7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
3D Unsteady Numerical Simulation of the Hydrodynamic of a Gas Phase
Polymerization Pilot Reactor
P. Fede*i, H. Neau**, O. Simonin*i and I. Ghouilai
University de Toulouse : INPT, UPS : IMFT :AllCe Camille Soula, F31400 Toulouse, France
SCNRS :Institut de M~canique des Fluides de Toulouse : F31400 Toulouse, France
i:INEOS: Innov~ne: CTL/PRO Ecopolis Lardra, F13117, Lardra, France
fedeiiimft.fr
Keywords: Fluidized bed, twophase flow, CFD, Euler multifluid approach
Abstract
Threedimensional numerical simulations of polymerization reactor have been carried out. The mean pressure drop
have been compared with experimental data from an existing industrial mediumscale pilot. The results show that a
standard pure slip wall boundary condition for the mean particle velocity leads to an underestimation of the fluidized
bed height. In contrast, a noslip wall boundary condition gives better predictions and the mean pressure distribution
is in better accordance with experimental data. The using of an existing model to account for nonspherical particle
shape shows that the bed height is increasing when nonspherical particle shape are considered.
Nomenclature
Roman symbols
L'D drag coefficient ()
g gravitational constant (i.s2)
Hbell bed height (In)
P, mean gas pressure (N.mn)
q~ mean particle agitation (mn.s )
Re, particle Reynolds number () _
UK,; mean velocity of phase k (m.s )
u< fluctuating velocity of phase k (m.s1)
Greek symbols
asvolume fraction of phase k ()
p, gas viscosity (kg.mng )
Pre density of phase k (kg.mn3)
,F, mean gasparticle relaxation timescale (s)
medical modeling of industrial fluidized bed is challeng
ing because of many complex phenomena taking place:
particlefluid interactions, particleparticle and particle
wall collisions, heat and mass transfers and chemical re
actions.
Since a few years the numerical modeling of flu
idized bed hydrodynamic has been extensively devel
oped. At the same time we have seen the strong de
velopment of the high parallel computing permitting to
perform numerical simulations of practical application
with big meshes (more than ten millions of cells). Then
these both developments allow now 3dimensional real
istic numerical simulations of industrial configurations
by using unsteady Eulerian multifluid approach.
Numerical simulations of industrial and pilotreactor
gassolid pressurized fluidized beds were carried out
with such an approach and have shown a good agree
ment with the qualitative knowledge of the flows but
further detailed experimental validations were needed
(Gobin et al. 2003). Indeed, such numerical approaches
are extensively used for circulating or dense gassolid
fluidized bed predictions but their assessments are usu
ally restricted to a comparison between the predicted and
the experimentally measured mean pressure drop at the
wall. Recently Fede et al. (2009) made a comparison be
tween numerical simulation and experimental data ob
tained by Positron Emission Particle Tracking (PEPT).
Subscripts
Y
p
gas
particle
ith component of a vector
Introduction
Gassolid fluidized beds are used in a wide range of
industrial applications such as coal combustion, cat
alytic polymerization or uranium fluoration. The nu
10,60m
0,222m
7m
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
Table 1: Powder properties and operating points.
RUN #1 #2 #3 #4
Particle properties
Solid mass (kg) 608.3 617.3 640.8 650.5
Density (kg/nl3) 825.6 825.6 825.6 825.6
Mean diameter (pm) 750 750 750 750
Gas properties
Density (kg/nt") 23.6 13.2 26.0 26.8
Viscositvx10" (Pa.s) 1.53 1.74 1.37 1.38
Pressure (bar) 20.4 12.6 21 21
Fluidiz. Vel. (nt/s) 0.48 0.48 0.40 0.56
Geometry and physical parameters
The pilot's geometry is shown by Fig. 1. The pilot re
actor is composed of a cdlindrical column with a diam
eter of 0.74ni and an height of 7nt. In the upper part
of the pilot we find an hemispherical dome of diameter
0.695ni following a widening of height 1.41ni with an
expansion angle of 12 Finally a chimney takes place
for gas outlet. We emphasized that in the real geome
try the chimney is cylindrical but for meshing reasons
we have represented the chimney with a squared section
(see Fig. 2) with an equivalent area.
