Vance Smiljanovski, Norbert Brehm

Transcription

1 ihe AMERCAN SOCETY OF MECHAMCAL ENGNEERS 'Three Park.Avanue,.New York. Y:, CFO LQUD SPRAY COMBUSTON ANALYSS OF A SNGLE ANNULAR GAS TURBNE COMBUSTOR Vance Smiljanovski, Norbert Brehm BMW Rolls-Royce AeroEngines, Eschenweg 11, D-1827 Dahlewitz NOTCE: This material may be protected by copyright law {Title_ 17, U.S. Code) ABST RACT n this paper CFD analysis of the steady two-phase turbulent combusting flow in a single annular low-nox combustor is presented. For this purpose the commercial code CFD-ACE (1998) was used, where Eulerian equations are solved for the gas phase and the liquid spray fuel droplets are treated in a Lagrangian frame of reference allowing for evaporation of droplets and providing source terms for the gas phase. The standard k-e model was used for turbulence and an assumed shape probability density function was used for the instantaneous chemistry in the conserved scalar combustion model. Thermal NOx is assumed to be the only source of NOx production and is decoupled from the gas phase reacting flow and calculated in a postprocessing step. The calculation is done on a block structured multi-domain computational grid. Particular attention has be paid to the detailed modeling of the fuel inject.or having multiple air swirler passages starting from the trailing edge of the air swirler vanes and utilizing up to 4 computational grid cells for the entire model. The model represents the single annular low-nox combustor for the BR7 aircraft engine family, which is based on a Rich Bum - Quick Quench - Lean Bum (RQL) concept. CFD analysis is done for high power reduced take off conditions and is compared with full annular rig test results for the temperature traverse and the integral ENOx. The results imply satisfactory prediction capability for the ENOx and the average radial temperature distribution. The prediction of the details of the temperature traverse is not satisfactory and will remain a challenge for the future. NOMENCLATURE AFR air to fuel mass ratio CF-D Computational Fluid Dynamics D,. turbulent diffusion coefficient ENOx emission index of nitrogenoxide F-C-1 fuel injector-combustor-interface L NOx omf PDF PrT P3 RQL RmF SMD ScT T3o T4 f g k x combustor length nitrogenoxide overall temperature distribution factor probability density function turbulent Prandtl number combustor inlet pressure Rich Bum- Quick Quench - Lean Bum radial temperature distribution factor Sauter mean diameter turbulent Schmidt number average combustor inlet temperature average combustor exit temperature mixture fraction mixture fraction variance turbulent kinetic energy spatial distance scalar dissipation rate equivalence ratio turbulent energy dissipation rate turbulent viscosity 1. NTRODUCTON The single annular combustion chamber for the BR.7 aircraft engine family was developed in an experimental NOx reduction program, which was based mostly on sector and full annular rig tests (Brehm et al., 1997a). Emissions reductions for the classical diffusion flame based RQL combustor were attained by consecutive development of the fuel injector and mixing air ports hardware. Further NOx reduction is envisaged by the double annular axially fuel staged combustor (Brehm et al., 1997b and Brehm et al., 1998). For future developments CFO shall help reducing the number of necessary but costly tests by predicting correct qualitative trends and choosing only the most promising hardware designs. Presented at the nternational Gas Turbine & Aeroengine Congress & Exhibition ndianapolis, ndiana - June 7-June 1, 1999 Downloaded From: on 12/1/217 Terms of Use:

2 Downloaded From: on 12/1/217 Terms of Use: m1m1ii1r111ir CFD support using the PACE code (Jones and Whitelaw, 1982) was already applied in the trend prediction of the ENOx reduction potential of two different single annular combustor standards (Brehm ct al., 1997a). However, because of a primary lack in the geometric modeling capabilities of the code, details of fuel injector geometries could not be modeled properly and the application to an axially fuel staged combustor geometry also failed. Since in the spray combustion process in gas turbine combustors the phenomena arc highly complex and nonlinearily interrelated there arc many possibilities to explain the prediction