CN111523194B - Modeling method of reaction rate and conversion rate regulation model under MIHA pure pneumatic operation condition - Google Patents
Modeling method of reaction rate and conversion rate regulation model under MIHA pure pneumatic operation condition Download PDFInfo
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Abstract
The invention relates to a modeling method of a reaction rate and conversion rate regulation model under the MIHA pure pneumatic operation condition, which builds an energy conversion model in a bubble breaker by analyzing a bubble generation process under the pure pneumatic condition; based on an energy conversion model and liquid circulation in the bubble breaker, calculating liquid flow, obtaining the energy dissipation rate and bubble size of a gas-liquid intensive mixing area, and finally obtaining a reaction rate and conversion rate calculation model. The method establishes a reaction rate and conversion rate regulation model under the pure pneumatic operation condition aiming at the MIHA, comprehensively reflects the influence of the reactor structure, the physical properties of the system and the operation parameters and the input energy on the reaction rate and the conversion rate, and can realize the guidance on the reactor design and the reaction system design of the MIHA and the guidance on the design of the efficient reactor structure and the reaction system.
Description
Technical Field
The invention belongs to the technical field of reactors and modeling, and particularly relates to a modeling method for a reaction rate and conversion rate regulation model under the MIHA pure pneumatic operation condition.
Background
For global environmental protection, the sulfur content of the marine fuel oil must be reduced, for example, the sulfur content of the open sea marine fuel oil must be reduced to 0.5%, so that it is imperative to replace the high sulfur residue fuel oil with the low sulfur distillate fuel oil. Most of the sulfur in crude oil is present in residuum, which is mainly distributed in aromatics, colloids and asphaltenes, with most of the sulfur being present in the form of five membered ring thiophenes and thiophene derivatives. The sulfur in the residuum is typically removed by breaking the C-S bonds of the residuum macromolecules via a hydrogenolysis reaction to convert the sulfur to hydrogen sulfide. Sulfur present in non-asphaltenes is easier to remove under hydrogenation conditions and can reach higher conversion depths. However, asphaltene is a macromolecule with the largest relative molecular mass, the most complex structure and the strongest polarity in the residual oil, so that sulfur in the macromolecule is difficult to remove, and the desulfurization rate in the residual oil hydrodesulfurization process is limited.
The conversion of sulfur-containing asphaltenes is critical during residuum hydrodesulfurization (hereinafter MIHA). The core of asphaltenes is a highly condensed, fused aromatic ring system. The condensed aromatic ring system is provided with alkyl and cycloalkyl structures with different numbers and sizes, is the component with the largest condensation degree in residual oil, contains heteroatoms such as S, N, O, metal and the like, and has complex morphology and molecular structure. In the residuum hydroconversion process, asphaltene mainly undergoes two reactions of cracking from macromolecules to small molecules and condensation from small molecule dehydrogenation polymerization to macromolecules. The invention takes asphaltene hydrodesulfurization reaction as a model reaction of residual oil hydrogenation process, and examines the influence of the structure of a reactor, physical properties and operation parameters of a system and input energy on the reaction rate and conversion rate in a bubble breaker.
Disclosure of Invention
The invention aims to provide a modeling method of a reaction rate and conversion rate regulation model under the MIHA pure pneumatic operation condition so as to study the influence of the reactor structure, the physical properties and operation parameters of a system and input energy on the reaction rate and conversion rate, thereby realizing the guidance on the MIHA reactor design and the MIHA reaction system design.
MIHA microbubbles can be formed in three ways: pure hydraulic, pure pneumatic and gas-liquid linkage. Under the pure hydraulic and pure pneumatic operation conditions, the energy required by the system operation and the formation of micro bubbles is completely provided by liquid mechanical energy or gas static pressure energy; under the condition of gas-liquid linkage operation, three modes can be adopted for forming MIHA micro-bubbles of a gas static pressure energy and a liquid machine, namely: pure hydraulic, pure pneumatic and gas-liquid linkage. Under the pure hydraulic and pure pneumatic operation conditions, the energy required by the system operation and the formation of the micro-bubbles is completely provided by the liquid mechanical energy. The invention discusses a modeling method of a reaction rate and conversion rate regulation model under a pure pneumatic operation condition, and the method comprises the following steps:
the hydrogen and asphaltene are respectively A and B, and the hydrogen and the asphaltene react in a stoichiometric ratio of 1:1, namely the macroscopic reaction rates of the hydrogen and the asphaltene are equal, and the reaction level of the B is 0.5; when the influence of the liquid-solid mass transfer resistance of the catalyst surface on the macroscopic reaction rate is not considered, the macroscopic reaction rate r of A under the unit volume of the catalyst A The definition according to this is simplified as follows:
wherein k is A Is the intrinsic reaction rate constant, k of A G Is the gas phase mass transfer coefficient, k L Is the mass transfer coefficient of liquid phase, P G Is the gas pressure in the bubble, a is the phase boundary area, H A Is Henry coefficient;is the average molar concentration of B in the reactor;
since the stoichiometric ratio of the A and B reactions is 1:1, k is A Intrinsic reaction rate constant k equivalent to B B The method comprises the steps of carrying out a first treatment on the surface of the Intrinsic reaction rate constant k of B at mass space velocity B 0 Expressed as:
the mass of the catalyst is converted into volume and the unit is simplified, and the actual intrinsic reaction rate constants k of A and B A And k B The expression is as follows:
wherein ρ is c And w c The catalyst density and the actual dosage, ρ L The density of the liquid is the density of the liquid, and T is the temperature of the system;
H A the expression is as follows:
wherein M is 1 Is the molar mass of the oil product of the liquid raw material;
conversion with B R T Initial concentration C B0 The related is that:
C B0 related to oil density, sulfur-containing asphaltene mass fraction and molar mass thereof, namely:
wherein M is B Is the molar mass of B, w B The mass fraction of B in the oil product;
R T and reaction rate r A And the number of moles N of B fed into the reactor per unit time L Regarding the fully mixed flow gas-liquid reactor under purely pneumatic operating conditions, the macroscopic reaction rates of B are considered to be equal throughout the reactor, i.e.:
R T ＝r A /N L (7)
wherein N is L Calculated in combination with formula (6) as follows:
taking N into account L And r A Is identical in units of (a); wherein Q is L For circulating flow of liquid, V C V is the volume of catalyst in the reactor C Density ρ of catalyst c In relation, the calculation is based on the following formula:
V C ＝ρ L S 0 H 0 (1-φ G )w c /ρ C (8)
wherein H is 0 For the initial level of the reactor, S 0 And reactor cross-sectional area phi G Is the gas content;
the following formulas (7) to (9) can be used:
substituting the above formula into formula (5) yields:
substituting equations (3) and (10) into equation (1) yields:
a reaction rate and conversion model was obtained based on the formulas (9) and (11).
