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Combustion characteristics of bluff-body turbulent swirling flames with coaxial air microjet Xiao Yang, Zhihong He, Shikui Dong, and Heping Tan Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.7b03048 • Publication Date (Web): 28 Nov 2017 Downloaded from http://pubs.acs.org on November 29, 2017

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Energy & Fuels

Combustion characteristics of bluff-body turbulent swirling flames with coaxial air microjet Xiao Yang, Zhihong He, Shikui Dong*, Heping Tan School of Energy Science and Engineering, Harbin Institute of Technology, Harbin 150001, People’s Republic of China

Abstract The interaction between the coaxial air microjet structure and the swirling flow is investigated numerically based on the non-premixed bluff-body stabilized swirling SM1 flame. The effects of microjet velocity and swirl number on flame combustion characteristics are analyzed. The modified standard k-ε model and the realizable k-ε model are considered for turbulent model using FLUENT software. The flamelet model is used with the GRI 2.11 detailed kinetic mechanism. In addition, the discrete ordinate model (DOM) and the weighted-sum-of-gray-gases (WSGG) model are used together to deal with the radiative heat transfer in the combustion process. Results of the modified standard k-ε model are in better agreement with the experimental data than the realizable k-ε model. Results show that the inner flame generates due to the existence of microjet configuration. High microjet velocity (such as 50 m/s) can reduce the soot formation and keep the inner flame high temperature region away from the microjet nozzle. The structure of the inner flame depends on the microjet velocity and is independent of the swirl number. Swirl number increases from 0.3 to 0.6, the primary peak temperature is enhanced by 48 K and the peak soot volume fraction is increased by 49% along the axis centerline. Keywords: Swirling flame; Air microjet; Velocity; Swirl number; Soot formation

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1. Introduction Nowadays, environmental contamination issue caused by fuels burning has been paid more and more attention. Combustion efficiency and pollutant discharge are two important indexes to evaluate the combustion process. Increasing the combustion efficiency and reducing emissions are the goal that human has been pursued all the time. With the advancement of technology, more and more advanced combustion technologies have been proposed, such as flameless combustion1, 2

, low temperature combustion3, 4, hydrogen enriched combustion5,

6

and microjet assisted

flames7-10, etc. Air microjet is defined as air injection from a small diameter nozzle.10 The microjet structure was firstly tested by Ganguly et al.7 and Sinha et al.8 in coaxial non-premixed flames. The experimental results of Ganguly et al.7 found that the increase of microjet velocity could shorten the flame length and reduce the flame fluctuations due to buoyancy. There was no distinction of microjet fluids between air and inert species. Measurements of Sinha et al.8 shown that the microjet could be an efficient tool to control the flame height and luminosity. Microjet didn’t change the overall heat transfer of the flame. The judicious microjet velocity was beneficial to combustion without an emissions penalty. After that, the microjet was introduced to Brookes and Moss configuration11 by Chouaieb et al.9, 10. The numerical results presented that an inner flame generated because of the microjet. The increase of hydrogen percentage of CH4-H2 mixture, resulted in mixing enhancement and lower soot emission. Simultaneously, low microjet diameter and high microjet velocity in CH4/air flame could achieve the same effect as increasing hydrogen percentage. On the other hand, Altay et al.12 had experimentally investigated that the air injection in the cross-stream and streamwise direction could be efficacious ways in controlling the thermoacoustic instabilities. Cao et al.13 used side microjets to control the flame structure to some extent. With the same fuel and air rates, flame length was shortest with two