failures of CFD. One very important parameter in the quality of CFD predictions are the boundary conditions. Generally the mathematical description of the combustion system is not complete because the relevant boundary conditions are not fully known to the accuracy level required to predict critical performance parameters and durability as stated by Mongia (1994). Specifically in gas turbine combustion chambers the fuel nozzle is of major importance for the flow- and mixrurefield in the primary combustion zone. As shown by Fuller and Smith (1993) the CFD results are very sensitive to the fuel nozzle boundary specification. Varying the turbulence level at the fuel nozzle-combustor interface they obtained strongly different pattern factors of the exit temperature traverse using, however, only 7 computational grid cells. For this reason and to obtain better prediction of conditions in the combustor primary zone younger calculations by Lai (1997) include the modeling of the entire three dimensional fuel nozzle passages and recent CFD modeling by Crocker et al. (1998) aims at the calculation from compressor exit to turbine inlet. Apparently, the primary combustion zone is strongly influenced by the near fuel injector flowfield. Modeling of the fuel injector-combustor-interface (F-C-1) could be done by taking profiles of velocities and scalar variables from an elaborate 2D axially symmetric calculation for the air swirler passages and the near fuel injector flow field as was done by Fuller and Smith (1994). This type of boundary conditions is not presented here in favor of the highest level of geometric modeling refinement, which consists of a complete 3D model including the different fuel injector air passages but without air swirler vanes. NOx emissions in an RQL based combustor are influenced mostly by the mixing of the hot combustion products from the rich primary zone with air jets from primary and secondary air ports. The quick mixing dominates the temperature distribution and hence the thermal NOx production. With respect to CFD the mixing process has to be modeled adequately and has to be resolved numerically well enough to capture physical turbulent diffusion and to avoid simple numerical diffusion because of coarse grids. For this reason a grid of about 4 cells is used. Although grid independence cannot be granted it utilized a much better resolution than other published works, e. g. by Lai (1997) and Tolpadi ( 199). n the following paragraphs the geometric and the physical and chemical models arc described within the framework of the CFD code. The boundary conditions are discussed in detail. CFD results for the 3D calculation are presented and are compared with finely resolved 2D axially symmetric calculations for the fuel injector exit plane. The experimental setup is outlined and the experimental data are compared with the CFD results. Some comments on the influence of the turbulent Prandtl and Schmidt numbers are given. 2. MODELS 2.1 Code Description For this study the commercially available code CFD-ACE (1998) was used. ts main numerical features and physical models used for the simulations are: block structured grids finite volumes-second order spatial differencing pressure correction/simplec algorithm fully implicit discretization Lagrangian treatment of droplets standard K-e turbulence model with logarithmic wall functions f, g mixture fraction and variance combustion model with presumed PDF liquid dilute spray model with C12H23 as fuel NOx production model based on oxygen equilibrium. The CFD-ACE code solves the gas equations in Eulerian form whereas the droplets are treated in a Lagrangian formulation with discrete trajectories. The spherical droplets evaporate according to the Uniform Temperature Model (Faeth, 1983) and interchange enthalpy, mass, and momentum with the gas phase and vice versa. The mean local gas temperature is calculated along the lines of the assumed PDF approach (f-g model) (Libby and Williams, 1994) by weighting the mixture fraction dependent thermodynamic equilibrium temperature with an assumed probability density function. This 2-pararneter function solely depends on the local average of the mixture fraction and its variance and was assumed to be a -function. n the transport equation for the variance g the