Industrial and Laboratory Pyrolyses - American Chemical Society

20 A n Industrial Application of Pyrolysis Technology: Lummus SRT III Module J. M . FERNANDEZ-BAUJIN and S. M . SOLOMON

Downloaded by CORNELL UNIV on October 18, 2016 | http://pubs.acs.org Publication Date: June 1, 1976 | doi: 10.1021/bk-1976-0032.ch020

C-E Lummus, Bloomfield, N.J.

The pyrolysis of hydrocarbons for the production of olefins and diole fins has been extensively investigated during the last four decades by industrial and academic institutions. Many investigators are still work ing on the development of kinetic models which can be used to predict yield patterns as a function of the operating conditions for a given hy drocarbon. Such models have not been completely accurate even for such a simple hydrocarbon as ethane.(1) Of course, the complexity of such a modelling technique is increased several fold when it is applied to heavier hydrocarbons and to a mixture of them, i.e., naphthas and gas oils. As a result, the development of techniques to predict industrial yield patterns generally is based on a combination of theoretical and empirical methods.

PYROLYSIS TECHNOLOGY - BRIEF REVIEW A hydrocarbon system undergoing pyrolysis is a most complex mixture of molecules and free radicals which react simultaneously with one another in a multitude of ways. Based on established theories supported by experimental data, the production of olefins and diolefins — that is, pyrolysis selectivity towards olefins and diolefins — has been found to be favored by short residence times and low hydrocarbon partial pres sures (2,3,4). Pyrolysis reactor selectivity has been expressed as a function of the following two parameters: 1. 2.

Residence Time, Θ Hydrocarbon Partial Pressure, P H C

The specific functions of residence time and partial pressure must des cribe the history of the feedstocks pyrolyzed and the history of the products produced in passing through the coil. 345

Albright and Crynes; Industrial and Laboratory Pyrolyses ACS Symposium Series; American Chemical Society: Washington, DC, 1976.

INDUSTRIAL AND LABORATORY PYROLYSES

Bulk residence time,

dVc V

0B

where

is defined by Equation (1)

Vj

(1)

Total volume of the pyrolysis coil,


Volume of the pyrolysis coil as a function of its length, and V

Volume of gas per unit time passing through the pyrolysis coil. This volume changes with coil length.

The bulk residence time, θβ, which is the more commonly used para meter to correlate pyrolysis selectivity, expresses only the length of time that a mass of gas spends in the pyrolysis coil. Bulk residence time is not adequate to represent the combined effects of temperature and chemical reaction which take place in the pyrolysis coil as witness by the failure to correlate the available experimental data as a function of bulk residence time especially when data from pyrolysis coils of dif ferent configurations and, therefore, different temperature profiles are considered. ( ) 3

In an effort to correlate the existing experimental data from bench scale, pilot plant and commercial reactors, the definition of average residence time, 0 ^ . shown in Equation ( 2 ) has been developed.

(2) where θ$

= Bulk residence time,

#b

= Gas residence time along the length of the pyrolysis coil

a

= Feedstock conversion along the length of the pyroysis coil, = Conversion of the feedstock at the pyrolysis coil outlet.


LumUS SRT III Module

FERNANDEZ-BAUJIN AND SOLOMON

Conversion, a , is empirically defined as a function o f the molecular weight o f the feedstock and the composition of the cracked gas. The average residence time, expressed as a function of feedstock conver sion, is a measure o f the average time the reaction products spend in the coil. As with residence time, the correlating definition of

hydrocarbon

partial pressure must consider the partial pressure history in the coil. F o r a given feedstock cracked at a fixed conversion and gas outlet


pressure in a pyrolysis coil o f a given configuration, the average hydro carbon partial pressure, PpjCA' * determined by: s

+

The pressure drop through the pyrolysis coil, and

+

The dilution steam-to-hydrocarbon ratio.

The arithmetic average o f inlet and outlet hydrocarbon partial pressure alone does not properly reflect the pressure history of the gas in the pyrolysis coil. The definition o f average hydrocarbon partial pressure, P H C A >

W

M

C

N

has been developed, is shown in Equation (3).

P

HCA

"

where PHC α

=

Hydrocarbon partial pressure along the coil length,

= Feedstock conversion along the length o f the pyroly sis coil, and

a

0

= Conversion o f the feedstock at the pyrolysis coil outlet.