The powder properties and operating points are given
in Table 1 corresponding to measurement performed on
the experimental pilot. Two operating pressure (RUN#1
& RUN#2) and two fluidization velocities (RUN#3 &
RUN#4) have been considered. In the experiment the
particulate phase is polvdispersed however the numeri
cal simulation have been carried out with monodisperse
particulate phase having a median diameter.
Mathematical Model
Three dimensional numerical simulations of dense flu
idized bed flows have been carried out using an Eule
rian nfluid modeling approach for fluidparticle turbu
lent polvdispersed flows developed and implemented by
IMFT (Institut de M~canique des Fluides de Toulouse)
in the NEPTUNE CFD V1.ll'c Ther version. NEP
TUNE_CFD is a multiphase flow software developed
in the framework of the NEPTUNE project, financially
supported by CEA (Commissariat g l'Energie Atom
ique), EDF (Electricit6 de France), IRSN (Institut de Ra
dioprotection et de Stiret6 Nucl~aire) and AREVANP.
The multiphase Eulerian approach is derived from joint
fluidparticle PDF equation allowing to derive the trans
port equations for the mass, momentum and agitation of
particle phases (Simonin 1996). In the proposed model
ing approach, separate mean transport equations (mass,
momentum, and fluctuating kinetic energy) are solved
for each phase and coupled through interphase transfer
0,74m
Figure 1: Sketch of the polymerization pilot reactor.
The PEPT technique allows to measure the mean par
ticle velocity inside a dense fluidized bed. Fede et al.
(2009) have shown that the wall boundary condition on
the particle velocity may have a strong influence on the
fluidized bed hydrodynamic. They showed that a noslip
wall boundary condition improves the numerical predic
tions of the mean particle vertical velocity component in
the near wall region. However the mean pressure drop
data was not experimentally measured and they cannot
investigate the effect of such a wall boundary condition
on the bed height
In the present study we consider the configuration and
operating conditions of an existing industrial medium
scale pilot employed for ethylene catalytic polymeriza
tion. Several pressure probes measure the gas wall
pressure during the polymerization process, allowing to
estimate the bed height. Numerical simulations have
been carried out in order to analyze the dependence
of the pilot reactor fluidized bed hydrodynamic to the
fluidization velocity, operating gas density and particle
shape. Also a systematic analysis of the effect of the
particle wallboundary condition is performed in order
to complete the analysis of Fede et al. (2009).
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
f
Z
X
Z
Figure 2: 3dimensional mesh composed of 217,668 hexahedra.
According to a large particle to gas density ratio
only the drag force is accounted as acting on the parti
cles. Hence the mean gasparticle interphase momentum
transfer is written:
terms.
In the following when subscript k = g we refer to
the gas and k = p to the particulate phase. The mass
balance equation is:
8 8
8tamcprc+8 agpgpU g,, =0 (1)
where as is the kth phase volume fraction, pa the den
sity and Us~i the ith component of the velocity. In (1)the
righthandside is equal to zero because no mass transfer
takes place.
The mean momentum transport equation takes the fol
lowing expression:
Vr7
Ip~i= Ig~i=appy
where the particle relaxation time scale writes
1 3 p, (vT I) D
Sp 4 pp dp
For solid volume fraction 0, > 0.3 Gobin et al. (2003)
proposed to compute the drag coefficient, CD, with
as*+ acPpag (2)
8x4
Ill= in d,WY, Cd,Erg)
+Ik,i + [ag1pL (U;,.. ;) + ep,23
where the Ergun's drag coefficient is given by:
CD,Erg = 200np
where ..' is the fluctuating part of the instantaneous
velocity of phase k, P, the mean gas pressure, gi the
ith component of the gravity acceleration and la y the
mean gasparticle interphase momentum transfer with
out the mean gas pressure contribution. Finally 8pc,ij is
for k g the molecular viscous tensor and for k p
the collisional particle stress tensor.
and Wen & Yu's correlation by
C ~~ (1+ 0.15Re .687) Q1.7 Re, < 1000
DWY 0.444Re, '> 1000
(7)
arPps dU,i + Us,3 d~i
81 8x,
In case of al, < 0.3 the Wen & Yu's correlation (7) is
employed. The particle Reynolds number is defined by
Rel, = as" , ) (8)
The mean gasparticle relative velocity, V,,;, is ex
pressed in terms of the mean gas velocity, mean particle
velocity and a drift velocity. The drift velocity accounts
for the correlation between the particle distribution and
the turbulent velocity (Balzer et al. 1996). In (2) the col
lisional particle stress tensor is derived in the frame of
the kinetic theory of granular media (Boelle et al. 1995).
