Modern Techniques for Optimization of Primary Reformer Operation. P.W. Farnell

Transcription

1 Modern Techniques for Optimization of Primary Reformer Operation P.W. Farnell

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3 1 S UMMARY This paper reviews the techniques available for monitoring the performance of the primary steam reformer within an ammonia, hydrogen or methanol production plant. The techniques used include tube wall temperature measurement, plant heat and mass balance reconciliation and reformer simulation. The combination of these techniques allows an accurate assessment of catalyst performance, reformer operation, tube life estimation and identification of operating limits. The results from these assessments can also be used to identify suitable reformer optimization and uprate strategies. Several case studies are presented which illustrate the combined use of these techniques in solving common issues encountered in the operation of primary steam reformers. 2 I NTRODUCTION The primary steam reformer can rightly be described as being at the heart of an ammonia, hydrogen or methanol plant, converting the hydrocarbon feed into the synthesis gas mixture. In terms of the chemical and physical processes it is the most complex operation within these plants. Furthermore, the primary steam reformer is the most expensive single item of capital expenditure, and is the dominant energy consumer in the plant. The steam reformer in the simplest terms is a heat exchanger, transferring heat from the hot fluegas to the cooler process gas within the tubes. However, upon closer examination primary reforming involves a combination of many differing processes. On the outside of the tube, transfer of heat by radiation occurs simultaneously with chemical reaction in the form of combustion. On the inside of the tube, transfer of heat and mass coupled with chemical reaction occurs, which requires additional involvement of kinetics and thermodynamics. Based on the severe conditions involved in the operation of a primary reformer it is not surprising that problems occur, or that reformers are sometimes operated at less than optimum conditions. Substantial quantities of data can be collected during operation and these data are invaluable in trending the performance of the furnace and of the catalyst and providing operating guidelines. However, these are not in themselves sufficient to diagnose either inefficient operation or the causes of poor reformer performance. Neither are they sufficient to make an assessment of optimum operating conditions for the reformer. In this case a full survey of the reformer must be carried out which involves the collection of all process data including the convection section performance, a full tube wall temperature survey and the analysis of these data by heat and mass balance reconciliation and full thermal modelling of the primary reformer itself. It is imperative that all of the above data are considered in a reformer survey aimed at investigation of either reformer performance or optimization. Only by using the combination of the techniques above can a full picture of the state of health of the reformer be built up. From this any underlying problem areas can be highlighted and remedial strategies developed. The analysis can also be used to assess if further optimization of the reformer operation is possible. Johnson Matthey Group 360W/109/3/REF 1

4 UBE WALL TEMPERATURE MEASUREMENT 3 T The tubes in a steam reformer operate close to the limits of the chosen materials in terms of stress induced as a result of high temperatures combined with large differential pressures across the tube wall. The maximum tube temperature in a primary reformer varies from 850 to 1000 C (1560 to 1830 F), which results in the tubes undergoing irreversible creep and experiencing a limited life before failure. Operation at tube wall temperatures above design can result in a rapid increase in the number of tube failures. A general rule of thumb is that an increase in tube wall temperature of 20 C (36 F) will halve the tube life. It is therefore important to measure tube wall temperatures accurately in order to prevent premature tube failure by overheating but also to avoid operating a furnace with tube temperatures so conservative that the full capacity of the furnace is not realized. The most commonly used device for tube temperature measurement is an infrared pyrometer. However, the infrared pyrometer is subject to one major source of error when used in a furnace. Infrared pyrometers cannot distinguish between radiation emitted by the target tube itself and radiation reflected by the tube but originating from the walls of the furnace, as shown in Figure 1. As a result, infrared pyrometers tend to read high by typically C (36-72 F) if left uncorrected. Figure 1: Background Radiation When measuring the temperatures of tubes with an infrared pyrometer an emissivity setting of 1.00 should be used even though the tube emissivity is below this value. If a tube has an emissivity of 0.80, then at least 20% of the radiation received by the pyrometer will be radiation reflected by the tube from the hotter walls. Therefore, the pyrometer in effect receives more than 1.25 times as much radiation as is emitted by the tube itself. The use of an emissivity of 1.00 compensates for this to a certain degree. A correction can then be made to the temperature measured with a pyrometer (with emissivity set to 1.00) as set out below. e.tt 4 = Tm 4 - (1 - e).tw 4 2