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microjets, and the flame was more stable with three microjets. From these studies, we can see that the potential of microjet in controlling flame characteristics and emissions. Meanwhile, swirling flows are exceedingly extensive in the field of combustion applications, such as gas turbines, internal combustion engines and gas furnaces and so on. The recirculation zone exists usually in the combustion process of mixing swirl air with fuel. The recirculation zone contributes to the enhancement of mixing process, which increases the flame stability and reduces the pollutant emissions. There are many studies on swirling combustion in the literature. Kewlani et al.14 had studied that the average size of the inner recirculation zone could be reduced by increasing the equivalence ratio. Kashir et al.15 predicted distinct hydrogen concentrations of propane/hydrogen bluff-body flames. The increase of flame length and the reduction of CO emission could be caused by increasing hydrogen concentration. In other studies, oxygen enrichment enhanced combustion efficiency, flame stability16 and reduced the CO emission17. Ilbas et al.18 predicted NOx values increased with the swirl number changed from 0 to 0.8 in the Harwell furnace. On the other hand, many researchers had focused on numerical models of simulating swirl combustion. Mardani et al.19 had compared eddy dissipation concept (EDC) model with transported probability density function (TPDF) model in a gas turbine model combustor. According to the results, both models could capture the main flow field structures. Recently, three chemical mechanisms based on the flamelet model were used to analyze the combustion and emission characteristics of a model combustor.20 In short, higher flame stability and lower emissions are present in the swirl combustion system. Even though the air microjet assisted flame has been investigated by experimental7, 8 or numerical9, 10 methods, its application to the swirl flame is not studied sufficiently. Meanwhile, the high momentum of coaxial air microjet has a damaging effect on the recirculation zone of


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swirl combustion system. But a reasonable air microjet configuration may be beneficial to the combustion process and emission. Therefore, a numerical study is implemented for investigating the interaction between the coaxial air microjet structure and the swirling flow in a swirl burner. In the present work, we focus on the Sydney swirl burner21-23. This burner has been investigated by many researchers to verify the turbulent model24 and combustion model25, study the flame characteristics26-28. Yang et al.24 reported two types of large eddy simulation (LES) and Reynolds-averaged Navier-Stokes (RANS) models simulation results of SMH1 flames. They found that the Cvar number was a vital factor in affecting the local extinction phenomenon for laminar flamelet model. The SM1 flame based on Sydney burner was employed by Luo et al.25 to validate a dynamic second-order moment closure combustion model of LES. Kashir et al.26 and Rohani et al.27 showed the characteristics of the H2 addition and different swirl numbers of SM1 flame by numerical methods. The instabilities of SM1 and SM2 flames were investigated by Ranga Dinesh et al.28 In this paper, the coaxial air microjet is added to the SM1 flame based on Sydney burner. In order to analyze the combustion characteristics of microjet assisted SM1 flame, the numerical models must be validated firstly. Accordingly, the realizable k-ε model and the modified standard k-ε model accompanied by the flamelet model are utilized for solving the SM1 turbulent flame in the first part. The numerical results are compared with the experimental data. Then, two variables of air microjet velocity and swirl number are numerically investigated to study the flame characteristics and analyze the potential ability of microjet.

2. Description of flame configuration and boundary conditions The Sydney swirl burner21-23 was tested to investigate the bluff-body stabilized swirling non-premixed flame. The burner has 3.6 mm diameter central fuel tube with a 50 mm diameter


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bluff-body surrounding it. The swirl air is injected through a 5 mm wide annulus around the bluff-body. The swirl burner is centrally located in an air co-flow with an area of 130 mm square. In order to investigate the effect of air microjet on the swirl burner, the SM1 flame is used as a reference case to verify the numerical model in current computation. An axisymmetric computational geometry has length and diameter of 350 mm and 130 mm respectively, as shown in Figure 1. The coaxial air microjet with a diameter of 1 mm is added to the Sydney swirl burner along the center of the fuel inlet. The dimension of air microjet is in conformity with the microjet assisted Brookes and Moss configuration reported by Chouaieb et al.9, 10. In order to give better agreement with the experimental data, fuel and swirl air inlet shift upstream 50 mm, such approach has been applied in previous work, such as Yang et al.24, Kashir et al.26 and Yang et al.29. 350 mm

(a)

Bluff-body Fuel

1.8 mm

30 mm

Swirl Air

65 mm

Outlet

Coflow Air

25 mm

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Axis

(b)

(c)

Y Fuel inlet

Fuel inlet Air microjet

X

Figure 1. Schematic diagram of flame configuration: (a) SM1 flame configuration; (b) Partial view of fuel inlet; (c) Partial view of air microjet assisted SM1 flame. There are four parameters that determine the flame structure: the axial velocity of central fuel jet Uf, the external ambient coflow air velocity Ue, the axial velocity of annular swirl air Us and the rotational velocity of annular swirl air Ws. The strength of swirl is gauged by the swirl number, defined as the ratio of the axial flux of angular momentum to the axial flux of axial


