Influence of stabilizer jets on combustion characteristics and NOx emission in a jet-stabilized combustor

c 846- Keywords: Jet-stabilized combustor Finite Volume method Reactive two-phase flow NO formation ese usti lence model is taken on for the flow predictions. An Eulerian–Lagrangian formulation is used for the is exte l boile eous pollutants. The NOx emission is known as a potentially hazardous pollutant in this regard. Thus, control of NOx emission from the combustion process is an important design criterion in modern gas turbine technology. In order to make available reliable design criteria along with suitable tools to meet the modern com- bustion technology requirements, early detailed analyses of differ- ent design variants are essential. A CFD-based analysis is capable of tal works to reveal the influence of some important parameters such as inlet swirl number, air–fuel ratio, and inlet pressure and temperature on the combustion characteristics using both gaseous and liquid fuels. Watanabe et al. [9] have carried out a numerical simulation on the liquid fuel spray investigating the combustion behavior regarding the soot and NO formations. Cameron et al. [10] and Richards and Samuelsen [11] have investigated experimentally the influence of significant factors such as swirl number, fuel effect, primary air and dilution air on turbulent mixing and combustion characteristics of a wall jet can combustor. Arghode and Gupta [12–14] have recently evaluated ⇑ Corresponding author. Tel.: +98 21 7749 1228; fax: +98 21 7724 0488. Applied Energy 92 (2012) 348–360 Contents lists available at lse E-mail address: [email protected] (F. Bazdidi-Tehrani). tion engines [1,2]. Particularly, in IC engines and gas turbine combustors, the fuel–air mixing process plays an important role in the combustion performance and pollutants emission [3,4]. Thermo-physical properties of fuel and air, fuel–air ratio, quality of liquid droplets distribution, fuel–air mixing quality, inlet veloc- ity of air and spray penetration are among the significant parame- ters that influence the performance of a combustion chamber. On the other hand, problems such as acid rain and ozone layer deple- tion have increased the concerns about detrimental effects of gas- Datta and Som [5] have studied a model gas turbine combustor numerically and two-dimensionally. They have investigated the influence of fuel volatility and spray parameters on combustion performance and NOx emission. Also, the effect of pressure and swirl condition on the gas turbine performance and emission has been studied. They have concluded that the pattern factor (PF) of exit temperature distribution is reduced with a decrease in initial sauter mean diameter (SMD). Cameron et al. [6], Heitor and White- law [7] and Cooper and Laurendeau [8] have performed experimen- x Stabilizer jets number Stabilizer jets location 1. Introduction Spray combustion of a liquid fuel bine combustion chambers, industria 0306-2619/$ - see front matter � 2011 Elsevier Ltd. A doi:10.1016/j.apenergy.2011.11.033 two-phase (gas-droplet) flow. The Discrete Ordinates method, adopting its S4 approximation is applied for thermal radiation modeling of the gas phase. It is demonstrated that an increase in axial distance of stabilizer jets from fuel injector results in NOx emission to decrease significantly and conversely it results in thermal power of combustor to enhance slightly. Also, an increase in number of jet holes (with invariable entrance air velocity) causes both the thermal power and NOx emission to enhance. NOx for- mation is shown to be more sensitive to location of stabilizer jet holes rather than its number. As the dis- tance between stabilizer jets and fuel injector increases from 40 mm to 60 mm and then 80 mm, uniformity of temperature profile is improved which could lead to better conditions at the combustor’s downstream section. This situation is valid for smaller number of stabilizer jets. An increase of stabilizer jets number from 4 to 6 and then 8 leads to an enhanced non-uniformity of temperature distribution towards the downstream. � 2011 Elsevier Ltd. All rights reserved. nsively used in gas tur- rs and internal combus- providing deeper insight into the transport processes encountered, particularly when a system is costly to build and to test. In the past years, manyworkers havemade an attempt to under- stand the various aspects of reactive two-phase flow of combustor. Accepted 13 November 2011 Available online 9 December 2011 Finite Volume staggered grid approach is employed to solve the governing equations. The eddy dissipation-finite rate model is adopted for the heat release simulation and the Realizable k � e turbu- Influence of stabilizer jets on combustion in a jet-stabilized combustor Hamed Zeinivand, Farzad Bazdidi-Tehrani ⇑ School of Mechanical Engineering, Iran University of Science and Technology, Tehran 16 a r t i c l e i n f o Article history: Received 29 June 2011 Received in revised form 29 October 2011 a b s t r a c t The main purpose of the pr stabilizer jets on the comb Applied journal homepage: www.e ll rights reserved. haracteristics and NOx emission 13114, Iran nt work is to investigate numerically the effects of number and location of on characteristics and NOx emission in a jet-stabilized combustor (JSC). A SciVerse ScienceDirect Energy vier .com/ locate/apenergy / App Nomenclature a empirical constant a absorption coefficient B Spalding number b empirical constant CD drag coefficient Cl constant number D diffusion coefficient Dh hydraulic diameter dp droplet diameter g gravity h enthalpy I turbulence intensity I radiation intensity Ib black body radiation intensity k turbulence kinetic energy k rate constant L characteristic length mp mass of particle MNO molecular weight of NO Nu Nusselt number turbulence kinetic energy NOx nitrogen oxides