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Adiabatic Laboratory Reactor Design and Verification Arno de Klerk† Fischer-Tropsch Refinery Catalysis, Sasol Technology Research and Development, P.O. Box 1, Sasolburg 1947, South Africa
An adiabatic laboratory fixed bed reactor design is presented. Such a reactor finds application in catalyst testing when final product quality must be evaluated and the reacting system has a high heat release. The design was modeled and verified experimentally with a nonreacting system. The model was used to suggest design improvements and analyze radial and axial heat loss contributions. It was found that the axial heat flux was 2 orders of magnitude more than the radial heat flux, but since the area for radial heat exchange was much larger, both contributed equally to the heat loss of the design. The reactor had an adiabatic zone of 0.8 m and 1 kW‚m-3 heat loss. This is equivalent to a 4 °C temperature loss for a hydrocarbon feed at a space velocity of 1 h-1. The design was also tested with a reacting system, and the adiabatic temperature rise during coal tar naphtha hydrogenation was within 2 °C of a comparable commercial unit. The adiabatic laboratory reactor therefore proved itself capable of mimicking commercial operation. Introduction Due to the high cost of new refineries, the refining industry regularly deals with the evaluation of new catalysts and processing conditions to improve the performance of existing units. The ability to scale-down refining processes for laboratory evaluation and to relate laboratory results to commercial operation are therefore of importance. Multiphase reactors are used for most refinery conversion processes, with trickle bed reactors playing a dominant role in hydroprocessing. There is consequently a large body of literature dealing with such reactors.1-3 When the process and feed characteristics are wellknown, new catalysts and operating conditions can be evaluated on a small scale, sometimes with only a single catalyst pellet. When significant changes in the nature of the catalyst, feed, or operating conditions are made, it is advisable to test real feedstock under industrial conditions, which may include adiabatic operation.4 In the refining industry it is especially important to evaluate catalysts in terms of final product quality. If the process has a high heat release, adiabatic testing becomes necessary, since commercial operation is usually adiabatic and temperature influences both catalyst activity and selectivity. In a questionnaire to probe industrial research on heterogeneously catalyzed processes,5 it was found that most companies made use of continuous flow fixed bed reactors, albeit not as only reactor type. Although catalyst bed isothermicity was not always ensured, no deliberate attempts were made to test adiabatically. This can be corroborated by the author, who conducted a pilot plant benching exercise in Europe and North America that involved 10 industrial facilities and observed adiabatic testing only at one of the facilities. In treaties on laboratory-scale reactors,6-11 the use and design of adiabatic reactors is seldom discussed in detail. The aim of this work was to design a fixed bed adiabatic laboratory reactor for refinery catalyst and † Tel.: +27 16 960-2549. Fax: +27 11 522-3517. E-mail:
[email protected].
process evaluations. The design, modeling, and verification of both the model and design will be discussed. Design For a laboratory reactor to be adiabatic there must be no heat transfer between the fluid and the equipment. In practice, there are two design requirements to be met for ideal steady state adiabaticity, namely, no radial heat loss and no axial heat exchange.12 The no radial heat loss criterion can be achieved by insulation (increased resistance to heat transfer) and temperature compensation (reduced driving force for heat transfer). The principle of insulating the reactor and then heating the insulation to the same temperature as the reactor13 remains an excellent way to minimize radial heat loss. It is not only practiced in adiabatic laboratory reactor design but also finds application in distillation studies. The no axial heat exchange criterion effectively requires that axial conduction must be small in relation to fluid heat transport. This is linked to the reactor rating (maximum temperature and pressure of the design) and hydrodynamic considerations for scaledown,8,14-15 like minimum Peclet-number, bed length required, and minimum column to particle diameter ratio. Lastly, there were practical and ergonomic design considerations. It was important that the heater response to heating and cooling had to be very similar. A problem that is often encountered in laboratory reactors is that they are so well insulated that heating is rapid, but cooling is very slow. This not only has implications for controller tuning, but also for the reaction. Attaining steady-state operation can be difficult if the reactor does not follow the change in adiabatic temperature closely. This may also have safety implications, since it becomes difficult recovering from a temperature runaway. It was therefore decided to design an adiabatic reactor that superficially appeared to have some heat loss. Instead of using a stationary insulating layer, the reactor was placed in a reactor sleeve that allowed a free flow of air by natural convection. This caused the bottom of the reactor to have some radial heat loss,
10.1021/ie050212a CCC: $30.25 © 2005 American Chemical Society Published on Web 05/20/2005
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Figure 2. Reactor module.
