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8/16/2019 A Comprehensive Approach to in-Situ Combustion Modeling. J.belgrave, Et Al.
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A Comprehensive Approach to In-Situ Combustion Modeling
J.D.M. Belgrave, R.G. Moore, M.G. Ursenbach, D.W. Bennion
The University
of
Calgary, Department of Chemical and Petroleum Engineering
ABSTRACT
Low temperature oxidation (L TO) has long been recognized as one
of
the dominant mechanisms controlling fuel availability in in-situ
combustion. Its effect on the physical properties of crude oils is also
well known. In spite
of
these fmdings, the prevailing conceptual model
of in-situ combustion still hinges on thermal cracking (in isolation)
ahead of the firefront, to provide sufficient fuel (coke) for propagation
of the reaction zone. Previous simulation studies, which purported to
include L TO as par t
of
the reaction scheme, have unrealistically
specified the reaction products as carbon oxides and water.
Furthermore, oil compositional changes due to oxidation have been
completely neglected.
This paper describes a unified pseudo-mechanistic reaction
model for mathematical modeling of in-situ combustion of Athabasca
bitumen.
The
model represents a consolidation of individual
experimental kinetic studies on thermal cracking and low temperature
oxidation
of
Athabasca bitumen, and reported data for the high
temperature oxidation of coke.
The
formulation is comprehensive in that
it allows bitumen to undergo density and viscosity increases, as well as
reduced reactivity to oxidation, with increased oxidation extent.
Hydrocarbon bypassing due to quenching of the combustion front is also
permitted with the proposed kinetic model.
The paper includes application of the reaction model in
numerical simulations of adiabatic combustion tube tests performed on
Athabasca bitumen. A significant feature of the model is its ability to
predict the dual oxidation uptake peaks associated with ramped
temperature oxidation experiments.
INTRODUCTION
The oil sands
of
Alberta, Canada collectively represent one
of
the
largest hydrocarbon deposits in the world
l
.
Cyclic steam stimulation, to
date, has been the most widely used technique for exploiting these
deposits. This technique is capable
of
recovering only 15
to
20
of
the
oil-in-place, and a follow-up process is required to improve recovery2.
Laboratory investigations
of
in situ combustion as
a
post-steaming process have been sufficiently encouraging
to
warrant
implementation
of
pilot studies by some lease operatorg3. However, the
transition from experimental and pilot stages to commercial
development has been virtually non-existent. This stems, in part, from
the fact that in situ combustion is not well understood, mechanistically.
A great deal of laboratory work has shown that frontal
advance and air requirement arc controlled by the kinetics
of
the
reactions in the vicinity
of
the burning front. Several studies
4
,5,6 have
reported on three major reactions which occur during fireflooding:
1)
thermal cracking, 2) liquid phase low temperature oxidation (L TO),
and (3) high temperature oxidation (HTO)
of
a solid hydrocarbon
residue. In their pioneering experimental effort, Alexander et
aU
concluded that of all the process variables which they studied, LTO
prior to high temperature burning had the greatest effect on fuel
availability. Poettmann et al. 8 estimated that L TO could increase the
fuel deposition by as much as 100 , and Lerner et al.
9
emphasized the
need to consider the effects of L
TO
in numerical simulations of
combustion processes.
In spite
of
these findings, published conceptual profiles
of
in
situ combustion still adhere to thermal cracking as the sole means of
98
fuel generation ahead of the reaction zone. In addition, most simulation
studies
o
l I 2 which considered LTO, have neglected associated chemical
changes such as increased viscosity and density
of
the oil as well as its
reduced reactivity to oxidation with increased oxidation extent.
The
main focus of this work was the development of a pseudo
component reaction model that is able to produce the oxidation related
phenomenon mentioned above, in combustion simulation studies
of
Athabasca oil sands. Individual thermal cracking and L TO kinetic
studies on Athabasca bitumen, and reported data for coke combustion
have been consolidated into such a model. In this paper, the
performance
of
the proposed reaction scheme is examined in numerical
simulations of a differential reactor undergoing an imposed ramped
temperature history, and two combustion tube tests (dry and superwet)
performed on Athabasca bitumen. Through this analysis an improved
quantitative description, and therefore understanding,
of
in situ
combustion has emerged.