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momentum.30 However, the geometric swirl number Sg (Sg=Ws/Us) is utilized for evaluating the swirl strength in this research.21 Flow conditions of SM1 flame are listed in Table 1. The operational pressure is 101325 Pa and the temperature of fuel and air inlet is 293 K. The outflow boundary condition type is used for outlet. For such fuel jet velocity, the mass flow rate is 2.221 × 10 −4 kg/s. With regard to different air microjet assisted SM1 flames, the fuel mass flow

rate is not changed to keep the flame stability. Meanwhile, the boundary conditions of coflow air and outlet are identical in all calculated cases. Table 1. Flow Conditions of SM1 Flame. Flame SM1

Fuel CH4

Us (m/s) 38.2

Ws (m/s) 19.1

Uf (m/s) 32.7

Ue (m/s) 20

Sg 0.5

3. Numerical models An approach to the numerical simulation is performed on the ANSYS FLUENT platform. The steady incompressible RANS equations are utilized for turbulent combustion flow. The conservation equations for mass, momentum, energy are numerically solved. 3.1 Governing equations. The equations in RANS can be written as follows:31 ∂ ( ρ ui ) = 0 ∂xi

(1)

 ∂ ∂p ∂  ∂ui ∂u j 2 ∂uk ( ρ ui u j ) = − ) − ρ ui′u′j  + + − δ ij µ( ∂xi ∂x j ∂xi  ∂x j ∂xi 3 ∂xk 

(2)

where −ρ ui′u′ is Reynolds stresses, related to the mean velocity gradients based on the Boussinesq hypothesis, which is given by:32

− ρ ui′u′j = µt (

∂ui ∂u j 2 ∂u + ) − δ ij ( ρ k + µt k ) ∂x j ∂xi ∂xk 3

(3)

where µt is the turbulent viscosity, given as:


6

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µt = ρ C µ

k2

ε

(4)

where k is the turbulence kinetic energy and ε is the turbulence dissipation rate.

3.2 Turbulence model. The standard k-ε model and the realizable k-ε model are used to solve the turbulent flow fields in current research. In standard k-ε model, there are two model transport equations for k and ε respectively determined as:33 ∂ ∂  µt ( ρ kui ) =  µ + ∂xi ∂x j  σk

µt ∂ ∂  ( ρε ui ) =  µ + σε ∂xi ∂x j 

∂u j  ∂k  − ρε  − ρ ui′u′j  ∂xi  ∂x j 

(5)

∂u j ε  ∂ε  ε2 ′ ′ − C2 ε ρ  − C1ε ρ ui u j  ∂xi k k  ∂x j 

(6)

The model constants C1ε , C2ε , C µ , σ k , σ ε have the default values of 1.44, 1.92, 0.09, 1.0, 1.3, respectively. In particular, Pope Correction34 is employed in the standard k-ε model in this paper. The C1ε constant is set to 1.60 instead of 1.44. This modified k-ε model has been proved to be more accurate in comparison with the experimental data in a round jet combustor.35 In contrast with the standard k-ε model, the turbulent viscosity and the transport equation for the dissipation rate are modified in the realizable k-ε model.32 The realizable k-ε model is preferable for rotating homogeneous shear flows and strong adverse pressure gradients. Therefore, the realizable k-ε model and the modified standard k-ε model are both utilized. The turbulent model of the minimum error solution compared with the experimental data is applied to calculate the air microjet assisted SM1 swirl flames.

v 3.3 Radiation model. The spectral radiation intensity I λ ( s ) at point s and direction Ω in an emitting-absorbing-scattering medium can be solved by the radiative transfer equation (RTE), which is given by:36


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κ dI λ ( s ) = −κ aλ I λ ( s ) − κ sλ I λ ( s ) + κ aλ I bλ ( s ) + sλ 4π ds