H. Zeinivand, F. Bazdidi-Tehrani numerically and experimentally colorless distributed combustion (CDC) as a new approach for high efficiency and ultra low pollution gas turbine combustor applications. Jo et al. [15] have investigated numerically the influence of inlet air temperature and equivalence ratio on spray penetration, mixing quality, subsequent burning, temperature distribution and NOx emission for a given C10H22 fuel in a wall jet can combustor (WJCC). They have shown that with an increase in inlet air temperature, the NOx level is reduced in the primary zone. But, its level is increased predominantly in the rest of the combustor. For larger inlet temperatures, higher flow veloc- ity and consequently deeper jet penetration have been observed. Sharma and Som [16] have numerically studied the effect of swirl strength and gas pressure on the spray penetration in a swirl sta- bilized combustor. Yan et al. [17] have used the Large Eddy Simu- lation (LES) approach in order to simulate the two-phase flow in a gas turbine combustor. They have reported that the number of pri- mary holes and fuel–air ratio have considerable influence on the turbulent reactive flow field, profiles of exit temperature, CO2 and O2 species. King et al. [18] have experimentally investigated a low NOx radial swirler with central fuel injection demonstrating that good flame stability without an excessive NOx penalty is achievable. They also have numerically evaluated the combustor with three different combustion models, namely flamelet, PDF transport and eddy dissipation and mentioned that although the flamelet model predicts the peak temperature better than the n refractive index p pressure PDF probability density function Pr Prandtl number q droplet size spread parameter Q droplets fraction with a diameter less than dp rs scatter- ing coefficient qr radiation heat flux qi generic source term Rz mass reaction rate R universal gas constant r position vector Re Reynolds number T temperature Tmax,e combustor outlet local maximum temperature Tmean,e combustor outlet mean temperature t time SJ stabilizer jet SJP stabilizer jets position s direction vector s/ scattering direction vector Sc Schmidt number SN N-th order discrete directions approximation Ugi instantaneous gas phase velocity up,i droplet velocity ui,j Velocity components u00i;j Velocity fluctuation components jiz diffusive flux of chemical species x, y, z radial, lateral and axial X size parameter Yi mass fraction of species i Greek symbols e dissipation rate of turbulence kinetic energy f distributed random number g, l, n direction cosines lied Energy 92 (2012) 348–360 349 others but the PDF transport with 2-step chemistry predicts the overall flame structure more reliably. Jones and Withelaw [19] Jones [20], McGurik and Taylor [21] and Pauel and Jones [22] have studied the application of numerical methods for the simulation of turbulence mixing, combustion and radiation in a gas turbine combustor. They have accentuated that the accuracy of simulation is dependent on the numerical scheme and also on the flow and its boundary conditions. Kyne et al. [23] have compared the performance of two separate combustion models describing the chemical interactions in an air spray combustor showing that the flamelet approach provided bet- ter results than the mixture fraction/PDF look up table equilibrium model. Foster et al. [24] have demonstrated the capabilities of CFD in the prediction of NOx emission of an aero engine by comparing their resultswith the experiments. Hampartsoumian et al. [25] have investigated the application of CFD in the spray combustion sys- tems. During the last two decades, considerable efforts have been aimed at developing radiative transfer computational modules for use in the CFD codes concerning a gas turbine combustor [26–29]. Many researchers have investigated [30] the influence of cross jets in a Rich-Burn, Quick-Mix, Lean-Burn (RQL) combustor, which was first introduced in 1980 as a strategy to reduce oxides of nitro- gen and later developed in the next generation of aero-propulsion engines. In RQL combustor, cross jets are applied for quick mixing of fuel and oxidizer or for dilution of exhaust gases. Howe et al. lt turbulent viscosity q density of gas qo density of particle r Stefan–Boltzman constant (5.672 � 10�8 W/m2 K4) ji absorption coefficient of i species s viscous stress tensor x weighting factor U phase function Subscripts b black body f fuel z species ox oxidizer p product / App [31] have numerically investigated the mixing characteristics of both reacting and non-reacting conditions in a configuration with simulating the quick mix region of RQL combustor. They have dem- onstrated that the jet-to-mainstream momentum flux ratio has a significant impact on the jet penetration depth while reaction ap- pears to reduce the penetration depth. They have also mentioned that outlet gas temperature and CO2 are increased with an increase in fuel–air ratio. Smith et al. [32] have numerically studied the effect of reducedflowarea onmixing andNOx emissions. They have shown that althoughmixing is unaffected by the reduction in the flow area, but NOx formation is reduced due to a shorter residence time. The main difference between an RQL combustor and a jet- stabilized combustor (JSC) is that a wall jet in the former has an important role in the flame stabilization and there is nomain swirling flow crossing over the stabilizer jets. However, in RQL combustor, cross jets have a significant role in the dilution of rich burn zone and also burning of remaining fuel in the lean burn zone such that the role of flame stabilization in this type of combustor is negligible. Accordingly, the penetration depth of jets in JSC is different from that in RQL combustor. Bauer et al. [33] have carried out various experimental measurements relating