Influences of Reactant Injection Velocities on Moderate or Intense Low

Sep 30, 2013 - The present work numerically investigates the effects of injection velocities or momenta per unit mass of the primary and secondary rea...
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Influences of Reactant Injection Velocities on Moderate or Intense Low-Oxygen Dilution Coal Combustion Zhenfeng Mei, Pengfei Li, Feifei Wang, Jianpeng Zhang, and Jianchun Mi* State Key Laboratory of Turbulence and Complex Systems and Department of Energy and Resources Engineering, College of Engineering, Peking University, Beijing 100871, People’s Republic of China ABSTRACT: The present work numerically investigates the effects of injection velocities or momenta per unit mass of the primary and secondary reactant streams on the moderate or intense low-oxygen dilution (MILD) combustion of pulverized coal. When the injection-nozzle diameters are changed, the inlet velocity of the primary stream with the inlet temperature of 313 K is varied between 26 and 99 m/s, while that of the secondary air (highly preheated) with the temperature of 1623 K is increased from 16 to 102 m/s. The modeling is verified by the measurements of Weber et al. in the International Flame Research Foundation (IFRF) furnace [Weber, R.; Smart, J. P.; van der Kamp, W. Proc. Combust. Inst. 2005, 30 (2), 2623−2629]. Results reveal that the primary reactant velocity exerts a stronger influence on the flame temperature and NO emission than the secondary air velocity. Considerable reductions of the maximal temperature (about 180 K) and NO emission (200 ppm) are obtained by raising the primary reactant velocity from 26 to 67 m/s. Under the MILD combustion, the sum of the contributions from the thermal-NO, prompt-NO, and N2O-intermediate routes only accounts for less than 2.5% to the total NO emissions. Although it is well-known that the fuel-NO route dominates the NOx emissions from coal combustion, our prediction shows that up to 10% is reduced through the NO-reburning mechanism. Moreover, only about 15% of fuel N is found to convert to fuel NO from the MILD combustion, which is much less than that from the conventional combustion.

1. INTRODUCTION Flameless oxidation (FLOX),1 high-temperature air combustion (HiTAC),2−4 and moderate or intense low-oxygen dilution (MILD) combustion5 may be grouped as the same type of recently developing combustion technologies for the purpose of simultaneously achieving low emissions (e.g., NOx) and high efficiency of industrial combustion (for convenience, we treat the three uniformly as “MILD combustion” in the present paper, although they are not exactly identical or equivalent6). In the past 2 decades, many experimental and numerical investigations have been conducted to find out effective ways to implement this technology in furnaces. In the MILD combustion of gaseous fuels (and also light oils), chemical reactions take place in a nearly entire volume of the combustion chamber and no flame is visible; both temperature and some species concentrations are very uniform; the net radiation flux is enhanced and combustion reactions are stable; NOx emissions are extremely low; and also there is no audible combustion noise.1−15 As the development of MILD combustion proceeds, this technology has been gradually applied to the pulverizedcoal combustion for high thermal efficiency and low NOx emissions.1−3 It is worth noting that, similar to the case of firing gaseous fuels, low-temperature gradients in the furnace are required for the occurrence of MILD combustion of pulverized coal. However, the appearance of MILD coal combustion is different: a high amount of char and ash present in coal makes coal flames distinguishable from the wall radiation and thus clearly visible.16,17 Thus far, several experimental and numerical studies have been performed on coal MILD combustion.16−22 Suda et al.18 applied high-temperature air to a pulverized-coal burner to investigate the effect of the air temperature on pulverized-coal ignition, burnout, and NOx emission in a furnace (1 m in © XXXX American Chemical Society