For the turbulence modeling we use a standard k e
model extended to the multiphase flows accounting for
additional source terms due to the interfacial interac
tions. For the dispersed phase, a coupled transport equa
tions system is solved on particle fluctuating kinetic en
ergy and fluidparticle fluctuating velocity covariance.
Mesh, boundary condition and numerical
parameters
The 3dimensional mesh is shown by Fig. 2. The mesh'
based on Ogrid technique, is composed of 217,668 hex
ahedra with approximately AJ:r Ay = 13nin and
aZ = 40nun. The time step is typically at 10's.
At the bottom (2 = 0), the fluidization grid is an in
let for the gas with imposed surfacic velocity (QSUS,S)
corresponding to the fluidization velocity 1 7. For the
particles this section is a wall. At the top of the flu
idized bed, we defined a free outlet for both the gas
and the particles. The walltype boundary condition is
friction for the gas. A recent study comparing 3D nu
merical simulations and experimental data from dense
fluidized bed have shown that the particle wall bound
ary condition is crucial for the numerical prediction of
the fluidized bed hydrodynamic (Fede et al. 2009). In
the present study two kinds of boundary condition for
the particulate phase have been tested. First a pure slip
boundary condition,
[U star l= o
wei~ l~l (9)
corresponding to particlewall elastic rebounds on a flat
wall. In (9) U,,,, is the tangential to the wall component
of mean particle velocity and Up,,, the normal to the wall
component of mean particle velocity. Fede et al. (2009)
have shown that noslip boundary'
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
gives numerical predictions in good accordance with ex
perimental measurement obtained by Positron Emission
Particle Tracking. Their analysis has been accomplished
by comparing mean particle velocity profiles at several
heights of the fluidized bed. However, Fede et al. (2009)
were not able to compare the full bed height because
of lack of pressure measurements. We emphasized that
the noslip boundary condition for the mean particle ve
locity and zero flux for the particle fluctuating kinetic
energy is very questionable but could represent elastic
bouncing on the wall with an isotropic angle distribu
tion. This situation could correspond to spherical par
ticles bouncing on very rough wall or to very irregular
particles bouncing on a smooth wall (Konan et al. 2009).
A numerical simulation is carried out as following.
First, at t = 0 the fluidized bed is filled of uniform solid
volume fraction according to the solid mass of experi
ment. A transitory phase takes place for t E [0, 20s]
corresponding to the destabilization of the fluidized bed.
The statistics are computed for t E [20s, 100s] in
suring a statistical convergence. Because of the mesh
and the physical time needed, the numerical simulation
have been performed on parallel computer with 32 cores
(Neau et al. 2010).
Effect of the operating fluidization velocity
In this section we study the effect of the fluidization ve
locity on the lwdrodynamic of the polymerization reac
tor. The physical parameters are given in Table 1 and we
compare RUN#3 and RUN#4. The numerical simula
tions have been performed with pure slip wall boundary
condition for the mean particle velocity and with noslip
boundary condition.
The Fig. 3 shows the wall distribution of the mean gas
pressure, for two fluidization velocities and for the two
kinds of wall boundary conditions. As expected, in the
upper part of the fluidized bed, the pressure wall distri
bution is linear corresponding to the hydrostatic pressure
of the gas. In the lower part (less than 4m) we observe
two different behaviors depending on the wall bound
ary condition. When noslip wall boundary condition is
used we have also a linear gas pressure distribution. In
contrast, for slip wall boundary condition the gas pres
sure distribution is slightly curved.
In a general manner, the fluidized bed height is de
fined as the height where the wall gas pressure distribu
tion shifts. Here the fluidized bed height is determined
by the following expression used by the industrial oper
ators
dP,
Hbed = *3 +(
where the pressure drops have been measured at several
[u~~art= 0
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
4 05
0 2
03 02 0 101
05
Co 4
w ~
k
c
03
I
r
t~c~
~"4 ~i
 IC
I
I.
bV 0 40m/s, shp
SV 0O 56m/s, shp
SV0 40m/s, noshp
SV 0 56m/s, noshp
01 02 03
Vo 40m/s, shp
SVO 56m/s, shp
VO 40m/s, no shp
SVO 56m/s, no shp
0 01 02 03
0 2
03 02 0 10
Figure 4: Effect of the fluidization velocity on the radial profile of the mean solid volume fraction (0, : Vf
0.40m/s and 0, #: Vf 0.56m/s). The empty symbols correspond to the slip wall boundary for the mean
particle velocity and blackfilled symbols to the noslip boundary.
heights: is in good accordance with experimental data.
