A Comprehensive Approach to in-Situ Combustion Modeling. J.belgrave, Et Al

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


2/10

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


3/10

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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No 1


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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


5/10


6/10

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

SPE Advanced Technology Series, Vol. I, No. I

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7/10

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

SPE

Advanced Technology Series Vol. 1 No.1


8/10

600

500

0.25

2

0

t;


9/10

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

SPE Advanced Technology Series, Vol. I, No. 1


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Documents

A Comprehensive Approach to in-Situ Combustion Modeling. J.belgrave, Et Al