It is another object of the present invention to provide a reaction rate and conversion rate regulation model under the pure pneumatic operating conditions of MIHA constructed by the above method.
It is a further object of the present invention to provide a reactor designed by the above process.
The structure of the reactor of the present invention can be seen in the patent CN106187660a previously filed by the inventor, and the description of the present invention is omitted. The invention utilizes the influence of constructed model reactor structure, system physical property and operation parameter and input energy on bubble scale, thereby carrying out relevant reactor structure parameter design according to the requirement.
The method establishes a reaction rate and conversion rate regulation model under the pure pneumatic operation condition aiming at the MIHA, comprehensively reflects the influence of the reactor structure, the physical properties of the system and the operation parameters and the input energy on the reaction rate and the conversion rate, and can realize the guidance on the reactor design and the reaction system design of the MIHA and the guidance on the design of the efficient reactor structure and the reaction system.
Drawings
FIG. 1 is a schematic diagram of a physical model of a bubble generation process under purely pneumatic conditions;
FIG. 2 is the operating pressure versus reaction rate r A Is a function of (1);
FIG. 3 is the operating pressure versus conversion R T Is a function of (1);
FIG. 4 is a graph of operating temperature versus reaction rate r A Is a function of (1);
FIG. 5 is the operating temperature versus conversion R T Is a function of (1);
FIG. 6 is a graph showing the difference in air supply pressure versus the reaction rate r A Is a function of (1);
FIG. 7 is a graph of differential air supply pressure versus conversion R T Is a function of (1);
FIG. 8 is ventilation Q G For reaction rate r A Is a function of (1);
FIG. 9 is ventilation Q G For conversion rate R T Is a function of (a) and (b).
Detailed Description
The technical scheme of the invention is further described below with reference to the attached drawings and the detailed description.
Example 1
It is generally believed that the basic structural unit of asphaltenes is a fused aromatic ring system as the core and incorporates a plurality of cycloalkane rings with a plurality of alkyl side chains of varying sizes on the aromatic and cycloalkane rings, which are interrupted by various heteroatom groups such as sulfur, oxygen, nitrogen, and the like, and which are complexed with metals such as vanadium, nickel, iron, and the like. The asphaltene molecule is composed of a plurality of (generally 4-6) structural units (or called unit sheets) with condensed ring aromatic ring systems as cores, wherein the structural units are connected by alkyl bridges or sulfur bridges with different lengths. Microstructure studies have shown that asphaltenes are typically single particles of semi-ordered, cumulative graphite unit cells formed from large molecular weight lamellar, condensed aromatic hydrocarbons stacked upon one another, while also containing small amounts of metalloporphyrin structures bound together by pi-electron interactions.