scalar dissipation rate X represents the consumption term and is, as commonly done, modeled by X = 2g elk. Heat radiation models or heat losses have not been included. The negligible effect of the thermal NOx production on the local density allows the decoupling of the computation of the main flow field and the NOx field. Therefore the computation of the thermal NOx field is performed as a postprocessing step whereas the velocity, enthalpy, k, e, f, and g distributions are kept constant. This is done by solving a turbulent convection - diffusion - reaction transport equation for NO assuming the same turbulent diffusion coefficient as for the mixture fraction variable. t is accounted for the turbulent variatios in the NO source term by using a density weighted -PDF on the mixture fraction variable. The mean value of the NO source term is then calculated from the integration of the PDF weighted NO source term over mixture fraction. An extended Zeldovich mechanism was used for the thermal NO formation. Besides quasi-equilibrium for the nitrogen atoms the reaction H+2 = OH+O is assumed to be in equilibrium yielding the concentrations of [OH) and [O]. Typical computation times for a 2 cells model on a RlOOOO SG workstation are about 3 hours. 2.2 Geometric Model Figure shows the overall mesh for a sector of the single annular combustor without fuel injector passages containing about 3 cells. The interface mesh shows, that the geometry is composed of simple cartesian as well as concentric blocks to better resolve the F-C-l. 2

3 All the other mass flow inlets are highlighted. They include heatshields, starter and Z-ring cooling films, as well as primary and secondary air mixing ports. 3. BOUNDARY CONDTONS The thermodynamic combustion chamber inlet boundary conditions, P3o = 1241 kpa, T3 = 842 K, and AFR=4.7, represent reduced take off conditions simulated in the full annular rig test. The mass flow boundary conditions for the cooling air inlets like heatshields and mixing holes are estimated by an in house code. The fuel injector exists of three different air swirlers, inner, outer, and dome air swirler, each one defined prim2rily by its effective area and its swirl angle. They are axially shifted and their shrouds can be further classified by their focuses. Additional swirling air is introduced by the heatshield inlets. The air mass flows for the swirlers are calculated from cold flow rig test measured effective areas, a given pressure differential across the combustor, and the conservation of the total air mass flow through the combustor. The circumferential velocities are simply calculated by the swirler vanes turning angles and the constant axial velocities and the air temperature equals T 3. nlet conditions for the turbulent kinetic energy are calculated using a turbulence intensity of % of the resultant velocity through an inlet. Fig. 1: singl e annular combustor geometry, mesh and, inlets Figure 2 shows the detailed 3 model grid for the fuel injector air passages matching the combustor grid at the Fl-C-1. The additional air passages are modeled by approximately more cells. n comparison to other publications (Lai (1997) uses 1 cells in total for the Allison 7KF combustor) or (Tolpadi (199) uses about cells in total for the GE/SNECMA CFM6 combustor) the single annular combustor is quit well resolved. Given the air mass flow rate the fuel mass flow is given by the AFR. Fuel spray boundary conditions are usually unknown and are very difficult to estimate. To better understand the complex two-phase flow a combination of measurements and simulation is usually applied for cold flow conditions, in which the droplet size and velocity distribution is measured as close to the F-C-1 as possible and later used as boundary conditions for the fuel spray model in the calculations. Experiment and calculation are then compared for locations further downstream (Tolpadi et al., 199). Other authors attempt to approximate the spray boundary conditions by iterating them until calculations and measurements match for locations funher downstream of the F-C-1 (Benjamin and Crocker, 1996). For reacting flow, however, the assessment of spray boundary conditions becomes even more difficult and spray conditions are usually assumed like in (Tolpadi, 199) or varied as parameters like