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Experimental and Modeling Study on the Flame Structure and Reaction Zone Size of Dimethyl Ether/Air Premixed Flame in an Industrial Boiler Furnace Yinhu Kang,† Xiaofeng Lu,*,† Quanhai Wang,† Xuanyu Ji,‡ Shanshan Miao,† Jie Xu,† Guangyu Luo,§ and Hai Liu§ †

Key Laboratory of Low-grade Energy Utilization Technologies and Systems (Chongqing University), Ministry of Education, Chongqing 400044, People’s Republic of China ‡ College of Mechanical and Power Engineering, Chongqing University of Science & Technology, Chongqing 401331, People’s Republic of China § Nanxi Boiler Co. Ltd. Yibin 644104, Sichuan, People’s Republic of China ABSTRACT: The flame structure and reaction zone size of the dimethyl ether/air premixed flame in a cylindrical furnace were experimentally and numerically investigated in this paper. Seven cases of flames with different operational parameters were involved to reveal the influences of excess air ratio and thermal load on the combustion behaviors. The simulations were conducted using the eddy dissipation concept (EDC) model with a reduced chemical kinetic mechanism including 39 species and 168 reversible reactions. The present work demonstrated that the fluid structure in the furnace consisted of the flame core area (FCA), recirculation zone I (RZ I), and recirculation zone II (RZ II). In addition, the fluid structure was coupled with heat and mass transfer phenomena in the furnace and had a considerable effect on them. The temperature in the FCA was mainly dependent on the excess air ratio. The temperature in RZ II was strongly affected by the thermal load. Moreover, the temperature at the furnace outlet was positively correlated with the thermal load. The species contents in the overall furnace were mainly dependent on the excess air ratio. The influence of the thermal load over the species contents was rather insignificant. Moreover, the intensified diffusion outside the flame zone resulting from the decrement of temperature could change the species distributions in part. Research on the reaction zone size indicated that either the decrease of excess air ratio or the increase of thermal load could result in an enlargement of the reaction zone. Additionally, the mean reaction rate of the dimethyl ether/air premixed flame was found to be independent of the thermal load. It was observed to be dependent on the excess air ratio. Finally, a functional expression between the mean reaction rate and the excess air ratio was developed in this paper.

1. INTRODUCTION In the past few decades, massive consumption of fossil fuels and the pollution problems have driven the development of environmental friendly and alternative fuels across the world. Dimethyl ether (DME, CH3OCH3) is the simplest ether, consisting of two methyl groups and one oxygen atom. It can be produced from the conventional energy resources with heavy pollution (for example, coal) in large amounts or exists as a byproduct of the chemical industry. DME has no C−C bonds in its molecular structure (and, therefore, has low soot emission1,2 when used in engines), a high cetane number of 55−60 (and, therefore, has excellent self-ignition performance1−3), a low boiling point of −25 °C in the atmosphere pressure (and, therefore, has almost instantaneous vaporization upon injection into the cylinder2,3 and low latent heat of vaporization1,2), and a high oxygen content of 35% by weight1 (and, therefore, has similar combustion behaviors to oxygenenriched combustion of the traditional gaseous fuels). As a kind of oxygenated fuel, DME can be used as an additive to traditional fuels to improve their performances,4 such as reducing the content of benzene, soot, and unburned hydrocarbons in the exhaust gas.5 As such, DME has the potential for energy conservation and emission reduction, and it is considered to be an attractive alternative fuel for the © 2013 American Chemical Society

replacement of petroleum, natural gas, and liquefied petroleum gas (LPG).6,7 Motivated by such considerations, many researchers have conducted extensive studies on the fundamental combustion behaviors of DME, including the laminar burning flame speed,8−14 ignition delay time,15,16 species profiles in the flame,17,18 detailed kinetic mechanism for DME’s oxidation and pyrolysis,19,20 combustion characteristics of the liquefied DME,21−23 cleaning properties and environmental behaviors,24−26 performance improvement of the conventional fuels if they are blended with a fraction of DME,27,28 security of the usage during the combustion process,29,30 etc. Currently, the most successful applications of DME are as an alternative fuel, fuel additive, or ignition enhancer in diesel engines. Previous research indicated that engines fueled by DME are characterized by smoke-free combustion, low emission of nitrogen oxides (NOx), reduction of noise generation, and higher thermal efficiency.3,4,31 Additionally, extensive studies have been done in an attempt to introduce DME into the fields of power generation and Received: July 14, 2013 Revised: October 21, 2013 Published: October 21, 2013 7054

dx.doi.org/10.1021/ef401337x | Energy Fuels 2013, 27, 7054−7066

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Figure 1. Schematic diagram of the medium-scaled gas combustion test platform.

household heating.31 These studies reported some superior combustion performances of DME over the traditional fuels. Hitachi of Japan had developed a 25 MW-scale and full-size DME combustor, which was composed of multiclusters of coaxial jet nozzles.32 Tokyo Electric Power Company conducted a study on the application of DME to a micro gas turbine for LPG. They reported that DME could obtain equivalent or superior performance on the gas turbine in comparison with LPG. 33 Lee et al.34,35 conducted an experimental test on a 60 kW-scale gas turbine combustion test facility fueled by neat DME and made a comparison to that fueled by methane. They found that DME was a very good fuel for power generation with low combustion instability and low NOx and CO emissions. Moreover, they proposed a design methodology for gas turbine fuel nozzles for DME to enable stable operation and low emissions of harmful gases. Chen et al.36 did a comparative study on the oxygen-enriched combustion characteristics of DME and LPG with different oxygen levels in the ceramic kiln furnace by a numerical approach. They found that DME produced no residual, no black smoke, and less carbon monoxide and nitric oxide in the combustion process. Meanwhile, the porcelain products had better gloss and higher quality. Besides, the application of DME could reduce the amount of oxygen needed in the furnace. Many successful applications were focused on engines and gas turbines, but tests and studies on other combustion systems such as industrial boilers were not studied as thoroughly.3,4,31 Compared with other types of boilers, the horizontal firetube boiler constitutes the largest share of small and medium industrial units and is the most prevalent type on the market for its intrinsic advantages, such as simple structure, small footprint, simplicity of operation and maintenance, low requirement for water treatment, and higher economical efficiency. In this type of boiler, the combustion chamber is cylindrical in shape. The shape and dimension of the combustion chamber have a significant influence on the combustion and heat transfer behaviors inside the furnace.

