Effect of Fuel-Evaporation Zone Length on Spray Combustion in a Gas Turbine Combustion Chamber

Transcription

1 ILASS Europe 2010, 23rd Annual Conference on Liquid Atomization and Spray Systems, Brno, Czech Republic, September 2010 Effect of Fuel-Evaporation Zone Length on Spray Combustion in a Gas Turbine Combustion Chamber W. Ahmadi *, M. Chrigui *, A. Sadiki * and J. Janicka * * Department of Mechanical and Processing Engineering, Energy and Powerplant Technology, TU- Darmstadt, Petersenstr. 30, 64287, Darmstadt Abstract The combustion, especially in gas turbine engines, is strongly affected by the fuel preparation process. This includes the atomization of the liquid fuel, dispersion and evaporation of droplets in a fuel spray and the mixing of fuel vapor and air. As the control, design and optimization tasks are repetitive and costly and involve a large number of configuration variations, CFD may allow to carry out repetitive studies with clearly defined boundary conditions at moderate economical costs. For that purpose, accurate numerical models that need less computational costs have to be made available. This work focuses therefore on RANS-based models with the aim to use the spray module developed by Sadiki et al. [4] to study the effects of the pre-vaporization zone length on the entrainment, evaporation degree and combustion of kerosene sprays in a gas turbine combustion chamber. Introduction In automotive engines or gas turbine combustors, the liquid fuel is supplied with varying physical and chemical properties to form a combustible mixture of fuel vapor and air. Focused on gas turbine combustion chambers, the success of some promising approaches, such as the LPP - or the RQL (rich quench-lean)-concept strategies, that can help to limit gas turbine emissions, depends on a suitable homogeneity of the air-fuel mixture in the reaction zone. To achieve this goal by means of numerical simulations, an accurate determination of droplet and vapor spatial distribution and a reliable control of the interaction between the evaporating and reacting spray with the surrounding turbulent gas flow are prerequisite. As pointed out in [1-30] a considerable amount of works have been done including diverse parameter studies (e. g. [1-7]). However, there are relatively few experimental and numerical results devoted to the effects of droplet evaporation process characteristics on spray combustion [1, 4, 8]. An accurate capture of the turbulence characteristics that strongly influence the spray dispersion is determinant for an accurate prediction of evaporation, that in turn, affects significantly the spray combustion along with the spray flame structure [1-13]. Although in many recent works some numerical calculations have been performed in which the effects of droplet vaporization on spray combustion have been pointed out, they dealt mostly with numerical codes in which the standard k-epsilon model was coupled to quasi-equilibrium evaporation models without a full consideration of turbulence modulation processes [6, 7, 10]. Thereby no reliable prediction of the induced turbulence effect could be expected. This is also the case in LES, where only the standard dissipative expression for twoway coupling is really available in existing formulations of subgrid scale turbulent kinetic equation (e.g. [5, 13,21]). With regard to kerosene sprays [3, 7, 22-25], a large amount of experimental research is being conducted to strive low NOx production by designing devices that permit rapid spray phase transition, mixing and combustion [22]. Baessler et al. [23] studied the NOx emission of premixed partially vaporized kerosene spray flame at atmospheric conditions and found out that dealing with lean conditions a reduction of NOx requires a prevaporization of the kerosene spray and should pass over 50% upstream of the flame. They realized that increasing the prevaporization zone length reduce the emissions. Rokke et al. [25] investigated venture LPP on mixing, atomization and evaporation behavior as well as emission, while Schaefer et al. [24] concentrated on the flashback phenomena. Nomura [26] studied experimentally a partially prevaporized spray burner with monodispersed ethanol droplets to investigate the interaction between fuel droplets and a flame. They investigated the effect of mean droplet diameter, and the entry length of droplets into a flame on the laminar burning velocity of partially * Corresponding author: ahmadi@ekt.tu-darmstadt.de 1

