7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
Influence of air temperature and evaporation zone length on evaporation and
combustion
Wahidullah Ahmadi, Mouldi Chrigui, Amsini Sadiki and Johannes Janicka
Technische Universitit Darmstadt, Dept. of Mechanical Engineering, Institute for Energy and Powerplant Technology ,
Petersenstr.30, 64287 Darmstadt, Germany
sadiki@tekt.tudarmstadt, ahmadi@ tekt.tudarmstadt.de
Keywords: Spray combustion, Evaporation zone length, EulerianLagrangian method, RANS
Abstract
Using an EulerianLagrangian RANS based procedure under a fully two way coupling a detailed numerical simulation of
kerosene spray combustion has been carried out in a partially premixed prevaporized three dimensional configuration. The
investigations were focused on the flame temperature profile dependency with respect to the length variation of the
prevaporization zone. For the combustion, an approach based on a modified BrayMossLibby model has been adopted to
account for the partially premixed combustion. First, the results have been compared to the experimental data for validation
purposes. A good agreement has been achieved. Then, a fundamental study has been performed by changing the droplet
diameter, the kerosene flammability limits and the location of the combustion initialization. Temperature variations and flame
flashback phenomena have been observed and analysed. All these investigations were performed for atmospheric pressure,
inlet air temperature of 900C and a global equivalence ratio of 0.7.
Introduction
The need of fundamental knowledge of spray formation
characteristics is vital for the development and improvement
of many engine applications. Thereby the droplet
vaporization process and the subsequent fuelair mixture
preparation are of great importance. Poor mixing of the fuel
and oxidant streams leads to mixture inhomogeneities that
affect combustion efficiency and result in enhanced
pollutants formation. In Lean Premixed Prevaporized (LPP)
combustor, the diminishing of the maximum generated
temperature and thereby the reduction of NOx production
are well known challenges. This paper deals with the lean
spray combustion in a chamber where chemical reaction
takes place far away from the injection nozzle. Several
approaches have been published in the literature and aim at
overcoming the flashback and reducing emissions [1]. These
approaches strive low NOx production by designing devices
that permit rapid spray phase transition, mixing and
combustion. Baessler et al. [2] 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. He realized that increasing the prevaporization
zone length reduce the emissions. Shaefer et al. [3] studied
flashback in Lean Premixed Prevaporized (LPP) combustion
of a turbulent kerosene flame. Rocke et al. [4] investigated
venturi LPP on mixing, atomization and evaporation
behavior as well as emission. Nomura et al. [5] 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 prevaporized sprays.
It is worth mentioning that the most data produced in the
literature for studying the NOx reduction and flame flashback
in a LPP context have been achieved for single phase
combustion. Numerical researches that involve two phase
flow, especially sprays under consideration of combustion are
still limited. Therefore we focus in this paper on the analysis
of the flame characteristics with respect to different boundary
conditions for the spray and the carrier phase.
The paper is organized as follows. In the subsequent section
an overview of the mathematical models used and numerical
procedures is provided. To evaluate the prediction capability
of the models proposed, the application configuration under
study will be introduced. In the next section, results obtained
are presented. Different parameter studies are then compared
and discussed, before concluding.
Mathematical models and numerical procedures
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, momentum and concentration. For the turbulence
description, the RNG model which was adjusted for
twophase 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 [6]. 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 104 [7]. Besides
the source terms due to the presence of particles, specific
turbulence modulation model has been introduced and
tested (see in [6] and therein quoted references). Thus a
full twoway 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
Markovsequence dispersion model [8]. 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 [9]. With regard to
evaporation, the socalled Uniform Temperature (UT)
model by Abramson and Sirignano [10] has been applied.
This model is based on the film thickness theory. It does
not consider any temperature variation in the interior of the
droplet (homogenous temperature). The UT model
describes the evolution of the droplet temperature and
diameter, i.e. evaporation rate and energy flux through the
liquid/gas interface.
Combustion modelling
Combustion takes place after fuel evaporation, and
occurs only where vapour and air mix. For partially
premixed prevaporized combustion, one realizes two main
features, namely inhomogeneity of the equivalence ratio
and the velocity of the flame propagation. In the frame of
this work, these two conditions are satisfied, 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. The laminar burning velocity is also
considered as a function of the equivalence ratio. Various
models based on premixed combustion have been extended
to account for such partially premixed features (e.g. [21])
as it will be presented in the following subsection.
