FLUID COKING USING HIGH THRUST FEED NOZZLES

Abstract

A process for converting a heavy hydrocarbonaceous feedstock to liquid products is provided comprising introducing the hydrocarbonaceous feedstock into a fluid coker comprised in part of a fluidized bed of heated coke particles, the fluidized bed having a high velocity core region of heated coke particles and a low velocity annular region of unreacted hydrocarbon and coke particles using a plurality of high thrust nozzles and reacting the hydrocarbonaceous feedstock with the heated coke particles in the fluid coker to produce the liquid products.

Description

FIELD OF THE INVENTION

This invention relates to a fluid coking process for converting a heavy hydrocarbonaceous feedstock to liquid products which uses high thrust feed nozzles for injecting feedstock into the circulating fluidized bed of heated coke particles.

BACKGROUND OF THE INVENTION

Fluidized bed coking (fluid coking) is a petroleum refining process in which heavy petroleum feeds, typically the non-distillable residue (resid) from fractionation or heavy oils are converted to lighter, more useful products by thermal decomposition (coking) at elevated reaction temperatures, typically about 480 to 590° C., (about 900 to 1100° F.) and in most cases from 500 to 550° C. (about 930 to 1020° F.). Heavy oils that may be processed by the fluid coking process include heavy atmospheric resids, vacuum resids, aromatic extracts, asphalts, and bitumen from oil sands.

The process is carried out in a unit with a large reactor vessel containing hot coke particles that are maintained in the fluidized condition at the required reaction temperature with a fluidizing gas (e.g., steam) injected at the bottom of the vessel. The heavy oil feed is heated to a pumpable temperature, typically in the range of 350 to 400° C. (about 660 to 750° F.), mixed with atomizing steam, and fed through multiple feed nozzles arranged at several successive levels in the reactor. The steam is injected into a stripper section at the bottom of the reactor and passes upwards through the coke particles in the stripper as they descend from the main part of the reactor above. The feed liquid coats the coke particles in the fluidized bed, which make up the emulsion phase of the fluidized bed. As the thermal cracking reactions proceed, the liquid is transformed to vapour, which must migrate from the emulsion phase into the bubble phase in order to exit the system.

Liquid yields in fluid coking can be increased by reducing the reaction severity, or the time that molecules are exposed to process temperature. The typical approach taken to reduce reactor severity is to reduce reactor temperature. However, the downside of reducing temperature is increased stripper and sore thumb fouling, which can lead to reduced run lengths. Another approach to reduce reactor severity is to decrease the exposure time at high temperatures by providing short vapour phase residence times.

Long hydrocarbon vapour residence times are the most likely contributor to higher than expected “gas make”, defined as C₄-components, in the fluid coking process. Vapour-liquid equilibrium suppression, coupled with less than adequate mass transfer between the emulsion and bubble phases, is the most probable mechanism responsible for high “coke make”, defined as the toluene insoluble solid by-product of the thermal cracking reaction. Both phenomena result in lower liquid yields, and preliminary estimates suggest that they can contribute to as much as 11 wt % liquid yield loss. Optimizing the rate of removal of vapour from the emulsion phase should reduce the overall hydrocarbon vapour residence time of the reactor, increase liquid yields, and reduce gas make. It is estimated that a 3-5 wt % liquid yield increase can be achieved through maximizing vapour recovery from the reactor dense bed.

Technologies that increase mass transfer between the emulsion and bubble phase and, thus, reduce the gas phase residence time and increase hydrocarbon vapour stripping, are required.

SUMMARY OF THE INVENTION

It has been discovered that the reactor section of a fluid coker is comprised of a dilute, upward-flowing stream of gas in the central (core) region of the reactor and a dense, downward-flowing, outer (annular) region of particles. This is due to the vaporized hydrocarbons rising primarily in the core. Thus, the core region has a high vapour and low solids concentration (solids lean) and the annular region has a low vapour and high solids concentration (solids dense).