2 pre-vaporized sprays. They found out that in spray streams of 0.7 and 0.8 in the total equivalence ratio, burning velocity increase with strain rates greater than 50 1/s. Most of the previous research dealt with the vaporization process at low temperatures conditions. Studies of the effect of the numerical evaporation models on the prediction of spray combustion are therefore rare. In particular numerical researches that involve two phase flow under consideration of combustion in partially prevaporization are still limited. Therefore this contribution aims to investigate the flame characteristics predictions under partially premixed combustion conditions by varying (1) the evaporation zone length and (2) the environmental temperature. Advanced models for turbulence, turbulence modulation [22], evaporation are combined with suitable combustion model based on an extension of the BML (Bray, Moss and Libby) theory [35]. To demonstrate the ability of the model combination employed in predicting the evaporating spray and droplet properties under combusting conditions, respective results are compared with available experimental data in a generic GT combustor [23].. Mathematical models Carrier phase description The carrier phase is considered as continuum phase and is described using the Reynolds averaging method. For this purpose, the governing transport equations have been solved for mass conservation, momentum and concentration. For the turbulence description, the RNG model which was adjusted for two-phase flows has been considered. Indeed the presence of droplets in the carrier phase may be a source for turbulence dissipation or production. The additional source terms which characterize the direct interaction of mass, momentum and species between particle and carrier gas are explained in detail in [4]. The volume variation of the carrier phase as consequence of the presence of kerosene droplets is neglected. This assumption is acceptable, since the droplet volume fraction is here below 10-4 [6]. Besides the source terms due to the presence of particles, specific turbulence modulation model has been introduced [4, 14, 22] and tested. Thus a full two-way coupling is ensured. Evaporation and dispersion models In the frame of this work, the droplets are captured using the Lagrangian procedure, in which all numerical droplets are tracked by solving their equations of motion that include only the drag and gravitation forces. In order to quantify the instantaneous fluid velocity seen by the droplets and its effect on the droplet distribution one should model the Root Mean Square (RMS) values of the fluid parcel velocity at the droplet location. This can be adequately done using a stochastic Lagrangian process. The model used in the frame of this work is the so called Markov-sequence dispersion model [7]. It is based on the computation of the fluid element instantaneous fluctuation along the particle trajectory using two correlation factors, namely the Lagrangian and Eulerian correlation factors denoting the time and spatial correlation functions, respectively. To avoid the phenomena of droplet immigration to locations having low pressure, a drift correction term has been considered [11]. During the computation two different models for vaporization, the uniform temperature model (eq) and the non-equilibrium evaporation model (neq), were used. The non-equilibrium evaporation model gave the better results; hence we used this model also for the numerical computation of kerosene combustion. Combustion modeling Combustion takes place after fuel evaporation, and occurs only where vapor and air mix. In the configuration under study, since the fuel and the oxidizer necessitate certain time for the mixing, the mixture forms a spatial variation of the equivalence ratio of the fresh gas. For the description of such a partially premixed flame, we base on the BML approach coupled to the mixing transport providing variable equivalence ratio. Accordingly, the laminar burning velocity has to be considered as a function of the equivalence ratio. The first extension of the BML model to partially premixed combustion was reported by Maltsev et al. [35]. Let us outline the main aspect of the modelling procedure. Following [35], the thermochemistry is implemented by means of a single reactive scalar c that is defined as a dimensionless temperature: T Tu c = T T b u (1) where subscript u ad b correspond to unburned and burnt flame side, respectively. Using assumption made in [35] and therein mentioned references for given expansion ratio 2