. EXPERIMENTAL BASIS FOR REACTION MODEL
For an explicitly correct kinetic representation
of
hydrocarbon cracking
and oxidation, an inordinately large number of chemical species would
have to be considered. Such a system would not be practical as it would
impose prohibitive computational demands on thermal reservoir
simulation. A pseudo component model offers the only useful
alternative. Furthermore, unification of the three classes of reactions
into a comprehensive model can only
be
achieved if kinetic studies on
the reactions are consistent in their fractionation or characterization of
the oil.
As
regards L TO and thermal cracking of Athabasca bitumen,
Adegbesad
3
and Hayashitani et al.
4
have reported pseudo component
kinetic data for these reactions. Both studies used bitumen which was
free
of
water and minerals.
Hayashitani thermally cracked Athabasca bitumen in a closed
system at 651 to 828
OF
[344 to 442
0c]
under an inert atmosphere.
The cracked liquid products were first separated into maltenes and an
asphaltenes-coke residue by filtration, using n-pentane as solvent.
Asphaltenes were next recovered from the residue by solution in
benzene. Hayashitani further fractionated the maltenes into light oils,
middle oils, and heavy oils by vacuum distillation.
Adegbesan used a stirred semiflow batch reactor to investigate
LTO of Athabasca bitumen in the 140 to 302 OF [60 to 150 0c]
temperature range, and at oxygen partial pressures of 7 to 324 psig [50
to 2233 kPa]. His characterization technique for the reaction products
duplicated Hayashitani's as far as separation of the maItenes and
asphaltenes-cokein n-pentane. Coke was defined as the bitumen fraction
insoluble in toluene.
The
maltenes were then separated in saturates,
aromatics, oils and resins by. a combination
of
solvent extraction and
chromatographic techniques.
In view of the difference in the methods used to analyze the
maltenes in the two studies (thermal as opposed to
solvent/chromatographic), consolidation of the kinetics data was
restricted to coke, asphaltenes, maltenes, and gas, as pseudo
components.
SPE Advanced Technology Series, Vol. 1, No.1
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100
397 ·C
• maltenes
80
• asphaltenes
E
•
60
~
f
40
:IE
20
•
•
0
0
2 4
6 8
10 12
Time hours)
Fig.
la -
Cracking model vs. Hayashitani s data for maltenes and
asphaltenes.
Reaction
Kinetics
Based on Hayashitani s data, the thermal cracking scheme
proposed in this paper assumes three fIrst order reactions:
MALTENES ASPHALTENES
1)
ASPHALTENES COKE
2)
ASPHALTENES GAS
(3)
If
kl> k2 and k3 are the rate constants respectively for the
above reactions, then differential equations describing the material
conversion can
be
written as
dCasp
dt
dC
gas
dt
4)
(5)
6)
with the temperature dependence of the rate constants being described
by the Arrhenius equation
~ e x p -EJRT
(7)
Using the technique described by Kalogerakis and Luus
l4
,
kinetic parameters
were
estimated for the thermal cracking scheme
specifIed above. In the order in which the reactions have been specifIed,
SPE Advanced Technology Series, Vol. I, No 1
25
397
·
• coke
20
• gas
C
15
:
if
10
:IE
5
o ~ ~ ~ ~ ~ ~ ~ ~
o
2 4 6
8
10 12
Time hours)
Fig.
Ib
- Cracking model vs. Hayashitani s data for gas and coke.
these parameters are:
Al
9.092 x 10
12
sec-
I
EI
2.347
X
10
5
kJ/kmol
A2
4.064 x 10
9
sec-
I
E2
1.772
x lOS
kJ/kmol
A3
1.362
x
10
9
sec-
I
E3
1.763 X 10
5
kJ/kmol
Figure 1 compares this cracking model versus Hayashitani s
data at 747
OF [397°C]
The percent
of
each component is on a mass
basis. At this and higher temperatures the agreement is quite good.
However, the initial increase in maltenes concentration above the
predicted curve becomes more prominent at lower temperatures,
indicating the need for another pseudo component.