in (Lai, 1997). For the calculations presented in this paper the spray boundary conditions are also assumed based on cold spray test measurements. The fuel spray model uses a Rosin Ramrnler drop size distribution function (Rosin and Rammler, 1933) characterized by a minimum drop size of O.Sµm, an SMO of 2µm, a maximum drop size of 4µm, and a drop size spread parameter of 2.. The droplets are divided into 4 different size ranges and are incroduced into the gas at 36 discrete circumferential injection points equally spaced at the fuel lip radius. The temperature of the fuel droplets is set to T fuel = 3 K and the velocity of the drops is estimated from 2 axially symmetric calculations for the air only. The air velocity in the near vicinity of the fuel exit passage is then defined to be the droplet velocity. Hence, there is no relative velocity between droplets and the air keeping mass and heat transfer between air and droplets to a minimum, which is a good approximation for small droplets. Fig. 2: 3D fuel injector model, inlet passages are dark 3.1 Comparison 2D vs. 3 Figure 3 shows some of the velocity vector field of the corresponding 2 axially symmetric calculation near the exit of the utilized 3 stream fuel injector. Furthermore it shows isocurves of the fuel air equivalence ratio cl> for values between. and 1., and the fuel spray trajectories for droplet diameters between 1 and 2 µm. The resulting Downloaded From: on 12/1/217 Terms of Use: 3

4 Downloaded From: on 12/1/217 Terms of Use: flowfield is highly complex. The fuel spray evaporates and between the airstreams and each air stream and the fuel vapor mixing-shear-layers are created producing large velocity as well as fuel vapor and temperature gradients. For these reasons and to produce a reference calculation the 3D analysis is done with the complete fuel injector model in three dimensions as shov.n in figure 2. The model starts from the swirler vanes trailing edges having the right radial position, except for the inner air swirler, which is shortened for convenience and which uses the velocity and turbulence profiles taken from the 2D calculation some distance upstream of the Fl-C-1. The prescription of the complete 3D distribution at the swirler vanes exit passage would be the next step of refinement but is not done here. Figure 4 shows nondimensional profiles of axial velocity versus nondimensional radius for both the 2D and the 3D calculation. Considering the small number of points in radial direction the 3D calculation captures all shear layers, which are finely resolved in the 2D calculation, quite well implying sufficient resolution of the main features. 4. EXPERMENT S Fig. 3: 2 calculation: velocity vectors, iso curves of equivalence ratio between. and 1. indicating the two flame branches, and spray droplet trajectories for droplet diameters between 1 and 2 µm. The recirculating flow produces a stagnation point region of large turbulent kinetic energy and gives the bulk flow an average cone angle. These details need many grid cells to be well enough resolved, what is still affordable in a 2D calculation however very expensive in 3D. Strong recirculation zones might be within the F-C-1 and a utilization of profiles of velocities and temperatures taken from the 2D calculation might not be straightforward to be used in case of strong backftow. Furthermore gradients would also need many grid cells for sufficient resolution of peak values and to guarantee mass conservation at the F-C-1. Finally the 2D axially symmetric calculation does not include the influence of the mixing jets on the recirculating flow field and the resulting cone angle correctly. 1. o.---"'=r;---""7""-:-""7""-;---;--i ::.:;,_...--'---+--:--r--;---;--;--1.8 P4-ll.-'----'---t1-2 calculation - 3 calculation.2 o.7...,,,...-'-- ; ;'--.,-1 --., ,-T"""-rl -'--l o.s!. l--+--l--:--;;;;--f i.2.4 i ;;=-+--'---'----' ; 1--i.3.2 J_---==--J.-_J_-+-1 i---;----;---,---;... i:--;::;::1" ;::::...-;:=::-F+-1f--. l..--'---.,,c...---'--'---'---''---' M U U U M M 1 axial velocity/u,., Fig. 4: radial profiles of axial velocity from 2 and 3 calculations at the fuel injector exit 13 i %h. <>--<> s9.97"1o h 12..., _,, _6 _ o. -- mean 8 ---l. angle [degrees],(duct height h) TtKJ ' L ; = 12 Fig. : temperature traverse, upper part: measured points for different radial positions given in % of duct height, lower part: contour plot; fuel injector is at the center of the segment; looking upstream The experiments were carried out in a 36 full-annular combustor rig. lnlet pressure p3 is restricted to about 12 kpa whereas inlet temperature T3 and AFR represented realistic engine conditions. The AFR measured by gas analysis was within 1% of the value adjusted on the rig. A sampling probe was located in the combustor exit plane measuring gas concentrations in radial positions, they are 11.7, 31.43, 1.38, 71.7, and percent of the exit duct height. The rake was traversed 4