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v

v v

∫ π I λ ( s, Ω′)Φ λ (Ω′, Ω)d Ω′ 4

(7)

where I bλ ( s ) is the blackbody spectral radiation intensity, κ aλ and κ sλ are the spectral v absorption coefficient and the scattering coefficient of the gas medium with particles, I λ ( s , Ω ′)

v v v is the incident spectral radiative intensity in Ω′ direction, and Φ λ (Ω′, Ω ) is the scattering phase function. In CH4-air combustion, soot consists of optically small particles, the scattering has generally been neglected. Consequently, the RTE can be determined as: dI λ ( s ) = −κ aλ I λ ( s ) + κ aλ I bλ ( s ) ds

(8)

To solve the RTE, the DOM based on the spatial discretization and angular discretization is carried out by using the finite volume method. More details of DOM have been given by Modest.37 The absorption coefficient κ a is calculated as the sum of the absorption coefficients of gas ( κ gas ) and soot ( κ soot ):

κ a = κ gas + κ soot

(9)

The weighted-sum-of-gray-gases model38 is used to calculate the absorption coefficient of gas.

κ gas

I   ln 1 − ∑ aε ,i (1 − e −κi pz )   = −  i =0 z

(10)

where z is the path length related to the geometry, p is the total partial pressure of carbon dioxide and water vapor, κ i and aε ,i are the constant absorption coefficient and the emissivity weighting factor of the ith gray gas. The weighting factor is given by:


8

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J

aε ,i = ∑ bε ,i , jT j −1

(11)

j =1

where bε ,i , j are the emissivity gas temperature polynomial coefficients. Smith et al.39 had fitted the most widely used WSGG model parameters. The absorption coefficient of soot species contributed to the mixture can be expressed by the following equation:40

κ soot = 1232.4 ρ sootYsoot 1 + 4.8 × 10 −4 (T − 2000) 

(12)

where ρ soot is the soot density, Ysoot is the soot mass fraction. The radiative source term S rad is the form of the RTE applied to energy equation. It is calculated as:

v S rad = −∇ ⋅ qr = κ a ( ∫

Ω= 4π

v I ( Ω)d Ω − 4π I b )

(13)

v where qr is the radiative heat flux.

3.4 Combustion model. The steady flamelet model is chosen for non-premixed combustion in this work. The GRI 2.11 detailed kinetic mechanism (49 species and 277 reactions)41 is incorporated into the turbulent flame. The mixture fraction space equations are solved, which are:26

ρχ ∂ 2Yi 2 ∂f 2 ρχ ∂ 2T 2 ∂f

2

−

1 cp

∑H ω i

i

+

+ ωi = 0

ρχ ∂c p (

2 c p ∂f

(14)

+ ∑ c p ,i i

∂Yi ∂T ) =0 ∂f ∂f

(15)

where Yi and ωi are the ith species mass fraction and reaction rate, f is the mixture fraction, the scalar dissipation χ is given by:


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χ( f ) =

2 2 a s 3( ρ ∞ ρ + 1) exp  −2 ( erfc −1 (2 f ) )    4π 2 ρ ∞ ρ + 1

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(16)

where as is the characteristic strain rate, erfc −1 is inverse complementary error function and

ρ ∞ is the density of the oxidizer stream. In the flamelet model, the turbulent flame is treated as an ensemble of discrete diffusion flamelets. The probability density function (PDF) is utilized for calculating the average scalar properties (mass fraction, temperature):27

φ=∫

∞

0

1

∫ φ( f , χ 0

st

) p( f , χ st )df d χ st

(17)

Note that the statistical independence assumption is adapted between f and χ st , therefore: p ( f , χ st ) = p ( f ) p ( χ st )

(18)

The β-shape PDF is used for the mixture fraction and the scalar dissipation. The mean value of scalar dissipation rate is modeled as:

χ st =

Cχ ε f ′2 k

(19)

where C χ has a value of 2.0. In addition, the average enthalpy H form of energy equation is also calculated:24 ∂ ∂  µ µt ∂ H  ( ρ ui H ) = )  + S rad ( + ) ∂xi ∂xi  σ σ t ∂xi 

(20)

where µ and µt are the laminar viscosity and turbulent viscosity, σ and σ t are the laminar Prandtl number and turbulent Prandtl number.