For the LTO data reported by Adegbesan, the following
. reactions are proposed:
MALTENES OXYGEN
ASPHALTENES
8)
ASPHALTENES
OXYGEN COKE
(9)
The kinetic parameters for this system
were
estimated from rate
equations similar to that described for the thermal cracking reactions.
However, oxygen concentration was specifIed in terms
of
its partial
pressure and the reaction order with respect to oxygen partial pressure
was allowed to vary in an unconstrained fashion. With respect to
hydrocarbon mass fraction the reactions were kept as fIrst order.
Respectively, the Arrhenius constants are:
6.819
X
10
3
8.673
X
10
4
2.133 X
10-
10
1.856
X
10
5
sec-I
Pa-
0
4246
kJ/kmol
sec-I Pa-4.7627
kJ/kmol
99
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100
135 ·C
• maltenes
eo
• asphaltenes
coke
c:
60
8
f
40
E
20
o L ~ ~ ~ ~ = = ~ = = ~ ~
o
1
2
3
4
5
6
Time (hours)
Fig. 2 - LTO model vs. Adegbesan s data at
135°C
for the conversion
of
the hydrocarbon reactants. Note that the frequency
factors inherently state reaction orders, with respect to oxygen partial
pressure,
of
0.4246 for maltenes oxidation and 4.7627 for asphaltenes
forming coke. Figures 2 and 3 show the experimental data at two
temperature levels, along with the predicted bitumen compositional
changes. Generally it was found that at temperatures of 275 OF [135
°C], or less, the predicted mass percentages 0 f the components agreed
very well with the experimental data. However, at higher temperatures
(Figure 3) it was not possible to predict the reduction
in
asphaltenes
content and increased coke synthesis which occurred at later reaction
times.
It
is
important to note here that we have not specified the
release
of
oxygen in the cracked product streams
of
reactions 1-3. This
stems from
our
experimental findings (unreported) which have shown
that thermal cracking
of
preoxidized bitumen does not regenerate the
molecular oxygen which was consumed in the additive LTO reactions.
Kinetic parameters for coke combustion were obtained from
the work
of
Thomas et al. 15. Using integral analysis, these researchers
studied the oxygen uptake of coke combustion with coke derived from
an oxidized Athabasca bitumen-water-sand mixture. From this
reference, coke combustion is first order with respect to both reactants,
and the reaction rate was given as:
gmolOzfhr
3.612 X 10-
6
(10)
x exp( -34763/RT) Cook.. P02
with C
eoa
in kg coke/m
3
bulk volume.
The in-situ combustion kinetic model offered above is
considered to be preliminary in nature, and there is much latitude for
refinement and optimization. As more consistent experimental data
becomes available, more intermediate reaction pathways and pseudo
components may be specified for thermal cracking as well as LTO.
The thermal cracking scheme we have proposed
is
based
solely on the experimental observation that the asphaltenes
concentration monotonically decreases (Figure
1
which facilitates the
determination
of
rate equations for predicting the monotonic increase
in
coke and gas formation. The change in maltenes concentration on the
100
100
80
60
40
20
150 ·C
•
maltenes
• asphaltenes
coke
•
•
o . . : : : : : : : : t : : : : ~ = = = 1 ~
0.0
0.5 1.0
1.5
2.0
2.5
Time (hours)
Fig. 3 - LTO model vs. Adegbesan s data at 150°C
other hand is not monotonic; it first increases and then decreases.
Maltenes involvement was there fore limited to that
of
a buffer supplier
of
asphaltenes.
With respect to high temperature oxidation, it
is
expected that
inclusion
of
the complete oxidation
of
maltenes and asphaltenes will
improve the representation of the combustion process. However, the
experimental data needed to furnish the stoichiometric coefficients and
kinetic parameters for these reactions remain unavailable.
Stoichiometry
Estimates for the molecular weights
of
the pseudo components
must be obtained if stoichiometric coefficients for the preceding
reactions are to be specified.
Bishnoi
et
al.
I6
presented characterization data for oil sands
bitumen, which included specific gravities and molecular weights
of
the
reported pseudo components. These data were compared with measured
specific gravities of the original bitumen and maltenes fraction, and
molecular weights were inferred for the oil components (see Table 1).