5 Downloaded From: on 12/1/217 Terms of Use: in the circumferential direction measuring at 34 angular positions. The inlet of the rig was represented by a prediffuser to provide a representative flowfield in the annuli of the combustor. The rig was instrumented with total pressure rakes, static pressure tappings and thermocouples for monitoring the inlet conditions. Turbine bleed was simulated by appropriate orifices at the combustor exit without interfering with the gas sampling. Emissions measurement equipment and calculation procedure is according to SAE Aerospace Recommended Practice 126. For calculating the mean emission-indices of the gaseous species, all radial lines of the sampling rake were linked together. To gain the temperature exit profile of the combustor, the lines of the rake were analysed separately and, by means of balancing the enthalpy, the exit temperature for individual positions were deduced. The exit temperature, called T4, is defined as the arithmetic mean temperature of all measured temperatures and has the value T.io=l681K. The values of OTDF=24.7% and RTDF=l4.1% result from the maximum measured temperature T m:ix= l 889K and the maximum of the circumferentially averaged temperatures, which amounts to l 799K, in nondimensional form (T-T4o)/(T4-T3). Figure consists of an upper and a lower part, both describing the "average 18' sector" from the 2 available sectors. The upper plot shows the discrete points versus angle for the five radial positions. The largest differences from the mean temperature are visible for the radial positions closest to the walls, while the other three radial positions show only small differences among each other. The lower part of figure shows the corresponding isoplot averaging the available data points. The ranges between 12 and l 9K are chosen for better visibility. Some further details of the temperature traverse will be discussed later. of the equivalence ratio for values between and having a maximum value of about 11, the fuel spray trajectories. and the isocurve of zero axial velocity in the fuel injector plane. Due to the high temperatures about the stoichiometric surface the spray evaporates quickly after entering the combustor at about the position of the first Z-rings. The zero velocity isocurve indicates the two recirculation zones, i. e. the central recirculation zone created along the centerline by the swirling flow inside the fuel injector and the dome recirculation zone, which is created by the sudden expansion of the nozzle air flow into the combustor.. CFO-ANALYSES. Fig. 7: velocity vectors in the fuel injector plane Fig. 6: shown are in the fuel injector plane: predicted distribution of equivalence ratio <?> for values between and, isocurve of zero axial velocity, and fuel spray trajectories '.,j -.!, j. 1 : The analysis shows a typical RQL combustor flowfield. The near fuel injector flow field, seen in figure 7, shows the predicted distribution Fig. 8: predicted temperatmo traverse 1 16COr 1= 1sco.fi 14 i4c1 i3j i3col 1?C i /.()(\.:... The flame position indicated by <l>=l shows how large!he primary combustion zone is, how close the flame can come to!he line!' walls, and gives a hint about the 3 geometry of the flame surface. Temperatures are functions of<?> and its variance ranging from values below the inlet air temperature because of the cooling effect of the evaporating