The present invention is directed to the use of high thrust feed nozzles to transport unreacted hydrocarbon and coke present in the annular region of the fluidized bed to the high velocity core region of the fluidized bed to improve hydrocarbon stripping, reduce the gas phase residence time, and increase liquid yields. Thrust is a mechanical force that is generated through the act of accelerating a mass of fluid. In other words, it is the reaction force created by the ejection of fluid from a nozzle at high velocity. The fluid pressure is related to the momentum of the fluid and acts perpendicular to any imposed boundary, which in this case is the fluidized solids in the reactor. The amount of thrust generated depends on the mass flow rate and the exit velocity of the fluid. High thrust can be achieved by either slightly accelerating a large mass of fluid, or greatly accelerating a small mass of fluid.

Prior art nozzles that are presently used in fluid cokers have a limited ability to transfer solids from the annular region of the fluidized bed to the upward flowing core of the fluidized bed. Thus, a process is provided herein for converting a heavy hydrocarbonaceous feedstock to liquid products, comprising:

• • introducing the hydrocarbonaceous feedstock into a fluid coker comprised in part of a fluidized bed of heated coke particles, the fluidized bed having a high velocity core region of heated coke particles and a low velocity annular region of unreacted hydrocarbon and coke particles; and
  • reacting the hydrocarbonaceous feedstock with the heated coke particles in the fluid coker to produce the liquid products;
    the feedstock being introduced into the fluid coker using a plurality of high thrust nozzles, said nozzles designed to transport unreacted hydrocarbon and coke from the low velocity annular region to the high velocity core region to improve hydrocarbon stripping, reduce gas phase residence time, and increase liquid products yields.

In one embodiment, the high thrust nozzles have a spray angle of about 3°-160° and a nozzle diameter between about 0.2 and 0.8″.

BRIEF DESCRIPTION OF THE DRAWINGS

In the accompanying drawings:

FIG. 1 is a simplified diagram of the reactor of a fluid coking unit useful in the present invention.

FIG. 2 shows the local pressure profiles along the length of a GEN2 feed nozzle when air and water are sprayed into ambient air, at air to liquid rations (ALRs) of 2.1 and 1.6 wt %.

FIG. 3 shows a drawing of a GEN3 feed nozzle having a diverging cloverleaf disperser.

FIG. 4 shows a drawing of a GEN4 feed nozzle, which consists of the same internal geometry as the GEN2 nozzle, but with slits at the nozzle tip.

FIG. 5 is a graph comparing measured axial thrust force (lb) with nozzle pressure (psig) for a variety of nozzles when spraying water only.

FIG. 6 shows the local pressure profile along the length of a Diffuser nozzle having a diverging/diffuser section when spraying air and water into ambient air at ALRs of 2.6 and 1.7 wt %.

FIGS. 7A and 7B show the differences in the jet plume for air-water mixtures exiting the GEN2 nozzle and the Diffuser nozzle, respectively.

FIG. 8 shows a drawing of the GEN1 nozzle, which consists of a simple constriction of 7° followed by a 3″ long straight section.

FIG. 9 is a graph of the measured axial thrust force (lb) as a function of the measured nozzle pressure for a variety of nozzles when spraying air and water.

FIG. 10 shows the local axial pressure profile along the length of a GEN2 nozzle with an additional diverging section at the exit when spraying air and water into ambient air at an ALR of 2.2 wt %.

DETAILED DESCRIPTION

The detailed description set forth below in connection with the appended drawings is intended as a description of various embodiments of the present invention and is not intended to represent the only embodiments contemplated by the inventor. The detailed description includes specific details for the purpose of providing a comprehensive understanding of the present invention. However, it will be apparent to those skilled in the art that the present invention may be practiced without these specific details.

The present invention is directed to the use of high thrust feed nozzles in a fluidized coking operation to push unreacted hydrocarbon and coke to the high velocity core region of the fluidized bed to improve hydrocarbon stripping, reduce the gas phase residence time, and increase liquid yields.

Thrust is a mechanical force that is generated through the act of accelerating a mass of fluid. In other words, it is the reaction force created by the ejection of fluid from a nozzle at high velocity (John and Keith, 2006). The fluid pressure is related to the momentum of the fluid and acts perpendicular to any imposed boundary, which in this case is the fluidized solids in the reactor. The amount of thrust generated depends on the mass flow rate and the exit velocity of the fluid. High thrust can be achieved by either slightly accelerating a large mass of fluid, or greatly accelerating a small mass of fluid.