The average hydrocarbon partial pressure, expressed as a function o f feedstock conversion, is a measure o f the average hydrocarbon partial pressure of the reaction products in the coil. Pilot plant, prototype and commercial data on pyrolysis selectivity col lected by Lummus have been correlated as a function o f the average residence time and the average hydrocarbon partial présure. A method o f correlation which has proved highly

successful is illustrated in

American Chemical Society Library 1155Industrial 16th and St.,Laboratory N.W. Pyrolyses Albright and Crynes; Washington» D.C. 20036 ACS Symposium Series; American Chemical Society: Washington, DC, 1976.



8

10 06

12

14

16

18

20

PSiA

0J

T


Coil 2 Avg.

± R

( 1 5 )

Where Es and Ep = Activation energies of primary and secondary reac tions. In evaluating Equations (14) and (15), average values have been used in order to consider the overall tubular reactor length. Based on the work published by Messrs. Hirato, Yoshioka, and Tanaka, (9) the most probable overall average values for (Es - Ep) have been estimated to about - 11000 Kcal/mol. This estimate has resulted from considering a variety of possible combinations of primary and secondary reactions which have been postulated to occur in thermal pyrolysis. Equation (14) has been applied to the two pyrolysis coils used in the experimental work previously discussed. The calculated average resi dence times, gas temperatures and hydrocarbon partial pressures for Coils 1 and 2 are presented in Table II. T A B L E II Average Parameters for Coils 1 and 2 Coil No.

θ , Sees

1

2

0.0893

0.1558

TA,°F

1468

1441.5

P ,Psia

13.6

A

HCA

12.5

From Equation (15),

A

2

0.9298



Then:

PiHCA2

(Kp/Ks) (Kp/Ks) 1901.2 (13.6) 1927.7


PiHCA2

A

2

A

1

(0.9298)

12.5 Psia

(16)

which is in agreement with the value shown in Table II and confirmed experimentally in Figure 3. This is a confirmation of the above discussed theoretical considerations. Therefore, Equation ( 15) which is based on theoretical considerations, defines conditions for identical selectivity. Shape of Pyrolysis Selectivity Lines The application of Equation (15) to pyrolysis coils operating at average residence times in the millisecond region leads to the important conclusion that the constant pyrolysis selectivity lines are not straight. A typical pyrolysis selectivity chart extended into the millisecond region is schematically shown in Figure 9. The above analysis of the factors affecting pyrolysis selectivity now permits the selection of a locus or conditions (i.e., a selectivity line) which will achieve a given pyrolysis selectivity and thus, a certain yield structure. Since this constant selectivity line covers a wide range of combinations of average residence times and average hydrocarbon partial pressures, the question remains "what other factors should be considered by the designer to select a point on this selectivity line which is optimum for large-scale commercial production of olefins?" To answer this question, it is necessary to consider the effects of simultaneous heat, mass and momentum transfer in the pyrolysis coil. SELECTION OF COIL C H A R A C T E R I S T I C S A pyrolysis coil of a given selectivity can be designed to lie at any point on the desired selectivity line. The selection of a coil operating at a lower residence time and a higher hydrocarbon partial pressure will result in a coil with a higher gas outlet temperature than a coil operating at a higher residence time and lower hydrocarbon partial pressure



LUTTIUS SRT

III

Module


A V E R A G E RESIDENCE TIME. SECONDS

ι

ι

ι

»

»

I

I

I

I

L

A V E R A G E H Y D R O C A R B O N P R E S S U R E , PSIA

Figure 9.

Schematic pyrolysis selectivity chart

on the same selectivity line. The implications of this higher outlet temperature are significant. A commercially acceptable pyrolysis coil design is constrained by the limitations in available metallurgy. For any selected material, there exists a maximum operating temperature beyond which tube life is sharply reduced. This limitation in metal temperature must be con sidered in setting the coil design. The coil with the higher gas outlet temperatures requires higher heat transfer coefficients and/or higher heat transfer surface-to-volume ratio in order to not exceed the tube metal temperature limitation. This is achieved by utilizing a coil with relatively small diameter outlet tubes. This design, however, results in other undesirable effects due to simul taneous heat, mass and momentum transfer. In order to clearly discuss these effects, it is interesting to refer to the theory of the relationships between momentum, heat, and mass transfer.