Once modified, the BrayMossLibby (BML) model is
a suitable combustion model for the simulation of the
multiphase combustion in this kind of configurations. This
achieved by extending the BML theory through the
coupling with the mixing transport according to [22]. The
transport equation for the progress variable, given by
equation (1), has been solved.
7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
9 (pc) 9(piic) _i \ ~
t + x x Hpcu +w
t x(1)
where c denotes the progress variable defined as:
T T
Tb T (2)
In equation (2) the indices u stand for the unburnt, and b for
burnt part of the flame, respectively. Equation (1) contains
also the mean reaction rate term c which is modelled as
following:
=uS (k32 D3 3/4 D
S0 (3)
where the constant CL is taken to 0.41 in all simulations and
the fractal dimension D is set to 7.7/3. In order to extract the
physical products, results of conditioned combustion
b
products must be multiplied by the probability P of
being behind the flame front.
Y4 ZPPPb (z (4)
In the context of the BML model this probability can be
directly related to the progress variable c, so that
Y(z,Pb)= Y z)c
S(5)
Assuming a bimodal PDF of c it can be shown that [12]
c (,, Z2)
Sb (,z11)
In addition to the Favreaveraged equations of mass,
momentum, and turbulent quantities, a Favreaveraged
mixture fraction equation z and a Favreaveraged
ff2
equation for its variance z have to be considered. The
Favreaveraged equation for the mixture fraction can be
written as follows [11]:
pz a(pu,zT ) a ( 8z .
ati + pD pu z +
9t 9x, 9x, x z
After neglecting the molecular diffusion terms, the final
closed equation for mixture fraction variance yields to:
a(z "2) __ __
+ p +g "2
at ax, ax S 8x, k
(8)
When droplets vaporization occurs due to the local sources
of fuel, the mixture fraction z is not a conserved scalar. It
Ss S,
results in two additional source terms ( "' and P)
"2
appearing in the transport equation of z and z The
S S,
source terms z'P and 'z p for the mixture fraction
(equation 7) and its variance (equation 8) are given in [13]
as follows:
m,, N,
",p 2 7 
S vi (9)
V (9)
s" z 12z)
(10)
Choice of surrogate for kerosene
Jet fuels are composed of hundreds of aliphatic and
aromatic compounds. Developing a reaction mechanism
under the consideration of all this components is a huge
task that is still under investigation. 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.[14]
used Ndodecane as a surrogate for jet fuel. Zhou et al.
[15] studied the thermal decomposition of ndodecane for
comparison with other fuels. Honnet et al. [16] used a
mixture of ndecane 80% and 1,2,4trimethylbenzene 20%
by weight to model jetA. In this study we used a detailed
chemical reaction mechanism for Ndodecane 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. For the
generation of the look up table a presumed pdf has been
considered. The laminar burning velocity for Ndodecane
at the mixture preheat temperature of 400 K has been
measured by Kumar et al. [17] and is given in Figure 1 as a
function of equivalence ratio. The flammability limits are
about 0.7 and 1.4 respectively for the lean and rich
compositions.
nDodecane/Air Mixtures
100 ...............=tK .
90
S80
70
6I
50 .
S40
30
20
0.6 0.7 0.8 0.9 I 1.1 1.2 1.3 1.4 1.5
Equivalcncc ratio I]
Figure 1. Laminar flame speed of ndodecane/air
mixtures with unburned mixture temperature of
360K [17]
Configuration and boundary conditions
7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
The geometry of the configuration for the premixing
and partially premixed prevaporization of kerosene droplets
is shown in Figure 2. The burner is composed of two parts:
the prevaporization 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 prevaporization after being accelerated by the
nozzle and it does entrain the dispersed phase [2]. The
droplets are being injected with a Sauter mean diameter of
50 m. 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 vapor takes
place along the distance L (Figure 2). 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 In/min (normalized 1/min) and its
temperature equals 900C. The kerosene nozzle uses 20
In/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 threedimensional CFDcode 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 blockstructured grid.
The velocitypressure coupling is accomplished by a
SIMPLE algorithm. The whole system is solved by the
SIPsolver. The Lagrangian equations for droplets are
discretized using first order scheme and solved explicitly.