6 :-:: spray up to equilibrium temperature at 4>=1. Especially very close to the cornbustor dome, the results compare well with the detailed 2D calculation, which is not shown here, showing high temperatures on the dome surface. Figure 7 shows the corresponding velocity vectors, which are not scaled with the velocity magnitude. They arc simply used to show the interaction of the primary air jets with the recirculating flow from the primary combustion zone and its penetration length. The result of the mixing process, the temperature traverse at the cornbustor exit, is shown in figure 8. The temperature range lies between 12 and J 9K and is chosen to fit the range of the experimentally obtained values. Two hot spots exceeding 19K are visible for this single sector. They stern from the incomplete mixing process between the hot combustion products and the secondary air jets. Furthermore the low liner wall temperatures because of the Z-ring cooling are visible. 11.7%h ,...---,,...--,---r---' ; e 1 14 CD 1.!!! '----'---'-_.JL L.--'--"'----'----' sector number (1 sector = 18 degrees) 6. COMPARSON 6.1 Temperature Traverse The previously shown temperature traverse in figure 8 was circumferentially averaged. The radial distribution is compared with the test results in figure 9, which shows nondirnensional temperatures, i. e. RTDF, versus percentage of the cornbustor exit duct height. g e CD c. E.!!! f-_.; -=,.;...:::: --+_...,r _, '1'6-"'""'----'-+----'----;>.-44-, sector number (1 sector = 18 degrees) 1o i- -2 t.. l, _,t...j.;t_...:-'-- --:T=""'--:Ex=---...; L"- : t ----:::J min ;! - 3 o ,,_. o---o T Ex,_. --'-+\--i t. T E mean i -4 o----<> max X. l----'..\ CFO - ht r-;--+-l;:==---1--t-li;t-11tl :..:.._-'---'--'----'---'--..::...-Lll height[%] Fig. 9: comparison of RTDF From the test results three curves are shown in figure 9, the minimum temperature for a constant radius T min the circumferential average temperature for a constant radius T mean and the maximum temperature for a constant radius T m ax The curve for T mean is extrapolated towards the liner walls to give the integrated mean T4 temperature. The CFD predicted RTDF distribution lies well between the minimum and maximum curves and cuts through the mean temperature curve four times. t does not have the typical parabolic shape with almost constant curvature as the test results but it shows one single peak at about 6% duct height versus 6% for the test curve. The maximum RTDF and the OTDF values for CFD are 18.6% and 4.8%, respectively, versus 14.1 % and 24.7% for the test results. The error for the OTDF is. like often reported in literature (Bain et al., 1992), quite high. 1.38%h -- CFD i - sec i - sec i - sec i - sec g e sector number (1 sector = 18 degrees) Fig.1: temperature vs. angle for 11.7% (upper}, 31.43% (middle}, and 1.38% (lower} duct height n the following figures and 11 a detailed comparison of the CFD predicted temperature with the test data for the entire circumference is plotted for all radial duct height positions. The measured exit temperatures are given by four curves denoted by sec.1-, sec. 6-1, sec. 11-1, and sec Each curve represents five sectors or 9. 6 Downloaded From: on 12/1/217 Terms of Use:

7 Downloaded From: on 12/1/217 Terms of Use: Hence, the curve denoted sec represents the angular positions between 18 and 21. n the upper part of figure 1 the radial position closest to the inner liner, 11.7% duct height, is shown. The periodic behavior is captured well by the calculation and the angular displacement is quite small. CFD predicts however much (about 2K) smaller minimum temperatures, especially at the interface of two sectors. A lack of precise boundary conditions for the last lower Z-ring might be the reason for the predicted thicker lower cooling film at the exit. n the middle part of figure 1, for 31.43% duct height, the periodic behavior as well as the temperature magnitudes are in good agreement. n the lower part for 1.38% duct height the largest differences between experiment and CFD can be found. The large temperature gradients between the hot and cold spots at the exit and the unmixedness is also visible in figure 8. Since the mixing structures are preserved up to the exit the periodic behavior is badly captured and the OTDF attains the large value. This can only be attributed to the poor mixing prediction of the turbulence model. 