diameter and 3 m in height). The temperatures of the combustion air were set to 623 or 1073 K, and they found delayed ignition, improved coal burnout, and decreased NOx emissions for the 1073 K case. Later, He et al.19 numerically studied the NO evolution with a simplified chemical reaction model, in which a low NO emission mechanism in the HiTAC was proposed. Their numerical results revealed that the HCN concentration distribution in the mixing region plays an important role in achieving low NO emission. Ponzio et al.20 investigated coal pellet combustion using a high-temperature (873−1073 K) oxidizer with varying oxygen concentrations (0−100%). The influences of the oxygen concentration on the ignition mechanism, the temperature inside the particle at the ignition, the mass lost in the ignition, and ignition time were analyzed and discussed. Zhang et al.21 conducted a series of experiments in an industrial-scale test facility, burning lowvolatile petroleum coke and an anthracite coal, with a specially designed primary air enrichment and preheating (PRP) burner. They found that rapid heating of the combustible mixture in the chamber facilitates pyrolysis and volatile matter release processes of coal particles, suppressing ignition delay and enhancing combustion stability. Weber et al.16 examined the fundamental and industrial application aspects of combustion of natural gas, heavy and light fuel oils, and coal in hightemperature preheated air. Importantly, a high-volatile coal Special Issue: 4th (2013) Sino-Australian Symposium on Advanced Coal and Biomass Utilisation Technologies Received: July 31, 2013 Revised: September 30, 2013

A

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Figure 1. Geometry of the experimental furnace of Weber et al.16

Table 1. Experimental Conditions of Weber et al.16 mass flow (kg/h) coal primary air secondary air

66 130 675

velocity (m/s)

temperature (K)

enthalpy (MW)

composition (wt %, wet)

0.58 26 65

313 1623

0.30

(1.49% N) was combusted, and the lowest NOx emissions were in the range of 160−175 ppm (at 3% O2). Following Weber et al.,16 Schaffel et al.22 successfully simulated the coal experiments. Both the chemical percolation devolatilization (CPD) model and the char combustion intrinsic reactivity model were adapted to Guasare coal combustion. Results showed that the velocity distribution and the temperature and oxygen distributions were quite accurately predicted for the MILD combustion of the high-volatile coal by the computational fluid dynamics (CFD)-based model. Then, Vascellari et al.23 examined an advanced model based on finite rate chemistry eddy dissipation concept (EDC) with both global and detailed kinetic mechanisms. Their comparisons of numerical simulations with experimental measurements indicated that advanced turbulence−chemistry interaction models (i.e., EDC) used with detailed kinetic mechanisms (i.e., DRM24) predict better the chemical and fluid dynamic behaviors of coal MILD combustion. However, they also suggested that the EDC model with global mechanisms (i.e., Jones and Lindstedt global multi-step combustion mechanism25) is preferred, owing to its lower computational cost and satisfactory results. All of the studies16−25 noted above indicated the advantages of MILD coal combustion, especially the character of low NOx emission. However, to offer a useful guidance for industrial applications, there is a need to investigate the influences of various injection conditions (e.g., inlet velocities and locations of the primary and secondary jets) on the establishment and characteristics of MILD coal combustion. The present study is to address partly this need by the Reynolds-averaged Navier− Stokes (RANS) modeling, instead of experimental measure-

O2 = 23; N2 = 77 O2 = 22; H2O = 9.5; CO2 = 12.5; N2 = 56; NO = 89 × 10−4

ments, which are not feasible or even impossible for many cases. More specifically, the main aim of the present work is to numerically study the effects of inlet velocities (v*sec and v*pri) of the primary air/pulverized-coal stream (at room temperature) and secondary (highly preheated) air on the MILD coal combustion occurring in the furnace of Weber et al.16 The velocities v*sec and v*pri are varied by changing the nozzle diameters. The RANS modeling with EDC is employed for this study; all simulations use the three-reaction kinetic mechanism and the global multi-step mechanism, respectively, for char surface oxidation and volatile combustion (see more details in section 2). The validation is performed using the measurements of Weber et al.16 Moreover, we also investigate the characteristics of the thermal-NO, prompt-NO, N2O-intermediate, fuelNO, and NO-reburning mechanisms for the MILD coal combustion.