The originality of this paper is to use DME as a neat fuel in an industrial boiler and to investigate the flame structure and reaction zone size of the DME/air premixed flame in a cylindrical furnace, which may provide some referential materials for the development of DME-fueled industrial boilers. In the present study, tests were carried out on a mediumscaled gas combustion test platform. Additionally, simulations were also conducted with a reduced chemical kinetic mechanism of DME to obtain the flame behaviors in the furnace.

2. EXPERIMENTAL DESCRIPTION 2.1. Medium-Scaled Gas Combustion Test Platform. The present work was aimed at studying DME’s flame behaviors in an industrial-scale boiler. However, with the mismatch between the flame behaviors of DME and the structure of existing infrastructures or apparatuses which are commonly designed for the traditional fuels, it is unreasonable to carry out the testing of DME on a traditional apparatus. First, the combustion chamber in a DME boiler would be smaller than that in the traditional one. This is because, on the one hand, DME has a shorter flame length than traditional fuels26 as it is easier to be oxidized and pyrolyzed during the combustion process due to the lower bonding energy of C−O bonds in its molecules. On the other hand, with the high oxygen content of 35% by weight in its molecule, DME needs less air for burning2 and has similar combustion behaviors to oxygen-enriched combustion of the traditional gaseous fuels. Hence the DME flame has a higher radiation intensity, which results in the need for less heat exchange area in the furnace. Second, as a result of the higher heat value per unit volume of DME/air mixture compared with that of the traditional fuel/air mixture,2,24 a DME boiler would produce less flue gas and require less convective heat transfer surfaces. Third, for the above two reasons, a DME boiler would be more compact and require less footprint area and steel consumption for manufacturing in comparison with the traditional types. As a result, a medium-scaled gas combustion test platform specially designed for DME, with the schematic diagram shown in Figure 1, was employed. It is composed of the main body, the gas supplying system, a burner, the cooling water system, and the measuring system. The main body is an assembly of three sections, each of which is a commonly used component in industrial boilers. Section 1 is 7055


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Additionally, to reveal the individual influences of excess air ratio α and thermal load Q on the flame behaviors, a series of operational parameters were designed for the experiments, as specified in Table 1.

composed of a cylindrical combustion chamber (internal diameter of 500 mm, length of 1500 mm), which is surrounded by a water-cooling jacket (internal diameter of 500 mm, external diameter of 800 mm, and length of 1500 mm). To sample gas or take measurements inside the furnace, 11 equidistant sampling ports were installed in both the horizontal and vertical directions, respectively. Meanwhile, four equidistant observation ports were also installed to provide visual access for observation. In Section 2, the flue gas flowed through 24 tubes (internal diameter of 44 mm, length of 1500 mm), which were immerged in the cooling water. Section 3 is a convention tube bank, which consists of 85 water tubes (5 rows × 17 columns, and each tube has an external diameter of 51 mm and a length of 700 mm). Both ends of each water tube were jointed to one of two rectangular water tanks (length 1800 mm × width 740 mm × height 150 mm) which were located in the upper and lower positions, respectively. Natural circulation of water occurred in Section 3 during normal operation. Cooling water in each section is fed separately. In general, the gas-fired boilers are fueled by gaseous fuels. However, since DME is usually pressurized into liquid for storage and shipment, the combustion test platform must be equipped with a gas supplying system to generate and provide sufficient gaseous fuel for burning. As shown in Figure 1, the gas supplying system consists of several DME cylinders, a carburetor, a pressure reducing valve, a filter, the pipeline network, and a few valves. A gas burner with the schematic diagram shown in Figure 2 was also employed. The openings of the electromagnetic valves of the burner

Table 1. Specifications of the Operational Parameters for the Experimentsa case

thermal load (kW)

excess air ratio

1 2 3 4 5 6 7

130 110 136 118 124 266 262

1.05 1.15 1.30 1.45 1.60 1.15 1.30

a

Data for the thermal load are calculated with formula 2 and those for the excess air ratio are calculated with formula 1.

The tests were conducted under five excess air ratio levels (α = 1.05, 1.15, 1.30, 1.45 and 1.60) and two thermal loads (the high load (about 270 kW) and the low load (about 130 kW)). The expressions for the excess air ratio α and the thermal load Q are shown in formulas 1 and 2 respectively. α=

YA /YF (YA /YF)st

(1)

where YF and YA represent the fuel and air mass fractions in the premixed mixture respectively. The subscript st designates the chemical equivalent state. Q = ṁ DME · LHVDME

(2)

where Q is the thermal load of the unit, ṁ DME is the mass flow rate of DME flowing into the burner, and LHVDME is the lower heat value of DME.

3. NUMERICAL CALCULATIONS 3.1. Computational Models. The simulations in this work were conducted using the commercial CFD package ANASYS FLUENT.38 Numerical calculations of the present problem included solutions of the Favre-averaged form of mass, momentum, energy, radiative intensity, species, turbulent kinetic energy k, and its dissipation rate ε transport equations available in FLUENT. In addition, the P − 1 radiation model39 with the weighted sum of gray gas model (WSGGM)39 was adopted to model the radiation of the flames. The realizable k − ε model40 with the standard wall function was employed to predict the turbulence in the combustion system. In this paper, the eddy dissipation concept (EDC)41,42 model was adopted to solve the turbulence/chemical interaction. The EDC model, which is a modification of the eddy dissipation (ED) model, can incorporate detailed kinetic mechanism into the calculations of turbulent reacting flows. The SIMPLE algorithm was used to solve the pressure-velocity coupling, and a second-order discretization scheme was utilized to solve all governing equations. 3.2. Simulation Details. Because of the axial symmetry of the furnace, two-dimensional triangular unstructured grids were generated over the whole computational domain. Moreover, the body-fitted meshing technique was employed. In the region near the burner jet and the flame base, finer grids were generated, while in regions far away from the flame base and the axis coarser ones were generated. In this way, a primary mesh with a total of 43222 cells was generated. After the grid-

Figure 2. Schematic diagram of the burner. were regulated to maintain the thermal load in the range of 60−318 kW. As shown in Figure 2, a porous diffuser plate, which was installed before the electrodes, not only mixed the fuel and air perfectly but also served as a high temperature ignition source to stabilize the flame base attached on its surface. During the measurement of gas temperature, several K type (measuring range of −40 to 1150 °C, precision accuracy of ±0.4%) and B type (measuring range of 600−1800 °C, precision accuracy of ±0.25%) thermocouples were employed. Generally, there exists a considerable margin of error in the measurement of temperature resulting from the radiation heat loss from the thermocouple to its surrounding, especially in the presence of a water-cooled wall which has a fairly low temperature. Hence all measured data in the present work were corrected using the approach proposed by Roberts et al.37 With regard to the gas measurement, a water-cooled stainless-steel probe was used to sample gases. A MGA 5 type infrared gas analyzer was employed to obtain the mole fractions of O2 and CO2 in the sampled gases. 2.2. Experimental Methodology and Operational Parameters. In the present work, experiments were carried out on Section 1, and dimethyl ether with purity of 99.7% was used. During the tests, the measurements of gas temperature and O2 and CO2 mole fractions were conducted at the 11 horizontally located sampling ports to obtain their axial distributions. Meanwhile, at each sampling port, the same measurements were also conducted at six equidistant locations along the radius to obtain the radial profiles. 7056


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Figure 3. Verification of grid-independency for case 2. (The upper is for the coarse grids with 43222 cells and the lower is for the fine grids with 73883 cells. The temperature distribution is shown in rainbow color. The streamlines, isocontours of RCO = 0.01 (the red dashed line) and XOH = 0.0005 (the red solid line) are superimposed on it.)