364


TRANSPORT PHENOMENA The transport phenomena which operate at a given point in a pyrolysis coil are related according to the following fluid analogies: (1) ρ/3

Sc*

K m ( M W ) _ hPr _ (F/S)gc G GC M/A

3

p

(F/S) = ΑΔΡ/S = *

ρ 2

ΔΡ

0

= _ΟΔΡ

4TTDL


^ f _ 0.023 2 (DG/μ) ·

(17) 2

(18)

4L

where: Km = Molal mass transfer coefficient MW = Fluid molecular weight G

= Mass velocity

h

= Heat transfer coefficient

Cp

= Specific heat at constant pressure

F/S = Drag force/unit surface area of pipe gc

= Universal gravitational constant

M

= Momentum flow

A

= Pipe cross sectional area

f

= Fanning friction factor

D

= pipe diameter

μ

= Viscosity

Sc

= Schmidt Number,

Pr

= Pradtl number, (CpM/k)

(μΙρφ)

M/A= G2/p G

= W/A

k

= Thermal conductivity

Φ

= Diffusivity

ΔΡ

= Pressure drop

L

= Axial pipe length

ρ

- Mass density

W

= Mass flow rate

Combining the above relationships, the following expressions are derived : 2/3

c S

2 / c

VKm(MW) (Muut = -c— Ρ h P r

μ = 0.023W°t,. D (TT/4)

02

a

(19)

a 8



LUTUUS SRT

III

Module

and APpgc 4L

0.023W μ · θ"(ιτ/4)" 18

=

0 2

(20)


From Equations (19) and (20), it is clear that the small diameter coil, while enhancing the heat transfer coefficient, also tends to increase the pressure drop and mass transfer coefficient. The effect of increased pressure drop per unit length has been demon strated in the pyrolysis selectivity experiment described above. This experiment showed that the shorter residence time coil, i.e. Coil 1, be cause of the higher pressure drop, and therefore higher average hydro carbon partial pressure, achieved the same selectivity as the larger dia meter, longer residence time, lower partial pressure coil. The effect of an increased mass transfer coefficient is significant in terms of coil coking which limits run length due to increase in metal temperature and pressure drop. COKING CHARACTERISTICS Coking in a commercial pyrolysis coil is a highly complex process which has not been modelled in precise mathematical terms.01> 1 2 , 1 3 ) However, as is often the case, a simplified model utilizing certain em pirically defined constants, has proven to adequately describe the coil coking phenomenon. (14)

The coil coking model postulates a two-step mechanism : 1.

Mass transfer of coke precursors from the bulk of the gas to the walls of the tube.

2.

The chemical reaction of coke precursors at the tube wall re sulting in the formation and deposition of coke. This simpli fied two-step process can be described as follows: Rm = Km (y—yp ^

=Kr(

y i l

^ )

(21) (22)

where ; Rm Rr Km Kr

= = = =

molal rate of mass transfer Rate of chemical reaction Mass transfer coefficient Reaction velocity constant



y yj Ρ R Τ

= = = = =

mole fraction of coke precursor in bulk fluid mole fraction of coke precursor at tube wall Total pressure Universal gas constant Absolute temperature at tube wall

The chemical reaction is assumed to be first order for the purpose of this derivation. However, the conclusions are shown to be independent of this assumption.


At any given location at the tube wall : Rm = Rr

(23)

Then:

Km(y-y:) = Kr if. RT 1

Kmy RT

+ K m

R = Rm = Rr = Kmy[l -

K

RT

m

\

(24)

+ Km

Where R = Rate of coke formation In a commercial pyrolysis coil, the temperature at the tube wall is con siderably higher than the bulk temperature of the gas. At these high temperatures, studies have indicated that the reaction velocity constant, Kr, is much higher than the mass transfer coefficient, Km. Since Kr > > Km Km ^ RT

0

+ Km]

J

(25) (26)

and R = ykm Therefore, the rate of coil coking is a mass transfer controlled process. From Equation (19),


LuiUUS

FERNANDEZ-BAUJIN A N D SOLOMON

Ρ _ coke d e p o s i t e d , K* W Day (D-2A)

SRT III Module (27)

0 , 8

Where W D Δ

1 8

= Total mass flow rate = Inside diameter of pyrolysis tube = Coke thickness

K*

= Constant which is a function of feedstock, cracking selectivity, dilution steam ratio, cracking severity


and other system properties. Then, the run length o f the pyrolysis coil, 0 R L can be predicted as follows: 0

R

L

=

Δ max. R

(

2

8

)

Where Δ max. is the coke thickness which coincides with the maximum allowable tube metal temperature or the maximum allowable pressure drop. Equations (19) and (27) indicate that the rate of coke formation is higher for a small diameter tube to the same extent that the rate o f heat transfer is greater for a small tube. Also the rate o f coke formation in the small tube increases more rapidly during the length o f a run because coke build-up in a small tube increases the mass transfer coefficient more rapidly than in a larger tube as indicated by the term ( ϋ - 2 Δ ) 1 , 8