to flow behavior and temper- ature distribution in a jet-stabilized model combustor. Kurreck et al. [34] have then compared their numerical simulations with the measured data of Bauer et al. [33]. They have applied the stan- dard k � e turbulence model for the flow predictions and the eddy dissipation model for the heat release simulation. They have also adopted an Eulerian/Lagrangian Method for calculating the two- phase flow. Bazdidi-Tehrani and Zeinivand [35] have numerically investigated JSC employing the presumed PDF model for turbu- lence–reaction interactions and Realizable k � e model for flow behavior predictions. They also have carried out an Sn approxima- tion for thermal radiation heat transfer. They have demonstrated that species and NOx predictions are in relatively good agreement with the experimental results of Bauer et al. [33] and concluded that the thermal radiation mode plays an important part in the pre- diction of NOx concentration. To date, no numerical or experimental studies have been specif- ically reported on the effects of number and location of stabilizer jets concerning a jet-stabilized combustor (JSC). In the present work, the influence of number and location of stabilizer jet holes on the combustion characteristics and NOx emission regarding JSC is investigated numerically. The eddy dissipation-finite rate model is adopted to model the heat release and also to take into consideration the turbulence–reaction interactions. The present re- sults, as much as possible, are compared with the available exper- imental data of Bauer et al. [33]. 2. General equations and applied models Turbulent reactive flows under steady-state and incompressible conditions are governed by Favre-averaging equations, repre- sented as in the following equations [36]: Mass @~qhuii @xi ¼ 0 ð1Þ Momentum @~qhujihuii @xi ¼ @ ~qhu00j u00i i @xi � @p @xj þ @~sij @xi þ ~qgj ð2Þ Energy @~qhhihuii @~qhh00u00i i D~p @~Jzi @uj 350 H. Zeinivand, F. Bazdidi-Tehrani @xi ¼ � @xi þ Dt � @xi þ sij @xi þ qi þ qr ð3Þ Species @~qhYzihuii @xi ¼ � @~qhY 00 zu 00 i i @xi � @ ~Jzi @xi þ _Rz ð4Þ where the viscous stress tensor, sij, is expressed using the following equation: sij ¼ lt @ui @xj þ @uj @xi � � � 2 3 ltdij @ui @xi � � ð5Þ and the diffusive flux of chemical species, z, is defined by the Fick’s law: �Jzi ¼ � lt Sck @Yz @xi ð6Þ The generic source term, qi, is modeled as in following equation: qi ¼ � l Pr @h @xi þ XNc z¼1 Pr Scz � 1 � � hz @Yz @xi " # ð7Þ The Realizable k � e turbulencemodel [37] is employed tomodel the viscous Reynolds stresses, expressed by Eq. (5). Further details including the reasons for choosing this particular model are pro- vided by the authors [35]. The Lagrangian method is applied for modeling of the liquid-phase. It is assumed that the liquid fuel is in- jected into the combustor as a fully atomized spray, comprising spherical droplets of different size. The motions of fuel droplets in the turbulent combustive (reactive) flow field are computed using a stochastic separated flow (SSF) [38]. The variations inmass, veloc- ity and temperature of the droplets are acquired from the relevant conservation equation on a Lagrangian frame as expressed below. Theequationofmotion for adroplet is definedas in equation [39]. dup;i dt ¼ 3 4 CD l qpD 2 p ðUgi � up;iÞRep þ gi ð8Þ where CD is given by Morsi and Alexander [40]. The effect of gas- phase turbulent eddies on the droplet motion is simulated using a stochastic representation of the gas velocity, with length and time- scales chosen to represent the local turbulent eddy properties [41,42]. The instantaneous gas phase velocity, Ugi , is acquired from following equation. Ugi ¼ Ui þ f ffiffiffiffiffiffi 2k 3 r ð9Þ where the fluctuating velocity component is derived from the tur- bulence kinetic energy under the consideration of isotropic turbu- lence and using a normal distributed random number, f. Droplet evaporation can be estimated by solving unsteady heat and mass transfer equations for droplet, with quasi-steady descrip- tions of the surrounding gas field. The evaporation rate of a droplet is estimated by means of following equation. dmp dt ¼ 2pdpqD lnð1þ BÞNu ð10Þ where Nu is derived from the Ranz–Marshal correlation [43]. Nu ¼ 2þ 0:6Re7=2p Pr1=3 ð11Þ However, the present simulation does not take into account the effects of droplet break-up, coalescence processes and a body force such as the gravity. The Rosin–Rammler distribution [44] with ten ranging classes from 10 to 70 lm is applied to perform the influ- ence of air-blast atomizer, which is fixed at the center of the model combustion chamber head. The Rosin–Rammler distribution is de- scribed by following equation. 1� Q ¼ exp�ðdp=XÞq ð12Þ lied Energy 92 (2012) 348–360 where Q is a fraction of the total volume containing droplets of diameter less than dp, and q is considered to be equal to 2.5 [45]. Table 1 represents the liquid fuel droplet size distribution. / App For the combustion reaction model, a one-step global reaction model is adopted, as depicted in following equation. C10H22 þ 312 O2 ! 