Figure 4. Unstructured grids and boundary conditions in the computational domain. (Dimensions are in millimeters.)

independency of the results was verified using a finer mesh with 73883 cells, which had been indicated by the fact that the results (including the flow structure, temperature distribution, iso-contours of RCO = 0.01 (definition in formula 3) and XOH = 0.0005) for the case 2 flame using the two different meshes were kept in high consistency (shown in Figure 3), the primary mesh was finally employed for solving all other cases. The generated mesh and boundary conditions are shown in Figure 4. For the setting of the boundary conditions, the burner inlet was set to a mass flow boundary condition, where the flux of the DME/air premixed mixture with a temperature of 300 K was specified. The outlet of the furnace was set to a pressure outlet boundary condition. Moreover, the refractory-lined area (its location is shown in Figure 1) was set to an adiabatic boundary condition. Other walls inside the furnace were set to constant-temperature boundary conditions. Specifically, the temperature of the water-cooled wall was set to 380 K, and temperature values of the diffuser plate and burner jet pipe were set to 1500 K (shown in Figure 4). The seven cases of flames tested in the experiments were simulated individually in this paper. Table 2 shows the

boundary conditions at the burner inlet for each case, which coincide with the operational parameters specified in Table 1. More recently, Fischer et al.19 and Curran et al.20 have developed a detailed kinetic mechanism for DME including 79 species and 351 reversible reactions. It was validated in a shock tube,15 a counterflow diffusion flame,43 the burner-stabilized premixed flames,17,18 and spherical bombs.8−14 Meanwhile, on the basis of their work, some of these researchers9,43 developed several reduced kinetic mechanisms for DME, which were verified to have good prediction accuracy as compared to the detailed chemical mechanism. Among these, a reduced chemical mechanism (including 39 species and 168 reactions) with fairly high accuracy and rapid convergence,9,44 developed by Qin et al.9 using the computational singular perturbation (CSP) method,45 was used for the present simulations. 3.3. Definitions of the Length and Width of the Reaction Zone. According to previous studies, the destruction of DME occurs mainly in the following path:19,46,47 CH3OCH3 → CH3OCH 2 → CH 2O → HCO → CO → CO2 7057


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mole fractions were kept within ±4.1%, ± 3.0%, and ±0.9% respectively. And the largest relative errors for them were 8.0%, 5.3%, and 1.8%, respectively. The largest error for the temperatures might be due to the heat emission from the thermocouples to the wall-cooled wall, although all measured temperatures were corrected.37 The predictions agreed well with the measurements. Therefore, the models herein could capture the flame behaviors fairly well, and they were employed to solve all other cases of flames in this work. 4.2. Fluid Structure in the Furnace. Before discussing the temperature and species behaviors in the furnace, it is reasonable to be aware of the fluid structure. This is because, as will be discussed later, the fluid structure was closely coupled with the temperature and species fields in the furnace and thereby affected them strongly. The fluid structure in the furnace is illustrated in Figure 8. High temperatures occurred in the central zone of the furnace, and the outside region around it was characterized by two huge recirculation zones. According to the fluid structure in the furnace, the computational domain can be divided into three zones (separated by the white dashed lines in Figure 8): the flame core area (FCA), recirculation zone I (RZ I), and recirculation zone II (RZ II). Among these, the FCA was located in the center of the furnace, where the high temperature flame occurred. RZ I was located roughly between the refractory-lined area (its location is shown in Figure 1) and the transverse section of the burner jet. RZ II was located in the region outside the FCA. Gases in these two recirculation zones both recirculated toward the flame base, and the recirculation intensity in RZ II was much larger than that in RZ I. In addition, compared with the high temperature flame zone, the RZs had relatively low temperatures (500−1200 K) because of their adjacency relation with the water-cooled wall and the obvious cooling effect. 4.3. Temperature Distribution Characteristics in the Furnace. In the present study, rather great similarities of the flame behaviors such as the fluid structure, temperature, and species distributions among the seven cases of flames were observed. Differences among them were mainly due to the variances of excess air ratio and thermal load. Hence in the later section, the case 2 flame is taken as an example to illustrate the typical flame behaviors first, and then the influences of excess air ratio and thermal load on the flame behaviors are discussed. 4.3.1. Typical Temperature Distribution Characteristics. As shown in Figure 6a, at each sampling port, the temperature of gases decreased monotonically along the radius. Meanwhile, in both the central zone (CZ, or FCA, radius r = 0−75 mm) and the outer annular zone (AZ, or RZs, radius r = 75−200 mm), the gas temperature decreased rather slowly along the radius, implying that there were rather uniform radial temperature

Table 2. Specifications of the Boundary Conditions for the Simulationsb boundary conditions of the burner inlet

low load, varying α

high load, varying α

case

thermal load

excess air ratio

ṁ (g/s)

YDME (%)

Yair (%)

T0 (K)

1

130

1.05

42.99

9.61

21.07

300

2 3 4 5 6

110 136 118 124 266

1.15 1.30 1.45 1.60 1.15

39.38 54.49 52.02 59.95 94.90

8.85 7.91 7.15 6.53 8.85

21.25 21.47 21.64 21.79 21.25

300 300 300 300 300

7

262

1.30

103.41

7.91

21.47

300

ṁ is the mass flow rate of the DME/air premixed mixture. YDME and Yair are the mass fraction of fuel and air in the premixed mixture respectively; the remainder mass fraction in the mixture is YN2. T0 is the initial temperature of the premixed mixture. b