71.7% case shows again good periodic behavior but still the influence from hot spots is visible from high maximum temperatures. The radial position closest to the upper liner (89.97% duct height) shows despite of a small angular shift good overall periodic behavior of the CFD calculation and reasonable temperature magnitudes. The known deficiencies of strongly differing temperatures still apply especially where mixing is poorly predicted but the qualitative periodic behavior of CFD prediction and test results is basically in agreement Although the detailed temperature traverse, e. g. the OTDF, prediction can be called not satisfactory, the prediction of the RTDF distribution looks quit promising. The strong influence of turbulence parameters on the OTDF will be discussed later. 6.2 ENOx The experimental results are primarily obtained for wet air and are later corrected to compare with dry air. The ENOx value for dry air is given to be 19.3 g NOx/kg fuel. 71.7%h 1 -cfd 22 r ,..---,---, ; - sec g 19!! 19 = ' 18 CD c.. 18 E 17,gi sec sec sec.16-2 c; it c:> "' )! z.e )( z jjj : >---<icfd = -== Experiment= sector number (1 sector= 18 degrees) x/l g 16!! 16 = 1 ' 1 CD 14 14,gi /, \ 11 : 89.97%h 1 -- CFD - sec.1- l - sec.6-1 ;, sec i - sec ' "'/ 1,.111 / r 1 :. J.! j llt1. ii sector number (1 sector= 18 degrees) J Fig. 11: temperature vs. angle for 71.7% (upper) and 89.97% (lower) duct height Figure 11 shows the temperature distributions for 71.7% duct height (upper part) and for 89.97% duct height Oower part). The '\.. Fig. 12: averaged ENOx along the nondimensional eombustor length Figure 12 shows the mean spatially averaged NOx emissions index along the nondimensional combustor axis spatial distance. Dry inlet air is assumed. While in the primary zone the NOx production proceeds with a smaller gradient the NOx source terms increase after mixing of the primary air jets indicated by the increase in the NOx gradient at about x/l =.4. Finally at about x/l =.8 the ENOx attains almost its exit value. The mixing is almost finished, temperatures are low, and the NOx production reactions occur to be frozen. The predicted exit ENOx of the single annular combustor is 21.3 g NOx/kg fuel, leaving an error of just 1%, which can be called satisfactory. 6.3 Comments on Numerical ssues n the turbulence model the mixing process is described by the turbulent diffusion and its coefficient Dr exceeding the molecular diffusion usually by orders of magnitude. t is well known, that the k-e turbulence model underpredicts the mixing intensity (see Bain et al., 1992 and Fuller and Smith, 1993). Turbulent fuel mixing and heat con- 7

8 ' Downloaded From: on 12/1/217 Terms of Use: duction are linked to the gradients of the mean mixture fraction and the mean temperature via the turbulent viscosity vt and the turbulent Prandtl and Schmidt numbers, which are of the oder of 1. For the calculation done so far both values are set arbitrarily to PrT = ScT = 1. Since they are not fixed values there is a certain degree of freedom in utilizing them. n (Tolpadi, 199) turbulent Schmidt and Prandtl numbers of. are used. Considering the linearized heat equation this would mean an increase of turbulent heat conduction by a factor of 2. Mixing and the clipping of temperature peaks would be enhanced and NOx production would be reduced. To investigate this effect a second calculation using PrT = Ser =.7 is performed. The qualitative temperature and NOx distributions are as presented above. The integral values of OTDF, R'DF, and ENOx are compared in table 1 showing improvement towards the test results. Especially the OTDF is reduced drastically. The relative errors are given in brackets. The maximum temperature reduces by about 1 OOK and the relative error of the OTDF drops by about 4%. The error in RTDF reduces to only 9.2%. Test CFD,Prr= 1 CFD, Prr=.7 OTDF[%] (6%) 3. (23.%) RTDFmax [%] (32%) 1.4 (9.2%) ENOx [g/kg] (1.4%) (-.3%) Tab. 1: Measured and computed OTDF, maximum RTDF, and ENOx, (relative errors in brackets) The error in ENOx drops by only 1%. The NOx concentrations at the exit depend not only on the nonlinear temperature dependent NOx production term but also on the residence time and the particle history along streamlines in the combustor. Since it is an integral value that is not compared point by point with the exit traverse the smoothing effect of integration leaves the ENOx to be less sensitive against the turbulence parameters than e. g. the OTDF. This example shows how sensitive the solution can be against turbulence parameters. The conclusion to be drawn is, that CFD shall preferably be used in a comparative manner until better estimations for turbulent mixing are available. For different hardware designs, as they were investigated e.g. for an axially fuel staged combustor in (Brehm et al., 1998), the geometric model and the grid resolution as well as the physical models should be kept constant to make the trend prediction more reliable. 