10CO2 þ 11H2O ðR1Þ The eddy dissipation/finite rate (ED/FR) model is applied to model the heat release and also to take into account the turbulence– reaction interactions. It computes both a chemical Arrhenius rate and a turbulent mixing rate based on the Magnussen–Hjertager [46] equation and then it chooses the lower of the two rates to be in- serted into the species transport equation. In addition, the temper- ature deduced from the enthalpy field is solved from a transport equation (i.e., energy equation). The mixing reaction rates of C10H22 can be considered to be proportional to the turbulence time scale (e/k) as well as to the smallest of the fuel, oxygen, or product concentration. ED/FR model embraces an entire range of Damköhler number [47] and it can be applied for both finite rate and infinitely fast chemistry. The Damköhler number, Da = st/sc, describes the ra- tio of the macroscopic turbulence time scale to the flame character- istic time. When it is less than unity, it means that the chemistry is slow compared with the turbulent flow processes. The mixing reac- tion rate for Eq. (R1) is then represented by: Rfu ¼ q ek min aY fu; a Yox s ; b YP 1þ s � � ð13Þ where s is the stoichiometric mass ratio of oxygen to fuel and empirical constants, a and b, are equal to 4.0 and 0.5. The kinetic reaction rate for Eq. (R1) can be considered as: Rkin�fuel ¼ 2:58e9 � exp �1:25e8RT � � ð14Þ Hence, the lower of Eqs. (13) and (14) determines the reaction rate of (R1). In the present work, the thermal and prompt NO formations are considered and calculated by the finite-rate chemistry. The ther- mal NO formation rate is estimated on the basis of the extended Zeldovich mechanism as follows: N2 þ O() NOþ N ðR2Þ O2 þ N() NOþ O ðR3Þ Nþ OH() NOþH ðR4Þ where a quasi-steady state assumption is taken on for the concen- tration of nitrogen atoms based on the work of Hanson and Salimian [48]. The concentration [O], is given by the partial equilibrium ap- proach as: ½O� ¼ 36:64T0:5½O2�0:5 expð�27123=TÞ ð15Þ According to De Soete [49], the prompt NO formation rate is repre- Table 1 Droplet size distribution [33]. d10% 20 lm d50% 32 lm d90% 50 lm H. Zeinivand, F. Bazdidi-Tehrani sented by following equation. d½NO� dt ¼ fckprompt½O2�a½N2�½C10H22� exp � ERT � � ð16Þ where a is the order of reaction and fc is a correction factor depen- dent on the air to fuel ratio and fuel type. Also, kprompt = 6.4 � 106 m3/gmol � s and E = 72,500 cal/gmol. The NO reduction is modeled by reactions proposed by Bowman [50]. The NO source term due to the formation/destruction of the thermal NO, prompt NO and NO reburning can be calculated by: SNO ¼ MNO d½NO�thermaldt þ d½NO�prompt dt � d½NO�reburning dt � � ð17Þ where k1 = 1 � 108 m3/gmol � s, k2 = 1.4 � 106e�550/T m3/gmol � s, k3 = 2 � 105 m3/gmol � s for 1600 K 6 T 6 2100 K. A joint two-vari- able probability density function (PDF) [35] in terms of temperature and oxygen concentration is employed. The mean formation rate, SNO, can be determined by: SNO ¼ ZZ SNOðT;YO2 ÞP1ðTÞP2ðYO2 ÞdT dYO2 ð18Þ where the shape of P1(T) and P2ðYO2 Þ is approximated by a b � PDF. In the liquid fuel combustion process, the thermal radiation mode is the dominant mechanism of energy transport to the sur- rounding surfaces. The radiative transfer equation (RTE) for an absorbing, emitting and scattering medium, at position, r, and in the direction, s, is given as [51]: dIðr; sÞ ds ¼ �ðaþ rsÞIðr; sÞ þ an2 rT 4 p þ rs 4p Z 4p 0 Iðr; s00ÞUðs � s00ÞdX00 ð19Þ In the present simulation, The Discrete Ordinates method (DOM), adopting its S4 approximation [52,53] is employed to calcu- late the thermal radiation. The DOM is based on the discretization of the RTE, according to a chosen number, Ndir, of discrete direc- tions cosines li, gi, ni, (always satisfied as: ni2 + gi2 + l2i ¼ 1) asso- ciated with their respective weights, wi, contained in the solid angle, 4p. In another word, the integral over the solid angle that appears in Eq. (19) is replaced by a weighted summation of inte- grand evaluated at discrete, conveniently selected, ordinates, si. The Gaussian quadrature provides the approximations.Z 4p /ðs; s00Þ IðsÞ dX ffi XN j¼1 xjIj/ðsj; siÞ ð20Þ where Ij is a shortcut for I(sj). Thus, the original integro-differential RTE becomes a system of N linear differential equations, an equa- tion for each discretized ordinate. Moreover, defining Si ¼ aIb þ rs4p X xjIj/ðsj; siÞ ð21Þ Bi ¼ aþ rs � rs4pxi/ðsi; s 0 iÞ ð22Þ the problem is changed to solve the following system of differential equations: dIi dl ¼ �BiIi þ Si 1 6 i 6 N ð23Þ where l is a length parameter that runs along each direction defined by si. The equation for energy source (or sink) per unit volume is de- scribed as in following equation. r � qr � a 4pIb � X j xjIj ! ð24Þ The Weighted Sum of Gray Gases Model (WSGGM) is taken onto compute the absorption coefficient with regard to the pressure, temperature and species variants. The total heat flux is then esti- mated by adding the heat flux of the gray gases after multiplying by certain weight factors. However, the total gas emissivity is eval- uated by the summation of a number of terms, each one being the lied Energy 92 (2012) 348–360 351 multiplication of a weighting factor and a gray emissivity. The total emissivity and absorptivity are computed from the following equa- tions [54]: e ¼ XNg i ae;iðTÞ 1� e�jiPS � � ð25Þ a ¼ XNg i aa;iðT; TxÞ½1� e�jiPS� ð26Þ where ae,i, aa,i, ji, s are weighting factors, absorption coefficient of the ith gray gas and path length, respectively. 3. Geometry and boundary conditions 60 mm downstream of the head plate. In order to study the influ- Fig. 2 depicts the three-dimensional geometry of the model combustion chamber applying a structured mesh throughout. The computational domain comprising a single-block is meshed with hexahedral cells. A mesh of the highest concentration is em- ployed at the inlet of fuel injector, jet holes and reaction zone. 4. Discretization procedure A Finite Volume [36] staggered grid approach is employed to solve the governing equations including mass, momentum, energy and species in addition to the turbulence transport, fuel combus- tion and radiation model equations. The second-order upwind Fig. 3 displays the independence of present results on the profiles tab ×Ø 352 H. Zeinivand, F. Bazdidi-Tehrani / Applied Energy 92 (2012) 348–360 60 (mm) y z D =8 0 (m m) S 4 Air-blast atomizer x ence of number of stabilizer jets, a modification in the current set up of the model combustor is necessary such that 6 and 8 stabilizer jets (with the same overall cross sectional area as the reference four jets) could be