The hydrogen and asphaltene are respectively A and B, and the hydrogen and the asphaltene react in a stoichiometric ratio of 1:1, namely the macroscopic reaction rates of the hydrogen and the asphaltene are equal, and the reaction level of the B is 0.5; when the influence of the liquid-solid mass transfer resistance of the catalyst surface on the macroscopic reaction rate is not considered, the macroscopic reaction rate r of A under the unit volume of the catalyst A (mol/m 3 cat.s) is simplified according to its definition as follows:
wherein k is A Is the intrinsic reaction rate constant, k of A G Is the gas phase mass transfer coefficient, k L Is the mass transfer coefficient of liquid phase, P G Is the gas pressure in the bubble, a is the phase boundary area, H A Is Henry coefficient;is the average molar concentration of B in the reactor;
since the stoichiometric ratio of the A and B reactions is 1:1, k is A Intrinsic reaction rate constant k equivalent to B B The method comprises the steps of carrying out a first treatment on the surface of the Intrinsic reaction rate constant k of B at mass space velocity B 0 Expressed as:
the mass of the catalyst is converted into volume and the unit is simplified, and the actual intrinsic reaction rate constants k of A and B A And k B Expressed as ((m) 3 /mol) 0.5 /s)：
Wherein ρ is c And w c The catalyst density and the actual dosage, ρ L The density of the liquid is the density of the liquid, and T is the temperature of the system; for the system of the invention, the catalyst is used in an amount of about 1% of the oil mass and the loading is 5%, so that the actual catalyst amount w c About 0.05% (w/w);
H A the expression is as follows:
wherein M is 1 Is the molar mass, g/mol, of the oil product of the liquid raw material;
conversion with B R T Initial concentration C B0 The related is that:
C B0 related to oil density, sulfur-containing asphaltene mass fraction and molar mass thereof, namely:
wherein M is B Is the molar mass of B, w B The mass fraction of B in the oil product;
R T and reaction rate r A And the number of moles N of B fed into the reactor per unit time L In a related aspect, the system of the invention is a fully mixed flow gas-liquid reactor under purely pneumatic operating conditions, and macroscopic reaction rates of B at all positions in the reactor are considered to be equal, namely:
R T ＝r A /N L (7)
wherein N is L Calculated in combination with formula (6) as follows:
taking N into account L And r A Is identical in units of (a); wherein Q is L For circulating flow of liquid, V C V is the volume of catalyst in the reactor C Density ρ of catalyst c In relation, the calculation is based on the following formula:
V C ＝ρ L S 0 H 0 (1-φ G )w c /ρ c (8)
wherein H is 0 For the initial level of the reactor, S 0 And reactor cross-sectional area phi G Is the gas content; for the system of the present invention ρ c Approximate to oil density ρ L Equal;
the following formulas (7) to (9) can be used:
substituting the above formula into formula (5) yields:
substituting equations (3) and (10) into equation (1) yields:
M B the value M is obtained in the calculation of example 3 according to the present invention, depending on factors such as the molar mass of the sulfur-containing asphaltenes, the polarity of the extraction solvent, the temperature, etc., which are approximately in the range of several hundred to 6000 B ＝1000g/mol。w B Depending on factors such as crude oil production area, w is set in the system of the invention B ＝4.0％。
Based on the above data, r in equation (11) can be determined by trial and error A . However, a large number of calculations show that under different working conditions (operating pressure, operating temperature, air supply pressure difference and ventilation), the denominator intrinsic reaction rate resistance term in the equation (11) corresponds to the final r A The impact of the size is small and therefore can be ignored in the actual calculation process.
In the above model, the gas phase mass transfer coefficient k G And a liquid phase mass transfer coefficient k L The method comprises the following steps of:
(1) Establishing a micro-bubble rising speed model under the MIHA pure pneumatic operation condition;
assuming that both the bubbles and the liquid in the reactor move vertically upwards, according to applicant's prior studies, the average rising velocity v of the bubbles in the reactor 32 Calculation based on the following formula:
wherein v is 0 、v G And v L The average diameters of the bubbles are d respectively 32 The rising velocity, apparent gas velocity and apparent liquid velocity of the bubbles in the infinitely large stationary liquid;
for v 0 Calculated based on the following formula:
wherein ρ is L Sum sigma L Respectively the density of liquid and the interfacial tension, mo is Morton number, d e Is equivalent diameter, K b The equivalent diameter and equation parameters are related to physical properties and are determined through experiments; for the MIHA system, c=1.4, n=0.8, due to the presence of various organic components;
d e ＝d 32 (ρ L g/σ L ) 1/2 (15)
K b ＝K b0 M O -0.038 (16)
wherein mu L Is hydrodynamic viscosity; for MIHA system, take K b0 ＝10.2；
For v G And v L Calculated based on the following formula:
v G ＝4Q G /πD 0 2 (17)
v L ＝4Q L /πD 0 2 (18)
wherein Q is G For the intake flow rate, Q L For circulating flow of liquid in bubble breaker, D 0 Is the reactor diameter;
(2) Establishing a gas phase mass transfer coefficient k under MIHA pure pneumatic operation condition G The model is as follows:
wherein d 32 Is the average diameter of the bubble sauter, m; t is t 32 S is the residence time of the bubbles in the reactor; d (D) G Is the gas-phase diffusion coefficient of gas in liquid, m 2 /s；
Wherein H is 0 For the initial level of liquid in the reactor, phi G The gas content in the bubble breaker;
gas diffusion coefficient D G Equation prediction based on Chapman-Enskog theory of motion is as follows:
wherein M is A And M B The molar masses of the gas and the liquid respectively, T is the temperature in the reactor, and P G Is the gas pressure in the bubble; neglecting the saturated vapor pressure of the liquid in the bubble, P G Approximately equal to the operating pressure P above the liquid level m ；v i For the molecular diffusion volume, the values can be calculated by reference to Fuller E N et al (Fuller E N, schettler P D, giddings J C.New method for prediction of binary gas-phase diffusion coefficients [ J)].Industrial&Engineering Chemistry,1966,58(5):18–27.)；
(3) Establishing a liquid phase mass transfer coefficient k under MIHA pure pneumatic operation condition L The model is as follows:
surface update time, k, defined according to Higbie permeation theory and velocity slip theory L Calculation based on the following formula:
wherein v is s Is the slip velocity between the bubble and its surrounding liquid, m/s; d (D) L Is the liquid phase diffusion coefficient, m 2 /s；
D L Calculation based on Stokes-Einstein correction formula:
in the above, D AB Is the diffusion coefficient of the gaseous solute A in the solvent B, cm 2 /s；M B Is the molar mass, g/mol, of the solvent B; mu (mu) B Is the viscosity of solvent B, cP; t is the system temperature, K; v (V) A Is that the gas solute A has normal boiling pointMolar volume in cm 3 Per mol, calculated approximately from van der Waals contrast equation for the actual gas:
wherein P is c 、T c 、V A，C Critical pressure (Pa), critical temperature (K) and critical molar volume (cm) of A respectively 3 /mol); p and T are the actual pressure and temperature of the system respectively; the parameter values of the hydrogen are respectively as follows:
P c ＝1.313×10 6 Pa，T c ＝33.19K，V A，C ＝64.147cm 3 /mol；
obtaining a gas phase mass transfer coefficient k based on the components (19), (23), (25) G And a liquid phase mass transfer coefficient k L 。
Average diameter d of bubble sauter in the above model 32 And the internal phase boundary area a of the reactor is obtained by the following way:
(1) Analyzing the bubble generation process under the pure pneumatic condition, and establishing an energy conversion model in a bubble breaker;
before the gas is not introduced, the bubble breaker is filled with a static reaction liquid. After the start of the gas introduction, due to the gas pressure P G And system operating pressure P m There is a pressure difference deltap between them, the hydrostatic energy of the gas will be transferred to the liquid, causing turbulence of the liquid, and the pressure of the gas itself will drop rapidly to the operating pressure within the MIHA. Due to the flow of the gas-liquid two phases, the gas-liquid flows out from the bubble breaker. For pneumatic operating conditions, the liquid flow rate Q L Much smaller than the gas flow Q G The energy required for the system to operate is almost entirely provided by the gas pressure energy.
Establishing a physical model diagram as shown in fig. 1:
the system liquid is assumed to be in closed cycle, i.e. the liquid amount does not change in the whole process. Part of the liquid will be forced into the bubble breaker external circulation line due to the ingress of gas. Setting the length of the bubble breaker as L (m) and the diameter as D 1 (m) cross-sectional area ofS 1 (m 2 )(S 1 ＝πD 1 2 /4). Nozzle diameter D N (m)。
The assumptions are made as follows:
(1) steady state operation, operating pressure P m Constant;
(2) the change of potential energy of liquid and the change of gas pressure in the bubbles caused by interfacial tension of the bubbles are ignored because of higher actual operation pressure;
(3) since the gas density is much smaller than the liquid, the kinetic energy of the input gas is ignored.
The energy balance under steady state conditions was performed using a bubble breaker as a control body. Under pneumatic conditions, the pressure is P G0 (Pa) and a volume flow of Q G0 (m 3 Gas inlet operating pressure of/s) is constant at P m The bubble breaker of (Pa) converts partial static pressure energy of gas released into liquid kinetic energy and bubble surface energy. The static pressure energy released by the gas is equivalent to the work W of the gas on the system G (W) according to the work definition:
Q G (m 3 s) is the gas flow rate in the bubble breaker, and for simplicity, it is assumed that the gas is an ideal gas within the scope of the present invention, and is obtained according to the ideal gas state equation:
in the formula (28), ρ G0 (Kg/m 3 ) And M A (Kg/mol) is the density and molar mass of the gas entering the breaker, respectively; r (8.314J/mol.K) and T (K) are the gas constant and the gas temperature, respectively.
Substituting formula (28) into formula (27) and integrating to obtain:
let the difference between the gas pressure at the gas inlet of the bubble breaker and the operating pressure be Δp (Pa), namely:
ΔP＝P G0 -P m (30)
since ΔP > 0, W is G < 0, i.e. the mechanical energy of the gas will be reduced after it enters the bubble breaker. Due to bubble breaker operating pressure P m Constant and relatively negligible liquid gravitational potential energy, so the reduced mechanical energy of the gas will be converted into liquid kinetic energy and bubble interface energy. The following relationship can be obtained from the formulas (29) and (30):
to the left of equation (31) is a reduction in gas static pressure energy (-W) G ) Namely, the energy source required by the system operation is adopted; the right two terms of equation (31) are respectively the liquid kinetic energy and the gas-liquid interfacial energy. Wherein ρ is L (Kg/m 3 ) Sum sigma L (N/m) liquid density and interfacial tension, respectively; u (U) L (m/s) the linear velocity of the liquid flowing from the disruptor; d, d 32 (m) is the average diameter of the bubbles Sauter flowing out from the bubble breaker; q is calculated according to mass balance G And Q is equal to G0 The following relationship is provided:
for the study of the present invention, ΔP < P m Thus, Q G ≈Q G0 . For convenience of description, the flow rates of the gas entering and exiting are referred to as Q G And (3) representing. Preliminary calculations indicate that the gas-liquid interface energy value is negligible relative to the liquid kinetic energy value. This term is ignored first and then checked by calculation. Thus, equation (31) can be reduced to:
(2) Calculating a liquid flow based on the energy conversion model and the liquid circulation in the bubble breaker;
according to the assumption of closed cycle, the flow of the inlet and outlet liquid is equal, so there is
Q L ＝U L S 1 (1-φ G ) (34)
The gas content phi in the bubble breaker G Calculated according to formula (21):
from the formula (34) (31):
obviously U L Is the apparent velocity of the gas-liquid mixture in the bubble breaker. Substituting equation (35) into equation (33) yields:
from equation (36), the liquid flow Q at the nozzle diameter due to gas input can be calculated L But the form is more complex, and the design is reasonably simplified according to the actual situation of the project. From equation (33):
calculations indicate that under the conditions studied in the present invention, Q L <<Q G . Equation (36) can be simplified as:
this gives:
in fact, from the ideal state equation, the following relationship exists:
substituting equation (40) into equation (39) yields:
from equation (41): bubble breaker cross-sectional area S 1 For liquid circulation flow rate Q L The effect is greater;
as a result of:
v in N Is the flow rate at the nozzle;
when V is N At a certain time, it is obtainable by the formulas (41) and (42):
when D is N At a certain time, it is obtainable by the formulas (41) and (42):
from formulas (35) and (41):
the above is based on Q under all pneumatic conditions L Is a rough calculation of (a). Further according to known V N Determining diameter D N (when D N V can also be obtained at a fixed time N )。
(3) Calculating the energy dissipation rate of the gas-liquid intensive mixing areaε mix ；
d 32 Energy dissipation ratio epsilon of gas-liquid intensive mixing area in bubble breaker mix Closely related. According to the first law of thermodynamics:
in the above, L mix The length of a gas-liquid intensive mixing zone in bubble breaking is m; lambda (lambda) 1 Is the ratio of the gas-liquid volume flow rate (lambda) 1 ＝Q G /Q L )。K 1 Is the ratio of the diameter of the bubble breaker nozzle to the diameter of the breaker (K 1 ＝D N /D 1 )。
Evans et al have derived L based on the principle of conservation of kinetic energy mix But is not applicable to the cases involved in the studies of the present invention and thus needs to be deduced again. The research of the invention considers that L mix Depending on the length of the decay of the highest flow rate of liquid in the bubble breaking zone until it disappears. The highest flow velocity of the liquid is in the attenuation process, and the central line speed U of the liquid is the same jm The attenuation law of (2) is not influenced by the disturbance of surrounding bubbles, and accords with the following attenuation law:
in equation (47), x is the horizontal distance from the bubble breaker core to the maximum speed. When U is jm Decaying to the apparent speed U of the gas-liquid mixture L At this point, the high velocity disappears, after which a uniform gas-liquid mixture stream will be formed. Thus, L mix Is U (U) jm ＝U L X value at that time. Namely:
the equation (48) is simplified to obtain:
substituting equation (49) into (46) and simplifying it yields:
combining (41) (45) and equation (50) to calculate ε mix ；
(4) Calculating the bubble size of micro bubbles in MIHA;
microbubbles d in MIHA 32 Calculating according to the following formula;
d max ＝0.75(σ L /ρ L ) 0.6 ε mix -0.4 (51)
d min ＝11.4(μ L /ρ L ) 0.75 ε mix -0.25 (52)
wherein d min Is the minimum diameter of the bubble; d, d max Is the maximum diameter of the bubble; mu (mu) L Is hydrodynamic viscosity;
(5) Calculating the phase boundary area of the micro gas-liquid system;
the phase boundary area is calculated according to the following formula:
example 2
This example specifically illustrates a reaction rate and conversion rate control model constructed based on the method of example 1.
The reaction rate and conversion regulation model obtained based on the modeling method of example 1 is as follows:
d e ＝d 32 (ρ L g/σ L ) 1/2 (15)
K b ＝K b0 M o -0.038 (16)
v G ＝4Q G /πD 0 2 (17)
v L ＝4Q L /πD 0 2 (18)
d max ＝0.75(σ L /ρ L ) 0.6 ε mix -0.4 (51)
d min ＝11.4(μ L /ρ L ) 0.75 ε mix -0.25 (52)
example 3
Based on the modeling method of example 1, it can be seen that for asphaltene hydrogenation reactions in the MIRA for purely aerodynamic conditions, the factors that determine the final conversion of asphaltenes include:
reactor structural parameters: diameter D of crusher 1 Diameter D of reactor 0 Diameter ratio K of nozzle to breaker 1 Reactor height H 0 ；
Operating parameters: operating pressure P m Operating temperature T, supply air flow Q G ；
Physical parameters: density ρ of residuum L Interfacial tension sigma of residuum L Dynamic viscosity of residue mu L ；
Energy parameters: the air supply pressure difference deltap;
intrinsic reaction rate: k (k) A ；
In order to optimize the MIHA structure and the actual operation, the present embodiment is based on the modeling method of example 1, and the operating pressure, operating temperature and supply are studied for specific reactor structure and reaction systemDifferential air pressure Δp and ventilation Q G Influence on the reaction rate and conversion. Because the physical parameters of the residual oil are mainly determined by the operation parameters, the structural parameters are limited by the design of practical devices, the invention focuses on discussing the operation parameters P of the reactor m T, ΔP vs macroscopic reaction rate r A Influence of asphaltene conversion RT.