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 underrelaxation 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. In this work the simulations
were performed using monodispersed particles. The overall
mesh for the single annular combustor is about 522 000
control volumes (Figure 3). 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
4 k3/2
e 0.41Ar
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
(2 = 0), since the injected air does not contain any fuel and
the variation of mixture fraction is originated only by the
produced vapor.
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.
Figure 2. Geometry of the
partially premixed
prevaporized combustor
[2]
Figure 3.Numerical Grid
Results and discussion
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 [18]. As operating condition for the reference simulation,
the carrier phase temperature is set to 900C, and the length
of the prevaporization zone equals 0.8 m. The number of
the numerical droplets are 160 000 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 4 shows the temperature distribution along the
axis of the combustion tube, thereby "0.0 m" represents the
position of the ignition source. The flame liftoff agrees
very well with the experimental measurements, whereas
the temperature maximum values show a 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 effected with
a considerable error. This fact was also confirmed by the
experimentalists.
a) Influence of the evaporation zone length
In order to study the influence of the variation of the
7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
prevaporization zone length on the combustion process,
two additional simulations were performed, with L=0.6m
and L=0.7m. The number of grid cells and tracked droplets
are plotted in Table 1. Experimentally the distance for the
prevaporization was adjusted by adding tube elements that
are interconnected by aluminium rings.
2000
16 00 " n m r n
1800
6100 
2400
60O
400 num.rfOnueq
200 ex *
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 04
L nm]
Figure 4. Temperature distribution along the
axis of the combustion tube
In contrast the mass of kerosene droplets which arrive
to the combustion tube and evaporate 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 exhibit less richer mixture whereas close
to the ignition source the concentration of vapor increases
and the mixture strides the flammability limit. Thus the
combustion process starts earlier than the reference one. It is
also to important to note that using shorter prevaporization
zone the maximum temperature increases of AT=100K.
This is to be explained by the fact that the mixture in the
combustion tube was slightly moved toward the
stochiometric 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.
2200 ii
2000 0 ..'' Co o t
1800 ..
1600 Q Ai . O
1400
S1000 numLO.6 
numLO.7 
800 numLO.8 n
600 expLO.6 O
400 a ; expLO.8 
200 '
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4
L [m]
Figure 5. Temperature profiles for different lengths of the
prevaporization zone (num. & exp.)
Figure 5 shows the results of the temperature profile for
different prevaporization length "L". One observes a clear
diminishing of the flame liftoff with decreasing value of
"L"; for reduction of A L=25% the flame reduces its
liftoff with A h=0.05m. Due to the reduction of the
prevaporization length "L", droplets have less time to
endure there, thus the evaporation degree is in turn
reduced.
Table 1. Grid cells and droplet number for different L
L [m] Grid cells Droplet Nbr.
[10001 [1000]
0.6 475 160
0.7 500 160
0.8 522 160
b) Influence of the droplet diameter
The boundary conditions provided by the
measurements are limited for monodispersed droplet with a
Sauter mean diameter of 50 m. The question that rises is
what would happen if droplets change the size
distribution? The diameter size variations happen by
changing the geometry of the nozzle or even by changing
the operating conditions. In gas turbine the Sauter mean
diameter has often a value of 20 / m [19]. Therefore
simulations with a relatively smaller diameter are worth.
Figure 6 presents the temperature profile for the case of a
Sauter mean diameter at the inlet equal 20 / m. One
observes a very high temperature at the beginning of the
combustion tube, which is originated to a flame flash back.
The small droplets (20 1 m) evaporate much faster than
these with 50 m. They are not able to reach the end of
the evaporation zone, thus the mixing takes place in the
early stage of the prevaporization zone exclusively. After
ignition the flame propagates close to the hot wire,
afterwards it come back to the prevaporization zone and
stabilizes there. Unfortunately there was no experimental
data for this class of droplet diameter; however it is of
particular interest to study the limit of the droplet diameter
at which flame flash back occur. One should mention here,
that the evaporation model at the same operating
conditions and same geometry has been already validated
[18].
"00 " . .... .
1600 D 
1400
I o
200D
1000 ..
sn
aoo o
400 ......