investigated. Also, the location of stabilizer jets could be altered from 60 mm to 40 mm and 80 mm from the air- blast atomizer, downstream of the head plate, so as to evaluate the influence of location. The liquid diesel fuel is injected at the head of combustor by means of an air-blast atomizer which is simulated with a central fuel inlet of inner diameter of 0.7 mm and an annular inlet for fuel atomization of inner and outer diameters of 1 mm and 3 mm, respectively. In order to solve the present problem numerically, the standard wall function is employed for the combustor walls treatment. The boundary conditions for the inlet and outlet of the computational domain are mass flow rate and pressure outlet, successively. The computations are sensitive to the inlet boundary values, particularly to the dissipation rate of turbulence which can- not be determined directly. Nevertheless, a first approximation for the k and e may be achieved from the assumed forms as in the fol- lowing equations [36]: ~kinlet ¼ 32 e U * inlet It� �2 ð27Þ ~einlet ¼ C3=4l ~k3=2inlet 0:07Dh ð28Þ For the radiation calculations, the walls are assumed to be gray of emissivity equal to 0.7. Further details on the present geometry, boundary conditions and fluid properties are presented in Table 2. Wall The present model combustion chamber (reference geometry), as represented schematically by Fig. 1, is chosen consistent with the experimental setup of Bauer et al. [33]. Its diameter and length are 80 mm and 400 mm, respectively. A part of the primary air en- ters the combustor through four circumferential perpendicular injection jet holes as stabilizer jets of 8 mm inner diameter, L= Fig. 1. Schematic view of model of CO2 concentration and axial velocity at z = 1.75D frommesh size. For the CO2 concentration, altering the number of cells from 3.7 � 105 to 8 � 105 (i.e., the finest grid) and also to 2.8 � 105 (i.e., the coarsest grid) maximum deviations of almost 2.17% and 3.6%, consecutively, occur. Thus, the grid of 3.7 � 105 cells is considered to be the optimum and it is applied throughout the present work. 5. Results and discussion In order to verify the accuracy of the present simulations, sev- eral comparisons are made between the present results and the available experimental measurements by Bauer et al. [33]. Some important local parameters such as temperature and species mass concentration including pollutants formation are considered for this purpose. Then, the influence of variation of stabilizer jets num- ber and location are assessed. 5.1. Verification of present simulations Fig. 4 shows the present temperature profiles across the com- bustion chamber diameter at two different axial positions. A Outlet ilizer jet holes 8 (mm) scheme is taken on for the space derivatives of the advection terms in all of the transport equations. The flow field pressure linked equations are solved using the SIMPLEC algorithm [55]. A pres- sure-based algorithm, on the basis of an Eulerian–Lagrangian for- mulation, is applied to simulate the spray combustion. The convergence criterion requirement is set to be 1 � 10�6 for energy and about 1 � 10�5 for the other terms of the transport equa- tions. This means that the maximum value of the normalized resid- uals of any equationmust be less than that of the set value. The grid spacing in the axial direction (z) is varied smoothly to reduce any fall of the accuracy of the discretization scheme, in such away that high- er concentration of nodes is allocated near the inlets, reaction zone and the walls. The influence of mesh size on the present computa- tions is examined extensively by considering different mesh sizes, in the range of 2.8 � 105–8 � 105 cells. As a typical set of results, 400 (mm) combustion chamber [33]. .5 = 60 u = 0 / App Table 2 Input conditions and fluid properties. Geometry Fuel inlet zone (mm) from r = 0 to r = 30 Atomization air (mm) from r = 0.5 to r = 1 Jets (mm) Diameter r = 8 at z Combustor radius (mm) 80 Combustor length (mm) 400 Inlet boundary conditions Liquid fuel Mass flow rate (kg/h) 1 Turbulence kinetic Energy (m2/s2) – Dissipation rate of Turbulence (m2/s3) – Temperature (K) 295 Pressure outlet (bar) 1 Walls No-slip condition: H. Zeinivand, F. Bazdidi-Tehrani reasonable agreement with a difference of less than 6% on average exists between the present results and the available experimental data at both values of z. At z = 1.225D, the temperature in the near wall region is higher than that in the central zone. The reason for this is that the core of combustor, between the axial distances of 0.07 and 0.15 m (i.e., between z = 0.875D and z = 1.875D), displays lower temperatures because of the injection of cold air jets. Also, the hot combustion products leave the primary zone through the space in between four jets without any strong mixing. At z = 1.75D, the temperature distribution becomes more uniform due to the mixing of cold flow of air through stabilizer jets and hot products of combustion. Further downstream, with an increase of mixing, the mean weighted average temperature decreases and reaches a constant value at the exhaust. Fig. 5 represents the radial distributions of CO2 and O2 concen- trations at two values of z. The oxygen rich zone is near the inter- section of four injection air jets (i.e., at z/D � 0.75). The CO2 concentration is slightly over-estimated by the present simulation, Composition (mass fraction) O2 0 N2 0 C10H22 1 Z X Y Fig. 2. Model combustion chamber’s g mm Atomization air Four jets 1.2 33.5 39 6 17,200 850 295 295 , v = 0, w = 0 lied Energy 92 (2012) 348–360 353 as compared with the experiment [33], for z = 1.225D. However, the present results show better capturing of the experimental data by moving further downstream toward the exhaust region (i.e., z = 2.8D). The concentration of CO2 is 5.8% on average at z = 2.275D, being very