The coil coking model has been tested in a number o f plants. Table III presents these data. The agreement between the predicted and the actual run length is very good, which is a confirmation o f the assump tions made in the development o f the model. S E L E C T I V I T Y LOSS D U R I N G R U N A further, extremely important characteristic o f a small outlet tube is that the pyrolysis selectivity decrease during the length o f a run will be greater than that of a large outlet tube. This results from the effect o f the deposited coke on the tube cross sectional area. A given thickness o f coke in a small tube increases the coil pressure drop more rapidly than in a large tube. As a result, the increase in average hydrocarbon partial pressure o f a small tube during the run length will cause a more significant decrease in selectivity than in the larger tube coil. Figure 10 illustrates this effect. The two coils studied in the selectivity experi-



TABLE ill


TEST OF THE PYROLYSIS COIL COKING MODEL Olefins Plant

Feedstock* »

A Β C D Ε F*

FRN FRN FRN FRN FRN LGO

Run Length, Days Plant Data Calculated 59 90 100 45 45 83

58 82 100 45 50 82

*Gas Oil Prototype Unit in Japan **FRN = Full Range naphtha LGO = Light Atmospheric Gas Oil

AVERAGE RESIDENCE TIME, SECONDS Λ

0

0.48

0.2

CI.52

(WEIGHT RATIO) 0.60 1 2

0.56

c

/c

-\ \ SOf

Λ \ %2

\

VEOF

0.1

^ \SOR\ 1

4

h

6

1

8

1

10

1— —ι r

12

1

14

1

ι

16 PSIA

r

|0.3 0.4 ,0.5 0.6 0.7 0.8 0.9 1.0 1.1 Kg/sq. cm. AVERAGE HYDROCARBON PARTIAL PRESSURE BASIS: KUWAIT NAPHTHA, LIMITING CRACKING SEVERITY

Figure 10.

Pyrolysis selectivity chart



LumUS

SRT III

Module

ment are compared in terms of end of run selectivity. The average selectivity of the larger outlet tube coil is greater than that of the shorter, small tube coil, despite the fact that the start of run selectivities are identical. In summary, not only will a large tube coil coke at a lower rate, but also the effect of whatever coke does form is less significant in terms of continued pyrolysis selectivity.


OPTIMUM COIL DESIGN Before discussing the characteristics of an optimum coil design, it is worthwhile to summarize the conclusions presented up to this point; 1.

A given pyrolysis selectivity can be achieved by designing for an average residence time/average hydrocarbon partial pressure combination which can be anywhere on a locus of points which describe a line of constant selectivity.

2.

The question facing the designer is where on this selectivity line is the point of optimum coil design. The choice is often between a short, small outlet diameter coil which emphasizes low residence time and a longer, large outlet diameter coil which emphasizes low hydrocarbon partial pressure.

3.

The short coil, small diameter tube option has the disadvantage of a higher rate of coke formation and a greater loss of cracking selectivity as the run progresses.

In addition, the capacity of the large outlet tube coil is greater than that of the short coil and as such, the quench complexity and costs are reduced. The capacity of the large diameter coil compared to the small diameter coil studied in the selectivity experiment was 4.3 times greater for the same operating selectivity. This higher coil capacity and simpler overall system result in a pyrolysis reaction module of a much lower cost for a fixed production of olefins. Having developed the arguments favoring a coil which achieves a given selectivity by emphasizing low hydrocarbon partial pressure, (longer, large outlet diameter), the remaining question is what are the specific characteristics of this coil type? Any optimum design should achieve a given result at minimum cost. This minimum cost is achieved by minimizing total heat transfer surface


370


T A B L E IV COMPARISON OF PROCESS CHARACTERISTICS* Coil No. Relative H y d r o c a r b o n Feed Rate

2 (SRT III Furnace) 4.31

Relative A v g . Residence Time

1.78

R e l a t i v e Pressure D r o p

0.81

Relative A v g . Hydrocarbon P a r t i a l Pressure


Relative C o i l C o k i n g Rate

0.92 0.75

* D e s i g n basis as d e s c r i b e d p r e v i o u s l y i n t h i s w o r k .

and minimizing the complexity o f the quench system. The large capacity, low cost quench system as used by Lummus is a by-product o f the large outlet diameter, high capacity coil and has been discussed elsewhere. 0 > 6 ) 5