2. COMPUTATIONAL DETAILS 2.1. Furnace Configuration and Operational Conditions of the Present Simulations. The present study simulates a set of coal combustions occurring in the furnace of Weber et al.16 Figure 1 shows the detailed furnace configuration and dimensions. The primary air streams carrying the pulverized coal issue from the two horizontal inlets into the furnace, while the secondary air enters the furnace from the central inlet. In the experiment of Weber et al.,16 the measurements were taken at seven traverse locations (see Figure 1). The experimental conditions and the characteristics of the Guasare coal are given in Tables 1 and 2, respectively. The burner was operated at 0.58 MW fuel input, and the primary air contained 23% O2 (by weight) with the temperature at 313 K. The secondary air temperature was preheated to 1623 K and contained 22% O2 (wet, by weight) and 89 ppm NO (wet, by weight). For the present simulations, the primary air with pulverized-coal discharges into the furnace at the inlet velocity B

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Table 2. Characteristics of the Guasare Coal of Weber et al.16 composition proximate analysis

ultimate analysis (dry and ash-free basis)

Table 4. Parameters for the CPD Devolatilization Model of Guasare Coal22

wt %

moisture (105 °C) volatile matter fixed carbon ash LCV (MJ/kg) composition

coal

char

2.9 37.1 56.7 3.3 31.74 volatile matter

C H O N S

81.6 5.5 10.7 1.5 0.6

92.6 1.3 4.0 1.7 0.4

72.51 9.10 16.3 1.3 0.8

Table 3. Guasare Coal Particle Distribution Parameters22 parameter

value

unit

42 300 10 1.36

μm μm μm

value

unit

initial fraction of bridges in coal lattice initial fraction of char bridges lattice coordination number cluster molecular weight side-chain molecular weight

0.5 0 5 300 30

kg/kmol kg/kmol

H2O, and O2) to the char surface is set at a value of 5 × 10−12 m3 K−0.75 s−1. Moreover, we take the char porosity = 74%, the swelling coefficient = 2, the particle emissivity = 0.8, and the reaction heat fraction absorbed by char particle = 30%. The dispersion of particles because of turbulence is taken into account by considering the stochastic tracking model (i.e., discrete random walk model). Besides, as suggested by Vascellari et al.,23 the Jones and Lindstedt global multistep combustion mechanism,25 shown in Table 5, is employed to simulate the volatile matter reacting in the gaseous phase, owing to its lower computational cost with respect to the detailed mechanisms. In the present simulations, the composition of the volatile matter is set as CH4/CO/H2 = 47/48/5%. Schaffel et al.22 employed the eddy break-up (EBU) model with the simplified two-step reaction mechanism of the volatile oxidation (C1.2H4.48O0.44 + 1.5O2 → 1.2CO + 2.24H2O and CO + 0.5O2 → CO2) and the simplest one-reaction kinetic mechanism of the char oxidation (C(s) + O2(g) → CO2(g)) and gained a satisfactory agreement with the measurement.16 On the other hand, Vascellari et al.23 demonstrated that the EDC model with the global multi-step combustion mechanism for the volatile matter and multiple surface reactions for char together are more accurate and suitable to reproduce the MILD coal combustion process. Thus, the latter models are preferred in the present computations. In the present work, both the formation and reduction of NOx are evaluated using the post-processing approach. The turbulent effects on the temperature fluctuations are also taken into account with the probability density function (PDF). The formation of NOx is attributed to four distinct chemical kinetic processes, i.e., thermal NO, prompt NO, fuel NO, and intermediate N2O. Also, the NOreburning mechanism is considered to predict the reduction of NOx. The modeling details can be found in refs 22 and 23, and only a brief description is provided below. 2.2.1. Thermal-NO Route. The thermal-NO route is determined by the extended Zeldovich mechanism,35 i.e.

varying between 26 and 99 m/s, while the secondary air is injected at the inlet velocity of 16.3 to 102 m/s. A highly volatile bituminous coal “Guasare” with 1.5% N (dry and ash-free basis) is combusted. 2.2. Computational Models. Because of the symmetry of the furnace system, only one-quarter of the geometry is employed to reduce the computational expense and a three-dimensional (3D) grid composed of about 1 190 000 hexahedral cells is constructed. The commercial software FLUENT 6.3 has been used for the present work. The SIMPLE algorithm method is used to solve the pressure−velocity coupling. To improve the accuracy of the simulations, the secondorder scheme is employed for pressure and the second-order upwind scheme is employed for momentum, turbulent kinetic energy, turbulent dissipation, species transport (i.e., O2, CO2, H2O, CO, CH4, H2, and NO) and energy. Solution convergence is obtained when (a) the residuals are less than 10−6 for the energy and 10−5 for all of the other variables and (b) the variations of the outlet temperature and velocity are allowed to be within 1.0 K and 0.1 m/s, respectively. The grid independency of the results is verified using a finer grid with 4 800 000 cells, and results (i.e., temperature and molar fractions of O2, CO2, CO, and NO) are kept in high consistency for different grids. To describe the particle size distribution of the pulverized Guasare coal used in the experiment,16 the Rosin−Rammler distribution function is adopted (see Table 3) and the 10−300 μm size range is