The Fig. 4 shows the mean solid volume fraction pro
files at different heights. As expected, we observe that
for Vf 0.56m/s the mean solid volume fraction is
lower than in case of Vf 0.40m/s. This effect leads
to higher bed expansion because in both case we have
nearly the same solid mass. Then increasing the fluidiza
tion velocity leads to an expansion of the fluidized bed
and consequently the mean solid volume fraction in a
given section decreases. The Fig. 4 shows that the dis
crepancies between both fluidization velocities are not
equivalent in a given section. Indeed, for pure slip wall
boundary condition, in the near wall region both flu
idization velocities lead to particle accumulation up to
ap 0.6. In contrast, in the center of the fluidized bed
for Vf 0.40m/s we have up 0.4 and a, 0.32
with Vf 0.56m/s. This trend is also observed when
= P,(z
SPg(z
=Pg(z
= (
= 1.3m)
= 2.3m)
=3.3m)
9.03m)
P, (z
Pg(z
Pg(z
Pg(z
0.3m)
1.3m)
2.3m)
0.3m) .
dPI
dP,
dPs
dP,
The Table 2 shows the comparison between the mean
pressure drops measured on the experimental pilot and
the ones from numerical simulations. As previously
mentioned increasing fluidization velocity leads to a
higher expansion of the bed and consequently to larger
value of the bed height in the model if a fixed mass
of solid is considered. We observe that noslip wall
boundary condition gives a linear pressure distribution
(dP1 dPi dPs) as observed experimentally. Also
the bed height, computed with (11), predicted by the nu
merical simulation with noslip wall boundary condition
z=0 )30m
1 06
04
0 2 t VO 56m/s, shlp
SVO 56m/s, no shlp
03 02 0 1 01 01 02 03
Table 2: Comparison of pressure drop measured in ex
periments and in numerical simulations Effect
of fluidization velocity.
Vf Numerical Simulation
Ex. Slip Noslip
dP1 (mbar) 33.24 42.80 37.30
SdPA (mbar) 29.20 41.62 34.65
SdP3 (mbar) 29.80 32.18 36.43
dPt (mbar) 152.0 150.2 150.4
aHbed (m) 5.24 4.17 4.46
dP1 (mbar) 28.84 42.30 34.35
SdPA (mbar) 26.97 39.20 30.93
SdP3 (mbar) 29.51 30.01 32.84
usdPt (mbar) 140.0 153.5 153.4
aHbed (m) 5.22 4.43 4.99
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
0O 30m
;i~7
a~~0~
C*, ,
..
..
i 1
03 02 01 60
V ) 40m/s, shp
eV ) 56m/s, shp
SV ) 40m/s, n sp
t ~ 6m/s, no shp
01 02 03
V,0 40m/s, shp
V,0 56m/s, shp
5 V0 40m/s, no
C0, 56m/s, no si
03 02 0 1 u0 01 02 03
z 330m
3
5
03 02 01 uO
*
**
t ~ 6m/s, no shp
01 02 03
,3
+ 2
03 02 0 1 00
t~yP;B
SV,0 40m/s, shp
SV,0 56m/s, shp
 V, (M6m/s, nosh
01 02 03
Figure 5: Effect of the fluidization velocity on the radial profile of the mean axial gas velocity normalized by the
fluidization velocity (0, : Vf=0.40m/s and 0, #: Vf 0.56m/s). The empty symbols correspond to
the slip wall boundary for the mean particle velocity and blackfilled symbols to the noslip boundary.
noslip wall boundary condition is used.
The Fig. 4 exhibits that the noslip wall boundary con
dition leads to lower mean solid volume fraction. This
point was expected because we have seen in Fig. 3 that
noslip wall boundary condition leads to larger value of
bed height. More interesting we observe the modifica
tion of the shape of mean solid volume fraction which
exhibits peaks localized in the near wall region. These
modifications can be explained by the modification of
the mass flux in near wall region by applying the noslip
wall boundary condition.