It has been suggested that carbon monoxide (CO) may be the last intermediate species in this process.48 Hence the technique proposed by Mei et al.48 was adopted herein to define the domain of the reaction zone. It was assumed that the contour of RCO (as expressed in formula 3) = 0.01 could approximately represent the border of the reaction zone. XCO R CO = XCO,max (3) where XCO and XCO,max designate the local CO mole fraction and its maximum value in the overall computational domain respectively. Definitions of the length L and width W of the reaction zone are shown in Figure 5. Note that the configuration of the flame in the furnace is cylindrical in shape, so the volume of the reaction zone can be approximately calculated as π V = LW 2 (4) 4

4. RESULTS AND DISCUSSION 4.1. Validation of the Numerical Results. To validate the simulation results, the predicted temperature and species profiles were compared with the measured data. Figure 6 shows the radial profiles for the temperature and O2 and CO2 mole fractions at sampling ports 2, 5, 8, and 11, respectively, and the axial profiles at radial positions of 0 mm and 150 mm for them are shown in Figure 7. Results of the error analysis indicated that the average relative errors for the predicted temperatures, O2 and CO2

Figure 5. Definitions of the length and width of the reaction zone. 7058


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Figure 6. Comparison of the measured and calculated temperature (a), O2 and CO2 mole fraction (b) distributions along the radius at sampling ports 2, 5, 8, and 11 for case 2. (Lines represent the predicted results and scattered symbols represent the measured data.)

Figure 7. Comparison of the measured and calculated temperature and O2 and CO2 mole fraction distributions along the axial distance at radius of 0 mm (a) and 150 mm (b) for case 2 respectively (lines represent the predicted results and scattered symbols represent the measured data).

Figure 8. Predicted temperature distribution and fluid structure for case 2.

downstream of the flame, this discrepancy became gradually smaller since adequate heat transfer, and mass mixing between these two zones occurred. Moreover, a thin layer adjacent to the water-cooled wall (radius r = 200−250 mm), having a sharp decrease of temperature, was observed in the present work. The temperature of the gas in this layer was close to that of the wall. The axial temperature profile at radius r = 0 mm in Figure 7a demonstrates that, after injection into the furnace, the premixed mixture was heated up nearly to the adiabatic temperature quickly by the heat release from the chemical reactions, while

profiles in the FCA and the RZs. However with the lower temperatures in the RZs compared with those in the FCA, which is mentioned above due to the RZs’ adjacent relation with the water-cooled wall and consequently more obvious cooling effect, there should be a considerable radial temperature gradient at the flame surface. As expected, this phenomenon was observed at the interface of the CZ and the AZ (i.e., the flame surface, r = 75 mm approximately). It is also found from Figure 6a that the temperature discrepancy between the CZ and the AZ upstream of the flame was fairly large. However 7059


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Figure 9. Predicted axial temperature distributions at radius r = 0 mm and 150 mm for cases 1−7 respectively. (Cases 1−5 designate excess air ratios of 1.05, 1.15, 1.30, 1.45, and 1.60 under the low load respectively; case 6 and case 7 designate excess air ratios of 1.15 and 1.30 under the high load, respectively.)

furnace. As we know, heat transfer in the furnace is dominated by thermal radiation.39 To simplify the understanding and discussion, the concept of volumetric heat loss by thermal radiation qv,rad, which is expressed in formulas 5 and 6,49 was adopted herein.

after that, the gas temperature began to decrease slowly along the axial distance in the high temperature flame zone (x = 50− 900 mm). In addition, a transition point (x = 900 mm) was observed at the end of the flame. The gas temperature decreased very fast downstream of this point. The temperature curve corresponding to radius r of 150 mm in Figure 7b indicates the axial temperature profile in the recirculation zones. Upstream of RZ II, the gas temperature decreased slightly at first. However downstream of RZ II, it increased gradually instead along the axial distance x. The gas temperature in RZ II reached that in the FCA at the furnace outlet. As discussed in the later section, the axial temperature profile in the RZs was the result of the coupling relationship between the temperature field and the fluid structure. The fluid structure (shown in Figure 8) in the furnace indicates that low temperature gases (400−500 K) adjacent to the water-cooled wall in two RZs flowed toward the flame base, resulting in the decrease of gas temperature upstream of RZ II, whereas downstream of RZ II sufficient heat transfer and mixing of gases between the FCA and RZ II led to the increase of gas temperature in this zone. 4.3.2. Effects of Excess Air Ratio and Thermal Load on Temperature Distribution. The axial temperature profiles at radius r of 0 mm and 150 mm for cases 1−7 are shown in Figure 9 respectively. Comparisons of the temperature curves for cases 1−5 can demonstrate the influence of excess air ratio on the temperature behaviors. It is found that the gas temperature in the flame zone (upstream of the intersection points between the green dashed line and the temperature curves in Figure 9) decreased significantly with the increase of excess air ratio. This is the result of a lower adiabatic flame temperature under the higher excess air ratio. However, in the post flame zone (downstream of the intersection point in Figure 9), the temperature discrepancies among the cases of different excess air ratios were very small. Meanwhile, the temperature at the furnace outlet seemed almost equal under different excess air ratios. This phenomenon will be explained in the next section. The temperature at the furnace outlet is the result of thermal equilibrium between heat absorption and heat emission in the

qv,rad = 4σK p(T 4 − T04)

K p = P ∑ XiK p, i k

(5)

(6)

where σ is the Stefan−Boltzmann constant. T and T0 are the local gas temperature and the water-cooled wall temperature respectively. Kp is a term accounting for the absorptions and emissions of the participating radiating species (including H2O, CO2, CO, and CH4). It is expressed in formula 6, where P is the total pressure of the flame, Xi and Kp,i are the mole fraction and Planck mean absorption coefficient of radiating species i respectively. Additionally, the term Kp,i is a decreasing function of temperature in the range above 700 K.50 First, under the precondition of equal thermal load, the increase of excess air ratio led to the decreases of flame temperature and partial pressures of the radiating species, which further led to the decrease of heat loss by thermal radiation according to formulas 5 and 6. Then the differential between the thermal output and the heat loss by thermal radiation, namely, the sensible enthalpy of the flue gas, increased instead. Second, the sensible enthalpy of the flue gas is the product of its mass, the mean specific heat, and the differential value between the flue gas temperature and the reference temperature. In addition, under equal thermal load, the case with a higher excess air ratio had more mass of flue gas. Therefore, the temperature of the flue gas, that is, the temperature at the furnace outlet under different excess air ratios might be equal. From another point of view, the radiation intensity is highly dependent on the flame temperature. Hence only when the gas temperature fell down below a fairly low value (about 1200 K), the heat loss by thermal radiation could be neglected and the gas temperature stopped decreasing. In this way, the same conclusion with regard to the independence of the temperature 7060


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Figure 10. Comparison of the fluid structure between the high load and the low load conditions. (for case 2 and case 6 only).