mean diameter maximum diameter minimum diameter spread parameter

parameter

O + N2 ⇔ N + NO

(1)

N + O2 ⇔ O + NO

(2)

N + OH ⇔ H + NO

(3)

It is worth noting that O2 and N2 distributions are provided by the previous combustion simulation, while the O and OH radicals are not contained in the kinetic mechanisms presented in Table 5. Thus, the equilibrium approaches are employed to determine the O and OH radical concentrations. 2.2.2. Prompt-NO Route. The actual formation of prompt NO involves a complex series of reactions and many possible intermediate species, and one of the accepted routes is given below.

divided into 20 size classes. The realizable k−ε turbulence model26 and the EDC27,28 are taken to consider the interaction between turbulent flow and chemical reaction; the P1 radiation model29 is used for radiation heat transfer. Besides, the absorption coefficient of the gas mixtures is assumed to be 1.5 m−1 and kept constant through the furnace volume, and the scattering is ignored. The CPD model30−32 is used to characterize the devolatilization behavior of a rapidly heated coal particle and to predict volatile rates, yield, and composition. The CPD parameters adopted by Schaffel et al.22 are shown in Table 4, which are also used in the present computations. After volatile matter is completely devolatilized, the char remaining in the coal particle reacts with the surrounding gas phase, which is usually composed of high CO2 and H2O and low O2 in coal MILD combustion. This implies that the simplest one-reaction mechanism of the char oxidation (C(s) + O2(g) → CO2(g)) used by Schaffel et al.22 may not be suitable for the low surrounding oxygen level. Thus, multiple surface reactions are selected for the combustion model, and the char surface reactions are shown in Table 5, including the char reactions with CO2 and H2O. The diffusion rate of the oxidant (CO2,

CH + N2 ⇔ HCN + N

(4)

N + O2 ⇔ NO + O

(5)

HCN + OH ⇔ CN + H 2O

(6)

CN + O2 ⇔ NO + CO

(7)

2.2.3. Fuel-NO Route. The fuel-NO route is originated from the element N in the char and volatile matter. The mass fractions of the char N and volatile N are presented in Table 2. Similar to Schaffel et al.,22 both the char N and volatile N are assumed to convert to HCN. C

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Table 5. Kinetic Mechanisms Used in the Computations (Units in m, s, kmol, J, and K)a

global multi-step combustion mechanism

a

β

A

reaction char multiple surface reactions

C(s) + 0.5O2 → CO C(s) + CO2 → 2CO C(s) + H2O → CO + H2 CH4 + 0.5O2 → CO + 2H2 CH4 + H2O → CO + 3H2 H2 + 0.5O2 → H2O H2O → H2 + 0.5O2 CO + H2O → CO2 + H2

5 6.35 1.92 4.4 3.0 6.8 1.255 2.75

× × × × × × × ×

10−3 10−3 10−3 1011 108 1015 1017 109

0 0 0 0 0 −1 −0.877 0

Ea 7.4 1.62 1.47 1.26 1.26 1.67 4.096 8.4

× × × × × × × ×

107 108 108 108 108 108 108 107

reaction order

reference

[O2] [CO2] [H2O] [CH4]0.5[O2]1.25 [CH4][H2O] [H2]0.25[O2]1.5 [H2]−0.75[O2][H2O] [CO][H2O]

33 and 34

25

Kinetic rate in Arrhenius form: k = ATβ exp(−Ea/RT), where R = universal gas constant.