Figure 1. Details of reactor and reactor sleeve.
thereby reducing the length of the adiabatic zone, but it resulted in a good cooling response. It had the added advantage of making reactor change out very easy. Details of the design are given in the Experimental Section. Experimental Section Equipment. The adiabatic laboratory reactor design is shown in Figure 1. The reactor consists of a 11/2 in. nominal diameter, schedule 160 stainless steel tube (48.3-mm o.d. and 34-mm i.d.), which is 1.6 m in length. Both ends of the tube are fitted with threaded boss heads, which are connected via standard tube fittings to the reactor system. The reactor is fitted with a 6-mm centered thermowell running the length of the reactor. Four 3-mm copper tubes, which also serve as thermowells, are clamped on the outside of the reactor tube. The reactor slides into a reactor sleeve, manufactured from a 21/2-in. nominal diameter, schedule 40 pipe (73.0mm o.d., 62.7-mm i.d.), which is 1.6 m in length. Four 6-mm stainless steel thermowells have been imbedded in the reactor sleeve. The reactor sleeve is attached to a framework by wings welded onto the sleeve at the top and bottom. Two semicircular metal plates welded to the bottom of the reactor sleeve support the reactor and it is the only place where the reactor and the reactor sleeve are in contact. The only other contact between the reactor and the framework is by the 6-mm stainless steel tubing connecting the reactor inlet and outlet to the feed and product modules of the laboratory reactor system. Four heaters, each 375 mm in length, are clamped on the outside of the reactor sleeve. Power to each heater was controlled separately. The thermocouples measuring the temperature of the heaters were positioned opposite the heaters in the thermowells of the reactor sleeve at positions 200, 600, 1000, and 1400 mm from the bottom. Thermocouples of similar length were placed in the copper thermowells on the outside of the reactor and aligned with those of the heaters. Six sliding thermocouples, 10 mm apart, were placed in the central reactor thermowell.
The inlet of the reactor module was connected a feed module and a preheater was installed in the inlet line to the reactor. The preheater had its own heater with independent temperature control. The outlet of the reactor module was connected to a product workup module.16 A schematic representation of the reactor module is given in Figure 2. Procedure. The reactor was loaded with 3-mm glass spheres (no catalyst was loaded). The preheater was operated at an outlet temperature of 250 °C and 3 MPa pressure. All four reactor heaters were operated in adiabatic mode, which requires the heater temperature to follow the temperature of the reactor wall. A C9-C11 Fischer-Tropsch-derived paraffin mixture and hydrogen were fed to the reactor at 460 g‚h-1 and 0.3 m3‚h-1 (normal), respectively. The reactor system was allowed to reach steady state. The thermocouples in the central reactor thermowell were used to record the temperature profile inside the reactor. Their position and temperature were noted and adjusted by 50 mm every 8-10 min. This ensured that one thermocouple duplicated the position between adjustments, which served as a cross-check. The thermocouples were not calibrated, but constant temperature measurement indicated that they varied less than 1 °C from each other. Repeat measurements of the temperature profile were made on different days and with different reactors of identical design. Modeling. Although laboratory reactor design can be done on a trial and error basis, such an approach tends to be costly. Modeling the design can not only help with the initial design but can also be used to test design tradeoffs and optimize the design for best approach to adiabaticity. The temperature profile in the reactor was modeled in terms of Fourier’s general equation, with additional terms for fluid flow and heat generation (eq 1).