The gaseous products from Hayashitani s cracking experiments
indicated an average molecular weight
of
29.0, and for coke,
Adegbesan reported a measured hydrogen/carbon ratio
of
1.13, which
gives a coke molecular weight of 13.13.
In thermal cracking, a unit mass
of
reactant produces a unit
mass
of
product. The stoichiometric coefficients for these reactions are
therefore determined from the ratios
of
the molecular weights
of
reactants and products. Thus the thermal cracking reactions, on a molar
basis, can be written as:
MALTENES - 0.372 ASPHALTENES
(11)
ASPHALTENES - 83.223 COKE
(12)
ASPHALTENES - 37.683 GAS (13)
The
amounts of oxygen which react with unit mass of
maltenes and asphaltenes were determined from Adegbesan s L TO data
by parameter estimation
l4
with the following oxygen uptake rate
SPE Advanced Technology Series, Vol.
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equation:
3.0
0.0
0.0
0 22 g 0 g Asphaltenes
0.27
g
0 g MaJtenee
•
•
•
•
0.5 1.0 1.5 2.0 2.5
•
Measured Oxygen Uptake Rate g/hr)
Fig.
4 -
Measured vs. predicted oxygen uptake rates.
r
2
dt
a
dtn wt
+
dt
b
dmc:oke
1 +b dt
3.0
14)
where a and b are the masses
of
oxygen that combines with maltenes
and asphaltenes respectively. Equation 14 reflects the fact that the total
oxygen uptake rate
is
due to both maltenes and asphaltenes conversion,
and that the total mass
of
oil in the system is increased by the additive
oxidation reactions. Figure 4 shows the estimated oxygen uptake ratios
for the two L TO reactions, as well as their ability to reproduce the
experimental data. Some scatter is evident, but the trend has been
adequately duplicated. Since the asphaltenes represent a partially
oxidized state they should have fewer oxygen addition sites than the
maltenes. This fact
is
observed in the lower oxygen uptake ratio.
Based on the L TO uptake ratios, and the assumed molecular
weights
of
the oil components, the molecularity
of
the reactions may
now be written:
MALTENES
+
3.431 OXYGEN
0.4726 ASPHALTENES
ASPHALTENES
+
7.513 OXYGEN -
101.539 COKE
15)
16)
The
general form
of
the coke combustion reaction
is
well
documented in the literature
l7
:
CH
[
2m+l
+ +
n
2m+2
n
m
0
-
CO
4 2 m+l 2
17)
1
n
+ CO + HzO
m+l
2
where m is the molar ratio of carbon dioxide to carbon monoxide
produced, and n is the hydrogen/carbon
of
the coke burned. Three
SPE Advanced Technology Series. Vol.
1
No. I
simulation runs are discussed in this paper.
For
these runs, a constant
value 8.96 was specified for the CO
2
/CO molar ratio, as the
stoichiometry
is
not appreciably affected by small changes to this
parameter. On a molar basis, and lumping the carbon oxides into CO ..
the coke burning reaction becomes:
CH1.13 + 1.2320
2
- COx + 0 5 6 5 ~ 0
18)
Heats
of
Reaction
There is little evidence in the literature which suggests that
thermal cracking
of
hydrocarbons is accompanied by any release
or
absorption
of
heat which significantly affects the combustion process .
Therefore the enthalpies
of
these reactions were assumed to be zero. To
obtain an estimate
of
the heat liberated by L TO and coke combustion,
we
referred to the publication by Burger and
Sahuqud.
For
the
stoichiometries
of
the oxidation reactions, they suggested heats
of
reactions for LTO and coke burning
of
the order
of
5.567 x
lOS
1.228
X
10
6
,
and 1.893
X
lOS Btu/Ibm mol [1.295
x
10
6
,
2.857
X
10
6
,
and
4.278 x 10
5
kJ/kmol]
of
maltenes, asphaltenes and coke respectively.
Bitumen Chemical Changes
t is evident that the reaction scheme specified above allows
oil components to
be
synthesized and/or consumed as a result
of
oxidation. This was taken one step further
in
that an attempt was made
. to determine the viscosity and density
of
the oil components, based on
laboratory measurements.