The Fig. 5 shows the radial profiles of mean gas veloc
ity normalized by the fluidization velocity. We observe
that the fluidization velocity does not modify the mean
gas velocity profile. In contrast, the wall boundary con
dition noticeably changes the mean gas velocity profile
because of the modification of the mean solid volume
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
Om
Vpo 40m/s, shp
Vpo 56m/s, shp
VpO 40m/s, no shp
VpO 56m/s, no shp
V,0 40m/s, shp
e V,0 56m/s, shp
V0 40m/sno shp
SV,0 56m/s, no shp
3 
r C~ t
o
r
03 02 0 1 0 01 02 03
r (m)
z=2 30m
Vpo 40m/s, shp
eVp 56m/s, shp
3 Vp 40m/s, no p
SVpO 56m/s, no sp
03 02 0 1 0 01 02 03
r (m)
z 330m
V,0 40m/s, shp
3 V0 40m/s, no sh
V,=0 5m/sno shp
03 02 0 1 0 01 02 03
r (m)
Figure 6: Effect of the fluidization velocity on the radial profile of the mean axial particle velocity normalized by the
fluidization velocity (0, : Vf 0.40m/s and 0, #: Vf 0.56m/s). The empty symbols correspond to
the slip wall boundary for the mean particle velocity and blackfilled symbols to the noslip boundary.
fraction observed in Fig. 4. Indeed, in a dense fluidized
bed the nearwall region exhibits downward solid mass
flux. The particles entrain the gas at their own veloc
ity and as the solid volume fraction is reduced, by no
slip wall boundary condition, the solid mass flux in the
nearwall region decreases and consequently the inten
sity of the downward gas velocity. The flow rate balance
implies that the gas velocity in the bulk is then also re
duced.
The Fig. 6 shows the radial profiles of the mean axial
particle velocity normalized by the fluidization velocity.
For z < 3m the fluidization velocity has no effect on
the shape of the radial profile of the mean axial particle
velocity.
Effect of the operating gas density (pressure)
The present section is dedicated to the influence of the
operating gas density on the hydrodynamic of the flu
idized bed. The modification of operating density corre
sponds to the modification of the operating pressure that
can occurs in industrial facilities.
The Fig. 7 shows the wall pressure distribution mea
sured for two operating gas densities. For p,=
2 : I l: /m3 the bed height is larger than in case of p,
13.2kg/m3 because the drag force is increased leading
to a fluidized bed expansion. We find also this trend in
Table 3 where the fluidized bed is computed with (11).
As previously, the numerical predictions are slightly im
proved by the using of the noslip wall boundary condi
tion for the mean particle velocity.
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
10 
S4
SV,=0 40m/s, slip
SV,=0 56m/s, slip
V,=0 40m/s, no slip
SV,=0 56m/s, no slip
~~p,=23 6kg/m3
Sp,=13 2kg/ms
p,=23 6kg/ms
<=p13 2kg/ms
*
*:.
16000
12000 16000
4000 8000
ppo (Pa)
12000
4000 8000
ppo (Pa)
Figure 3: Effect of the fluidization velocity on the wall
distribution of the mean gas pressure (0:
Vf 0.40m/s and 0: Vf 0.56m/s).
The empty symbols correspond to a slip wall
boundary for the mean particle velocity and
blackfilled symbols to the noslip boundary.
The reference pressure po is taken at z =
8.5m.
Effect of particle shape
In this section we propose to analyze the effect of non
spherical particle shape on the hydrodynamic of the
dense fluidized bed. According to Loth (2008), non
spherical particle shape can be taken into account in
the drag law by introducing two parameters: fshape and
Shape. These two coefficients represent the asymptotic
behaviors of drag law for Re, < 1 and Re, > 1 re
spectively. Following Loth (2008), the modified particle
Reynolds number writes
Re shape Re(12)
shape
where Re, is the particle Reynolds number of spheri
cal particles and is given by (8). The drag coefficient is
modified according to
24 (+lie"g i (3
CD = shape E .5e067 17 1)
This drag coefficient is used instead of (7) in case of
Re* < 1000. To investigate the effect of nonspherical
particle shape two numerical simulations have been car
ried out with the same operating condition correspond
ing to RUN#1 but in one case Cshape = shape 1.
(spherical particle) and in other case Cshape 2.05 and
Figure 7: Effect of the operating gas density on the
wall distribution of the mean gas pressure
(0, : p, 2 : al: _/m3 and 0, #: p,
13.2kg/m3). The empty symbols correspond
to the slip wall boundary for the mean par
ticle velocity and blackfilled symbols to no
slip boundary.
fshape 1.05 (nonspherical particle shape). These val
ues have been chosen according to Loth (2008). Based
on the previous analysis noslip wall boundary condi
tions have been applied.