There are three main factors that affect thrust: friction effects, axial momentum loss and thrust loss due to the pressure difference between the nozzle exit plane and the background. When friction is considered, it is best to have nozzles with large exit angles. However, the axial momentum losses increase as the angle increases since a higher percentage of the exiting flow will be non-axial.

Cruz (Cruz, N., “Interactions between Supersonic Gas Jets and Gas-Solid Fluidized Beds”, MSc Thesis, The University of Western Ontario, 2009) found that the thrust force of gas in a convergent-divergent nozzle used to attrition particles in a fluidized bed had a strong relationship with the particle grinding efficiency. During the attrition process, particles are entrained into the gas jet and are accelerated to high velocity where they collide with the particles in the dense phase of the fluidized bed at the tip of the jet plume and cause particle breakage to occur. This concept can be used to enhance the movement of solids from the annular section of the reactor to the core region in the feed zone by using feed nozzles that produce a high thrust force.

FIG. 1 is a simplified diagram of the reactor of a fluid coking unit. The reactor coking zone 10 contains a fluidized bed 11 of heated seed coke particles, heated to a temperature sufficient to initiate the coking (thermal cracking) reactions, into which the feedstock is added. The feedstock contacts the coke particles and reacts, and deposits a fresh coke layer on the hot fluidized coke particles circulating in the bed. The fluidized bed of coke comprises a dense bed surface 34, which is static, a dilute core region 32, which is upward flowing, and a dense annular region 30, which is downward flowing.

The feed is injected through multiple high thrust nozzles located in feed rings 12a to 12f, which are positioned so that the feed with atomizing steam enters directly into the fluidized bed of hot coke particles in coking zone 11. Each feed ring consists of a set of high thrust nozzles (typically 10-20, not designated in FIG. 1) that are arranged around the circular periphery of the reactor wall at a given elevation with each nozzle in the ring connected to its own feed line which penetrates the vessel shell (i.e. 10-20 pipes extending into the fluid bed). These high thrust feed nozzles are typically arranged non-symmetrically around the reactor to optimize flow patterns in the reactor according to simulation studies although symmetrical disposition of the nozzles is not precluded if the flow patterns in the reactor can be optimized in this way. There are typically 4-6 feed rings located at different elevations although not all may be active at any one time while the unit is working.

Steam is admitted as fluidizing gas in the stripping section 13 at the base of coker reactor 10, through spargers 14 directly under stripping sheds 15 as well as from lower inlets 16. The steam passes up into stripping zone 13 of the coking reactor in an amount sufficient to obtain a superficial fluidizing velocity in the coking zone, typically in the range of about 0.15 to 1.5 m/sec (about 0.5 to 5 ft/sec). The coking zone is typically maintained at temperatures in the range of 450 to 650.degree. C. (about 840 to 1200.degree. F.) and a pressure in the range of about 0 to 1000 kPag (about 0 to 145 psig), preferably about 30 to 300 kPag (about 5 to 45 psig), resulting in the characteristic conversion products which include a vapor fraction and coke which is deposited on the surface of the seed coke particles.

The vaporous products of the cracking reactions with entrained coke particles pass upwards out of the reaction zone 11, through a phase transition zone in the upper portion 17 of the vessel and finally, a dilute phase reaction zone at the inlets of cyclones 20 (only two shown, one indicated). The coke particles separated from the vaporous coking products in the cyclones are returned to the fluidized bed of coke particles through cyclone dipleg(s) 21 while the vapors pass out through the gas outlet(s) 22 of the cyclones into the scrubbing section of the reactor (not shown). After passing through scrubbing section which is fitted with scrubbing sheds in which the ascending vapors are directly contacted with a flow of fresh feed to condense higher boiling hydrocarbons in the reactor effluent (typically 525° C.+/975° F.+) and recycles these along with the fresh feed to the reactor. The vapors leaving the scrubber then pass to a product fractionator (not shown). In the product fractionator, the conversion products are fractionated into light streams such as naphtha, intermediate boiling streams such as light gas oils and heavy streams including product bottoms.

The coke particles that pass downwards from the dense bed 11 to stripper section 13 comprising sheds 15 are partially stripped of occluded hydrocarbons in the stripper by use of a stripping gas, usually steam, which enters via spargers 14. The stripped coke particles are passed via line 25 to a heater (not shown) which is operated a temperature from about 40 to 200° C., preferably about 65 to 175° C., and more preferably about 65 to 120° C. in excess of the actual operating temperature of the coking zone and recycled back to the fluid coking unit.