Figure 2. Traverse distributions of numerical and experimental axial velocities at x ≤ 4.97 m. This HCN mechanism proposed by Smoot and Smith36 is described as follows:

2.2.4. Intermediate-N2O Route. This path for NO is predicted by the mechanism of Melte and Pratt,37 which is

char N → HCN

(8)

N2 + O + M ⇔ N2O + M

(12)

volatile N → HCN

(9)

N2O + O ⇔ 2NO

(13)

2.2.5. NO-Reburning Route. This mechanism occurring in the gaseous phase is based on the model proposed by Kandamby et al.,38,39 i.e.

The conversion rates of char N and volatile N via reactions 8 and 9 are dominated by the burning rate of char (combustion model: multiple surface reactions) and the devolatilization rate of the volatile matter from coal particles into the gas phase (devolatilization model: CPD), respectively. Then, HCN is subject to two competitive reaction paths, i.e.

HCN + O2 → NO + ...

(10)

HCN + NO → N2 + ....

(11)

NO + CH 2 → HCN + OH

(14)

NO + CH → HCN + O

(15)

NO + C → CN + O

(16)

The NO reduction on the char surface is also considered with the model by Levy et al.,40 i.e., NO reburning on the char surface D

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Figure 3. Traverse distributions of numerical and experimental temperatures at x ≤ 4.97 m.

d[NO] = − 2.27 × 10−3[NO]pe−142737/ RT CsABET dt

This feature is captured by the present simulation. It is also evident that the predicted temperature takes the local minima at z ≈ 0.28 m for x ≤ 1.32 m. This is expected because the primary air jets carrying pulverized coal are not preheated, with the inlet temperature of only 313 K, so that their heating inside the furnace takes time or distance to complete. However, unexpectedly, the measured temperature does not show the similar; perhaps the experimental primary jets were highly asymmetric (therefore yielding the high measured temperature) or the measurements themselves were quite flawed. In addition, Figure 3 demonstrates that the present simulations agree better with the measurements than the previous simulations for x ≤ 2.05 m. Similar to the temperature, both the present and Vascellari et al.’s23 predictions of concentrations of O2 and CO are somewhat different from the experimental data (see Figures 2−6). This is likely related to the ignition delay in the numerical simulations, so that the local temperature and CO concentrations are lower, while the O2 molar fraction is higher, than the experimental data at traverses 1 and 2. Surprisingly, the simulation of Schaffel et al.22 appears to perform better on the ignition than both the present simulation and that of Vascellari et al.23 This would be unexpected from the fact that Schaffel et al.22 employed the simplified two-step reaction mechanism of the volatile oxidation and the simplest onereaction kinetic mechanism of the char oxidation, while the more reliable multi-step combustion mechanism for the volatile matter and multiple surface reactions for char were used in the

(17)

where Cs is the concentration of particles and ABET is the pore Brunauer−Emmett−Teller (BET) surface area.

3. EXPERIMENTAL VALIDATION OF THE RANS MODELING To validate the modeling, Figures 2−6 show the comparisons made between the present computations and previous experiments16 of the MILD combustion of pulverized coal. The previous modeling results22,23 are also included for comparison. In Figure 2, the traverse profiles of the measured and calculated velocities at x ≤ 4.97 m are presented. Clearly, the predicted velocity on the centerline, from both the present and previous simulations,22,23 is generally higher than the measured velocity16 at x ≤ 1.32 m. Both Schaffel et al.22 and Vascellari et al.23 attributed this difference to the poor laser Doppler anemometer (LDA) measurements in the central jet. Nevertheless, the present simulation overall predicts the velocity field quite well at x ≥ 2.05 m in the furnace. Figure 3 shows the traverse distributions of the temperature at x ≤ 4.97 m. Schaffel et al.22 obtained the maximal temperature of around 1800 K, while it exceeds 2000 K from the prediction of Vascellari et al.23 The corresponding value from the present modeling is 1950 K, which is in between the above two. As shown by Weber et al.,16 the measured maximum temperature occurs somewhere between x = 0.44 and 0.735 m. E

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Figure 4. Traverse distributions of numerical and experimental O2 concentrations at x ≤ 4.97 m.