3 Tb Tu τ = Tu the algebraic relation for the mean density can be derived as ρu ρ = 1 + τ c ɶ (2) A Favre averaged transport equation for the single reactive scalar is given by ( ρcɶ ) ( ρuc ɶɶ) " " j j ( ρc u ) w j c + = + ɺ (3) t x x where the last term on right side expresses the mean chemical source term and the first term, that is the scalar flux vector, is modelled by a well known linear gradient diffusion approximation [35] including a turbulent Prandtl number. For the mean reaction rate term, an algebraic closure based on the flame surface density approach and on the assumption that the flame geometry has a fractal character following [34] is applied: 3/ 2 D 3 3/ 4 2 D ν 1 1/ 4 ɺ k c u L L ε ε ɶ ɶ (4) 0 w = ρ s C c ( c) 0 Where the constant C L is taken to 0.41 in all simulations and the fractal dimension D set to 7.7/3. s L is the laminar burning velocity that together with the expansion ratio depends in partially premixed flames on the equivalence ratio that can be directly related to the mixture fraction, which takes zero value at pure oxidizer and unity at pure fuel. We include this fact extending the BML theory through the coupling with the mixing transport and the chemistry governing the jet fuel. This fuel is composed of hundreds of aliphatic and aromatic compounds. Developing a reaction mechanism under the consideration of all these components is a huge task that is still challenging. One should use therefore a surrogate for kerosene in order to reduce the size of the reaction mechanism. The major components of kerosene are alkanes, aromatics and alkenes, thus surrogates are mainly based on these. Huang et al.[31] used N-dodecane as a surrogate for jet fuel. Zhou et al. [32] studied the thermal decomposition of n-dodecane for comparison with other fuels. Honnet et al. [33] used a mixture of n-decane 80% and 1,2,4-trimethylbenzene 20% by weight to model jet-a. In the present study we used a detailed chemical reaction mechanism for N-dodecane as surrogate for kerosene, describing the combustion process. It involved 57 species and 281 reactions. The Lewis number is set to the unity and the strain rate equals 100/s. In order to take into account the turbulence-chemistry interaction a presumed Beta-pdf has been considered in calculating the mean thermochemical quantities required. The laminar burning velocity for N-dodecane is obtained fitting the experimental data measured at the mixture preheat temperature of 400K by Kumar et al. [34]. The flammability limits are about 0.7 and 1.4, respectively, for the lean and rich compositions. Thus the laminar burning velocity and the expansion ratio are functions of the mean mixture fraction and its variance which are described by the following Favre averaged transport equations: ( ρuɶɶ iz) z " " ρ zɶ ɶ + = ρd ρui z + Sɶ zɶ, p (5) t xi xi xi The last term introduced the contribution of droplet evaporation. As in [3] the turbulent scalar flux term will be modelled by a linear gradient assumption. After neglecting the molecular diffusion terms, the final closed equation for mixture fraction variance yields to: ( ) "2 "2 "2 ρ z ρui z νt z νt zɶ zɶ ε "2 + = ρ + 2ρ cρ z + Sɶ (6) "2 t x z, p i x i Sct 1 x i Sct 2 x i x i k 3

