The Swedish and Finnish National
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Entrained flow Black Liquor Gasification –
Considerations for improvement of CFD reactor model
predictions
MAGNUS MARKLUND1*, PER CARLSSON1 and RIKARD GEBART1
1
Energy Technology Centre in Piteå
Corresponding author. Tel.:+46 911 23 23 85
E-mail address: magnus.marklund@etcpitea.se
*
ABSTRACT
Black liquor gasification has been proposed as an alternative to the Tomlinson boiler for
recovery of energy and chemicals in chemical pulping processes. The process is under
intense development and a commercial breakthrough is likely within the next few years.
Hence, there is a large interest in developing engineering tools that can be used for design
and optimisation of the process in arbitrary scale. To that end, a comprehensive CFDbased model has been developed that includes all important mass and heat transfer
mechanisms between the gas phase and black liquor droplets. The model has been
compared to experiments in a semi-industrial scale entrained flow gasifier operated by
the technology vendor Chemrec. In order to get good agreement between the model the
CH4 and H2S concentration had to be prescribed in order to get reasonable predictions for
the main gas components CO, CO2 and H2. The objective with the current work is to try
to explain this deficiency in order to make the CFD model of the reactor less dependent
of prescribed model parameters and thereby improve the predictions for a wide range of
operating conditions. Based on the calculations performed in this work and the
experiences from the DP-1 black liquor gasification plant, it is concluded that the
common Jones and Lindstedt mechanism is not suitable for modeling the current process
in the temperature range 1200 K< T < 1400 K.
Keywords: Black liquor, gasification, CFD, modelling
1.
INTRODUCTION
Black liquor (BL), a valuable by-product of the chemical pulping process, is an important
liquid fuel in the pulp and paper industry. It consists of approximately 30% inorganic
cooking chemicals along with lignin and other organic matter separated from the wood
during chemical pulping in a digester. It is imperative to recover both chemicals and
energy from the black liquor, which is traditionally done in Tomlinson Kraft recovery
boilers. Since 2005, Chemrec has been operating a development plant [1] for the novel
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alternative recovery concept: Pressurized entrained flow high temperature black liquor
gasification. This plant (named DP-1) is located at ETC in Piteå, Sweden, and now have
an accumulated operating time of approximately 12 000 h. The DP-1 plant is
schematically described in Figure 1 where the main parts are; the refractory lined 3 MWth
oxygen blown gasification reactor; the quench; and the gas cooler. In the reactor,
preheated black liquor (~140 °C) is injected through a centrally placed spray burner
nozzle at the top of the gasifier, together with oxygen and a small portion of nitrogen
(about 20 % of the oxygen mass flow rate). The reactor operates at an nominal elevated
pressure of 30 bar(g) and at a temperature of about 1000 °C. In the reactor the black
liquor droplets are gasified to form H2O, CO2, CO, H2, H2S and CH4 and a liquid
inorganic smelt. This smelt consists of the inorganic cooking chemicals used in the pulp
process and needs to be recovered. This recovery begins in the quench located
downstream directly after the gasification reactor. Here, the gas and the smelt are rapidly
cooled by water spray nozzles and separated so that the smelt is dissolved in water at the
bottom of the quench forming green liquor (GL), which is returned to the pulp mill for
further recovery treatment. Downstream the quench is the gas cooler, which is design to
recover heat by condensation of the water vapor in the gas and at the same time remove
carry-over particulates and condensable hydrocarbons that were not trapped in the
quench. To remove the sulphur components in the syngas (mainly H2S) a novel method
based on short time contactors have been used [2].
Figure 1. Schematic drawing of the main components making up the entrained flow
black liquor gasifier (courtesy of Chemrec)
In previous work, gas samples were withdrawn from the hot gasification reactor [3] using
a water-cooled sampling probe [4]. During the experimental campaign several operating
conditions, such as temperature, residence time and pressure were varied. Furthermore, in
a recent study [5] these experimental results were compared to Computational Fluid
Dynamics (CFD) predictions when the oxygen to black liquor equivalence ratio was
varied. One of the conclusions found from that study was that the CH4 and H2S
concentration had to be prescribed in order to get reasonable predictions for the main gas
components.
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The objective with the current paper is to explain the resulting deficiency in predicting
the concentrations of the minor species found in [5] and to consider different alternatives
in order to make the CFD model of the DP-1 reactor less dependent of prescribed model
parameters and improve the predictions for a wide range of operating conditions.
2.