CpF
WCp ∂T ∂T ) k32T +Q)0 ∂t A ∂z
(1)
For a nonreacting system at steady state, the time dependence and heat generation terms become zero. The equation was written in cylindrical coordinates, and small deviations from radial symmetry in the actual reactor were disregarded (eq 2). The temperature profile
k
(
)
WCp ∂T ∂2T 1 ∂T ∂2T + + 2 )0 2 r ∂r A ∂z ∂r ∂z
(2)
was solved numerically by rewriting eq 2 as a finite difference equation. The model took into account that the thermal conductivity in different zones of the reactor was different and that flow was restricted to specific areas. The axial nodes were equally spaced, but the
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(e) Nodes in the reactor inlet line were given a fixed temperature corresponding to the preheater outlet temperature for which the system was to be modeled. (f) Nodes in the reactor outlet line were required to have the same temperature as just inside the reactor (∂T/∂z ) 0). The air flow rate in the reactor sleeve was not measured. For modeling purposes the buoyancy of the air due to heating was used. The pressure differential was used to calculate the air flow rate that would satisfy the force balance.
Figure 3. Position of nodes for modeling.
radial nodes were not equally spaced (Figure 3). The Supporting Information gives more detail on the node positions and thermal conductivity of the zones. The model tried to represent the actual design. The following boundary conditions were used in the model: (a) Nodes on the centerline of the reactor were required to have a temperature profile with no discontinuity, which required the first derivative of temperature with respect to radial position to be zero (∂T/∂r ) 0). (b) External nodes that were in contact with the air surrounding the equipment were given a fixed temperature value corresponding to ambient conditions. (c) External nodes that corresponded to a heater element were required to have the same temperature as the outside of the reactor sleeve (∂T/∂r ) 0). One point per heater was given a fixed temperature corresponding to the temperature on the outside reactor wall in the center of the heating zone. This mimicked the actual adiabatic reactor design where the heater temperature was controlled at the value measured on the outside of the reactor wall. It also implied that a heater was not seen as a constant temperature entity but allowed for small changes in axial temperature (less than 1 °C). (d) External nodes in the wings of the reactor sleeve that are used to attach it to the framework were required to have a constant rate of temperature loss (∂2T/∂r2 ) 0). This was an important boundary condition, since the wings were the only place where the reactor module was attached to the structure, apart from the inlet and outlet of the reactor itself.
Results The temperature profiles measured on the adiabatic laboratory reactors showed some variation. The sample standard deviation of each axial temperature was calculated and the maximum sample standard deviation was 11 °C, close to the bottom of the reactor. The average sample standard deviation was 3 °C. The average temperature profile measured in the laboratory adiabatic reactor was correlated to the model results for the system. Due to differences in the ambient conditions on the dates that data were collected, the temperature data was adjusted to have the same value at an axial position 0.8 m from the top of the reactor (Figure 4). The profile was not changed, because a constant value was added or subtracted. It was found that the end-effects at the extremities of the reactor were not modeled well, but the model gave a good description of the section of interest in the reactor. The model description of the axial section from 0.2 to 1.4 m had an average absolute error of 1 °C and standard deviation of 1 °C. The errors were not systematic, showing no model bias. It can therefore be said that the model was successfully validated and that it could be used to probe deviations from adiabatic behavior in the real system. It has already been noted both radial and axial heat flows can result in nonadiabatic behavior. By changing the properties of the air gap to that of a perfect thermal insulator, the temperature profile for a reactor with no radial heat loss was obtained (Figure 4). In this way the model could separate the contributions of axial and radial heat loss. An ideal adiabatic reactor at steady state operation with a nonreacting system will be
Figure 4. Calculated and experimental temperature profiles for a nonreacting system, with axial distance reflecting the distance along the reactor tube.