Figure 11. Predicted distributions for the turbulent kinetic energy and the turbulent viscosity (a), the effective diffusion coefficients of O2 and CO2 (b) for case 2.

first, as shown in Figure 9, the high load and low load cases of the same excess air ratio had an equal temperature at the end of the flame zone. Therefore, with more amount of flue gas, the high load case emitted a higher amount of heat than the low load case. Second, under equal excess air ratio, the flame temperature (shown in Figure 9) and partial pressures of the radiating species (shown in Figure 12) in the flame zone for the high load case were basically equal to those for the low load case. Then the term Kp, and subsequently the volumetric heat loss by thermal radiation qv,rad for the high load case and those for the low load case were equal as well, according to formulas 5 and 6. As a result, the high load case needed a larger thermal radiating volume and had a longer flame length.

at the furnace outlet and the excess air ratio can be obtained also. Moreover, under equal thermal load, the case with a higher excess ratio had a slightly longer flame length. In addition, the discrepancies of gas temperature in RZ II among the cases of different excess air ratios were very small, and the case with a higher excess air ratio had a slightly lower temperature in this zone. On the basis of comparisons between the temperature curves for the high load cases (case 6, 7) and those for the low load cases (cases 2 and 3) in Figure 9, it can be concluded that the high load case had a longer flame length than the low load case under the precondition of equal excess air ratio. This is because 7061


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Figure 12. Predicted axial profiles of O2 (a) and CO2 (b) mole fractions at radius r of 0 mm and 150 mm for cases 1−7 respectively.

the RZs (RZ I and RZ II), the gas recirculation enabled such a long residence time that gases in these two zones were basically burnt out. Hence the O2 content in the RZs was lower than that in the FCA, and the CO2 content in the RZs was higher than that in the FCA. Therefore, considerable radial species gradients were observed at the flame surface. Moreover, although the radial diffusion could reduce the radial species gradients at the flame surface in part, the effective diffusion coefficients for O2 and CO2 (shown in Figure 11b) in the upstream region were very small, implying that the effect of diffusion on reducing the radial species gradients was insignificant in this region. However, in the downstream region, the effective diffusion coefficients increased drastically. The intensified diffusion resulted in the homogeneous species distributions in this region. The axial species profiles at radius r of 0 mm in Figure 7a demonstrate that, in the FCA, the O2 content decreased and the CO2 content increased gradually along the axial distance. In addition, a transition point existed at the axial position of about 900 mm. The species contents varied slowly upstream of this transition point, while downstream of the transition point, the variation of species concentrations increased. Note also that the axial position of the species curves’ transition point was highly consistent with that of the temperature curve’s transition point (shown in Figure 7a) which was at the end of the flame zone. This phenomenon implied that it was the coupling relationship between the temperature and species diffusion that resulted in the intensified diffusion and larger variation in species concentrations downstream of the transition point, but not the chemical reactions because of the discrepancy between the border of the reaction zone (shown in Figure 13) and the position of the transition point. The reason will be presented below. The diffusion flux of species i in the species transport equation can be expressed as

In addition, the high load case had a higher temperature at the furnace outlet in comparison with the low load case. This is because the thermal output under the high load (about 270 kW) was approximately twice that under the low load (about 130 kW). However, the flame volume under the high load was less than twice that under the low load (shown in Figure 10). In addition, as discussed previously, the volumetric heat loss by thermal radiation under the high load case was nearly equal to that under the low load case. Hence the total amount of heat loss by thermal radiation under the high load was less than twice that under the low load. Therefore, the high load case had a higher temperature at the furnace outlet than the low load case, under the precondition of equal excess air ratio. It can also be found from Figure 9 that the gas temperatures in RZ II (represented by the temperature curves at radius r of 150 mm) under the high load (case 6, 7) were higher than those under the low load (case 2, 3). The reason is that first, heat transfer, mixing and diffusion were more intense under the high load condition. Second, as shown in Figure 10, RZ I shrank and RZ II enlarged and moved upstream under the high load condition, implying that more high temperature gases were entrained from the FCA to RZ II. As a result, the high load case had higher temperatures in RZ II. In summary, the gas temperature in the FCA was mainly determined by the excess air ratio. The temperature in RZ II was strongly affected by the thermal load. In addition, the fluid structure could affect the temperature behaviors partly for its coupling relationship with the temperature field. 4.4. O2 and CO2 Distribution Characteristics in the Furnace. 4.4.1. Typical Distribution Characteristics for O2 and CO2. Figure 6b shows that, at each sampling port, the O2 content decreased and the CO2 content increased monotonically along the radius. In addition, being similar to the temperature behavior, the radial gradients for O2 and CO2 mole fractions were rather small in the CZ (or FCA, radius r = 0−75 mm) and the AZ (or RZs, radius r = 75−200 mm), while at the interface of the CZ and the AZ (i.e., the flame surface, r = 75 mm approximately), the species curves had considerable radial gradients. This phenomenon can also be explained by the fluid structure in the furnace. As illustrated in Figure 8, gases in the FCA flowed out of the furnace directly after burning, and the residence time was very short, which resulted in the relatively high O2 content and low CO2 content in this zone. However, in

⎛ μ ⎞ ∇T Ji = −⎜ρDi ,m + t ⎟∇Yi − DT , i Sct ⎠ T ⎝

(7)

where the first term on the RHS of eq 7 is the diffusion caused by concentration gradient. The second term is the Soret diffusion,45,51 which is caused by temperature gradient. The first 7062


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Figure 13. Reaction zone size visualizations for cases 1−7 respectively; the temperature (upper) and CO mole fraction (lower) distributions are shown in rainbow colors, isocontours of XOH = 0.0005 (the white solid lines) and RCO = 0.01 (the white dashed lines) are superimposed.

on the RHS of eq 9 is the turbulent diffusivity,45,52 which is correlated with the turbulence and accounts for the diffusion caused by eddy mixing. In practical turbulent flows, the turbulent diffusion flux overwhelms the laminar one. ρ, μt and Sct are the density, the turbulent viscosity, and the turbulent Schmidt number of the fluid respectively. Cμ is a function of the mean strain and rotation rates, the angular velocity of the system rotation, and the turbulence fields (k and ε).40 In the flame zone, the high gas temperature resulted in the high viscosity53 μ, which further led to the fairly low turbulent kinetic energy (shown in Figure 11a), turbulent viscosity μt (formula 10) and effective diffusion coefficient Deff,i (formula 9). Thus the diffusion in the flame zone was not sufficiently intense. However in regions outside the flame zone with a sharp

term can be rewritten in terms of the effective diffusion coefficient Deff,i as ⎛ μ ⎞ −⎜ρDi ,m + t ⎟∇Y = −ρDeff, i∇Yi Sct ⎠ ⎝