4 When droplets vaporization occurs due to the local sources of fuel, the mixture fraction z is no more a conserved scalar. It results in two additional source terms ( S ɶ z ɶ, p and Sɶ ) appearing in the transport equation of zɶ z "2, p and z "2. Their expressions are given in [12]. Configuration, numerical setup and boundary conditions The geometry of the configuration for the premixing and partially pre-vaporization of kerosene droplets is shown in Figure 1. The burner is composed of two parts: the pre-vaporization zone and the combustion tube. Within the first one, the kerosene is fed to an ultrasonic nozzle. The carrier phase (air) is heated by a set of sinter metal plates. It enters the pre-vaporization zone after being accelerated by the nozzle and it does entrain the dispersed phase [23]. The droplets are being injected with a Sauter mean diameter of 50 µm. They are subjected to heated environment; thereby they change their physical state and evaporate. The kerosene droplets are initialized with 90% slip velocity of the fluid element. The mixing between the air and the vapour takes place along the distance L (0.5-1m). Downstream the mixing zone, a second nozzle (54mm) is used to increase the flow velocity and prevent flame flash back. The main air volume flow rate was 300 ln/min (normalized l/min) and its temperature equals 90 C. The kerosene nozzle uses 20 ln/min additionally air for the amelioration of the spray dispersion during the injection. The mixture is ignited by a hot wire ring. In the wake of this ignition source a stable flame develops and spreads over the cross section further downstream. A three-dimensional CFD-code (FASTEST3D) in which the equations for the gas phase are solved by finite volume method has been used. The diffusion terms are discretized with flux blending schemes on a non orthogonal block-structured grid. The velocity-pressure coupling is accomplished by a SIMPLE algorithm. The whole system is solved by the SIP-solver. The Lagrangian equations for droplets are discretized using first order scheme and solved explicitly using the LAG3D-code. The source terms for the gas phase are computed in each cell with the contributions of all relevant droplets. The interaction between the continuous and the dispersed phase consists in couplings between two codes. After the convergence of gas phase, the gas variables are kept frozen and all the droplets representing the entire spray are injected in the computational domain. Due to the presence of the droplets source terms, the conventional residuals are characterized by a jump after each coupling. To avoid oscillations, an additional under-relaxation technique should also be employed for droplet source terms. The droplet injection is based on a stochastic approach by considering the droplet mass flux and the droplet size distributions obtained from the experimental measurements at the inlet near the nozzle exit. The overall mesh for the single annular combustor is about control volumes (Figure 1). The inlet conditions for the turbulent kinetic energy are calculated using a turbulence intensity of 10% of the resultant velocity through the inlet. The distribution of the dissipation rate is estimated using the expression ε = C 3/ 4 µ 3/ 2 k 0.41 r (7) Here the turbulent length scale was assumed to be equal to the hole s diameter or inlet s opening. The mixture fraction boundary conditions are set to zero at all inlets ( z ɶ = 0 ), since the injected air does not contain any fuel and the variation of mixture fraction is originated only by the produced vapour. Since there was no temperature variation at the inflow, the progressive variable was set to zero at all inlet boundaries except at the position of the hot wire ring for the ignition or stabilization and insurance of continuous combustion. The global equivalence ratio equals 0.7. Results and Discussion First validation in water spray In a first step detailed boundary and inlet conditions concerning the spray distribution and the droplet sizes are required to ensure the required quality of the predictions. therefore we used water in our initial numerical simulations and compared them with experimental data. The code was valided, regarding dispersion and evaporation in ambient conditions, using the measured data of water spray (see figure 2). 4