The viscosity
of
the original bitumen and its maltenes fraction
were independently measured and the data curve fitted to give the
following viscosity-temperature relationships:
1/
= 0.48267
X
10-
6
exp 7685.2{f)
19)
-bitumen
0.19359 X 10-4 exp 5369.2{f)
20)
Based on the assumed molecular weights, the asphaltenes viscosity
temperature relationship was then inferred from the mixing rule
to give
~ s p =
4.892 X 10-
25
exp 33147{f)
22)
Bitumen viscosity at 212
OF [100°C]
as a function
of
asphaltenes mole fraction is shown plotted in Figure
5.
Asphaltenes by
themselves are observed to be an essentially immobile component.
Furthermore, any increase in the asphaltenes content of the bitumen is
accompanied by a significant increase in the overall bitumen viscosity.
Similarly, density increases due to oxidation are
accommodated by the model simply by assigning different densities to
the oil pseudo components.
The
determined specific gravities for the
asphaltenes and maltenes were respectively 1.1580 and 0.9832.
Extending this concept, reduced volatility with increased oxidation
extent is effected by giving the asphaltenes a zero gas/oil equilibrium
K-Value.
101
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TABLE 3 - RELATIVE PERMEABILITY DATA FOR RUNS 1 AND 2
Oil/Water
Gas/Oil
Sw
krw
krow
S.+Sw
krg
kro
0.05000 0.00000
1.00000
0.07000 0.10000 0.00000
0.10000
0.00039
0.88581
0.16000 0.08615 0.00316
0.15000 0.00156
0.77855
0.21000
0.06632
0.01262
0.25000
0.00625 0.58478
0.31000 0.03711 0.05050
0.35000
0.01406 0.41869
0.41000 0.01881 0.11362
0.45000
0.02500 0.28028
0.51000 0.00829 0.20199
0.65000 0.03906
0.08651
0.61000 0.00296 0.31562
0.75000 0.05625
0.03114
0.71000
0.00073 0.45449
0.85000 0.07656
0.00346
0.80000
0.00011 0.60106
0.95000 0.10000
0.00000 0.95000 0.00000 0.89080
1.00000 0.12656
0.00000
1.00000
0.00000 1.00000
TABLE 4 - RELATIVE PERMEABILITY DATA FOR RUN 3
Oil/Water
Gas/Oil
Sw
krw
krow S.+Sw
Krg
krog
0.11000 0.00000
1.00000 0.12000 0.10000
0.00000
0.15000 0.00013 0.87891
0.16000
0.08615 0.00316
0.20000
0.00052 0.76562
0.21000
0.06632 0.01262
0.30000
0.00208 0.56250 0.31000 0.03711
0.05050
0.40000
0.00468 0.39063
0.41000
0.01881 0.11362
0.50000
0.00833 0.25000
0.51000
0.01029 0.20199
0.60000 0.01302
0.14063 0.61000 0.00400
0.31562
0.70000 0.01875 0.06250
0.71000
0.00105 0.45449
0.80000 0.02552
0.01562 0.80000
0.00021 0.60106
0.85000 0.02929
0.00391 0.95000
0.00000
0.89080
0.90000
0.03333 0.00000
1.00000
0.00000
1.00000
1.00000
0.04218 0.00000
TABLE 5 - MODEL INITIALZATION FOR SIMULATIONS
Run 1
Run 2
Run 3
Sand pack length (m) 0.250 1.83 1.83
Sand pack diameter (m)
0.0508 0.0994 0.0994
Number o axial grid blocks
3
36 36
Initial water saturation 0.0500
0.1180 0.1130
Initial oil saturation 0.2500
0.8820 0.8870
Initial gas saturation 0.7000
0.0000 0.0000
Pressure (kPa)
4100 4100 5500
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600
500 0.25
2
0
U
CD
5
400 0.20
...
-
II)
0
I
g
5
m
300 0.15
II
...
II
S
.
E
j
...... ../
a.
II
200 0.10
::l
l-
e:
\
/ ~
II
\
I,
oxygen
\
100 0.05
0
\
\
/
0
.....