3.5 Soot model. Combustion of hydrocarbon fuels is often accompanied by the formation of soot which makes the combustion efficiency declined. Meanwhile, soot has great power to affect the radiative heat transfer as well as the combustion process. Therefore, the one-step soot formation


10

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Energy & Fuels

model42 is utilized in turbulent swirl flames. This soot model has been successfully applied in previous studies considering combustion systems of CH4-air.9, 10, 43 The model only solves the soot mass fraction transport equation, which is given as: ∂ ∂ µt ∂Ysoot ( ρ uiYsoot ) = ( ) + ℜ soot ∂xi ∂xi σ soot ∂xi

(21)

where Y soot is soot mass fraction, σ soot is turbulent Prandtl number for soot transport. ℜsoot , net rate of soot generation, is equal to soot formation minus the soot combustion:

ℜsoot = Cs p fuelφ r e− E / RT − min [ℜ1, ℜ2 ]

ℜ1 = AρYsoot

ε k

ε  Y  Ysootυ soot ℜ2 = Aρ  ox      υ soot   Ysootυ soot + Y fuelυ fuel  k

(22)

(23)

(24)

where C s is soot formation constant, p fuel is fuel partial pressure, φ is equivalence ratio, r is equivalence ratio exponent, E/R is activation temperature, A is constant in the Magnussen model,

Yox and Y fuel are the mass fractions of oxidizer and fuel, υ soot and υ fuel are the mass stoichiometries for soot and fuel combustion.

3.6 Numerical methods and grid independency. The 2D, axisymmetric, pressure based and steady solution algorithm is implemented in FLUENT 15.0. The pressure-velocity coupling is accounted for SIMPLE method. The PRESTO! Scheme is used for pressure. Other governing equations all use the second upwind scheme. The convergence criterion is 10-6 for energy and soot, and 10-3 for all other equations. Grid quality has great influence on the numerical results. In order to get accurate simulation and reduce computation time, grid independency must be executed to acquire preferable grid


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Energy & Fuels

structure. The non-uniform, orthogonal grid is used, with a high resolution near the fuel and air inlet. Figure 2 shows the computational grid of SM1 flame with 82300 cells. A comparison of the axial temperature distribution between 5100 to 115700 grid cells is shown in Figure 3. With the encryption of the grid, the temperature distribution becomes closer. It’s clear that the axial temperature distribution is distinct in 5100 grid cells. Nevertheless, after 82300 grid cells, the temperature difference does not exceed 2% compared with 115700 grid cells. As a result, the grid mesh composed of 82300 cells is selected as the preferable grid structure. 60

Y (mm)

30

0 -50

0

50

100

X (mm)

Figure 2. Computational grid. 2000 5100 cells 20550 cells 82300 cells 115700 cells

1600

Temperature (K)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

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1200 800 400 0

0

50

100

150

200

250

300

350

Axial distance (mm)

Figure 3. Grid independency solutions.

4. Results and discussion In this section, the SM1 flame is calculated by using two turbulent models to validate the


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accuracy of numerical simulations. We provide the contrasts of axial and rotational velocities, temperature and mixture fraction between the numerical results and the experimental data. After that, the combustion characteristics of air microjet assisted swirl burner are numerically investigated. First of all, we attach importance to probe the characteristics of microjet assisted swirl burner by changing the microjet velocity. Subsequently, different swirl numbers are simulated to analyze the combustion features.

4.1 Validation of numerical model. Recirculation zone is an extremely important feature of the swirl flame. Just as measured by the experiment21, there are two recirculation zones in SM1 flame. Therefore, the precise simulation of two recirculation zones is the main criterion for the numerical methods. Figure 4 presents the comparison of axial velocity at different axial positions of X=6.8, 20, 40, 70 mm. On the whole, the simulation results of two turbulent models are in good agreement with the experimental data. There is slight difference in axial velocity distributions of X=40 and 70 mm between the two turbulent models. At the locations of X=6.8, 20, 40 mm, negative axial velocity all appears. These negative axial velocity regions belong to the recirculation zone within the range of 43 mm downstream of the bluff-body observed in the experiment. At the location of X=70 mm, there is a small negative axial velocity of modified standard k-ε model near the axis centerline area. This is the initial section of the second recirculation zone that begins to develop. In this turbulent model, two recirculation zones can be exactly captured, as shown in Figure 8. The first recirculation zone stagnates at about 45 mm downstream of the bluff-body, and the second recirculation zone extends from 68 mm to 106 mm while 65 mm to 110 mm within measurements21.