Deff, i = Di ,m + μt = ρCμ

k2 ε

(8)

μt ρSct

(9)

(10)

where Di,m is the laminar diffusion coefficient for species i. It is a physical parameter of species and correlated with the molecular weight, molecular structure, and temperature. The second term 7063


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where k̅ is the mean reaction rate constant. [X i] is concentration of species i. υ1 and υ2 are reaction orders for DME and O2 respectively. It is hypothesized that, under the condition of higher excess air ratio, there was more oxidizer participating in the combustion process, so the mean reaction rate Rnet increased despite the lower flame temperature. As a result, the reaction zone shrank. Moreover, comparisons between the high load cases (case 6, 7) and the low load cases (case 2, 3) demonstrate that, under equal excess air ratio, the high load case had a longer and larger reaction zone in comparison with the low load case. This is because, under equal excess air ratio, the gas temperature (shown in Figure 9), fuel, and oxidizer mole concentrations (shown in Figure 12), and subsequently the volumetric reaction rate Rnet (formula 11) in the reaction zone for the high load case, were basically equal to those for the low load case. Consequently, the high load case with more mass of mixture needed more reaction volume and had a longer length of the reaction zone. As discussed previously, the mean reaction rate Rnet of the global reaction R1 was nearly independent of the thermal load, and it was only dependent on the excess air ratio α. The relationship among the mean reaction rate Rnet, the volume of the reaction zone V, and the thermal load Q can be expressed approximately as48

decrease of gas temperature, the effective diffusion coefficient increased drastically, which led to the intensified diffusion in this region. In addition, the Soret diffusion was high in regions with a high temperature gradient, such as the flame surface. Therefore, downstream of the flame zone (i.e., the post flame region), the O2 content in the FCA decreased because O2 in this zone was transported to the RZs, and the CO2 content in the FCA increased because CO2 in the RZs was transported to this zone. At the same time, the intensified diffusion in the post flame region resulted in an increase of O2 content and the decrease of CO2 content downstream of RZ II, as shown in Figure 7b. 4.4.2. Effects of Excess Air Ratio and Thermal Load on O2 and CO2 Distributions. The axial profiles for O2 and CO2 mole fractions at radius of 0 mm and 150 mm for cases 1−7 are shown in Figure 12. With the increment of excess air ratio, the O2 content in the overall domain of the furnace increased, and the CO2 content decreased instead. In addition, it is interesting to note that the differential value between the O2 (or CO2) content in the flame zone and that in the RZs increased with the increment of excess air ratio. Moreover, comparisons between the curves for the high load cases (cases 6 and 7) and those for the low load cases (cases 2 and 3) illustrate that the transition point of the species curves moved downstream with the increment of thermal load, which implied that the high load case had a longer flame length. In addition, O2 and CO2 contents in the flame zone and those in regions outside it seemed independent of the thermal load. They were more likely to be dependent on the excess air ratio. 4.5. The Reaction Zone Size of DME/Air Premixed Flame. Generally, in high temperature flames, hydrocarbon molecules break up into smaller species such as CH3, CH2O, HCO, CO, etc. quickly. Then these intermediate species are oxidized by H and OH further, forming the final products of CO2 and H2O. Hence the highly active radicals such as H and OH can indicate the domain of the reaction zone.54−59 For example, Katta et al.54 used the isocontour of XOH = 0.0005 to indicate the border of the flame surface, and it agreed well with the blue luminous zone in the experimental observation. In the present study, the isocontours of RCO = 0.01 and XOH = 0.0005 were both employed to indicate the position and domain of the reaction zone. The temperature and CO mole fraction distributions for cases 1−7 are shown in Figure 13 respectively. The isocontours of RCO = 0.01 (the white dashed lines) and XOH = 0.0005 (the white solid lines) are superimposed on the color maps. It is found that the reaction zones obtained from the two methodologies overlapped with each other fairly well, so the results of the reaction zones were reasonable. Comparison of cases 1−5 (in Figure 13) indicates that under equal thermal load, an increase of the excess air ratio resulted in a decrease of the length of the reaction zone and a shrinkage of its volume. This is consistent with the findings of Mei et al.48 ́ and Verissimo et al.57 The reason is that for the global reaction mechanism of DME: CH3OCH3 + 3O2 = 2CO2 + 3H 2O

Q = R netV MWDMELHVDME

where MWDME and LHVDME are the molecular weight and lower heat value of DME respectively. In the above formula, the reaction zone volumes V obtained by the RCO = 0.01 methodology were used to calculate the mean reaction rates for different cases. The results are shown in Figure 14.

Figure 14. Calculated mean reaction rates for cases 1−7.

Figure 14 indicates that the influence of thermal load on the mean reaction rate was insignificant, which in turn offers an endorsement for the verification of the former discussions. Using the regression analysis method, a functional expression between the mean reaction rate Rnet and the excess air ratio α was developed in this paper.

(R1)

the mean volumetric reaction rate Rnet can be expressed as R net =

d[XCH3OCH3] dt

= − k ̅[XCH3OCH3]υ1 [XO2]υ2

(12)

R net = 2.9747 + 0.0176 exp(α /0.2469) (1.05 ≤ α ≤ 1.60, R net is in mol/m 3·s)

(11) 7064

(13)


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Notes

Although formulas 12 and 13 were obtained based on the assumption of a global reaction mechanism, neglecting many intermediate species and the detailed chemical mechanism, they can adequately satisfy the accuracy of engineering applications. Formula 13 can be employed to estimate the mean reaction rate, the volume of the reaction zone, the flame length, etc. of the DME/air premixed flame within the range of excess air ratio α = 1.05−1.60 at any thermal output. It can be used in engineering applications, such as the design for DME-fueled industrial boilers, fire protection aspects, etc.

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The present research was supported by the Key Science and Technology Project of Yibin, Sichuan, China (Grant No. Finance and Education Bureaus of Yibin, (2012) 19), Specific Research Fund for the Doctoral Program of Higher Education of China (Grant No. 20100191120017), and Scientific Research Foundation for the Returned Scholars, Ministry of Education of China. Special thanks are due to Prof. X. Lu and Dr. Q. Wang for providing insightful comments on the manuscript.