0.00
0
2
3
4
5
6
Time (hours)
Fig. 6a - Calculated oxidation behaviour
of
Athabasca bitumen:
oxygen uptake.
APPLICATION OF THE REACTION SCHEME
Combustion
Tube
Simulator
A general purpose mathematical model of combustion tube
reactors
18
which rigorously includes the annular region surrounding the
sand pack was used in the simulations described below. Our experience
has been that radial heat transfers to and from the sand pack can
significantly affect experimentally observed temperature levels, and
must
be
accounted for during simulations of combustion tube
experiments. The coded mathematical treatment of in-situ combustion
in the sand pack is essentially that described
by
Coats
19
•
Combustion Simulations
Three simulation runs were performed to investigate the
oxygen uptake characteristics
of
the reaction scheme, and
to
evaluate
its feasibility as a predictor of fireflooding frontal advance. Rock and
fluid properties
are
given
in
Tables 1 to 4, and model initialization data
are presented in Table 5. Note that all non-condensible gas components
other than oxygen have been lumped and referred to in Table 1 as gas.
Run 1 was a simulation of a differential reactor undergoing
ramped temperature oxidation, of the kind reported by Burger and
Sahuquet
5
. The
combustion tube model was programmed to raise the
temperature
of
the tube wall by 180
°F/hr
[100 °C/hr] while air (21.0%
02)
was injected at a rate of 2.828 scf/hr [0.081 std-m
3
/hr]. With the
small axial length of the grid specified (see Table 5) temperature and
oxygen concentration gradients were still present. Therefore only the
conditions in the first grid block
are
reported, and are plotted in
Figure 6.
Two successive oxidation peaks
are
observed in Figure 6a.
The first occurs at about 1.25 hrs and becomes a maximum at
2.0
hrs.
During this period maltenes
are
converted to asphaltenes which can
be
seen to increase in Figure 6b. As the maltenes become almost depleted,
the oxygen uptake rate drops to a minimum around 2.25 hrs, indicating
that the bitumen has become less reactive. With further increases in
temperature, and reduced maltenes content, asphaltenes
are
oxidized to
produce a substantial amount of coke (whose initial appearance
is
delayed). The thermal cracking reactions also assist in the conversion
of
asphaltenes to coke. Finally, coke
is
entirely consumed around 1110
OF [600°C]. The second oxygen uptake peak is therefore due to the
oxidation of asphaltenes and coke combustion. It is worthwhile
mentioning that these dual oxidation peaks have been experimentally
observed
5
,
and that the temperature at which the calculated first peak
104
c::
1.0
100
0
asph. mole
n
CD
fraction
I
u:
0.8
~
80
I
\
0
/ \
e
:::iE
.
( / )
/
E
2
0.6
\
60
-
OJ
II
e-
a
\
coke
II
1:.
I
..loI:
a.
\.......
0
/)
0.4
40
)
oil sat.
I..
-
\
to
(
:J
e:
:2
~
(/)
...
0.2
\
20
:J
-
\
CD
n
\
0
0.0
0
0
2
3
4 5 6
Time (hours)
Fig.
6b
- Calculated oxidation characteristics of Athabasca bitumen:
oil composition/saturation and residual coke.
reaches its maximum (518 OF [270 °CD agrees very well with those
published data.
This simulation was repeated without any LTO reactions, and
the results plotted in Figure 7. There is a single oxidation peak, and
only a small quantity
of
coke
is
deposited. It does not appear feasible
that this amount coke can support an advancing combustion front. It is
also important to note that coke is formed at a much higher temperature
without LTO.
Run 2 simulated a dry enriched air (94.78 % 02) combustion
tube test which has been reported as Test 206 by Belgrave
18
. The
stabilized air injection rate was 1.955 scf /hr [0.056 std-m
3
/hr]. This run
was performed as a test of the ability of the proposed reactions
to
duplicate the combustion front velocity as well as fuel and oxygen
consumption.
As was also the case with run
3,
the model initialization
procedures closely followed the experiment, since at ignition fluid
saturations are not uniformly distributed.