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Energy & Fuels

60

Experimental data[21] k-ε (realizable) k-ε (modified standard)

40

X=6.8 mm

Axial velocity (m/s)


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X=40 mm 40 20 Experimental data[21] k-ε (realizable) k-ε (modified standard)

0 -20

X=20 mm

20

-20

40


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0

10

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30

40


0 -20

0


10

20

30

40


Figure 4. Axial velocity profiles at various axial locations. The rotational velocity profiles at distinct axial positions are illustrated in Figure 5. Note that the realizable k-ε model and modified standard k-ε model maintain high level of consistency. With the increase of distance from the swirling air inlet plane, the peak rotational velocity at distinct axial positions increases. The max rotational velocity occurs at the axial position of X=57 mm. The rotational velocity distributions of predicted can also coincide with the experimental data as given in Figure 5. Generally speaking, the velocity values are in accord with the experimental data extremely and the two turbulent models can predict the velocity fields of SM1 flame with high quality.


14


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Rotational velocity (m/s)

40 X=6.8 mm

20 10 0 -10

0

10

20

30

40

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30

10 0 -10

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X=40 mm

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Radial distance (mm) Rotational velocity (m/s)

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X=20 mm

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Radial distance (mm) Rotational velocity (m/s)

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Rotational velocity (m/s)

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30

40

40


30

X=70 mm

20 10 0 -10

0


10

20

30

40


Figure 5. Rotational velocity profiles at various axial locations. Figure 6 indicates the comparison of temperature profiles at axial locations (X=10, 20, 40, 55 mm). The temperature profiles are slightly over-predicted for X=10 mm. The same phenomenon had occurred in Luo et al.25 by using LES. For X=20 mm, there is some fluctuation between the simulated values and the experimental data. Agreement can be seen between the numerical simulation of the modified standard k-ε model and the experimental data23. The temperature distribution of the realizable k-ε model has large error compared with the experimental results, and the high temperature region moves downstream. As a consequence, the modified standard k-ε model is more efficient than the realizable k-ε model in predicting the temperature fields.


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Energy & Fuels

2000

2000 X=20 mm

Temperature (K)

Temperature (K)


500 0

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X=40 mm 1500

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Temperature (K)

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X=55 mm 1500

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Figure 6. Temperature profiles at various axial locations. Mean mixture fraction distributions for dissimilar axial locations (X=10, 20, 40, 55 mm) are depicted in Figure 7. The predicted mixture fraction distributions are related to the temperature profiles, and the results of modified standard k-ε model are closer to the experimental values than the realizable k-ε model. At the axis centerline, the prediction of mean mixture fraction is somewhat higher than the experimental data apart from X=55 mm of the modified standard k-ε model. It’s evident that the mixture fraction is slightly underestimated for radial 4 mm to 25 mm at X=10 mm and 20 mm. Kashir et al.26 had got the identical simulation results by using the Reynolds Averaged Simulation method.


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0.8

Mean mixture fraction

1.0 X=10 mm

0.6 0.4 0.2 0.0

0

10

20

30

1.0

0.6 0.4 0.2 0.0

40

Experimental data[23] X=20 mm k-ε (realizable) k-ε (modified standard)

0.8

0

Radial distance (mm) 0.8 Experimental data[23] k-ε (realizable) k-ε (modified standard)

0.6

X=40 mm

0.4 0.2 0.0

10

20

20

30

40

0.8

30

40


0.6

X=55 mm

0.4 0.2 0.0

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Radial distance (mm) Mean mixture fraction


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0


10

20

30

40


Figure 7. Mean mixture fraction profiles at various axial locations. From the above comparisons of axial and rotational velocities, temperature and mean mixture fraction, numerical results of the modified standard k-ε model agree well with the experimental data generally. Unfortunately, there is slight deviation of simulation by comparison with experiment. Nevertheless, the numerical method is sufficient to provide accuracy for the simulation. As a consequence, the modified standard k-ε model accompanied with the flamelet model are adopted for investigating the combustion characteristics of SM1 flame assisted with an air microjet.