close to the experiment. Maximum devia- tions of present CO2 concentration from experimental data at z = 1.225D and z = 2.275D are 6% and 1.8%, respectively. On the other hand, the O2 concentration is moderately under-predicted by the present computations at z = 1.225D. The maximum O2 con- centration presently predicted at z = 1.225D is equal to 14% which is almost 10.7% less than that of the experiment. The presently pre- dicted O2 concentration at z = 2.275D shows a maximum deviation of 3.6% from the experiment. At z = 2.275D, the present concentra- tion of O2 is about 11% on average displaying excellent conformity with the experimental measurements. Fig. 6 illustrates the variations of NO pollutant concentration with axial distance along the central axis of model combustion chamber. The present predictions of NO are fairly close to the 0.2315 0.2315 0.7685 0.7685 0 0 eometry with a structured mesh. y (m 12 z=1.75D / App A xi al v el oc it 2 4 6 8 10 /s) 14 16 18 20 280000 cells 370000 cells 570000 cells 800000 cells z=.175D (a) 354 H. Zeinivand, F. Bazdidi-Tehrani experiment [33]. However, at the axial distance of z � 0.75D, fresh air injected by the stabilizer jet holes reduces the NO concentration locally. In the downstream section, hot products of combustion zone (with rich NO concentration) are mixed with fresh air so that NO concentration in the rest of combustor remains constant. The predicted NO concentration in the exhaust region (i.e., z = 2.8D) is approximately equal to 9 ppm showing reasonable agreement with the experimental data. 5.2. Evaluation of thermal and flow behavior Fig. 7 depicts the most important regions of the model combus- tion chamber. Temperature distribution along with streamlines at the center-line section, y = 0, are demonstrated. Two major recircu- lation zones are appeared at approximately z = 0.03–z = 0.05 m, due to the presence of cross jets aimed at stabilizing the flame. Also, a stagnation point in the central region of cross jets (i.e., z � 0.06 m) is observed. The recirculation zone enhances mixing of burned and burning gases with the incoming air and fuel. Thus, Radius (m) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 0 Fig. 3. Independence of present computations from mesh size Radius (m) Te m pe ra tu re (m ) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 600 800 1000 1200 1400 1600 1800 Present study Experimental [33] z=1.255D Fig. 4. Profiles of temperature at CO 2 Co nc en tra tio n 0.07 0.08 0.09 0.1 0.11 280000 cells 370000 cells 570000 cells 800000 cells (b) 0.12 lied Energy 92 (2012) 348–360 a mechanism of continuous ignition is established and combustion can be sustained over wide ranges of pressure, velocity, and fuel/air ratio. It is also observed that a reversed flame is formed up just be- fore the injection jets due to the existence of a recirculation zone. This reversed flame is extended downstream through the region in between the discharge of four injection air jets. Also, the hot gases leave the primary zone through this region without any strong mixing. The core of combustor, between the axial distances of z = 0.07 m and 0.15 m represents lower temperatures due to the injection of cold air jets. In the downstream section ðzP 0:14 mÞ, the hot and cold gases are mixed together creating a combustion gas with an average temperature of approximately 1200 K remaining almost constant throughout that section. Tem- perature distribution in the further downstream is more uniform. 5.3. Influence of number of stabilizer jets As mentioned in Section 3, an increase in the number of stabi- lizer jets from 4 to 6 and 8 while keeping the entrance air velocity Radius (m) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 0.06 : (a) axial velocity and (b) CO2 concentration at z = 1.75D. Radius (m) Te m pe ra tu re (K ) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 600 800 1000 1200 1400 1600 1800 Present study Experimental [33] z=1.75D two different axial positions. Co nc en tr at io n (% ) 10 15 20 Present study (O2) Experimental (O2) [33] Present study (CO2) Experimental (CO2) [33] z=2.275D Co nc en tr at io n (% ) 10 15 20 Present study (O2) Experimental (O2) [33] Present study (CO2) Experimental (CO2) [33] z=.1225D H. Zeinivand, F. Bazdidi-Tehrani / Applied Energy 92 (2012) 348–360 355 5 constant by maintaining the same overall cross sectional area, necessitates a decrease in the diameter of jets from 8 mm to 6.5 mm and 5.65 mm, consecutively. Fig. 8 displays the tempera- Radius (m) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 0 Fig. 5. Profiles of O2 and CO2 concen z/D N O (pp m) 0 1 2 3 4 5 0 2 4 6 8 10 Experimental [33] Present study Fig. 6. Variation of NO concentration with axial distance (center-line, x = y = 0). Axial d R ad iu s ( m) 0 0.1 -0.04 -0.02 0 0.02 0.04 300 400 500 600 700 800 900 1000 110 Fig. 7. Temperature (K) distribution along wit ture variation of air along a stabilizer jet flow path starting from the hole entrance on the periphery at a radius of 40 mm (i.e., r/ R = 1) to the stagnation point on the central axis of combustor (i.e., x = y = 0, r/R = 0), and for three different number of jet holes, namely, 4, 6 and 8. The jets location (i.e., distance between stabi- lizer jets and fuel injector) is kept fixed at 60 mm. An increase of stabilizer jets number from 4 to 6 and then 8 results in the en- trance air temperature to increase noticeably. This in turn leads to the pre-heating of air which has a positive influence on the com- bustion of fuel and oxidizer in the reaction zone. Fig. 9 demonstrates the radial temperature distributions for the model combustor with 4, 6 and 8 stabilizer jet holes, at various val- ues of z. In the vicinity of jet holes (i.e., at z = 1.225D), an increase in the jet holes number from 4 to 6 and then 8 causes the temper- ature near the wall region to also