Example 1

Current commercial fluid coking feed nozzles are designed to atomize the bitumen at the nozzle exit through shear from the high velocity and rapid decompression of the atomization steam upon exiting the nozzle. This decompression happens both axially and radially.

One such coker nozzle is described in detail in Canadian Patent No. 2,224,615, and is referred to herein as TEBM-2b with circular exit, or GEN2 nozzle. The GEN2 nozzle consists of a series of converging, diverging, and converging sections. The pressure drop across the exit of the GEN2 coker feed nozzle is on the order of 70 psi. The flow exiting the nozzle consists of bubbles dispersed in the liquid phase and the large decompression from the resultant pressure drop at the exit causes an explosive expansion of the bubbles, resulting in a phase inversion where the flow changes from liquid continuous in the nozzle to gas continuous in the jet, with liquid droplets and ligaments distributed in the gas stream.

FIG. 2 shows the local pressure profiles along the length of a GEN2 feed nozzle when air and water are sprayed into ambient air, at air to liquid ratios (ALRs) of 2.1 and 1.6 wt %. The increase in pressure along the diverging section of the nozzle is consistent with subsonic flow. FIG. 2 shows a significant pressure drop at the exit of the nozzle, which causes the gas to expand rapidly in both the axial and radial direction.

In this example, the GEN2 nozzle and the 1.25GEN2, which is the same as the GEN2 nozzle except all of the dimensions are scaled up so that the throat area is 25% larger than the GEN2 nozzle, were tested in order to measured axial thrust force (lb) as a function of the nozzle pressure. In addition to the GEN2 nozzles, three commercially available fan spray nozzles, referred to herein as Nozzle B, Nozzle C and Nozzle D, and a curved throat fan nozzle used in the FCC process, described in detail in U.S. Pat. No. 6,199,768, referred to herein as CTF, were tested in this example. A GEN3 nozzle, which consists of the same internal geometry as the GEN2 nozzle but contains a diverging cloverleaf disperser at the tip of the nozzle, was also tested. A drawing of a GEN3 nozzle is shown in FIG. 3 and described in more detail in U.S. Pat. No. 8,999,246. 1.1GEN3 and 1.25GEN3 nozzles are the same as the GEN3 nozzle, but all of the dimensions are scaled up so that the throat area is 10% and 25% larger than the GEN3 nozzle, respectively.

Finally, FIG. 4 shows a drawing of a GEN4 nozzle, which consists of the same internal geometry as the GEN2 nozzle but with slits at the nozzle tip. The GEN 4 nozzle is described in more detail in U.S. Pat. No. 9,889,420. The 1.1GEN4 is the same as the GEN4 nozzle, but all of the dimensions are scaled up so that the throat area is 10% larger than the GEN4 nozzle.

Experiments were conducted with the aforementioned feed nozzles having different equivalent throat diameters and exit angles by spraying water into open air over a range of liquid flow rates and nozzle pressures. Table 1 shows a summary of the nozzles that were tested and their specifications. The nozzles were mounted on a stand that allowed them to move freely in the axial direction. The reaction thrust force was measured using a 3000 lb thru-hole compression load cell, which was mounted on the nozzle conduit and was compressed between two plates while the nozzle was spraying.

TABLE 1
Summary of Nozzle Specifications Tested with Water
		Throat
		Diameter	Spray Angle
Nozzle	Description	(inches)	(°)
Nozzle B	Fan spray nozzle	0.344	50
Nozzle C	Fan spray nozzle	0.297	65
Nozzle D	Fan spray nozzle	0.297	120
GEN2	TEBm-2b* with circular exit	0.512	11
GEN3	TEBm-2b* with cloverleaf	0.512	23
	disperser
1.1GEN3	TEBm-2b* with cloverleaf	0.537	23
	disperser
1.25GEN3	TEBm-2b* with cloverleaf	0.572	23
	disperser
CTF	Curved throat fan nozzle	0.27	40
1.1GEN4	TEBm-2b* with four slits in the	0.557	50
	exit
*Base et al. (1999)