5 Figure 1. Geometry of the partially premixed prevaporized combustor (left) and its numerical Grid (right) Achievements for kerosene spray Based on the numerical computation of water we investigated the prevaporization and combustion of kerosene along the pre-vaporization and combustion zone. As part of the validation procedure, the results for the evaporation, dispersion and entertainment of the dispersed phase were compared to experimental data and published in [39]. As operating condition for the reference simulation, the carrier phase temperature is set to 90 C, and the length of the pre-vaporization zone equals 0.8 m. The number of the numerical droplets are within one coupling. When plotted, the properties of the dispersed phase featured a smooth profile, i.e. the statistics were completely reliable. Increasing the droplet number did not change the results but enhanced the computing time. Figure 2. Comparison between numerical and experimental results of the Arithmetic Mean Diameter dp 10 (red color) for water, Surface Area Mean Diameter dp 20 (blue color), Volume Mean Diameter dp 30 (green color), and the Sauter Mean Diameter dp 32 (mauve color) at different distance s from the inlet 610 mm (left diagram) and 810 mm (left diagram). Main air temperature is 20 C Figure 3 shows the temperature distribution along the axis of the combustion tube, thereby 0.0 m represents the position of the ignition source. The flame lift-off agrees very well with the experimental measurements, whereas the temperature maximum values show a T of ca. 250 K. This deviation is not originated from the negligence of radiation, since the experimental data were corrected and the effect of heat losses due to radiation were accounted for. Nevertheless, one should mention at this stage, that the temperature measurements were performed using thermocouple elements and the data plotting showed very high fluctuation (of ca. 450 K) with high frequency. This is of course a strict limitation of the thermocouple, and therefore the temperature measurements were affected with a considerable error. 5

6 Figure 3. Temperature distribution along the axis of the combustion tube Figure 4 shows the field of the mixture fraction. The stoichiometric contour corresponds to a value of One observes a lean mixture between the kerosene vapor and the oxidizer in the second part of the configuration (combustion tube). The mixing is mainly influenced by the evaporation degree of the kerosene droplets, which in turn is affected by the carrier phase temperature as well as the length of the pre-vaporization zone. The mixing is concentrated around the centerline of the configuration; close to the wall the mixture fraction equals zero. This is due to the dispersion and entrainment of the dispersed phase. Figure 4. Mixture fraction field Figure 5. Progressive variable field The progressive variable, which distinguishes between the fields of burned and unburned gases, is presented in Figure 5. One observes nicely the form and the position of the flame front. The maximum values of the progressive variable are located along the axis; the flame lift-off that could be easily quantified by the flexion point of the temperature profile in figure 3 equals ca m. The progressive variable has zero values in the vicinity of the wall because of the poor mixing in the region, i.e. the vapour concentration is very low and the mixture lies out of the lower flammability domain. Influence of the evaporation zone length 6

7 In order to study the influence of the variation of the pre-vaporization zone length on the combustion process, one additional simulations was performed, with L=0.6m. The number of grid cells and tracked droplets are plotted in Table 1. Table 1. Grid cells and droplet number for different lengths, L L [m] Grid cells [1000] Droplet Nbr. [1000] Experimentally the distance for the pre-vaporization was adjusted by adding tube elements that are interconnected by aluminum rings. Figure 6 shows the results of the temperature profile for different pre-vaporization length L. One observes a clear diminishing of the flame lift-off with decreasing value of L ; for reduction of L=25% the flame reduces its lift-off with h=0.05m. Due to the reduction of the pre-vaporization length L, droplets have less time to endure there, thus the evaporation degree is in turn reduced. In contrast the mass of kerosene droplets which arrive to the combustion tube and evaporation there increases. This phenomena influences the mixing process very strongly and a new field of mixture is obtained, namely the first part of the configuration exhibits less richer mixture whereas close to the ignition source the concentration of vapour increases and the mixture