increase by almost 11.2% and 18.9%, respectively. Whilst, the temperature in the central zone is decreased by nearly 7.7% and 15.3% as the jet number is raised Radius (m) -0.03 -0.02 -0.01 0 0.01 0.02 0.03 0 5 trations at two axial positions. from 4 to 6 and then 8. A similar situation arises at z = 1.75D. In the downstream section (i.e., at zP 2:275D), due to the mixing of hot flow on the sides (i.e., supplied from the reaction zone) and cold flow of central region, the temperature distributions be- come more uniform. Fig. 10 shows, with the aid of streamlines, how an increase in the number of stabilizer jets enhances moderately the dimension of recirculation zone in the region just before the location (i.e., at z = 0.06 m) of jets. A relatively enlarged recirculation zone reason- ably augments the rate and volume of mixing between reactant istance (m) 0.2 0.3 0.4 0 1200 1300 1400 1500 1600 1700 1800 1900 2000 h streamlines at center-line section, y = 0. and product and it aids to increase the total reaction rate of fuel and oxidizer. The maximum reaction rate for combustors with 4, 6 and 8 stabilizer jets is computed to be equal to 0.08, 0.13 and 0.145 kmol/m3 s, respectively. When the number of jets is raised, the difference between the temperature of cold flow of central re- gion and hot flow of side region becomes larger and this leads to an enhanced non-uniformity of temperature distribution towards the downstream, as also displayed in Fig. 9. From another point of view, as depicted by Fig. 11, with an in- crease of stabilizer jets number, the fraction of flow moving in the central region increases drastically and the cold flow moving to- ward the downstream concentrates more in the central region. At z = 1.225D, the peak of axial velocity profile in the central region increases by almost 22.3% and then 38.7% as the jets number aug- ments in turn from 4 to 6 and then 8. Further downstream, at z = 2.275, the velocity profiles become comparatively more uni- form and it is also observed that as the jets number is raised the uniformity of velocity distribution declines, in line with Fig. 9. Fig. 12 illustrates the distributions of NO concentration for three different numbers of jet holes of 4, 6 and 8, at various axial dis- tances. At all of the present values of z, increasing the number of jets leads to an increase of NO concentration. This is particularly true in the near wall region where higher temperatures exist, as r/R Te m pe ra tu re (K ) 0 0.2 0.4 0.6 0.8 1 300 350 400 450 500 550 600 650 4 SJ 6 SJ 8 SJ z=0.75D Fig. 8. Temperature variation of air along a jet flow path from entrance to stagnation point. Te m pe ra tu re (K ) 1000 1200 1400 1600 1800 4 SJ 6 SJ 8 SJ z=1.75D Radius (m) Te m pe ra tu re (K ) -0.02 0 0.02 800 1000 1200 1400 1600 4 SJ 6 SJ 8 SJ z=2.275D Radius (m) Te m pe ra tu re (K ) -0.02 0 0.02 600 800 1000 1200 1400 1600 1800 2000 4 SJ 6 SJ 8 SJ z=.1225D Fig. 9. Influence of stabilizer jets number on temp 356 H. Zeinivand, F. Bazdidi-Tehrani / Applied Energy 92 (2012) 348–360 Radius (m) -0.02 0 0.02 800 Te m pe ra tu re (K ) -0.02 0 0.02 800 1000 1200 1400 1600 1800 4 SJ 6 SJ 8 SJ z=2.8D Radius (m) erature distribution at various axial distance. 5.4. Influence of stabilizer jets location Fig. 13 depicts the radial temperature distributions for three dif- ferent stabilizer jets locations (i.e., the distance between stabilizer jets and fuel injector) of 40, 60 and 80 mm, at two values of z. The number of jet holes is kept constant at 4. At z = 1.225D, as the jets position alters from 40 to 60 mm and then 80 mm the temperature of gases near the wall region of combustor rises by nearly 15% and 38%, respectively, and at the same time the temperature in the cen- tral zone decreases significantly by almost 13.5% and 50%. How- ever, the temperature in the downstream (i.e., z/D = 2.8) shows a slight decrease with an increase in the stabilizer jets position from 40 to 80 mm. It can also be seen that as the distance between sta- bilizer jets and fuel injector becomes larger, the uniformity of tem- perature profile is improved, as compared with z = 1.225D, which could lead to better conditions at the combustor’s downstream section. Fig. 14 demonstrates the influence of stabilizer jet holes loca- tion on the NO concentration distribution, at two values of z. As the distance between jets and fuel injector increases from 40 to R ad iu s (m ) 0 0.01 0.02 0.03 0.04 4 SJ R ad iu s (m ) 0.01 0.02 0.03 0.04 6 SJ H. Zeinivand, F. Bazdidi-Tehrani / Applied Energy 92 (2012) 348–360 357 0 u s (m ) 0.02 0.03 0.04 shown in Fig. 9. At z = 1.225D, the maximum NO concentration raises from 13 to 27 ppm as the jets number augments from 4 to 8. Similarly, at z = 2.8D, the maximum NO concentration moves up from 9 to 17 ppm. It is also observed that the uniformity of NO distribution lessens as the number of stabilizer jets increases from 4 to 8. This situation is similar to that of the temperature dis- tribution, as discussed in Figs. 9 and 10. Axial distance (m) R ad i 0 0.02 0.04 0.06 0 0.01 8 SJ Fig. 10. Streamlines in the recirculation zone for various number of stabilizer jets. Radius (m) A x ia l v el oc ity (m /s) -0.02 0 0.02 0 5 10 15 20 25 30 4 SJ 6 SJ 8 SJ z=1.255D Fig. 11. Influence of stabilizer jets number on ve 80 mm, the NO concentration decreases significantly. For all the three jet locations under study, NO concentration has a uniform distribution in the downstream section (z = 2.8D). However, the NO concentration drops from about 26 to 6 ppm as the jets location alters from 40 to 80 mm. Table 3 shows the influence of number of stabilizer jet holes on the combustion characteristics such as thermal power and NO emission at the exhaust of the model combustor (i.e., z = 5D). An increase in the jets number causes both the NO concentration (see Fig. 12) and thermal power to enhance. The