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absolutely isothermal. The difference between the profile with no radial heat loss and isothermal behavior gave the axial heat loss contribution. The axial heat loss of the design was about 2 kW‚m-2. Similarly, the difference between the profile of the real system and that with no radial heat loss gave the radial contribution. The radial heat loss of the design was about 10 W‚m-2. However, despite the apparent large difference between axial and radial heat loss when expressed in terms of heat flux, the actual heat loss contribution was about equal. If only the center part of the reactor system (0.4-1.2 m) was considered, without all the end-effects increasing heat loss, the heat loss of the design was an order of magnitude less, the axial heat loss being 0.3 kW‚m-2 and radial heat loss being 3 W‚m-2. Discussion Axial Heat Loss. The results clearly showed that the axial heat flux is much higher than the radial heat flux. The design of an adiabatic reactor is consequently sensitive to the aspect ratio. Not much can be done to reduce the reactor diameter beyond the requirements dictated by hydrodynamics and mechanical integrity. Yet, some aspects of the design can be manipulated, like the material of construction, catalyst dilution material, and dimensions of the thermowell. By understanding the relative contribution of each, it is possible to optimize the design with respect to axial heat loss. There is a reasonable body of literature dealing with thermal conductivity in packed beds.17-21 On the basis of the values reported by Sharma and Hughes,20 it was not expected that the catalyst bed would contribute significantly to axial heat transfer, since its thermal conductivity is 2 orders of magnitude less than that of the reactor. Using the model, the properties of the reactor and thermowell were changed to perfect insulators. Calculations confirmed that the catalyst bed contributed little to axial heat transfer. Nevertheless, dilution of the catalyst bed with material that has a high thermal conductivity, like silicon carbide, should rather be avoided. Inert material with low thermal conductivity, like silica, is preferable for catalyst dilution during adiabatic testing. Axial heat transfer was mainly by the reactor wall, with the thermowell contributing to a lesser extent. Stainless steel is a poor conductor of heat, which helps reduce axial heat flow. A slight improvement can be obtained by using 310 stainless steel, since it has the lowest thermal conductivity of the commonly used grades of stainless steel. Radial Heat Loss. It was found in practice that insulation materials, such as glass wool, were not much better than using air but had the disadvantage that it made reactor cooling more sluggish. The cold air entering the air gap caused some cooling at the bottom of the reactor. The model nevertheless showed that once the air reached the correct temperature, radial heat transfer was limited. A slight bias could be introduced in the temperature compensation to reduce radial heat flow, but this was not done. The principle used to limit radial heat flow was not only based on insulation but also on reducing the driving force for heat flow, namely, the temperature difference. The heaters were controlled separately to maintain a zero temperature difference between the temperature at the outside reactor wall and the reactor sleeve. The
temperature measurements were point measurements, and this implied that the number of points at which the temperature difference could be controlled was equal to the number of heaters. It was therefore not surprising that the model showed the radial heat loss to be sensitive to the number of heaters used and that the law of diminishing returns applied. By improving the thermal conductivity of the reactor sleeve, temperature effects on the reactor sleeve could be further diminished, thereby reducing the overall temperature difference between the sleeve and the reactor. Modeling Limitations. During the verification of the model, it was noted that the model described the temperature profile at the extremities of the reactor poorly. Improvements in the design that are suggested by reducing such end-effects, could be modeling artifacts. This is an important limitation of the model. Another possible limitation of the model is related to the heat transfer coefficients at the interfaces. The model assumes that the thermal conductivity of the different media is the main resistance to heat flow and that the transition from one material to another material imposes no additional resistance. The model also tacitly assumes that the dynamics of the flow components (flow inside the reactor and in the air gap) are so fast that conductive heat transfer determines the steadystate temperature profile. If this is not true, the radial heat loss component found in practice, which is seemingly well-described by the model, may actually be poorly modeled. For example, when the resistance to heat transfer in the air gap is dominated by the heat transfer coefficient, the