strides the flammability limit. Thus the combustion process starts earlier than the reference one. It is also important to note that using shorter pre-vaporization zone the maximum temperature increases of T=100K. This can be explained by the fact that the mixture in the combustion tube was slightly moved toward the stoichiometric value since the amount of kerosene vapour at the second part of the configuration increased too. The experimental data presented similar results of the temperature enhancement. One observes also that the combustion process takes place earlier compared to the reference measurement. Figure 6. Temperature profiles for different lengths of the pre-vaporization zone (num. & exp.) Conclusion This work studied the influence of the evaporation zone length on the temperature evolution of partially premixed prevaporized kerosene spray combustion. A good agreement with experimental data could be achieved. The decreasing of prevaporization zone length involved diminishing of the flame lift-off. This was originated by the reduction of the evaporation degree, which generated different mixture fields, so that the combustion process started earlier. It was also observed that using shorter pre-vaporization zone the maximum temperature increased of T=100K. Furthermore the flame lift-off showed a very good agreement with the experimental measurements, whereas the temperature maximum values showed a T of ca. 250 K. 7

8 Acknowledgments The authors acknowledge the support from the European Space Agency and the FAUDI foundation. References [1] Gökalp, I., Chauveau, C., Simon, and O., Chesneau, X., 1992, Mass transfer from Liquid fuel droplets in turbulent flow, The combustion Institute. [2] Sornek, R.J., Dobashi, R., and Hirano, T., 2000, Effect of Turbulence on Vaporization, Mixing, and Combustion of Liquid-fuel Sprays, Combustion and Flame 120, pp [3] [3] Klose, G., Schmehl, R., Meier, R., Meier, G., Koch., R., and Wittig, S., 2000, Evaluation of advanced two-phase flow and combustion models for predicting low emission combustors, Proc. of ASME Turbo Expo, May 8th-11th, Munich, Germany. [4] Sadiki A., Chrigui M, Maneshkarimi M. R., and Janicka, J., 2004, CFD-Analysis of Conjugate Effects of Turbulence and Swirl Intensity on Spray Combustion in a Single Gas Turbine Combustor, ASME GT Vienna, Austria. [5] Sankaran, V., Menon, S., 2001, LES of spray combustion in compressible high Reynolds number swirling flows, Proc. of the 2nd International Symp. on Turbulence and Shear Flow Phenomena, Stockholm, Sweden. [6] Krebs, W., Gruschka, U., Fielenbach, C., and Hoffmann, S., 2001, CFD-analysis of reacting flow in an annular combustor, Progress in Computational Fluid Dynamics, 1, Nos 1/2/3, pp [7] Zamuner, B., Gilbank, P., Bissieres, D., and Berat, C., 2002, Numerical Simulation of the reactive twophase flow in a kerosene/air tubular combustor, Aerospace Sci. and Tech., 6, pp [8] Abramzon, B., Sirignano, W.A., 1989, droplet vaporization Model for Spray Combution Calculations, Int. J. Heat Mass Transfer, 32, pp [9] Miller, R.S., Harstad, K., and Bellan, J., 1998, Evaluation of equilibrium and non-equilibrium evaporation models for many-droplet gas-liquid flow simulations, Int. Journal of Multiphase Flow 24, pp [10] Hasse, C., Peters, N., 2001 Flamelet Modeling of DI Disel Engine Combustion, International Multidimensinal Engine Modeling Users, Group Meeting. [11] M. Chrigui, A. Sadiki, F. Z. Batarseh, I. Roisman and C. Tropea Numerical and experimental study of spray produced by an airblastatomizer under elevated pressure conditions, Proceedings of ASME Turbo Expo, GT , 2008 [12] Hollmann, C., Gutheil, E., 1998, Flamelet-Modelling of Turbulent Spray Diffusion Flames based on Laminar Spray Flame Library, Combust. Sci. and Tech., 135, pp [13] Menon, S., 2004, CO Emission and Combustion Dynamics Near Lean-Blowout in Gas Turbine Engines, ASME GT , ASME TURBO EXPO,Vienna, 2004 [14] Chrigui, M., Ahmadi, G., and Sadiki, A., 2004, Study on Interaction in Spray between Evaporating Droplets and Turbulence Using Second Order Turbulence RANS Models and a Lagrangian Approach, Progress in Computational Fluid Dynamics, Special issue. [15] Godsave, G.A.E Studies of the combustion of drops in a fuel spray: the burning of single drops