7. SUMMARY AND CONCLUSONS CFD analysis of the steady two-phase turbulent combusting flow in the BR7 aircraft engine combustor has been performed. A 3 geometric model of about 4 grid cells including the fuel injector air passages has been used. Comparison of a fully resolved 2D axially symmetric calculation for the near fuel injector flowfield with the coarser 3D calculation shows sufficient resolution of the main flow characteristics. Plots of equivalence ratio, fuel spray trajectories, and velocity vectors for the nozzle plane of the combustor, as well as the combustor exit temperature traverse were shown. Comparison with a full annular rig test for reduced take off conditions was performed. Average RTDF values show satisfactory agreement while OTDF values are generally overpredicted. Details of the temperature traverse are captured only qualitatively with respect to periodicity. The agreement for the ENOx between CFD prediction and measurement is reasonable leaving errors of less than 1%. Slight changes in the turbulent mixing model done by variation of the turbulent Prandtl and Schmidt numbers show a large impact on the values of RTDF, OTDF, and ENOx coming closer to the measured data. This sensitivity implies the preferable usage of CFD as a comparative tool between different hardware designs only. The main emphasis has to be on keeping the models, e. g. the constants (Prandtl, Schmidt numbers), unchanged and not trying to fit the combustor exit profiles to measured profiles. REFERENCES: Bain. D. B., Smith, C. E., Holdeman. J. D. 1992, CFD Mixing Analysis of Jets njected from Straight and Slanted Slots into Confined Crossftow in Rectangular Ducts.., AAA Benjamin, M.A., Crocker, D.S., 1996, "Spray Characterization of a Relatively High Flow Simplex Atomizer using Experiment and CFD.., AAA , 32nd AAA/ASMFJSAE/ASEE Joint Propulsion Conference, 1-3 July 1996 Brehm, N., Baker. SJ., Jones, S.P. 1997a. "A Three Step Nox Reduction Programme: Achievements with the Single Annular Low-NOx Combustor for the BR7 engine Family", ASME 97-GT-14 Brehm. N., K.au, H.P., Jones. S.P., 997b. "The BR71, an environmentally friendly engine... CASl-AMual Conference 97, 28-3 April 1997, Toronto Brehm. N., Schilling, Th., Mack, A. Kappler, G. 1998, "NOx Reduction in a Fuel Staged Combustor by Optimization of the Mixing Process and the Residence Time", RTO (AGARD) Symposium on Gas Turbine Engine Combustion, Emissions. and Alternative Fuels; Lisbon October 1998 CFD Research Corporation. 1998, "CFD-ACE Users Manual" Crocker, D. S.. Nickolaus. D., Smith, E. S , "CFO Modeling of a Gas Turbine Combuscor from Compressor Exit to Turbine nlet" ASME 98-GT-184 Fuller. E. J., and Smith. C. E., 1993, "ntegrated CFD Modeling of Gas Turbine Combustors". AAA Paper Fuller, E. J., and Smith, C. E., "CFD Analysis of a Research Gas Turbine Combustor Primary Zone", by AAA Jones, W. P., Whitelaw, J. H., 1982, "Calculation Methods for Reacting Turbulent Flows: A Review.., Combustion and Flame 48: 1-26 Lai, M. K., 1997, "CFD Analysis of Liquid Spray in a Gas Turbine Combustor'', ASME 97-GT-39 Mongia, H. C and Future Direction", AAA , "Combustion Modeling in Design Process: Application Rosin, P., Rammler, E. 1933, "The Laws Governing the Fineness of Powdered Coal", nst. of Fuel, 1933, pp Tolpadi, A.C., 199, Calculation of Two-Phase Flow in Gas Turbine Combustors", Journal of Engineering for Gas Turbines and Power, October 199, Vol. 117/69-73 Tolpadi, A.C., Burrus, D.L. Lawson, R.J..199, "Numerical Computation and Validation of Two-Phase Flow Downstre:U of a Gas Turbine Combustor Dome Swirl Cup... Journal of Engineering for Gas Turbines and Power, October 199, Vol. 117n4-712 Libby, P.A. \villiams, F. A. 1994, "Turbulent Reacting Flows", Academic Press Faeth G. M., 1983, "Evaporation and Combustion of Sprays". Prog. in Energy and Combustion Sci. Vol. 9, pp