present simulation and that of Vascellari et al.23 This difference may reflect that the reaction rates predicted by the simplified kinetic mechanisms are generally higher than those by multistep kinetic mechanisms (see ref 14), so that the ignition delay is compensated in the former case. This has also been noted in ref 22. Despite the above, the variation trends of those quantities, overall, agree reasonably well with the experimental data. Particularly, no substantial difference of the CO distribution between the modeling predictions and the measurements can be seen in the region near the furnace exit, and this reveals that the combustibles are almost completely burnt out at x = 4.97 m (traverse 7). NOx emissions are also modeled, and their comparison to the measurements is presented in Figure 6. The NO peak from Vascellari et al.23 is nearly 2000 ppm compared to the experimental value of about 900 ppm. The present simulation and also that of Schaffel et al.22 predict well, even though the predicted location (near traverse 3) differs from the experimental location (near traverse 2). Significantly, quite close to the measurement (320 ppm), the predicted NO emission at the furnace exit is 322 ppm from the present simulation and 333 ppm from Schaffel et al.22 Moreover, the satisfactory predictions of the velocity, temperature, and concentrations at x = 4.97 m (Figures 2−6) ensure that the predicted NO emission at the furnace outlet is quite reliable. Above all, the present predictions of the velocity, temperature, and species concentrations are generally in good

agreement with the measurements of Weber et al. 16 Accordingly, the RANS modeling using the EDC model with the Jones and Lindstedt global multi-step mechanism25 can approximately characterize the MILD combustion of pulverized coal and, therefore, be used to investigate appropriately the effects of the inlet velocities of the primary and secondary streams on the flame characteristics.

4. EFFECTS OF INJECTION VELOCITIES OF THE PRIMARY AND SECONDARY JETS In the experiment of Weber et al.,16 the inlet velocities of the primary and secondary air jets were taken as v*pri = 26 m/s and v*sec = 65 m/s, respectively. With the two velocities, these authors achieved the MILD coal combustion and, therefore, obtained a more homogeneous temperature field than that normally achieved in the conventional combustion. The present study has interest to see whether the MILD combustion mode can still maintain when the injection velocities of the primary and secondary air jets are varied. 4.1. Effect of the Inlet Velocity of the Secondary Air (v*sec). To investigate the v*sec effect on coal MILD combustion (see Table 6), v*sec is varied from 16.3 to 102 m/s by decreasing the inlet diameter (Dsec) from 0.25 to 0.1 m. The mass flow rate (675 kg/h) and temperature (1623 K) of the secondary air jet are identical to those of Weber et al.16 Also, the parameters of the primary air are kept the same as those in the experiment.16 F

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Figure 5. Traverse distributions of numerical and experimental CO concentrations at x ≤ 4.97 m.

To check the flow process in the furnace, Figure 7 shows the velocity vector field in the x−z plane for cases 1−5. Because the velocity vector varies dramatically in both magnitude and direction, we found it better to display the velocity magnitude by the vector color than by the length, as performed in Figure 7. It is evident that the primary and secondary jets issue into the furnace individually from the separate nozzles and entrain their surrounding flue gases (low oxygen concentration) immediately downstream from the nozzle exits. Because of the low initial temperature of the primary air and pulverized coal (313 K), before the ignition, the primary air is diluted and heated by the entrained surrounding hot flue gases. The dilution level is important for the establishment of MILD combustion, as demonstrated in refs 12 and 13 for the CH4 combustion. Because the primary air mass is much less than that of the secondary air, the pulverized coal is far from burning out before the confluence of the primary and secondary air streams. The primary and secondary streams are separated in sufficient distance and, therefore, produce two reaction zones apart; i.e., one is dominated by the primary (air + pulverized coal) jet, and the other is dominated by the secondary air, as indicated late in Figure 9. Moreover, from Figure 7, in the later segment of the furnace, the presence of the back wall leads to a strong largescale internal flue gas recirculation. As v*sec is decreased, the recirculation weakens and its effective region shifts downstream. The presence of the large-scale recirculation may help