The
simulations started from
uniform water and oil saturation and zero gas saturation (Table 5). Gas
was the injected at the top
of
the vertically oriented sand pack until the
model reached the experimental operating pressure given in Table 5.
Next, gas was flowed through the core until a continuous gas phase
saturation had been established. At this point the injection end of the
core was heated to 752
OF
[400°C]
in
run 2 and 572
OF
[300
0c]
in
run 3. Enriched air injection was then started. Temperature profiles for
run 2 at model times of 12.0 and 15.9 hrs are shown
in
Figure 8.
Model run time at ignition was 8.20 hrs.
The
experimental and
calculated (solid lines) leading edges of
the combustion front are
in
good agreement. Behind the fronts, however, the calculated temperature
profiles were higher than those obtained by experiment. This
discrepancy
in
radial heat losses is due to a less than adequate
description
of
the average thermophysical properties
of
the annular
region surrounding the combustion tube. These properties, which are
very dependent on experimental conditions
18
and are also are affected
by tube operation
,
were not manipulated during these simulations.
Up
to
15.9 hrs, 44 % of the tube had been traversed in the
simulation, and at this time the calculated average coke and oxygen
consumed were 1.44
Ibm/ft3
[23.09 kg/m3] and 64 scf/ft
3
[64
std-m
3
/m
3
],
respectively.
The
stabilized experimental values for the test
were 1 47lbm ft
3
and 56.4 to 62.9 scf/ft
3
[23.6 kg/m3 and 56.4 to 62.9
std m
3
/m
3
]
Figure 9 shows that the produced fluids were also
in
good
agreement.
Run 3 simulated a superwet combustion tube test that has been
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600
500
0.25
2
0
t;
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6
.
-
-
s
CD
Q.
E
{
800 1 ;: :: : :: :== == == ;1
sao
400
200
• 12.0 hours
15 9 hours
o
~ ~
__ __ ~ ~ ~ ~ J
0 0
0 4
0 8
1.2
1.6
2.0
Distance (meters)
Fig. 8 - Experimental and numerical temperature profiles for dry
enriched-air combustion test.
i
000 1.2
&
gas
•
oil
000
• water
~
0.9
8
3000
It
8
I
0.6
Q:
2000
i
=a
0.3
.1lI
1000
a
~
0 0.0
0
4
8
12
16
20
Time (hours)
Fig. 9 - Experimental and numerical produced fluids for dry enriched-
air combustion test.
the dominant mechanism controlling fuel availability for the in-situ
combustion process. Thermal cracking, in isolation, does not generate
sufficient fuel for high temperature combustion propagation.
(5) The dual oxidation uptake peaks, associated with ramped
temperature oxidation tests, and the delay in coke formation have
been reproduced by the reaction model, and attributed to
significant differences in reactivity between oxygen and
individual components which make up the oil. This implies that
in-situ combustion cannot
be
adequately simulated using a single
component oil system.
(6)
The reaction scheme presented here is capable
of
predicting
experimentally determined frontal velocity, and oxygen and fuel
requirements.
106
6
.
-
-
CD
Q.
E
{
r ~ ~
sao
400
200
•
0
0 0 0 4
•
0 8
• experiment
-model
•
1.2
1.6
Distance (meters)
2.0
Fig. 10 - Experimental and numerical temperature profiles
superwet test at
3 0
hrs after ignition.
c:
1.0
30
2
'0
i
0.8
25
CD
15
coke
E
:: E
\ /\
20
.6
CD
:
\
s
15
i
0.4
asph mole
/ \
1:
fraction
,
~
-----,
\
10
c:
:8
as
0.2
CD
...
8
a
5
0
0.0
0
0.2 0.4 0.6
0.8
1.0
Distance (meters)
of
Fig. 11 - Spatial variation
of
grid block variables for superwet test at
3 0 hrs after ignition.
a
b
NOMENCLATURE
mass of oxygen that combines with unit
mass
of
maltenes
mass
of oxygen that combines with unit
mass
of
asphaltenes
frequency factor for reaction r
frequency factors for reactions
1, 2,
and
3
mass fraction
activation energy for reaction r
activation energies for reactions
1, 2,
and
3
rate constant for reaction r
gas relative permeability
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