NO concentration and thermal power for the case of six jet holes are almost 34% and 20%, respectively, more than the case of four stabilizer jets. These figures for eight jet holes against a combustor with four jets (as ref- erence number) show a rise of about 70.2% and 29.2% for NO con- centration and thermal power, successively. It is mainly due to an enhanced mixing and also a relatively enlarged recirculation zone (see Fig. 10). Moreover, a growth in the jets number leads to the difference between maximum temperature and mass weighed average temperature at the exhaust (i.e., Tmax,e–Tmean,e) to increase noticeably, as also discussed in Figs. 9 and 10. This in turn causes the temperature profile uniformity to decline (see Fig. 9). Finally, a slight raise in the CO2 concentration with jet number is noticed from Table 3. A x ia l v el oc ity (m /s) -0.02 0 0.02 0 5 10 15 20 25 4 SJ 6 SJ 8 SJ z=2.275D Radius (m) locity distribution at various axial distance. Radius (m) N O (pp m) -0.02 0 0.02 0 5 10 15 20 25 30 35 4 SJ 6 SJ 8 SJ z=1.75D Radius (m) N O (pp m) -0.02 0 0.02 0 5 10 15 20 25 30 35 4 SJ 6 SJ 8 SJ z=2.275D Radius (m) N O (pp m) -0.02 0 0.02 0 5 10 15 20 25 30 35 4 SJ 6 SJ 8 SJ z=2.8D Radius (m) N O (pp m) -0.02 0 0.02 0 5 10 15 20 25 30 35 4 SJ 6 SJ 8 SJ z=.1225D Fig. 12. Effect of stabilizer jet holes number on NO concentration distribution at various axial distance. Radius (m) Te m pe ra tu re (K ) -0.02 0 0.02 600 800 1000 1200 1400 1600 1800 2000 SJP 40mm SJP 60mm SJP 80mm z=1.255D Radius (m) Te m pe ra tu re (K ) -0.02 0 0.02 1000 1100 1200 1300 1400 SJP 40mm SJP 60mm SJP 80mm z=2.8D Fig. 13. Influence of stabilizer jets position (SJP) on temperature distribution at two values of z. 358 H. Zeinivand, F. Bazdidi-Tehrani / Applied Energy 92 (2012) 348–360 n NO / App Radius (m) N O (pp m) -0.02 0 0.02 0 5 10 15 20 25 30 35 SJP 40mm SJP 60mm SJP 80mm z=1.255D Fig. 14. Effect of stabilizer jets position (SJP) o Table 3 Comparison of model combustor’s thermal power and pollutants concentration at various number of stabilizer jets, at the exhaust. Stabilizer jets number 4 6 8 NO (ppm) 9.21 12.4 15.68 Thermal power (W) 3203 3859.4 4138.9 Tmax,e–Tmean,e (K) 21.7 46 51.6 CO2 concentration (%) 5.7 5.95 6.31 H. Zeinivand, F. Bazdidi-Tehrani From Table 4, as the distance between stabilizer jet holes and fuel injector goes up from 40 to 80 mm the NO concentration de- creases significantly (see Fig. 14) and the thermal power increases slightly at the exhaust. As the distance reduces from 60 (as refer- ence location) to 40 mm, the NO concentration enhances by about 182% and the thermal power decreases by almost 3.5%. Similarly, a reduction from 80 to 40 mm brings about a noticeable rise of approximately 420% in the NO concentration and a slight drop of 6.9% in the thermal power at the combustor’s exhaust. It appears from the present results that the NO formation is more sensitive to the location of stabilizer jets rather than its number. Also, a neg- ligible change in the CO2 concentration is noted. Furthermore, the present simulations display that with an increase in the distance, the difference between maximum and mean temperatures at the exhaust decreases moderately. This means that as the distance be- comes larger, in as much as the flame stabilization does not expe- rience any trouble, it could lead to better conditions at the exhaust. 6. Concluding remarks 1. As the number of stabilizer jets is raised from 4 to 6 and then 8 (with invariable entrance air velocity), the difference between the temperature of cold flowof central region andhot flowof side region becomes larger and this leads to an enhanced non-unifor- mity of temperature distribution towards the downstream. 2. Increasing the number of jets leads to an increase of NO concen- tration at various axial distances. This is particularly true in the model combustor’s near wall region where higher temperatures exist. At the exhaust, the NO concentration and thermal power for the case of six stabilizer jets are almost 34% and 20%, respec- tively, more than the case of four jets. These figures for eight jet holes against four jets (as reference number) show a rise of about 70.2% and 29.2% for NO concentration and thermal power, successively. Radius (m) -0.02 0 0.02 0 10 concentration distribution at two values of z. Table 4 Comparison of model combustor’s thermal power and pollutants concentration at various stabilizer jets locations, at the exhaust. Stabilizer jets position (mm) 40 60 80 NO (ppm) 26 9.21 5 Thermal power (W) 3093.8 3203 3323.37 Tmax,e–Tmean,e (K) 23.9 21.7 20.3 CO2 concentration (%) 5.73 5.71 5.72 N O (pp m) 20 30 40 50 SJP 40mm SJP 60mm SJP 80mm z=2.8D lied Energy 92 (2012) 348–360 359 3. In the downstream (z = 2.8D), as the stabilizer jets location is increased from 40 to 60 and then 80 mm, the uniformity of tem- perature profile is improved, as compared with z = 1.225D, which could lead to better conditions at the combustor’s down- stream section. 4. 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J Heat Mass Transfer 1984:699–706. [53] Viskanta R, Mengüc MP. Radiative heat transfer in combustion systems. Prog Energy Combust Sci 1987;13:97–160. [54] Hottel HC, Sarofim AF. Radiative transfer. New York: McGraw-Hill; 1967. [55] Vandoormal JP, Raithby GD. Enhancements of the SIMPLE method for predicting incompressible fluid flows. Numer Heat Transfer 1984;7:147–63. Influence of stabilizer jets on combustion characteristics and NOx emission in a jet-stabilized combustor 1 Introduction 2 General equations and applied models 3 Geometry and boundary conditions 4 Discretization procedure 5 Results and discussion 5.1 Verification of present simulations 5.2 Evaluation of thermal and flow behavior 5.3 Influence of number of stabilizer jets 5.4 Influence of stabilizer jets location 6 Concluding remarks References