enhance the oxygen dilution and the uniformity of the temperature and species distributions. Note that, even in case 5 with the lowest v*sec, the inlet momentum of the secondary air jet is about 3.25 times the total momentum of the primary jets. Because the central secondary air jet is much stronger than the side primary jets, the latter is “sucked” toward the centerline and eventually merges into the former. As v*sec or the secondary air jet momentum is increased (Figure 7), the confluent points apparently move upstream, therefore weakening the dilution process of each jet. On the other hand, the higher momentum of either the primary or secondary jet is expected to entrain more flue gases and, thus, speed up the large-scale recirculation, enhancing the dilution of reactants. That is, increasing v*sec has two competitive opposite effects on establishing the environmental condition for MILD coal combustion. Also worth noting is that, when the injection velocity is varied by changing the injection nozzle diameters, a method similar to that for firing natural gas,12,13 the effective separation between the primary and secondary jets is altered as well. For instance, to increase v*sec from 16 to 102 m/s, the secondary nozzle diameter has to be decreased from Dsec = 0.25 to 0.1 m, and consequently, the effective separation between the secondary and primary nozzles has to be enlarged from 0.141 to 0.216 m. These variations delay the confluence of the primary and secondary streams and, thus, the occurrence of MILD combustion, which is confirmed below in Figure 10. G

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Figure 6. Traverse distributions of numerical and experimental NO concentrations at x ≤ 4.97 m.

Table 6. Injection Conditions with Different Diameters of the Secondary Air Nozzle (Dsec) experiment16 case 1 case 2 case 3 case 4 case 5

secondary secondary secondary secondary secondary secondary

air air air air air air

mass flow (kg/h)

temperature (K)

velocity, v*sec (m/s)

inlet diameter, Dsec (m)

675 675 675 675 675 675

1623 1623 1623 1623 1623 1623

65 102 65 40 25.4 16.3

0.125 0.100 0.125 0.160 0.200 0.250

mass flow rate (kg/s) of the flue gas entrained (mr) by the fuel and oxidant jets must be equal to that of the recirculated gas downstream. The recirculated gas downstream can be wellidentified by the negative value of the local x velocity. Thus, mr at different x is calculated by

Both Schaffel et al.41 and Mi et al.13 have noted the effect of changing the nozzle diameters on the nozzle separation and found the important role played by the separation for the occurrence of MILD combustion in combustors of different configuration. Hence, it is significant to investigate in future the impacts on MILD coal combustion of the locations of the primary and secondary jets (or their separation) and other important parameters (e.g., injection angle). To quantify the dilution level, we employ the recirculation rate (Kv) first defined by Wünning and Wünning,1 viz. mr Kv = ma + mf (18)

m r (x ) =

∬A(x) ρvx(y , z)dydz

(19)

where A(x) is the area for vx(y,z) < 0. Figure 8 shows the recirculation rate Kv at traverses 1−6 for cases 1−5. The relationship between Kv and v*sec is nearly linear; when v*sec is increased from 16 to 102 m/s, Kv increases by about 1.5 times. That is, the entrainment of the surrounding flue gases is definitely enhanced by increasing v*sec. Moreover, as expected, the growth rate of Kv increases with x. Figure 9 shows the effect of v* sec on temperature distributions in the x−z plane. This diagram suggests that

where mr is the mass flow rate (kg/s) of the entrained flue gas and ma + mf denotes the injection total mass flow rate (kg/s) of the injecting reactants (i.e., air and fuel). According to the mass conservation law, at each cross-section of the chamber, the H

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Figure 7. Velocity vector fields in the x−z plane of the furnance flow for cases 1−5.

suppressed over there. Obviously, the secondary high-temperature region is located at the border or confluent between the primary and secondary jets. It can be inferred that, from Kv = 1.8 at x = 1.32 m (in the secondary reaction zone) for case 5 with v*sec = 16.3 m/s, the oxygen volume fraction is lower than 11%. In other words, in the secondary reaction zone, there is a low oxygen environment and, therefore, the maximal temperature is also suppressed. It is worth noting that increasing v*sec from 16.3 to 102 m/s results in a significant increase of Kv from 1.8 to 4.0 (Figure 8) but a small decrease (