s1c">eled by diffusion-kinetic relations [7]. To account for
polydispersity, the clusters of fresh coal particles were
split into ~100 fractions according to Rosin-Rammler
size distribution law, with this the co al milling dispersity
(specified as mass fraction of remainder on 90 m sieve)
was taken as R90 = 0.15 and polydispersity degree =
1.2.
To predict the nitric oxides emission in the vortex fur-
nace, at the stage of post-processing of main solution
data, the transport equations for the concentrations of
{NO, HCN, NH3} species have been solved. With this
the following NOx formation mechanisms are taken into
account: a) extended “thermal” N2 oxidation mechanism
of Ya.B.Zeldovich; b) “prompt” NO formation in reac-
tion of N2 with hydrocarbon radicals following C. P. Fe-
nimore mechanism; c) “fuel” NO formation mechanism.
Also the factors of NO reduction are accounted: – via
“reburn-mechanism” in reaction with CHi radicals; – via
heterogeneous consumption of NO on the surface of coke
particles.
3. Boundary Conditions and Numerical
Algorithm
Because the studied vortex furnace has a symmetry plane
SXY (in the section zsymm=3 m) between two nozzles (see
Figure 1), the mirror symmetry of the flow is assumed in
relation to this plane, thus to save computational re-
sources only half of the furnace volume was considered
in simulations. The bounding dimensions of computa-
tional domain were taken as: xmax=D0= 6 m, ymax= 36 m,
zmax= zsymm= 3 m. The upper nozzle angle of inclination to
horizon (XZ plane) was =2. The unstructured com-
putational grid built for finite-volume discrimination of
equations consists of 351784 hexahedral cells. The
“sticking” boundary conditions are set for velocity com-
ponents at walls, and “enhanced wall treatment” method
[8] is applied for near-wall turbulence modeling. A uni-
form profile of mean flowrate velocity is prescribed at
inlet sections of rectangular nozzles, with this the inten-
sity of inlet turbulen t pulsations is assumed equal to 5%.
The known Marsha bound ary conditions [7] f or the ra-
diation energy transport equation have been set at walls,
with emissivity of the furnace walls taken as w = 0.8 and
the temperature of superheated water-vapor inside the
heat-exchange tubes presumed at 390℃. Also to set
boundary conditions for radioactive and convective heat
transfer at the heat-exchange surfaces, the integral heat
transfer coefficient w through the walls (from ash- de-
Copyright © 2013 SciRes. EPE
I. S. ANUFRIEV ET AL.
308
position layer surface to water-vapor inside the tubes)
has been prescribed at different heat-exchange screens
(walls) in the furnace cooling chamber. Its values speci-
fied at these surfaces, as well as the average values of the
screens thermal efficiency av obtained in computations,
are given in Table 1.
The mass flow rate of pulverized coal loaded per com-
puted half-section of the vortex furnace volume was
Gcoal=3.75 kg/s, the overall air excess coefficient at fur-
nace inlets was set typically for the vortex furnace design
as 1.15. With this the primary stream of coal-air mixture
at inlet temperature of 180℃ was loaded through the
upper nozzle only, while the secondary dry air stream at
320 C was supplied through th e bottom nozzle. Also the
ratio of gas-phase flow rates through the upper and the
bottom nozzles =Gupper/Gbottom was taken as =3 – as it
has been revealed in computations, this parameter influ-
ences the aerodynamic structure formed by two tangen-
tially injected jets insid e the vortex combustion chamber.
For numerical approximation of convective terms of
Favre-averaged momentum equations the second-order
upwind scheme [8] has been applied. The numerical so-
lution at each “global” iteration for the gas phase equa-
tions is obtained according to PISO algorithm [9] for
pressure-velocity decoupling, while the interphone ex-
change of mass, momentum and heat transfer is ac-
counted following the Particle-Source-In-Cell method
[10].