sleeve and heater configuration may be less important than suggested by the model. This could seriously detract from the usefulness of the model, and the assumption was checked experimentally. A heat sink, consisting of a metal rod that connected the sleeve to the framework supporting the reactor, was welded to the middle of the reactor sleeve. The reactor design and operation were not altered otherwise. It was found that this affected the temperature profile in the reactor (Figure 5), as would be expected from the model description. The heat loss by conductive heat transfer split the adiabatic zone of the reactor into two smaller adiabatic zones. The assumption that thermal conductivity is the main resistance to heat transfer is therefore a reasonable one. Lastly, it should be noted that the model was derived without taking emissive heat loss into account. The model is consequently not valid at very high temperatures where heat transfer by radiation becomes significant. Verification of Adiabaticity in Practice. The ultimate aim of an adiabatic laboratory reactor design is to provide a tool for catalyst evaluation that mimics industrial adiabatic reactor operation well. The heat loss characteristics of the laboratory reactor were translated into numbers that would be meaningful for catalyst evaluation. The 0.8-m reactor zone that was considered adiabatic started 0.4 m from the top of the reactor. The overall heat loss was about 1 kW‚m-3, which was equivalent to a 4 °C drop in temperature when a hydrocarbon was processed at a space velocity of 1 h-1. The performance of the laboratory reactor was compared to that of a commercial coal tar naphtha (CTN) hydrogenation unit in the Sasol Synfuels refinery at Secunda, South Africa. The CTN hydrogenation unit consists of two reactors. The first reactor contains a
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Figure 5. Experimental temperature profile with heat sink in the middle of the reactor sleeve.
single catalyst bed and the second reactor contains two catalyst beds with an inter bed quench. This was mimicked in the laboratory by using three adiabatic laboratory reactors, which also allowed inclusion of interbed hydrogen addition as quench. The laboratory reactors were loaded with the same catalysts and in the same proportions as were loaded commercially (combination of commercial Ni/Mo, Co/Mo, and Ni/Co/Mo catalysts). The feed was obtained from the CTN hydrogenation unit and contained 15% olefins, 25% paraffins, and 60% aromatics, which included 4% oxygen as phenolics, 0.5% nitrogen, and 0.5% sulfur. The laboratory reactors were operated at similar inlet conditions as the commercial unit. The overall adiabatic temperature rise of the laboratory reactors was 170 °C, while that of the commercial unit was 172 °C. The adiabatic laboratory reactor design therefore proved itself capable of mimicking commercial operation. (Details of this test work can be found in the Supporting Information.) Conclusion The two criteria for adiabaticity, namely no radial heat loss and no axial heat exchange, were both found to be important in the design of a laboratory reactor. Although the axial heat flux was 2 orders of magnitude more than the radial heat flux, the area available for radial heat exchange was much larger, causing both to contribute equally to heat loss. By modeling the reactor design and verifying the model experimentally, the model could be used to suggest improvements to the design. The reactor design had an adiabatic zone of about 0.8 m, with an overall calculated heat loss of about 1 kW‚m-3. This is equivalent to a 4 °C temperature loss when a hydrocarbon feed is passed through the reactor at a space velocity of 1 h-1. The adiabaticity of the laboratory reactor was also tested with a reacting system. The operation of a commercial coal tar naphtha hydrogenation unit was mimicked, and it was found that the laboratory reactor was capable of reproducing the adiabatic temperature rise within the expected margin of error.
Acknowledgment All work was done at Sasol Technology Research and Development, and permission to publish this work is appreciated. Supporting Information Available: Test work done to verify adiabaticity in practice is discussed in more detail, providing information on the temperature profiles (Figures S1 and S3), laboratory reactor system (Figure S2), and feed conversion (Table S1). Additional information on the modeling is also given with respect to node positions, materials, and thermal conductivity values (Tables S2 and S3). This material is available free of charge via the Internet at http://pubs.acs.org. Nomenclature A ) cross-sectional area for flow, m2 Cp ) heat capacity, J‚kg-1‚K-1 k ) thermal conductivity, J‚s-1‚m-1‚K-1 Q ) heat generation, J‚s-1 r ) radial distance from center, m T ) temperature, K t ) time, s W ) mass flow rate, kg‚s-1 z ) axial distance, m F ) density, kg‚m-3
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Received for review February 21, 2005 Revised manuscript received April 11, 2005 Accepted April 13, 2005 IE050212A