5. CONCLUSIONS Flame behaviors, including the fluid structure, temperature, and species distributions, and the reaction zone size of the dimethyl ether/air premixed flame in the furnace were experimentally and numerically investigated in this paper. Seven cases of flames with different operational parameters were studied to reveal the influences of excess air ratio and thermal load on the combustion behaviors. The experiments were carried out on a medium-scaled gas combustion test platform. Numerical simulations were conducted using the EDC model with a reduced chemical kinetic mechanism consisting of 39 species and 168 reversible reactions. The main findings of this work are as follows: (1) The fluid structure in the furnace consisted of three parts: the flame core area (FCA), recirculation zone I (RZ I), and recirculation zone II (RZ II). The temperature and species fields were coupled with the fluid structure and were influenced by it partly. (2) The temperature in the FCA was mainly dependent on the excess air ratio; the case with a higher excess air ratio had a lower temperature level in this zone. Thermal load had a strong effect on the temperature in the RZs (RZ I and RZ II); the high load case had a higher temperature level in the RZs in comparison with the low load case. Moreover, the temperature at the furnace outlet was positively correlated with the thermal load, while its correlation with the excess air ratio seemed insignificant in our observations. In addition, the fluid structure could influence the temperature behaviors in the furnace partly for its coupling relationship with the temperature field. (3) Species concentrations were highly correlated with the excess air ratio; the case with a higher excess air ratio had a higher O2 content and a lower CO2 content in the overall furnace whereas the influence of thermal load over the species contents was insignificant. In addition, in regions outside the flame zone, the intensified diffusion resulting from the decrement of temperature could change the species distributions in part. (4) Investigation of the reaction zone size indicated that either the decrease of excess air ratio or the increase of thermal load could lead to an enlargement of the reaction zone. Additionally, the mean reaction rate of the dimethyl ether/air premixed flame was shown to be independent of the thermal load. It was only dependent on the excess air ratio. Finally, a correlation between the mean reaction rate and the excess air ratio was developed in this paper. It can be used to estimate the mean reaction rate, the volume of the reaction zone, the flame length, etc. of the DME/air premixed flame within the range of excess air ratio α = 1.05−1.60 at any thermal output.

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AUTHOR INFORMATION

Corresponding Author

*Tel./Fax: +86 023-65102475. E-mail: xfl[email protected]. 7065

NOMENCLATURE AZ = annular zone CFD = computational fluid dynamics CSP = computational singular perturbation CZ = central zone DME = dimethyl ether (CH3OCH3) ED = eddy dissipation EDC = eddy dissipation concept FCA = flame core area LPG = liquefied petroleum gas RZ = recirculation zone RHS = right-hand side WSGGM = weighted sum of gray gas model Deff,i = the effective diffusion coefficient for species i, m2/s Di,m = laminar diffusion coefficient for species i, m2/s DT,i = Soret diffusion coefficient for species i, kg/m·s Ji = diffusion flux of species i, kg/m2·s Kp = a quantity accounting for absorptions and emissions of all the radiating species, m−1 Kp,i = Planck mean absorption coefficients of the radiating species i, m−1·atm−1 k = turbulent kinetic energy of the fluid, J/kg k̅ = mean reaction rate constant for reaction R1, its unit is determined by the reaction order, and are in mol, m, s LHVDME = lower heat value of DME, 31630.4 kJ/kg L = length of the reaction zone, m MWDME = molecular weight of dimethyl ether, kg/mol ṁ DME = mass flow rate of DME flowing into the burner, kg/s P = total pressure of the flame, atm Q = thermal load of the unit, kW qv,rad = volumetric heat loss by thermal radiation, W/m3 Rnet = the mean volumetric reaction rate for reaction R1, mol/(m3·s) r = radius of the cylindrical combustion chamber, mm Sct = turbulent Schmidt number, dimensionless T = temperature, K V = volume of the reaction zone, m3 W = width of the reaction zone, m X CO = local mole fraction of CO in the system, dimensionless XCO,max = the maximum mole fraction of CO in the whole computational domain, dimensionless Xi = mole fraction of species i, dimensionless XOH = mole fraction of species OH, dimensionless [Xi] = concentration of species i in the system, mol/m3 x = axial distance of the cylindrical combustion chamber, mm YA = mass fraction of air in the premixed mixture, dimensionless dx.doi.org/10.1021/ef401337x | Energy Fuels 2013, 27, 7054−7066

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(31) Fleisch, T. H.; Basu, A.; Sills, R. A. J. Nat. Gas Sci. 2012, 9, 94− 107. (32) Kobayashi, N.; Inoue, H.; Koizumi, H. Proc. ASME Turbo EXPO 2003, GT2003−38410. (33) Tsuchiya, T.; Okamoto, M. Proc. Asian Congress Gas Turbines 2005, ACGT2005−066. (34) Lee, M. C.; Seo, S. B.; Chung, J. H.; Joo, Y. J.; Ahn, D. H. Fuel 2008, 87, 2162−2167. (35) Lee, M. C.; Seo, S. B.; Chung, J. H.; Joo, Y. J.; Ahn, D. H. Fuel 2009, 88, 657−662. (36) Chen, R.; Wang, H. T.; Wang, H. Proc. Eng. 2012, 31, 934−940. (37) Roberts, I. L.; Coney, J. E. R.; Gibbs, B. M. Appl. Therm. Eng. 2011, 31, 2262−2270. (38) ANSYS FLUENT Documentation; Ansys, Inc.: Canonsburg, PA, 2011. (39) Modest, M. F. Radiative Heat Transfer; Academic Press: New York, 2003. (40) Shih, T. H.; Liou, W. W.; Shabbir, A.; Zhu, J. Comput. Fluids 1995, 24, 227−238. (41) Rehm, M.; Seifert, P.; Meyer, B. Comput. Chem. Eng. 2009, 33, 402−407. (42) Magnussen, B. F.; Hjertager, B. H. Symp. (Int.) Combust. 1977, 16, 719−729. (43) Zheng, X. L.; Lu, T. F.; Law, C. K.; Westbrook, C. K.; Curran, H. J. Proc. Combust. Inst. 2005, 30, 1101−1109. (44) Ju, Y.; Xue, Y. Proc. Combust. Inst. 2005, 30, 295−301. (45) Law, C. K. Combustion Physics; Cambridge University Press: Cambridge, U.K., 2006. (46) Curran, H. J.; Pitz, W. J.; Westbrook, C. K.; Dagaut, P.; Boettner, J. C.; Cathonnet, M. Int. J. Chem. Kinet. 1998, 30, 229−241. (47) Wang, Y. L.; Holley, A. T.; Ji, C.; Egolfopoulos, F. N.; Tsotsis, T. T.; Curran, H. J. Proc. Combust. Inst. 2009, 32, 1035−1042. (48) Mei, Z.; Mi, J.; Wang, F.; Zheng, C. Energy Fuels 2012, 26, 3257−3266. (49) Lock, A. J.; Briones, A. M.; Qin, X.; Aggarwal, S. K.; Puri, I. K.; Hegde, U. Combust. Flame 2005, 143, 159−173. (50) Lee, B. J.; Chung, S. H. Combust. Flame 1997, 109, 163−172. (51) Yang, F.; Law, C. K.; Sung, C. J.; Zhang, H. Q. Combust. Flame 2010, 157, 192−200. (52) Gordon, F. G.; Thévenin, J. D. Flow Turbul. Combust. 2012, 88, 451−478. (53) Turns, S. R. An Introduction to Combustion: Concepts and Applications; McGraw-Hill Companies, Inc.: New York, 2012. (54) Katta, V. R.; Stouffer, S.; Roquemore, W. M. Combust. Flame 2011, 158, 1149−1159. (55) Nogenmyr, K. J.; Kiefer, J.; Li, Z. S.; Bai, X. S.; Aldén, M. Combust. Flame 2010, 157, 915−924. (56) Theron, M.; Bellenoue, M. Combust. Flame 2006, 145, 688− 702. (57) Veríssimo, A. S.; Rocha, A. M. A.; Costa, M. Energy Fuels 2011, 25, 2469−2480. (58) Yan, B.; Li, B.; Baudoin, E.; Liu, C.; Sun, Z. W.; Li, Z. S.; Bai, X. S.; Aldén, M.; Chen, G.; Mansour, M. S. Exp. Therm. Fluid Sci. 2010, 34, 412−419. (59) Juddoo, M.; Masri, A. R. Combust. Flame 2011, 158, 902−914.