Numerical
In the current work the CFD reactor model as described by Marklund et. al. [6] is
considered. The numerical model is implemented in the commercial Ansys CFX code via
subroutines and uses the coupled Eulerian-Lagrangian approach to model the gasification
process of black liquor droplets. For the continuous gas phase, the Reynolds Averaged
Navier-Stokes equations (RANS) and energy equation are solved using an Eulerian
description and the k-ε turbulence model with standard wall functions. Further details
around the CFD model description can be found elsewhere (e.g. [5] and [6]).
An important pre-processing step in CFD modeling of thermal conversion processes is
the droplet composition modeling of the considered fuel droplets. Based on the elemental
composition, the droplet composition modeling leads to the definition of an idealized
model fuel composition and a corresponding heat of formation of the modeled species.
This translation of chemical elemental analysis data and thermodynamic data into model
information must be consistent with known experimental data.
Regarding the gas phase species, there are numerous of different combustion mechanisms
for hydrocarbons, and particularly for CH4, available in the literature. One of the most
cited one is the GRI 3.0 mechanism [7]. In the current work this mechanism is used as a
reference, in comparison to two alternative mechanisms. The mechanisms considered are
the commonly used 4 step, 6 species global reaction mechanism developed by Jones and
Lindstedt [8] which was used in the previous modelling work [6] and the 26 step 15
species mechanism by Bilger et al. [9] (see Table 1).
The entrained flow gasifier can roughly be characterized by two separate regions; the
flame region (where oxygen is present) and the post flame region. The post flame
regions, i.e. locations where all the oxygen are depleted, make up the majority of the
reactor volume. In this work, 1-D calculations were performed in the chemical kinetics
program Cantera [10] for two different approaches; an isothermal Plug Flow Reactor
(PFR) which represents the post flame region and an opposed jet oxygen-methane flame,
which represent the flame region. The opposed jet calculation were done at 27 bars of
pressure with pure oxygen and methane (strain rate a = 290 s-1) and at 1 bar with a diluted
(Ar) methane oxygen flame (a = 170 s-1).
The PFR calculations performed at 27 bar and isothermal conditions in order to simulate
regions in the gasification reactor where the temperature is still relatively high but all
oxygen is depleted. The inlet gas composition was set to a typical syngas composition
(10% CH4, 20%, CO2, 20% CO, 25% H2O, 25% H2) and the effect of varying
temperature (1200, 1350, 1500 and 1650 K) were investigated.
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Table 1. Mechanism from Bilger et al. [9]
REACTION
A
b
E
1 CH4 + H ↔ CH3 + H2
2.20E+04
3.00
8750
2 CH4 + OH ↔ CH3 + H2O
1.60E+06
2.10
2460
3 CH3 +O → CH2O+H
6.80E+13
0.00
0
4 CH3 + OH ↔ CH2 + H2O
1.50E+13
0.00
5000
5 CH3 + H ↔ CH2 + H2
9.00E+13
0.00
15100
6 CH2 + H ↔ CH + H2
1.40E+19
-2.00
0
7 CH2 + OH → CH2O + H
2.50E+13
0.00
0
8 CH2 + OH → CH + H2O
4.50E+13
0.00
3000
9 CH + O2 → HCO + O
3.30E+13
0.00
0
10 CH2 + O ↔ CO + H + H
3.00E+13
0.00
0
11 CH2 + O ↔ CO + H2
5.00E+13
0.00
0
12 CH2 O + OH → HCO + H2O
3.43E+09
1.18
-447
13 CH2 O + H → HCO + H2
14 HCO + M → CO + H + M
15 HCO + H ↔ CO + H2
2.19E+08
1.60E+14
4.00E+13
1.77
0.00
0.00
3000
14700
0
16 HCO + O2 ↔ CO + HO2
3.30E+13
-0.40
0
17 CO + OH ↔ CO2 + H
1.51E+07
1.30
-758
18 OH + H2 ↔ H2O + H
1.17E+09
1.30
3626
19 H + O2 ↔ OH + O
5.13E+16
-0.82
16507
20 O + H2 ↔ OH + H
1.80E+10
1.00
8826
21 H + O2 + M → HO2 + M
3.61E+17
-0.72
0
22 OH + HO2 → H2O + O2
7.50E+12
0.00
0
23 HO2 + H ↔ OH + OH
1.40E+14
0.00
1073
24 OH + OH ↔ H2O + O
6.00E+08
1.30
0
25 H + OH + M ↔ H2O + M
1.60E+22
-2.00
0
26 HO2 + H ↔ H2 + O2
1.25E+13
0.00
0
Rn 21 & 25 TBEs: H2O: 18.6, CO2: 4.2, H2: 2.86, CO: 2.1, N2 : 1.26
b
Rate constants are in form kf=A T exp (-E/RT) Units are moles, cm, s, K, cal/mol
Notice that Rn. 3, 7, 8, 9, 12, 13, 14, 21, 22 are irreversible
3.