The general calculation conditions are as follows:
diameter D of crusher 1 =0.02m; ratio K of diameter of bubble breaker nozzle to breaker diameter 1 ＝0.5；
Density ρ of residuum L ＝800kg/m 3 ；
Interfacial tension sigma of residuum L The fitting formula is as follows:
σ L ＝[31.74-0.04775(T+273.15)]×10 -3 (N/m)；
dynamic viscosity of residuum mu L The fitting formula is as follows;
(1) The effect of operating pressure on reaction rate and conversion;
the calculation conditions are as follows:
ventilation Q G =80l/h; operating pressure P m =10 to 20MPa; air supply pressure difference Δp=6 MPa; gas temperature t=500℃.
Macroscopic reaction Rate r of operating pressure on Sulfur-containing asphaltenes A As shown in FIG. 2, the system operating pressure P is singly varied m Only affects the mass transfer process of the air film, the mass transfer rate of the liquid film is hardly affected, when P is m When the gas film mass transfer resistance increases, the reaction is not easy to proceed, but at the same time, the macroscopic reaction rate r A Approximately in a linear increasing trend. This reflects: (1) For sulfur-containing asphaltene hydrodesulfurization reactions, increasing the operating pressure will increase the reaction rate; (2) The gas film mass transfer resistance increases with the increase of the operating pressure, but the influence on the reaction rate is small and can be ignored approximately.
Operating pressure versus sulfur asphaltene conversion R T As shown in FIG. 3, it can be seen that the sulfur-containing asphaltene conversion R T With operating pressure P m Is approximately linear.
(2) Influence of the operating temperature on the reaction rate and conversion;
the calculation conditions are as follows:
ventilation Q G =80l/h; operating pressure P m =14 MPa; air supply pressure difference Δp=6 MPa; the gas temperature t=400 to 500 ℃.
Macroscopic reaction Rate r of operating temperature on Sulfur-containing asphaltenes A The effect of (2) is shown in figure 4.
Operating temperature versus sulfur asphaltene conversion R T As shown in fig. 5, it can be seen that increasing the operating temperature favors the conversion of sulfur-containing asphaltenes.
(3) Influence of the gas supply pressure difference deltap on the reaction rate and conversion rate;
the calculation conditions are as follows:
density ρ of residuum L ＝800Kg/m 3 The method comprises the steps of carrying out a first treatment on the surface of the Operating pressure P m =14 MPa; the air supply pressure difference delta P=1 to 10MPa; gas temperature t=450℃.
Macroscopic reaction rate r of air supply pressure difference to sulfur-containing asphaltene A The influence of (a) is shown in FIG. 6 (ventilation 80L/h), and the air supply pressure difference DeltaP is r A The influence of (a) is mainly due to the change of mass transfer resistance (i.e., the inverse of the volumetric mass transfer coefficient) on both sides of the liquid film.
Pressure difference of gas supply to conversion rate of sulfur-containing asphaltene R T As shown in fig. 7, the air supply pressure difference increases, the bubble breaker energy dissipation rate increases, the bubble diameter decreases, but the asphaltene conversion rate decreases. Since the present invention is purely pneumatic in operation, an increase in the air supply pressure difference tends to accelerate the flow of liquid in the reactor, which produces two results: on the one hand, the enhanced turbulence of the liquid in the bubble breaker leads to an increased energy dissipation rate and a decreased bubble; on the other hand, the residence time of the bubbles in the reactor is shortened, resulting in a reduction in the gas content, and the latter effect is more pronounced, as a result of which the gas-liquid phase interfacial area is reduced.
(4) Ventilation Q G Influence on reaction rate and conversion rate;
the calculation conditions are as follows:
ventilation Q G =1 to 100L/h; operating pressure P m =14 MPa; the air supply pressure difference delta P=0.1-10 MPa; gas temperature t=500℃.
Ventilation Q G Macroscopic reaction Rate r for Sulfur-containing asphaltenes A The effect of (2) is shown in figure 8;
ventilation Q G Conversion to sulfur-containing asphaltenes R T The effect of (2) is shown in figure 9.
Claims (4)
1. The modeling method of the reaction rate and conversion rate regulation model under the MIHA pure pneumatic operation condition is characterized by comprising the following steps:
the hydrogen and asphaltene are respectively A and B, and the hydrogen and the asphaltene react in a stoichiometric ratio of 1:1, namely the macroscopic reaction rates of the hydrogen and the asphaltene are equal, and the reaction level of the B is 0.5; when the influence of the liquid-solid mass transfer resistance of the catalyst surface on the macroscopic reaction rate is not considered, the macroscopic reaction rate r of A under the unit volume of the catalyst A The definition according to this is simplified as follows:
wherein k is A Is the intrinsic reaction rate constant, k of A G Is the gas phase mass transfer coefficient, k L Is the mass transfer coefficient of liquid phase, P G Is the gas pressure in the bubble, a is the phase boundary area, H A Is Henry coefficient;is the average molar concentration of B in the reactor;
since the stoichiometric ratio of the A and B reactions is 1:1, k is A Intrinsic reaction rate constant k equivalent to B B The method comprises the steps of carrying out a first treatment on the surface of the Intrinsic reaction rate constant k of B at mass space velocity B 0 Expressed as:
the mass of the catalyst is converted into volume and the unit is simplified, and the actual intrinsic reaction rate constants k of A and B A And k B The expression is as follows:
wherein ρ is c And w c The catalyst density and the actual dosage, ρ L The density of the liquid is the density of the liquid, and T is the temperature of the system;
H A the expression is as follows:
wherein M is 1 Is the molar mass of the oil product of the liquid raw material;
conversion with B R T Initial concentration C B0 The related is that:
C B0 related to oil density, sulfur-containing asphaltene mass fraction and molar mass thereof, namely:
wherein M is B Is the molar mass of B, w B The mass fraction of B in the oil product;
R T and reverse toResponse rate r A And the number of moles N of B fed into the reactor per unit time L Regarding the fully mixed flow gas-liquid reactor under purely pneumatic operating conditions, the macroscopic reaction rates of B are considered to be equal throughout the reactor, i.e.:
R T ＝r A /N L (7)
wherein N is L Calculated in combination with formula (6) as follows:
taking N into account L And r A Is identical in units of (a); wherein Q is L For circulating flow of liquid, V C V is the volume of catalyst in the reactor C Density ρ of catalyst c In relation, the calculation is based on the following formula:
V C ＝p L S 0 H 0 (1-φ G )w c /p c (9)
wherein H is 0 For the initial level of the reactor, S 0 For the cross-sectional area of the reactor, phi G Is the gas content;
the following formulas (7) to (9) can be used:
substituting the above formula into formula (5) yields:
substituting equations (3) and (11) into equation (1) yields:
a reaction rate and conversion model is obtained based on the formulas (10) and (12).