YF = mass fraction of fuel in the premixed mixture, dimensionless α = excess air ratio, dimensionless ε = dissipation rate of the turbulent kinetic energy, m2/s3 μ = molecular viscosity, Pa·s μt = turbulent viscosity, Pa·s ρ = density of fluid, kg/m3 σ = Steffan−Boltzmann constant, = 5.67 × 10−8 W/(m2·K4) υ1 = reaction order for DME in reaction R1, dimensionless υ2 = reaction order for O2 in reaction R1, dimensionless

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REFERENCES

(1) Park, S. H.; Lee, C. S. Prog. Energy Combust. Sci. 2013, 39, 147− 168. (2) Arcoumanis, C.; Bae, C.; Crookes, R.; Kinoshita, E. Fuel 2008, 87, 1014−1030. (3) Pfeifer, C.; Kuhn, D.; Class, A. G. Fuel 2013, 104, 116−127. (4) Rakopoulos, D. C.; Rakopoulos, C. D.; Giakoumis, E. G.; Dimaratos, A. M. Energy 2012, 43, 214−224. (5) McEnally, C. S.; Pfefferle, L. D. Proc. Combust. Inst. 2007, 31, 603−610. (6) Adachi, Y.; Komoto, M.; Watanabe, I.; Ohno, Y.; Fujimoto, K. Fuel 2000, 79, 229−234. (7) Liang, C.; Ji, C.; Gao, B.; Liu, X.; Zhu, Y. Energy Convers. Manage. 2012, 58, 19−25. (8) Daly, C. A.; Simmie, J. M.; Wurmel, J.; Djeballi, N.; Paillard, C. Combust. Flame 2001, 125, 1329−1340. (9) Qin, X.; Ju, Y. Proc. Combust. Inst. 2005, 30, 233−240. (10) Huang, Z.; Wang, Q.; Yu, J. Fuel 2007, 86, 2360−2366. (11) Chen, Z.; Wei, D.; Huang, Z.; Miao, H.; Wang, X.; Jiang, D. Energy Fuels 2009, 23, 735−739. (12) Zhao, Z.; Kazakov, A.; Dryer, F. L. Combust. Flame 2004, 139, 52−60. (13) Wang, Y.; Holley, A. T.; Ji, C.; Egolfopoulos, F. N.; Tsotsis, T. T.; Curran, H. J. Proc. Combust. Inst. 2009, 32, 1035−1042. (14) Jaap de, Vries; William, B. L.; Zeynep, S.; Curran, H. J.; Petersen, E. L. Fuel 2011, 90, 331−338. (15) Tang, C. L.; Wei, L. J.; Zhang, J. X.; Man, X. J.; Huang, Z. H. Energy Fuels 2012, 26, 6720−6728. (16) Cook, R. D.; Davidson, D. F.; Hanson, R. K. Proc. Combust. Inst. 2009, 32, 189−196. (17) Kaiser, E. W.; Wallington, T. J.; Hurley, M. D.; Platz, J.; Curran, H. J.; Pitz, W. J.; Westbrook, C. K. J. Phys. Chem. A 2000, 104, 8194− 8206. (18) McIlroy, A.; Hain, T. D.; Michelsen, H. A.; Cool, T. A. Proc. Combust. Inst. 2000, 28, 1647−1653. (19) Fischer, S. L.; Dryer, F. L.; Curran, H. J. Int. J. Chem. Kinet. 2000, 32, 713−740. (20) Curran, H. J.; Fischer, S. L.; Dryer, F. L. Int. J. Chem. Kinet. 2000, 32, 741−759. (21) Toshio, M.; Hiroumi, S.; Yuji, W.; Ritsu, D. Fuel 2011, 90, 2508−2513. (22) Suh, H. K.; Park, S. H.; Kim, H. J.; Lee, C. S. Fuel 2009, 88, 1070−1077. (23) Park, S. H.; Kim, H. J.; Lee, C. S. Fuel 2010, 89, 3001−3011. (24) Semelsberger, T. A.; Borup, R. L.; Greene, H. L. J. Power Sources 2006, 156, 497−511. (25) Kazuhiro, H.; Toshio, M.; Kenji, A.; Masataka, A. Fuel 2011, 90, 493−498. (26) Hwang, C. H.; Lee, C. E.; Lee, K. M. Energy Fuels 2009, 23, 754−761. (27) Song, K. H.; Nag, P.; Litzinger, T. A.; Haworth, D. C. Combust. Flame 2003, 135, 341−349. (28) Chen, Z.; Qin, X.; Ju, Y.; Zhao, Z.; Chaos, M.; Dryer, F. L. Proc. Combust. Inst. 2007, 31, 1215−1222. (29) Mogi, T.; Horiguchi, S. J. Hazard. Mater. 2009, 164, 114−119. (30) Mogi, T.; Shiina, H.; Wada, Y.; Dobashi, R. J. Loss Prev. Process Ind. 2013, 26, 32−37. 7066


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