RESULTS
In this section, results from different plug flow reactor (PFR) calculations are presented
along with some results obtained by the CFD reactor model.
In Figure 2 the evolution of methane is presented as function of residence time in an
isothermal PFR and a total residence time of 5 s. It can be noted that for the GRI
mechanism, methane starts to reform at about 1350 K where only 3% has been converted
into other species (mainly CO and H2). At 1500 K, about 30% has been reformed and at
the highest considered temperature (1650 K) almost all methane is reformed. On the other
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hand, the JL mechanism shows a significantly higher reformation rate. At 1350 K, about
70 % of the methane has been reformed, which should be compared to 3% for the GRI.
However, at the highest temperature (1650 K) the agreement is much better between the
JL and GRI mechanisms. Considering the Bilger mechanism, no conversion of methane
in the investigate temperature range was found.
CH4
0.1
← T= 1200 K
← T= 1350 K
0.09
0.08
← T= 1200 K
0.07
← T= 1500 K
Mol fraction
0.06
0.05
0.04
0.03
← T= 1350 K
0.02
0.01
0 -4
10
-3
10
-2
10
-1
10
0
10
← T= 1500 K
← T= 1650 K
1
10
log (Residence time / s)
Figure 2. Profiles of CH4 molar fraction vs. residence time at different temperatures
in an isothermal plug flow reactor at 27 bar. Solid line: GRI-Mech. Dashed line:
Bilger et al. Dotted Line: Jones et al. The inlet molar gas composition was CH4:
10%, CO2: 20% CO: 20%, H2O: 25%, H2: 25%. The inlet temperatures were:
T=1200, 1350, 1500 and 1650 K
The conversion of the methane has an immediate effect on the major gas components as
can be seen in Figure 3. Note that, the GRI and Bilger show very good agreement up to
1350 K.
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CO2
CO
0.36
0.2
← T= 1200 K
← T= 1650 K
0.34
← T= 1500 K
0.18
0.32
← T= 1350 K
0.16
← T= 1350 K
← T= 1200 K
← T= 1500 K
← T= 1500 K
0.28
Mol fraction
Mol fraction
0.3
0.14
← T= 1650 K
← T= 1500 K
0.12
← T= 1650 K
0.26
← T= 1500 K
← T= 1350 K
0.1
← T= 1200 K
0.24
← T= 1350 K
0.08
0.22
0.2 -4
10
← T= 1650 K
← T= 1200 K
← T= 1200 K
-3
10
-2
-1
10
10
0
10
10
← T= 1500 K
0.06 -4
10
1
-3
10
log (residence time / s)
-2
-1
10
10
0
10
1
10
log (residence time / s)
H2
H2O
0.38
0.32
← T= 1650 K
0.36
← T= 1650 K
← T= 1500 K
0.34
← T= 1500 K
0.3
← T= 1350 K
← T= 1350 K
0.32
0.28
← T= 1500 K
← T= 1200 K
0.28
Mol fraction
Mol fraction
0.3
0.26
← T= 1200 K
← T= 1200 K
0.26
← T= 1200 K
← T= 1200 K
0.24
← T= 1200 K
0.24
← T= 1500 K
← T= 1350 K
← T= 1650 K
0.22
← T= 1350 K
0.2
0.22
← T= 1500 K
← T= 1500 K
← T= 1650 K
0.18 -4
10
-3
10
-2
10
-1
10
0
10
10
0.2 -4
10
1
log (Residence time / s)
-3
10
-2
10
-1
10
0
10
1
10
log (Residence time / s)
Figure 3. Profiles of CO, CO2, H2 and H2O molar fraction vs. residence time at
different temperatures in an isothermal plug flow reactor at 27 bar. Solid line: GRIMech. Dashed line: Bilger et al. Dotted Line: Jones et al. The inlet molar gas
composition was CH4: 10%, CO2: 20% CO: 20%, H2O: 25%, H2: 25%.The inlet
temperatures were: T = 1200, 1350, 1500 and 1650 K
Predicted temperature (K) contours obtained by the CFD reactor model by Carlsson et al.
[5] are showed in Figure 4. As can be seen, the temperature is below 1400 K in a majority
of the reactor.
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Figure 4. Temperature (K) distribution obtained from CFD calculations [[5]]. From
left to right: λ = 0.40, 0.44 and 0.45. The dotted lines indicate where the temperature
contours coincides along the centre axis for the different cases.