2. The method according to claim 1, wherein the gas phase mass transfer coefficient k G And a liquid phase mass transfer coefficient k L The method comprises the following steps of:
(1) Establishing a micro-bubble rising speed model under the MIHA pure pneumatic operation condition;
assuming that both the bubbles and the liquid in the reactor move vertically upward, the average rising velocity v of the bubbles in the reactor 32 Calculation based on the following formula:
wherein v is 0 、v G And v L The average diameters of the bubbles are d respectively 32 The rising velocity, apparent gas velocity and apparent liquid velocity of the bubbles in the infinitely large stationary liquid;
for v 0 Calculated based on the following formula:
wherein ρ is L Sum sigma L Respectively the density of liquid and the interfacial tension, mo is Morton number, d e Is equivalent diameter, K b Is a physical property parameter; for the MIHA system, c=1.4, n=0.8;
d e ＝d 32 (ρ L g/σ L ) 1/2 (16)
K b ＝K b0 Mo 0.038 (17)
wherein mu L Is hydrodynamic viscosity; for MIHA system, take K b0 ＝10.2；
For v G And v L Calculated based on the following formula:
v G ＝4Q G /πD 0 2 (18)
v L ＝4Q L /πD 0 2 (19)
wherein Q is G For the intake flow rate, Q L For circulating flow of liquid in bubble breaker, D 0 Is the reactor diameter;
(2) Establishing a gas phase mass transfer coefficient k under MIHA pure pneumatic operation condition G The model is as follows:
wherein d 32 Is the average diameter of the bubble sauter, t 32 For the residence time of the bubbles in the reactor, D G Is the gas phase diffusion coefficient of the gas in the liquid;
wherein H is 0 For the initial level of liquid in the reactor, phi G The gas content in the bubble breaker;
gas diffusion coefficient D G Equation prediction based on Chapman-Enskog theory of motion is as follows:
wherein M is A And M B The molar masses of the gas and the liquid respectively, T is the temperature in the reactor, and P G Is the gas pressure in the bubble; when the saturated vapor pressure of the liquid within the bubble is ignored,P G approximately equal to the operating pressure P above the liquid level m ；v i Is the molecular diffusion volume;
(3) Establishing a liquid phase mass transfer coefficient k under MIHA pure pneumatic operation condition L The model is as follows:
surface update time, k, defined according to Higbie permeation theory and velocity slip theory L Calculation based on the following formula:
wherein v is s D is the slip velocity between the bubble and the surrounding liquid r Is the liquid phase diffusion coefficient;
D L calculation based on Stokes-Einstein correction formula:
wherein D is AB Is the diffusion coefficient of gaseous solute A in solvent B, μ B Viscosity of B, V A The molar volume of A at normal boiling point is calculated approximately according to van der Waals contrast equation of actual gas:
wherein P is c 、T c 、V A，C The critical pressure, critical temperature and critical molar volume of A are respectively; p and T are the actual pressure and temperature of the system respectively;
obtaining the gas phase mass transfer coefficient k based on (20), (24), (26) G And a liquid phase mass transfer coefficient k L 。
3. The method of claim 1, wherein the in-reactor phase interface area a is obtained by:
(1) Analyzing the bubble generation process under the pure pneumatic condition, and establishing an energy conversion model in a bubble breaker;
under purely pneumatic operating conditions, the liquid flow rate Q L Gas flow rate Q G Before the gas is not introduced, the bubble breaker is filled with static reaction liquid; assuming that the system liquid is in closed cycle, namely the liquid amount does not change in the whole process; part of the liquid is forced to enter the outer circulation pipeline of the bubble breaker due to the entering of the gas; setting the length of the bubble breaker as L and the diameter as D 1 Cross-sectional area S 1 ＝πD 1 2 4; nozzle diameter D N ；
The assumptions are made as follows:
(1) steady state operation, operating pressure P m Constant;
(2) the change of potential energy of liquid and the change of gas pressure in the bubbles caused by interfacial tension of the bubbles are ignored because of higher actual operation pressure;
(3) since the gas density is much smaller than the liquid, the kinetic energy of the input gas is ignored;
taking a bubble breaker as a control body, and performing energy balance under a steady-state condition; under pneumatic conditions, the pressure is P G0 The volume flow is Q G0 Is constant at P m When the bubble breaker is used, part of static pressure energy is released by the gas and is converted into liquid kinetic energy and bubble surface energy; the static pressure energy released by the gas is equivalent to the work W of the gas on the system G According to the work definition, it can be seen that:
QG is the gas flow in the bubble breaker, and assuming that the gas is ideal gas, the gas is obtained according to an ideal gas state equation:
in formula (29), ρ G0 And M A The density and the molar mass of the gas entering the crusher are respectively; r and T are the gas constant and the gas temperature respectively;
substituting formula (29) into formula (28) and integrating to obtain:
let the difference between the gas pressure at the gas inlet of the bubble breaker and the operating pressure be Δp, namely:
ΔP＝P G0 -P m (31)
since ΔP > 0, W is G < 0, i.e. the mechanical energy of the gas will be reduced after it enters the bubble breaker; due to bubble breaker operating pressure P m Constant, neglecting the gravitational potential energy of the liquid, so that the reduced mechanical energy of the gas is converted into liquid kinetic energy and bubble interface energy; thus, the following formulas (30) and (31) can be obtained:
to the left of equation (32) is the reduction of the static pressure energy of the gas, i.e., -W G The method comprises the steps of carrying out a first treatment on the surface of the The right two terms of the equal sign of the equation (32) are respectively liquid kinetic energy and gas-liquid interface energy; wherein ρ is L Sum sigma L Liquid density and interfacial tension, respectively; u (U) L Is the linear velocity of the liquid flowing out of the disrupter; d, d 32 An average diameter of the air bubbles Sauter flowing out from the air bubble breaker; q is calculated according to mass balance G And Q is equal to G0 The following relationship is provided:
due to DeltaP <)＜P m Thus Q G ≈Q G0 The method comprises the steps of carrying out a first treatment on the surface of the Neglecting the gas-liquid interface energy value relative to the liquid kinetic energy value, the equation (32) is simplified to:
(2) Calculating a liquid flow based on the energy conversion model and the liquid circulation in the bubble breaker;
since the liquid entering and exiting the crusher is in closed cycle, i.e. the flow rates of the liquid entering and exiting are equal, the liquid entering and exiting are:
Q L ＝U L S 1 (1-φ G ) (35)
the gas content phi in the bubble breaker G Calculated according to formula (22):
from the formula (35) (32):
U L substituting equation (34) for the apparent velocity of the gas-liquid mixture in the bubble breaker, it is possible to obtain:
from equation (37), the liquid flow Q at the nozzle diameter due to gas input can be calculated L From equation (34):
under purely pneumatic operating conditions, Q L ＜＜Q G Then equation (37) is reduced to:
this gives:
from the ideal state equation, the following relationship exists:
substituting equation (41) into equation (40) yields:
from equation (42): bubble breaker cross-sectional area S 1 For liquid circulation flow rate Q L The effect is greater;
as a result of:
v in N Is the flow rate at the nozzle;
when V is N At a certain time, it is obtainable by the formulas (42) and (43):
when D is N At a certain time, it is obtainable by the formulas (42) and (43):
from formulas (36) and (42):
thereby completing the Q under the pure pneumatic condition L Is determined by the estimation of (a);
(3) Calculating energy dissipation rate epsilon of gas-liquid intensive mixing area mix ；
According to the first law of thermodynamics:
in the above, L mix The length of a gas-liquid intensive mixing zone in bubble breaking is the length of a gas-liquid intensive mixing zone; lambda (lambda) 1 Lambda is the ratio of the gas to the liquid volume flow 1 ＝Q G /Q L ；K 1 K is the ratio of the diameter of the bubble breaker nozzle to the diameter of the breaker 1 ＝D N /D 1 ；
L mix In connection with the length of the decay of the liquid maximum flow velocity in the bubble breaking zone until it has disappeared, the liquid maximum flow velocity during its decay, the central line velocity U jm The attenuation law of (2) is not influenced by the disturbance of surrounding bubbles, and accords with the following attenuation law:
in equation (48), x is the horizontal distance of the bubble breaker core to the maximum speed; when U is jm Decaying to the apparent speed U of the gas-liquid mixture L At this time, the high velocity disappears, after which a uniform gas-liquid mixture stream will be formed; thus, L mix Is U (U) jm ＝U L The value of x at that time, namely:
the reduction of equation (49) yields:
substituting equation (50) into (47) and simplifying it yields:
combining (42) (46) and equation (51) to calculate ε mix ；
(4) Calculating the bubble size of micro bubbles in MIHA;
microbubbles d in MIHA 32 Calculating according to the following formula;
d max ＝0.75(σ L /ρ L ) 0.6 ε mix -0.4 (52)
d min ＝11.4(μ L /ρ L ) 0.75 ε mix -0.25 (53)
wherein d min Is the minimum diameter of the bubble; d, d max Is the maximum diameter of the bubble; mu (mu) L Is hydrodynamic viscosity;
(5) Calculating the phase boundary area of the micro gas-liquid system;
the phase boundary area is calculated according to the following formula:
4. a reactor designed by the process of any one of claims 1-3.
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