The Fig. 8 shows the axial mean gas pressure distribu
tion measured at the wall. As expected the nonspherical
particle shape leads to a fluidized bed height larger than
in case of spherical particles. Table 4 summarizing the
pressure drop measured at the wall and the bed height
also exhibits this trends. This increasing of fluidized
bed results from increasing drag force by nonspherical
Shape.
The radial distribution of mean solid volume fraction
is shown by Fig. 9. In each section, the averaged solid
volume fraction is reduced because the bed height is in
creased. However, we observe that, in all sections, the
shapes of the radial profiles are very similar. It means
that, in our particular case, the particle shape modifies
the bed height but not the hydrodynamic of the fluidized
bed. We also find this trend in Fig. 10 showing the ra
dial profiles of mean axial velocity particle distribution.
The modification of the particle shape does not modify
the mean particle velocity distribution in a section of the
fluidized bed.
of gas density.
p E Numerical Simulation
x. Slip Noslip
dPI (mbar) 25.77 42.44 36.33
SdP, (mbar) 26.58 39.73 34.06
SdP (mbar) 24.01 29.00 35.19
dPt (mbar) 133.0 142.1 142.5
Hbed (m) 5.52 4.14 4.35
SdPI (mbar) 27.89 43.15 38.20
SdP, (mbar) 29.99 41.68 35.81
SdPs (mbar) 26.59 34.96 38.15
dPt (mbar) 128.4 136.3 136.3
Hbede (m) 4.86 3.71 3.94
Table 4: Comparison of pressure drop measured in ex
periments and in numerical simulations Effect
of particle shape.
Numerical Simulation
Ex. Spherical Nonspherical
dPI (mbar) 25.77 36.33 32.61
dPi (mbar) 26.58 34.06 30.20
dP3 (mbar) 24.01 35.19 29.00
dPt (mbar) 133.0 142.5 143.2
Hbed (m) 5.52 4.35 4.98
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
0O 30m
z 130m
~~ 07~~
06+
,,
~.~..
,, I
X
Cc
etc,
`
r J
~ACtC r
~utttt~x
0 2
SSpherical partcles
Nonspherical partcles
03 02 0 1 00 01 02 03
r (m)
z=2 30m
0 2
SSpherical partcles
Nonsphericalpartles
03 02 0 1 00 01 02 03
r (m)
z 330m
0 7
Figure 9: Effect of nonspherical particle shape on the radial profile of the mean solid volume fraction. The wall
boundary for the mean particle velocity is noslip.
Table 3: Comparison of pressure drop measured in ex
periments and in numerical simulations Effect
7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
0O 30m
=130m
2
x
* I
o
B
03 02 O 1
SSpherical partcles
SNonspherical partcles
01 02 03
Figure 10: Effect of nonspherical particle shape on the radial profile of the mean axial particle velocity. The wall
boundary for the mean particle velocity is noslip.
Conclusions
Threedimensional unsteady numerical simulations of
polymerization reactor have been performed. The mean
pressure distribution has been extracted in order to com
pare with available experimental data. The results show
that a noslip wall boundary condition improves the nu
merical predictions of the hydrodynamic of the polymer
ization reactor in terms of bed height. The influence of
operating conditions has been investigated. As expected,
it has been shown that increasing fluidization velocity
leads to higher fluidized bed height. However, the gas
velocity radial profiles are not strongly modified. In con
trast the particle velocity radial distribution is affected
by the modification the solid volume fraction and then
the mass flux near the wall. The numerical simulations
have shown that increasing the operating gas density
leads to an expansion of the fluidized bed. The effect of
nonspherical particle shape has been investigated. The
results shows that nonspherical particle shapes lead to
higher bed because of larger drag coefficient. The gas
pressure distribution at the wall exhibit more linear pro
file.
Acknowledgements
This work was granted access to the HPC resources
of CINES under the allocation 2010026012 made by
GENCI (Grand Equipement National de Calcul Inten
sif).
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7th International Conference on Multiphase Flow,
ICMF 2010, Tampa, FL, May 30 June 4, 2010
0. Simonin. Combustion and turbulence in twophase
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PPo (Pa)
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