FIG. 5 shows a plot of the measured axial thrust force (lb) as a function of the nozzle pressure for the various nozzle geometries. The axial thrust force was highly correlated with the nozzle pressure. FIG. 5 shows that nozzles that provide the same nozzle pressure, with the same exit area, operating at the same liquid flow rate, result in different axial thrust measurements due to the difference in nozzle geometry. For example, Nozzle C and Nozzle D both have an orifice diameter of 0.297″, however, Nozzle C produces a spray with an angle of 65°, whereas Nozzle D produces a spray with an angle of 120° and when operating at a nozzle pressure of 400 psig. The measured axial thrust force of Nozzle C was approximately 10 lb greater than the measured axial thrust force of Nozzle D. The larger spray angle resulted in a smaller axial thrust force as more of the thrust force was directed in the radial direction. Nozzles B-D are BETE nozzles designed to be operated with liquid only and, therefore, produced higher thrust forces than the GEN2, GEN3, GEN4 and CTF nozzles, which are designed to be operated with both gas and liquid. However, even when operating with only liquid, noticeable differences in the thrust force were observed with changes in nozzle geometry. For example, the GEN2 and GEN3 nozzles have the same exit orifice diameter, and when operating at a fluid pressure of 157 psig the GEN3 nozzle, which has a diverging nozzle exit, resulted in an axial thrust force that was approximately 7 lbs greater than the axial thrust force of the GEN2 nozzle that has a converging nozzle exit.

In summary, the results in FIG. 5 show that the thrust force is highly correlated with the fluid pressure upstream of the nozzle. In addition, nozzle spray angles and exit geometries can be optimized to produce higher thrust forces.

Example 2

In order to maximize the axial thrust force and reduce the expansion of the jet in the radial direction, a supersonic nozzle with a diverging/diffuser section was designed to accelerate the fluid axially in the nozzle exit, prior to injection into the fluidized bed (hereinafter referred to as the “Diffuser nozzle”). The Diffuser nozzle consists of the same internal geometry as the GEN2 nozzle but without the final constriction at the nozzle tip. The diffuser section now resulted in a much narrower jet plume. The phase inversion occurs within the nozzle, and the fluid acceleration through the nozzle will increase in the axial direction. In addition, a supersonic nozzle maximizes the velocity of the jet at a much larger cross sectional exit area compared to a subsonic nozzle.

FIG. 6 shows the local pressure profile along the length of a Diffuser nozzle when spraying air and water into ambient air at ALRs of 2.6 and 1.7 wt %. The pressure decrease along the diverging section of the nozzle indicates that the flow is supersonic. It can be seen that the fluid exits the nozzle at the same pressure as the ambient air, which reduces thrust loss and results in minimum radial expansion.

FIGS. 7A and 7B are photographs showing the differences in the jet plume for air-water mixtures exiting the GEN2 nozzle (FIG. 7A) and the Diffuser nozzle (FIG. 7B). The images clearly show that the spray plume is much narrower with the Diffuser nozzle, since all of the driving pressure has been used to accelerate the flow in the axial direction.

Experiments were conducted using feed nozzles with different equivalent throat diameters and exit geometries by spraying air and water into open air over a range of liquid flow rates and nozzle pressures. Table 2 shows a summary of the nozzles that were tested and their specifications. The nozzles were mounted on a stand that allowed them to move freely in the axial direction. The reaction thrust force was measured using a 3000 lb thru-hole compression load cell, which was mounted on the nozzle conduit and was compressed between two plates while the nozzle was spraying.

TABLE 2
Summary of Nozzle Specifications for Nozzles Tested with Air and
Water
		Throat
		Diameter
Nozzle	Description	(inches)	Exit Geometry
GEN1	Simple constriction with	0.512	Converging with
	long circular exit		circular exit
GEN2	TEBm-2b with circular exit	0.512	Converging with
			circular exit
GEN3	TEBm-2b with cloverleaf	0.512	Diverging with
	disperser		cloverleaf shaped
			exit
1.25GEN2	TEBm-2b with circular exit	0.572	Converging with
			circular exit
1.25GEN3	TEBm-2b with cloverleaf	0.572	Diverging with
	disperser		cloverleaf shaped
			exit
Diffuser	Constriction followed by	0.516	Diverging with
	long diffuser		circular exit
1.2Diffuser	Constriction followed by a	0.562	Diverging with
	long diffuser		circular exit
1.5Diffuser	Constriction followed by a	0.628	Diverging with
	long diffuser		circular exit

A drawing of a GEN1 nozzle is shown in FIG. 8 and consists of a simple constriction of 7° followed by a 3″ long straight section. The 1.2Diffuser nozzle and 1.5Diffuser nozzle are the same as the Diffuser nozzle shown in FIG. 6, but all of the dimensions are scaled up so that the throat area is 20% and 50% larger than the Diffuser nozzle, respectively.