1
3200
0.9
2880
0.8
2560
0.7
2240
0.6
1920
0.5
1600
0.4
1280
0.3
960
0.2
640
0.1
320
0
0
T/K
Mass fraction
PFR results for oxygen-methane opposed flow jet flames at 27 bar and 1 bar are
presented in Figure 5 and Figure 6, respectively. Notable from the temperature profiles
and methane concentrations is that, at 1 bar the combustion is initiated earlier for the GRI
mechanism compared to the Bilger mechanism.
0.002 0.004 0.006 0.008
0.01
z (m)
0.012 0.014 0.016 0.018
0
0.02
Figure 5. Profiles of temperature and mass fraction in an oxygen-methane opposed
jet flame at 27 bar. GRI-Mech (solid line) and Bilger et al. (dashed line). No
marker; Temperature, □; O2, ◊; CH4, ○; CO2, +; CO, *; H2O, × ; H2
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0.4
3200
0.36
2880
0.32
2560
0.28
2240
0.24
1920
0.2
1600
0.16
1280
0.12
960
0.08
640
0.04
320
0
0
0.002 0.004 0.006 0.008
0.01
z (m)
0.012 0.014 0.016 0.018
T/K
Mass fraction
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0
0.02
Figure 6. Profiles of temperature and mass fraction in an argon diluted oxygenmethane opposed jet flame at 1 bar. GRI-Mech (solid line) and Bilger et al. (dashed
line). No marker; Temperature, □; O2, ◊; CH4, ○; CO2, +; CO, *; H2O, × ; H2
4.
DISCUSSION
The methane concentration is an essential output from modeling gasification since it
strongly influences the total gas composition and the expected efficiency of the
downstream handling of the product gas. From earlier findings [5] and the results
presented above it is evident that the Jones and Lindstedt (JL) mechanism is not suitable
for modeling black liquor gasification in the current process concept (1200 K< T < 1400
K) due to the high methane reformation rates resulting in this temperature range. This
particular result should come as no surprise since it was actually derived for combustion
conditions of hydrocarbons. In fact, when studying combustion in a counter flow jet
flame, the JL mechanism performs fairly well given its simplicity. However, the rate of
reformation of methane becomes over predicted when looking at gasification conditions
in the temperature range 1200 K< T < 1400 K, at least compared to the GRI mechanism
and in reactor model validation experiments in DP-1 [5]. This explains why the previous
work with the DP-1 reactor CFD model where not able to predict the methane
concentration accurately [5].
The current results further show that the Bilger mechanism is not suitable for gasification
conditions when the temperature is expected to exceed 1400K in the post flame region.
For the major gas components, the GRI and Bilger show very good agreement up to
about 1400 K. However, at higher temperatures the reduced rates of methane reformation
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in the Bilger mechanism would most certainly over predict the methane content
compared to what is found in the experiments for the DP-1 case [5].
5.
CONCLUSIONS
Based on earlier findings and the calculations performed in this work around the DP-1
black liquor gasification plant, the following main conclusions are drawn:
•
•
•
The earlier CFD model by Marklund et al. [6], which is based on the Jones and
Lindstedt (JL) mechanism is unable to predict the content of CH4 and H2S in real
gasifiers.
The temperature predictions of the DP-1 reactor show that the temperature is in
the interval 1000 K< T < 1400 K in the majority of the reactor.
Comparisons of the JL mechanism with the GRI mechanism that has been shown
to agree well with kinetic experiments with methane show that the main weakness
of the JL-mechanism under reducing conditions (gasification) is that the rate of
reaction for methane is overpredicted in the temperature range that is typical for
black liquor gasification (1000 - 1400 K)
In future work, the focus will be on implementation of a suitable mechanism for the
formation and conversion of both CH4 and H2S in the CFD reactor model.
6.
REFERENCES
[1]
Lindblom M. and Landälv I., Status of the Swedish National Black liquor
Gasification (BLG) Development Program, TAPPI EPE Conference 2006.
TAPPI, Atlanta
[2]
Öhrman O., Johansson A-C., Lindblom M., H2S removal with short time
contactors in a pressurized black liquor gasification plant. International Chemical
Recovery Conference, Williamsburg, USA, March 29-April 1, 2010.
[3]
Carlsson P., Wiinikka H., Marklund M., Grönberg C., Pettersson E., Lidman M.,
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(12): 4025–34
[4]
Wiinikka H., Carlsson P., Granberg F., Löfström J., Marklund M., Tegman R., et
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[5]
Carlsson P., Marklund M., Furusjö E., Wiinikka H., Gebart R., Experiments and
mathematical models of black liquor gasification – influence of minor gas
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components on temperature, gas composition, and fixed carbon conversion.
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[6]
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[7]
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[8]
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[9]
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[10] Goodwin D., http://sourceforge.net/projects/cantera (last visited 23.09.2010)
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