4.2 Effect of air microjet velocity. Experimental study8 had illustrated that the microjet can alter the flame length and emissions in buoyant non-premixed flames. The hydrodynamic nature is the main impact of the microjet in the non-swirl burner. But, in the swirl burner, the effect of microjet on combustion characteristics has not been studied effectively. In this section, we dedicate to analyze the effects of a series of microjet velocities on the combustion characteristics.


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The different numerical cases are detailed in Table 2.

Table 2. Simulation Details of Different Microjet Velocities Cases Case I II III IV V VI VII

Vj (m·s-1) 0.1 1.4 3 10 50 100 115

mj (kg·s-1) 9.425×10-8 1.319×10-6 2.827×10-6 9.425×10-6 4.712×10-5 9.425×10-5 1.084×10-4

ms (kg·s-1)

mf (kg·s-1)

3.96×10-2

2.221×10-4

Figure 8 depicts the axial velocity contours and streamlines for different air microjet velocities of the microjet assisted swirl burner. We can clearly see that there are two recirculation zones in Figure 8a, which is highly consistent with the experiment. When the microjet is introduced to the SM1 swirl burner, the hydrodynamic characteristics of the flame can be changed by microjet velocity. As displayed in Figure 8b, the third recirculation zone exists in the vicinity of the microjet inlet with low microjet velocity. As a matter of fact, the mass flow of microjet is not sufficient for the near fuel flow, which generates a return of the downstream fluid. Figure 8c is the partially magnified view of the third recirculation zone in Figure 8b. The third recirculation zone has the range of 0.4 mm downstream of the microjet inlet at 0.1 m/s microjet velocity, and it disappears while the microjet velocity reached 1.4 m/s, as presented in Figure 8d. After that, the second recirculation zone becomes smaller as the microjet velocity is raised. The second recirculation zone disappears when the microjet velocity beyond 115 m/s. It is detrimental to the combustion under this circumstance. Consequently, the higher microjet velocity should be avoided in practical applications.


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Energy & Fuels

Figure 8. The axial velocity contours and streamlines for different microjet velocities. Figure 8c is the partially magnified view of the markings in Figure 8b.


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Temperature profiles along the axial centerline of various microjet velocities assisted swirl burner are illustrated in Figure 9. The inner flame grows in the microjet downstream region by comparison with the SM1 flame. As exhibited in Figure 9, the temperature near the microjet nozzle rises with the increase of the microjet velocity when the velocity is less than 1.4 m/s, and the temperature distribution along the central axis diminishes first and then increases to the primary peak temperature. Owing to the presence of third recirculation zone, the mixing process proximity to the microjet nozzle is enhanced, and the increased microjet air promotes the methane combustion reaction to occur quickly. Therefore, the temperature near the microjet nozzle rises sharply. When the velocity is more than 1.4 m/s, the temperature in the vicinity of the microjet nozzle decreases with the increase of velocity. After the velocity reaches 10 m/s, the temperature in the vicinity of the microjet nozzle is the same as the temperature of the air microjet inlet. In this velocity range (Vj > 1.4 m/s), the axis temperature distribution has a second peak value because of the inner flame. Actually, the larger air microjet momentum leads to the combustion reaction to occur farther away from the microjet nozzle. Hence, the temperature near the microjet nozzle drops to microjet inlet temperature. As the velocity increases further, the inner flame peak temperature position is farther away from the microjet inlet, and the inner flame peak temperature increases as well. When the velocity exceeds 50 m/s, temperature between the primary peak temperature position and its upstream 30 mm appears to decline compared with low microjet velocity. The primary peak temperature can be changed when the velocity is beyond 115 m/s. At the same time, the second recirculation zone has disappeared. Thus, high microjet velocity is not adopted. Figure 10 plots the inner flame peak temperature locations of different microjet velocities.