4. Computational Results
Typical flow structure in the studied vortex furnace is
represented in Figure 2 in the form of velocity magni-
tude is contours in XY-section (z = 1.6 m) across the noz-
zle center (zoomed view of only the vortex chamber and
diffuser is shown). The following features of the vortex
furnace aerodynamics can be seen, such as the “glove-
scheme” flow structure when the upper inlet jet (injected
tangentially from the nozzle) evolving alongside the
main vortex flow in the combustion chamber is being
crossed and surrounded by the upstream flow moving out
of this chamber towards the diffuser part of the furnace.
Near the center of vortex combustion chamber (but shifted
in upper-right di rect ion) a region of vortex core is clearly
Table 1. Average heat transfer coefficients (w) and thermal
efficiencies (av) at heat-exhange surfaces.
Heat-exhange
surface location Specified value
of w , W/(m2K) Computed value
of av
Screen “A” in section z=3 m 250 0.590
Two-sided screen “B”
in section z=0 200 0.483
Other screens (front , rear and
ceiling in the cooling chamber) 150 0.415
seen. The upstream flow inside the cooling chamber de-
monstrates noticeable spatial no uniformity (such as
Coanda effect).
This Coanda effect can be also seen from Figure 3
where the temperature field in the new vortex furnace is
demonstrated. High-temperature level inside the vortex
combustion chamber (reaching 1880℃ in the near-wall
reaction zone, see Figure 3) provides th e stable regime of
liquid slag removal. This is also confirmed from Figure
4 where the profile of temperature averaged over hori-
zontal XZ-sections is depicted along furnace height (solid
line in Figure 4). Also the dashed line there indicates the
profile of maximum temperature values along furnace
height. With this the estimated heat release rate per unit
volume inside the combustion chamber is ~800 kW/m3.
The sharp downward trough in averaged temperature
profile (at ~7 m height) in Figure 4 indicates the cold jet
stream evolved from the upper nozzle (see also Figure 3).
Then in the cooling chamber the temperature decreases
rapidly along the furnace height, and the temperature
Figure 2. Contours of velocity magnitude in the new vortex
furnace (section across nozzle center z=1.6 м), m/s.
Figure 3. Temperature field in the new vortex furnace
(section across nozzle center z = 1.6 м),.
Copyright © 2013 SciRes. EPE
I. S. ANUFRIEV ET AL. 309
field becomes more uniform as the difference between
maximum and averaged values becomes smaller. This
improved performance appears due to efficient heat
transfer from the furnace volume to heat-exchange
screens. Indeed the profiles of the screens thermal effi-
ciency (defined as =Qres/Qinc where Qres is the re-
sulting heat flux at wall surface and Qinc – the incident
heat flux) shown in Figure 5 demonstrate rather good
thermal efficiency values of screens in the cooling
chamber. The dashed line there denotes values at the
middle of screen “A” (mounted in section z=3 m) and the
solid line – at the middle of two-sided screen “B” (in
section z=0). Values av averaged over each screen sur-
face area (shown in Table 1) also indicate the range 0.41
< av < 0.59 which is better than a typical level of ~0.4
for most furnaces.
Figure 4. Averaged (solid line) and maximum (dashed line)
temperature distribution along vortex furnace height, C.
Figure 5. Heat-exhange screens thermal efficiency distri-
bution along furnace height in the cooling chamber; dashed
line – screen “A”, solid line – screen “B”.
The contours of resulting heat fluxes (at heat-exchange
screens) are shown in Figure 6. The profiles of O2, CO,
NO concentrations averaged over horizontal XZ-sections
are plotted in Figure 7 along furnace height.
The following integral parameters have been obtained
in the vortex furnace outflow section (before exhaust
flue): mean temperature Tmean= 980℃, maximum tem-
perature Tmax= 1055℃, coke burnout incompleteness
coefficient q4~1.3%, averaged volumetric concentrations
[O2]= 2.47%, [CO]= 14 ppm, [NO]= 307 ppm. It is seen
that NOx emission level remains contained within ac-
ceptable limits – notwithstanding the high temperature
level inside the vortex combustion chamber – mainly due
to the effect of recirculation of combustion products
Figure 6. Heat fluxes absorbed at the surface of heat-ex-
change screens, kW/m2.