200
0 Ms O a015 2 0a25 03 035 04
L Iml
Figure 6. Comparison of the temperature profiles
for droplet diameter 50 u m (refO), 20 u m (dp20)
and exp. (50 P m)
c) Influence of the flammability limit
The flammability limits of fuels depend generally on
different parameters e.g. temperature, pressure etc. The
flammability limits are not absolute, they depend also on
the type and strength of the ignition source, therefore a
7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
slight variation on the flammability limits of dodecane,
which feature different volatility than kerosene is worth to
investigate. The upper and lower limits of the kerosene
surrogate were set to 0.65 and 1.45, respectively.
Figure 7 illustrate the influence of enlarging of the
flammability limits. It is to remark that the flame liftoff has
been reduced with Ah =0.02m, whereas the temperature
profile remained unchanged. Since the new laminar velocity
has values different than zero for mixture starting from an
equivalence ratio equal 0.65, the chemical reaction was able
to arise for a leaner composition and than the combustion
started earlier than the reference case.
The upper flammability limit did not manifest any
influence on the results, because we are dealing with a lean
combustion.
2200
2000
1800
1600 
1400
1200 
000 
WO m f
400 Cmff exp. 
200 I I
S 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4
L [m]
Figure 7. Effects of the flammability limit
enlarging (refO 1 ) on the temperature distribution
d) Influence of the ignition location
Statistically steady turbulent flames require
stabilization mechanisms [11]. Instead of providing
continuous ignition using a heat source, which was
numerically performed by setting the boundary condition
for the progress variable to 1 at the inlet close to the hot
wire ring, one has changed the aspect using spark ignition.
Here it is very important to study where the spark should be
initialized. For that the boundary condition of the progress
variable is set to zero in all inlets and three locations for the
initialization of the progress variable were chosen (see
Figure 8). It is of particular importance to note that the
spark ignition, i.e. progress variable initialization, could not
be placed arbitrarily, namely two observations have been
made. 1) Flame flash back occurs when the progress
variable was initialized behind the hot wire i.e. within
prevaporization zone. 2) The combustion process did not
take place when the progress was initialized at a position
higher than 0.2m the hot wire forwards.
Figure 9 shows the temperature profile for the two
cases of initializations. When flash back happens the
temperature rises up to 1400K at 0.0 m. Figure 10 displays
the progress variable field in a center plane crosssection.
Due to the high value of the progress variable, one clearly
observes the flame flash back and thereby the temperature
jump in the prevaporization zone. In case of initialization
between the hot wire and a distance of 0.2m forwards, one
finds no difference to the reference simulations.

Conclusions
This work studied the influence of the evaporation
degree on the temperature evolution of partially premixed
prevaporized kerosene spray combustion. The numerical
simulations were performed in a RANSspray context
7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
module embedded in an Eulerian Lagrangian approach. The
BrayMossLibby (BML) combustion model was extended
for the application of the LPP concept. The numerical
simulation of the flame liftoff showed a very good
agreement with the experimental measurements, whereas
the temperature maximum values showed a A T of ca. 250
K.
The decreasing of prevaporization zone length
involved diminishing of the flame liftoff. 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 prevaporization zone the maximum temperature
increased of AT=100K. The experimental data presented
similar results for the temperature enhancement.
By changing the droplet Sauter mean diameter from 50
p m to 20 p m, one remarked a flame flashback. On the
other hand the enlarging of the flammability limits provoked
a flame liftoff reduction of A h =0.02m, whereas the
temperature profile remained unchanged. An augmentation
of the upper flammability limit did not manifest any
influence on the results as dealing with a lean combustion.
The fractal dimension "D", which is a model parameter in
the BML combustion model may demonstrate an important
influence on the prediction of the temperature profile. This
aspect and further investigations are left for future work.
Acknowledments
For financial support we gratefully acknowledge the ESA
consortium, the Deutsche Forschungsgemeinschaft (DFG)
through the Sonderforschungsbereich 568 (project A4) and
the FAUDI Stiftung through the project Nr. 75.
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2200 i i i i
200
1860
1400
1200 
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S80 num.re 
n00 num.refsposi
num.refopos2 
400  exp. 
200 I  I ;, I
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7th International Conference on Multiphase Flow
ICMF 2010, Tampa, FL USA, May 30June 4, 2010
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