T o further illustrate these conclusions, the process characteristics o f the C o i l N o . 1, small diameter, and C o i l N o . 2, S R T III heater, previously discussed are compared in Table I V and i n Figure 10. SUMMARY The average hydrocarbon partial pressure and the average residence time are the factors which affect pyrolysis heater selectivity. The available data o n pyrolysis selectivity have been successfully correlated as a function o f these two parameters. The temperature and partial pressure profiles along the pyrolysis coil are considered in the definition o f average residence time and average hydrocarbon partial pressure. The results o f this work have shown that, while the average residence time and hydrocarbon partial pressure are both key factors in determining pyrolysis selectivity, the hydrocarbon partial pressure is somewhat more important than previously realized by investigators. It has been experimentally shown that the pyrolysis selectivity o f two coils o f different geometry, tube diameters and temperature profiles can be identical provided their average residence times and hydrocarbon partial pressures fall on the same selectivity line. The same conclusions have also been drawn from kinetic considerations. This approach has been used to extend the pyrolysis selectivity lines into the millisecond region. A mathematical model for calculating the rate o f coking i n the pyrolysis


20.

FERNANDEZ-BAUJIN

AND

Lumus

SOLOMON

SRT III Module

coil has been discussed. These models have been confirmed with performance data from commercial pyrolysis reactors. The pyrolysis selectivity correlation and the coking models have been combined with momentum heat and mass transfer models to design pyrolysis coils. This application has led to the SRT III pyrolysis reactors which emphasizes low hydrocarbon partial pressure by employing a coil design with large diameter outlet tubes.


The approach of a high capacity, high selectivity and lower complexity reactor is viewed to best meet the needs of the olefins plants of the 1980's. ABSTRACT The hydraulic and kinetic factors affecting pyrolysis coil performance are defined. These factors are correlated to quantify their effects on pyrolysis yields and coking tendencies for a given feedstock. The correlation is supported by theoretical and experimental considerations. Based on momentum, heat and mass transfer, the effect of coil diameter on module performance is discussed. A logical result of the above considerations is the Lummus SRT® III pyrolysis module. ACKNOWLEDGEMENT The authors express their appreciation to Dr. K.W. Li of Lummus for his assistance in reviewing and offering invaluable advice on the contents of this paper. REFERENCES 1.

Williams, K.D. and H.G. Davis, "Mechanistic Studies of Ethane Pyrolysis at Low Pressures," American Chemical Society, Philadelphia Meeting, April, 1975.

2.

Dorn, R.K. and M.J. Maddock, "Design Pyrolysis Heater for Maximum Profits," Hydrocarbon

Processing, November, 1972.

3. Fernandez-Baujin, J.M. "Factors Affecting Pyrolysis Selectivity," Safety and Reliability of Large Single Train Ethylene Plants, Unpublished Report, New York, May, 1974.


371

372


4. Maddock, M.J., "SRT Heater Design and Engineering Characteri stics", Safety and Reliability of Large Single Train Ethylene Plants, Third Ethylene Seminar-Lummus, Unpublished Report, New York; May, 1972.


5. Brooks, M.E. and J. Newman, "Gas Oil Cracking — The Problems That Had to be Solved," VIII World Petroleum Congress, Moscow, U.S.S.R., June, 1971. 6. Fernandez-Baujin, J.M., and A.J. Gambro, "Technology and Econ omics for Modern Olefins Plants," VI Interamerica Congress of Chemical Engineering, Caracas, Venezuela; July, 1975. 7. Chambers, L.E. and W.S. Potter, Hydrocarbon Processing, p. 121, January, 1974; p. 95, March, 1974; p. 99, August, 1974. 8. Smith, J.H., Chemical Engineering Kinetics, McGraw-Hill Book Company, New York, 1970. 9. Hirato, M., Yoshioka, S., and M. Tanaka, "Gas Oil Pyrolysis by Tubular Reactor and its Simulation Model of Reaction," Hitachi Review 20:8,326, 1971. 10. Bennett, C.O. and J.E. Myers, "Momentum, Heat and Mass Trans fer," McGraw-Hill Book Company, New York, 1962. 11. Chen, J. and M.J. Maddock, "How Much Spare Heater for Ethylene Plants?" Hydrocarbon Processing, May, 1973. 12. DeBlick, J.L. and A.G. Goossens, "Optimize Olefin Cracking Coils," Hydrocarbon Processing, March, 1971, p. 76. 13. Mol, Α., "How Various Parameters Affect Ethylene Cracker Run Lengths," Hydrocarbon Processing, July, 1974. 14. Maddock, M.J., "Coking on Pyrolysis Heaters," Private Communi cation, Lummus, November, 1971. 15. Chen, J. and W. Vogel, "Fouling of Transfer Line Exchangers in Ethylene Service," 74th National Meeting, A.I.Ch.E., New Orleans, La. March, 1973.


Industrial and Laboratory Pyrolyses - American Chemical Society

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