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The inner flame peak temperature location (L) can be correlated with the microjet velocity parameter, obtained by exponential fitting. The fitting formula is presented in Figure 10. 2400 Vj= 0m/s Vj=0.1m/s Vj=1.4m/s Vj=3m/s

Temperature (K)

2000 1600

Vj=10m/s Vj=50m/s Vj=100m/s Vj=115m/s

1200 800 400 0 0.1

1

10

100

Axial distance (mm)

Figure 9. Temperature profiles along the axial centerline of various microjet velocities. Inner flame peak temperature location downstream the microjet , L (mm)

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Energy & Fuels

20 Peak temperature locations, L Exponential fitting line

15

10

5

L = 20.56 − 20.63exp( −V j 56.94)

0 0

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Vj (m/s)

Figure 10. Inner flame peak temperature locations of various microjet velocities. The mean mixture fraction distributions are depicted in Figure 11. Mean mixture fraction profiles are obviously decreased with the increase of velocity. The temperature distribution is directly concerned with the mean mixture fraction profile. It is slightly increased of mean mixture fraction between the primary peak temperature position and its upstream 30 mm, which is related to the declined of temperature as shown in Figure 9 when the velocity exceeds 50 m/s. Low mixture fraction means enhancement of mixing, which is advantageous to combustion


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Energy & Fuels

process.

Vj=0m/s Vj=0.1m/s Vj=1.4m/s Vj=3m/s Vj=10m/s Vj=50m/s Vj=100m/s Vj=115m/s

1.0


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0.8 0.6 0.4 0.2 0.0 0.1

1

10

100

Axial distance (mm)

Figure 11. The mean mixture fraction profiles along the axial centerline of various microjet velocities. In what follows, attentions are given to the CO and CO2 mass fractions profiles of different microjet velocities. Figure 12 and 13 reveal the radial profiles of CO and CO2 mass fractions at different axial positions respectively. The predicted CO and CO2 distributions for Vj = 10 m/s are in accord with the flame without microjet. However, a slight difference and a significant difference of CO distribution can be seen for Vj = 50 m/s and Vj = 100 m/s comparing with no microjet. The CO2 distribution is not changed at downstream locations (X>75 mm) except Vj = 100 m/s. Due to the inner flame, augmentation of the CO and CO2 mass fractions is observed near the axial centerline for X=10 mm. In the combustion process, increased CO and CO2 distributions are related to the enhanced mixing process caused by microjet, which promotes the combustion reaction. In general, suitable microjet velocity can make the combustion more efficient and not change the overall CO and CO2 formations.


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0.05 X=10 mm

0.04


0.03 0.02 0.01 0.00 0

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CO mass fraction

CO mass fraction

0.05

X=40 mm

0.04

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0.005 X=150 mm

X=75 mm

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CO mass fraction

CO mass fraction

10


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0.00 0


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0.004 Vj=0m/s Vj=10m/s Vj=50m/s Vj=100m/s

0.003 0.002 0.001 0.000 0


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Figure 12. CO mass fraction profiles at axial locations of various microjet velocities. 0.10 X=10 mm

CO2 mass fraction

CO2 mass fraction

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CO2 mass fraction

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CO2 mass fraction

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Energy & Fuels

X=150 mm 0.08 Vj=0m/s Vj=10m/s Vj=50m/s Vj=100m/s

0.06 0.04 0.02 0.00 0


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Figure 13. CO2 mass fraction profiles at axial locations of various microjet velocities. Soot emission caused by incomplete combustion of flames is detrimental to the environment.


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Energy & Fuels

Simultaneously, the intense absorption of soot particles has a great effect on the radiation heat flux to the combustor.43 Thus, to make the combustion favorable, it’s necessary to reduce the soot formation. Figure 14 indicates the centerline soot volume fraction profiles of different microjet velocities. It’s apparent that there is no difference of soot volume fraction in comparison with the flame without microjet while the microjet velocity is less than 10 m/s. When the microjet velocity is greater than 10 m/s, the soot volume fraction profiles on the axis centerline move downstream, and the peak soot volume fraction gets lower. In general, the reducing of soot formation is ascribed to the mixing enhancement caused by microjet. 0.4

Soot volume fraction (ppm)

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Vj=0m/s Vj=10m/s Vj=50m/s Vj=100m/s Vj=115m/s

0.3

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0.1

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Axial distance (mm)

Figure 14. Soot volume fraction profiles along the axial centerline of various microjet velocities. Practically, the appropriate microjet velocity (10 m/s < Vj

research 1..14 - American Chemical Society

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