FIG. 9 shows a plot of the measured axial thrust force (lb) as a function of the measured fluid pressure. The axial thrust force was highly correlated with the fluid pressure. FIG. 9 shows that nozzles that provide the same fluid pressure, with the same throat diameter result in different axial thrust measurements due to the differences in nozzle geometry. For example, the GEN1, GEN2 and GEN3 all have the same throat diameter, however, at a nozzle pressure of 275 psig, the measured axial thrust force of the GEN3 nozzle was approximately 11 lb greater than the GEN2 nozzle and 14 lb greater than the GEN1 nozzle. The diverging cloverleaf disperser located on the tip of the GEN3 nozzle reduces the radial expansion of the fluid and results in a higher axial thrust force. The nozzles that provide the largest axial thrust force are the 1.5Diffuser and the 1.2Diffuser. These nozzles have a larger throat and have a long diffuser section located at the end of the nozzle, which results in a decrease in pressure along the nozzle outlet, which means that supersonic velocities are achieved. The axial thrust through these nozzles has been maximized since the pressure at the nozzle exit is equal to the ambient pressure and therefore, all of the driving pressure has been used to accelerate the flow in the axial direction.

Example 3

Another nozzle geometry that would maximize the axial thrust force would be to add a diverging section to the GEN2 nozzle in order to accelerate the fluid to supersonic velocities before exiting the nozzle. FIG. 10 shows the local axial pressure profile along the length of a GEN2 nozzle with an additional diverging section at the exit when spraying air and water into ambient air at an ALR of 2.2 wt %. The pressure along the outlet diffusing section decreases, indicating that supersonic velocities were achieved. The axial thrust through this nozzle has been maximized since the pressure at the nozzle exit is equal to the ambient pressure and therefore all of the driving pressure has been used to accelerate the flow in the axial direction.

From the foregoing description, one skilled in the art can easily ascertain the essential characteristics of this invention, and without departing from the spirit and scope thereof, can make various changes and modifications of the invention to adapt it to various usages and conditions. Thus, the present invention is not intended to be limited to the embodiments shown herein, but is to be accorded the full scope consistent with the claims, wherein reference to an element in the singular, such as by use of the article “a” or “an” is not intended to mean “one and only one” unless specifically so stated, but rather “one or more”. All structural and functional equivalents to the elements of the various embodiments described throughout the disclosure that are known or later come to be known to those of ordinary skill in the art are intended to be encompassed by the elements of the claims. Moreover, nothing disclosed herein is intended to be dedicated to the public regardless of whether such disclosure is explicitly recited in the claims.

Claims

1. A process for converting a heavy hydrocarbonaceous feedstock to liquid products, comprising: the feedstock being introduced into the fluid coker using a plurality of high thrust nozzles, said nozzles designed to transport unreacted hydrocarbon and coke from the low velocity annular region to the high velocity core region to improve hydrocarbon stripping, reduce gas phase residence time, and increase liquid products yields.

introducing the hydrocarbonaceous feedstock into a fluid coker comprised in part of a fluidized bed of heated coke particles, the fluidized bed having a high velocity core region of heated coke particles and a low velocity annular region of unreacted hydrocarbon and coke particles; and

reacting the hydrocarbonaceous feedstock with the heated coke particles in the fluid coker to produce the liquid products;

2. The process as claimed in claim 1, whereby the high thrust nozzles have a spray angle of about 3°-160° and a nozzle diameter between about 0.2″ and 0.8″.

3. The process as claimed in claim 1, wherein the high thrust nozzle comprises a diverging section at the tip of the nozzle.

4. The process as claimed in claim 1, wherein the high thrust nozzle comprises a converging section followed by a diverging section at the tip of the nozzle.