of fuel Ln: Proceedings of the Fourth Symposium (International) on Combusion. Combustion Institute, Baltimore, MD, 818±830. [16] Spalding, D.B. 1953a. The combustion of liquid fuels. In: Proceedings of the Fourth Symposium (International) on Combustion. Combustion Institute, Baltimore, MD, 847±864. [17] Hubbard, G.L., Denny, V.E., Mils, A.F., Droplet evaporation: effects of transients and variable properties. Int. J. Heat Mass Transfer 18, 1003±1008 [18] Aggarawal, S.K., Tong, A.Y., Sirignano, W.A., A comparison of vaproization models in spray calculations. AIAA J. 22 (10), 1448±1457 [19] Bellan, J., 2000, Perspectives on Large Eddy Simulation for Sprays, Issues and Solutions, Atomization and Sprays, 10, pp [20] Jackson, R., Davidson, B.J., An equation set for non-equilibrium two phase flow, and an analysis of some aspects of choking, acoustic propagation, and losses in low pressure wet steam. Int. J. Multiphase Flow 9 (5), [21] Bellan, J., Summerfeld, M.: Theoretical examination of assumptions commonly used for the gas phase surrounding a burning droplet. Comb. Flame 33, , 1978 [22] Chrigui, M. and Sadiki, A, Prediction performance of a thermodynamically consistent turbulence modulation for multiphase flows, 3rd International symposium on two-phase flow modeling and experimentation, Pisa, Italy,

9 [23] S. Baessler, K. Moesl and T. Sattelmayer: NOx Emissions of a premixed partially vaporized kerosene spray flame J. of engeneering for gas turbine and power, July 2007, Vol. 129 / [24] O. Schäfer, R. Koch, and S. Wittig: Flashback in Lean Prevaporized Premixed Combustion: Non swirling Turbulent Pipe Flow Study J. Eng. Gas Turbines Power, July 2003 Volume 125 [25] N. A. Rokke and A. J. W. Wilson Experimental and Theoretical Studies of a Novel Venturi Lean Premixed Prevaporized (LPP) Combustor J. Eng. Gas Turbines Power July 2001 Volume 123 [26] M. Chrigui, M. Hage, A. Sadiki, J. Janicka, A. Dreizler Experimental and Numerical Analysis of Spray Dispersion and Evaporation in a Combustion Chamber accepted for publication in Atomization and Spray 2009 [27] Crowe CT, Troutt TR, Chung JN (1996): Annu Rev Fluid Mech 28:11 43 [28] H. Nomura, M. Hayasaki and Y. Ujiie Effects of fine fuel droplets on a laminar flame stabilized in a partially prevaporized spray stream Proceedings of the Combustion Institute Volume 31, Issue 2, January 2007, Pages [29] Oefelein, J.C., Aggarwal, S.K., 2000, Toward a Unified High-Pressure Drop Model for Spray Simulation, Proceedings of the Summer Program 2000, Center of Turbulence Research, [30] Zhu, G.S., Aggarwal, S.K., 2001, Transient Supercritical droplet Evaporation with Emphasis on the Effects of Equation of State, Int. J. of Heat and Mass Transfer, 43, pp [31] H. Huang, D. R. Sobel, L.J. Spadaccini, AIAA (2002), [32] P. Zhou, O.L. Hollis, B.L. Crynes, Ind. Eng. Chem.Res. 26 (1987) 852. [33] S. Honnet, K. Seshadri, U. Niemann and N. Peters: A surrogate fuel for kerosene Proceedings of the Combustion Institute Volume 32, Issue 1, 2009, Pages [34] K. Kumar and C. Sung, Laminar flame speeds and extinction limits of preheated n-decane/o2/n2 and ndodecane/o2/n2 mixtures. Combustion and Flame Volume 151, Issues 1-2, October 2007, Pages [35] Maltsev, A., Sadiki, A., Janicka, J.:Coupling of Extented BML-Model and Advanced Turbulence and Mixing Models in Prediction Partially Premixed Flames. Progress in CFD, an Int. Journal (4), pp , [36] M. Chrigui, A Sadiki, J. Janicka, M. Hage, A. Dreizler: Experimental and numerical analysis of spray dispersion and evaporation in a combustion chamber, Atomization and Sprays, 19(10): , [37] A. Sadiki, M. Chrigui, J. Janicka and M.R. Maneshkarimi: Modelling and Simulation of Effects of Turbulence on Vaporization, Mixing and Combustion of Liquid-Fuel Sprays, Flow, Turbulence and Combustion : [38] W. Ahmadi: Numerische Simulation der Vorverdampfung in Abhängigkeit von der vormischstrecke und der Lufttemperatur, Diplomarbeit (Master Thesis), TU Darmstadt, [39] M. Chrigui, A. Zghal and J. Janicka Droplet behiavior with a LPP ambiance Proceeding of the 4 th International Conference on Advances in Mechanical Engineering and Mechanics 2008 [40] Smiljanovski V. and Brehm N. CFD liquid spray combustion analysis of a single annular gas turbine combustor ASME paper 99-GT