosendahl, L.; Hvid, S. Co-firing straw with coal in a swirl-stabilized dual-feed burner: Modelling and experimental validation. Bioresour. Technol. 2010, 101, 4169−4178. (17) Chen, L.; Ghoniem, A. Simulation of Oxy-Coal Combustion in a 100 kWth Test Facility Using RANS and LES: A Validation Study. Energy Fuels 2012, 26, 4783−4798. (18) Hashimoto, N.; Kurose, R.; Hwang, S. M.; Tsuji, H.; Shirai, H. A numerical simulation of pulverized coal combustion employing a tabulated-devolatilization-process model (TDP model). Combust. Flame 2012, 159 (1), 353−366. (19) Kewlani, G.; Shanbhogue, S.; Ghoniem, A. Investigations into the Impact of the Equivalence Ratio on Turbulent Premixed Combustion Using Particle Image Velocimetry and Large Eddy Simulation Techniques: “V” and “M” Flame Configurations in a Swirl Combustor. Energy Fuels 2016, 30, 3451−3462. (20) Spitz, N.; Saveliev, R.; Korytnyi, E. Prediction of Performance and Pollutant Emission from Pulverized Coal Utility Boilers. In Electric Power Generation, Transmission and Efficiency; Lefebvre, C. M., Ed.; Nova Science Publishers Inc.: New York, 2007; pp 121−170. (21) Fluent, A. Fluent 16.2 Theory Guide; 2015. (22) Yamamoto, Y.; Potthoff, M.; Tanaka, T.; Kajishima, T.; Tsuji, Y. Large-eddy simulation of turbulent gas-particle flow in a vertical channel: effect of considering inter-particle collisions. J. Fluid Mech. 2001, 442, 303−334. (23) Yin, C.; Yan, J. Oxy-fuel combustion of pulverized fuels: Combustion fundamentals and modeling. Appl. Energy 2016, 162, 742−762. (24) Chen, L.; Yong, S. Z.; Ghoniem, A. F. Oxy-fuel combustion of pulverized coal: Characterization, fundamentals, stabilization and CFD modeling. Prog. Energy Combust. Sci. 2012, 38 (2), 156−214. (25) Kobayashi, H.; Howard, J.; Sarofim, A. Coal Devolatilization at High Temperatures. Symp. (Int.) Combust., [Proc.] 1977, 16, 411−425. (26) Niksa, S.; Lau, C. Global rates of devolatilization for various coal types. Combust. Flame 1993, 94, 293−307. (27) Beer, J.; Lee, K. The effect of the residence time distribution on the performance and efficiency of combustors. Symp. (Int.) Combust., [Proc.] 1965, 10, 1187−1202. (28) Baum, M. M.; Street, P. J. Predicting the combustion behaviour of coal particles. Combust. Sci. Technol. 1971, 3 (5), 231−243. (29) Ku, X.; Li, T.; Løvås, A. T. Eulerian−Lagrangian Simulation of Biomass Gasification Behavior in a High-Temperature Entrained-Flow Reactor. Energy Fuels 2014, 28, 5184−5196.

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DOI: 10.1021/acs.energyfuels.6b02225 Energy Fuels XXXX, XXX, XXX−XXX