Figure 7. Distributions of O2, CO, NO concentrations (by
volume), averaged over horizontal XZ-sections, along vortex
furnace height.
Copyright © 2013 SciRes. EPE
I. S. ANUFRIEV ET AL.
Copyright © 2013 SciRes. EPE
310
towards reaction zones inside the vortex chamber. Also
the improved thermal efficiency of heat-exchange screens
in the cooling chamber has been demonstrated.
5. Conclusions
The numerical simulation of 3-D aerodynamics and pul-
verized brown coal combustion processes in the new
vortex furnace configuration with dual-port loading has
been performed, and detailed flow field information in-
cluding the flow structure, the fields of temperatures,
radiated heat fluxes, species and dispersed phase concen-
trations have been obtained. With this it has been shown
that the integral heat engineering and ecological parame-
ters of new vortex furnace meet the ranges corresponding
to engineering practice.
6. Acknowledgements
Computations have been performed with the use of CFD
package FLUENT at the supercomputer clusters
NKS-160 and NKS-30T (SSCC SB RAS, Novosibirsk,
Russia). The work has been supported by Russian Minis-
try of Education and Science (Agreement No. 8187).
REFERENCES
[1] V. V. Salomatov, D. V. Krasinsky, Yu. A. Anikin, I. S.
Anufriev, O. V. Sharypov and Kh. Enkhjargal, “Experi-
mental and Numerical Investigation of Aerodynamic
Characteristics of Swirling Flows in a Model of the Swirl-
ing-type Furnace of a Steam Generator,” J. of Engineer-
ing Physics and Thermophysics, Vol. 85, No. 2, 2012, pp.
282-293.doi:10.1007/s10891-012-0651-8
[2] N. V. Golovanov, V. E. Nakoryakov, A. P. Burdukov, V.
V. Salomatov and A. A. Dorozhkov, “Vortex Furnace,”
Russian Federation patent No.2042084, 1995 (in Rus-
sian).
[3] T.-H. Shih, W. W. Liou, A. Shabbir, Z. Yang and J. Zhu
“A New k- Eddy-Viscosity Model for High Reynolds
Number Turbulent Flows – Model Development and Va-
lidation,” Computers&Fluids –Vol. 24, No. 3, 1995, pp.
227-238. doi:10.1016/0045-7930(94)00032-T
[4] B. E. Launder and D. B. Spalding, “Lectures in Mathe-
matical Models of Turbulence,” London (England): Aca-
demic Press, 1972.
[5] B. F. Magnussen, “The Eddy Dissipation Concept,” IEA,
1989.
[6] T. F. Smith, Z. F. Shen and J. N. Friedman “Evaluation of
Coefficients for the Weighted Sum of Gray Gases Mod-
el,” Proceedings of XX-th National ASME-AIChE Heat
Transfer Conference, Milwakee, USA, August 2-5, 1981.
[7] E. P. Volkov, L. I. Zaichik and V. A. Pershukov, “Model-
ling the Solid Fuel Combustion,” Moscow: Nauka, 1994
(in Russian).
[8] FLUENT 6.3 User’s Guide, Fluent Inc., 2006.
[9] R. I. Issa, “Solution of Implicitly Discretized Fluid Flow
Equations by Operator Splitting,” Journal of Computa-
tional Physics, Vol. 62, No. 1, 1986, pp. 40-65.
doi:10.1016/0021-9991(86)90099-9
[10] C. T. Crowe, M. P. Sharma and D. E. Stock, “The
Particle-Source-In-Cell (PSI-CELL) model for
gas-droplet flows,” ASME Journal of Fluids Engi-
neering, Vol. 99, No. 2, 1977, pp. 325-332.
doi:10.1115/1.3448756