5 where : Tt - true temperature (K / R) Tm - measured temperature with pyrometer emissivity = 1.00 (K / R) Tw - average background temperature (K / R) e - emissivity of tube Synetix occasionally uses a gold cup pyrometer, which is not prone to the same errors as an infrared pyrometer. A gold cup pyrometer consists of a radiation detector located in a gold plated hemisphere mounted on a water cooled probe. It is the most accurate method of temperature measurement but is cumbersome to use, as the probe has to be placed on the tube wall and its use is limited by the length of the probe, general access and available portholes. The main use of the gold cup pyrometer in Synetix is as a confirmation of the infrared pyrometer results. The subject of tube wall temperature measurement in primary steam reformers is covered in substantial depth in Reference 1. RIMARY REFORMER SURVEY 4 P The primary reformer survey is at its most basic level, a data collection exercise. The major portion of the plant survey is the collection of a full set of tube wall temperature data. The survey is carried out using either a gold cup pyrometer or an infrared pyrometer, or in certain circumstances a combination of both. If the survey utilizes an infrared pyrometer then the background temperatures are also obtained for each measurement location, and the corrected temperature is obtained as described above. At the same time as the tube wall survey the four major gas streams listed below (as a minimum) are sampled and analysed for composition. Hydrocarbon feed Reformer exit gas Fuel (mixed or separate constituents) Fluegas (exit radiant section and at stack) The operating process data for the reformer are collected concurrently with the tube wall temperature survey with as much additional process information as possible. This includes but is not limited to: Hydrocarbon feed flowrate Recycle hydrogen flowrate Reaction steam flowrate Reformer inlet temperature Reformer inlet pressure Reformer exit temperature Reformer exit pressure Combustion air temperature Fuel temperature Fuel flowrate Fluegas temperature at stack Convection section tubeside flowrates Convection section tubeside inlet and exit temperatures. 3

6 As noted elsewhere, if more plant data can be made available then there will be an increase in the accuracy of the plant simulations based on the plant data. In some cases certain plant data are not available. This must be recognized during the survey and alternate data collected which will allow a determination of the missing data. EAT AND MASS BALANCE RECONCILIATION 5 H Heat and mass balance programs (flowsheeting programs) have seen a step change in their availability and usability over the past few years. There are now many commercially available general flowsheeting packages with easy to use interfaces for operation on Personal Computers. Prior to this period in time powerful flowsheeting packages were limited to the large chemical companies employing the required specialists to produce and support this proprietary software. Over many years ICI developed its own in house flowsheeting software known as FLOWPACK. Synetix has taken this standard product and developed many specialist unit operations for use in the analysis of the performance of ammonia, hydrogen and methanol plants. These models have been augmented with highly accurate physical properties and thermodynamics which allow for further accuracy in the modelling of these plants. A complete flowsheeting tool is the best means of assessing the plant data from a reformer survey as the model can be built up with both the process streams and the furnace side streams, for both the radiant and convection sections of the primary reformer. The whole primary reformer is modelled as this allows the total thermal performance of the reformer to be determined for both the tube side and the furnace side. The heat and mass balance analysis itself may not determine why a primary reformer is under performing, but it will confirm this to be the case and provide a consistent set of plant data for further analysis by the primary reformer simulation program. The process flow data collected from the primary reformer survey are used in the developed flowsheet model of the plant along with the temperatures and pressures at the various key points in the process. The model is then run to determine the ideal performance of the primary reformer in terms of process gas exit composition and the fuel requirement to close the heat balance around the whole reformer. From this first analysis, the model can be developed further such that the plant data are matched. This is achieved by allowing certain key variables in the heat and mass balance model to float, such as the steam flow, approach to equilibrium and fuel flow in order to obtain a best fit to the plant data. The approach taken is to manipulate these key variables to minimize the errors between the plant data and the heat and mass balance. Using this approach highlights any inconsistencies in the plant data and results in a reconciled set of plant data. The more plant data that are available, the more accurate the reconciliation becomes as there is less freedom for data manipulation. OMPUTER SIMULATION OF PRIMARY REFORMERS 6 C The steam reforming process has been well described in many previous papers and publications. In order to simulate the performance of the primary reformer in detail it is necessary to consider the physical and chemical processes occurring both inside and outside the reformer tubes. Data from a number of sources have been combined by Synetix into a single complex computer simulation model to predict the performance of primary steam reformers. The program uses kinetic models for the hydrocarbon steam reforming reactions which cover the full range of steam reformer feedstocks and catalysts. The kinetic models have been validated on full size single tube reformers and in instrumented tubes within operating primary reformers. 4

7 Tubeside heat transfer is modelled based on experimentally derived heat transfer correlations, which relate the heat transfer coefficient to the reforming catalyst parameters. Standard heat transfer correlations do not adequately predict the heat transfer properties of modern multi-hole steam reforming catalysts, therefore, this testing is of great importance. The most advanced and realistic method of predicting the furnace side heat transfer to date is one whose methodology was described by Roesler in Reference 2. This model divides the overall radiative heat flows into two colours : radiation which interacts with the carbon dioxide and water molecules in the fluegas (beam radiation) and the radiation which passes through the fluegas with no interaction (window radiation). The method is suitable for use in all furnace geometries such as top-fired, side-fired, terraced-wall and bottom-fired. The only data input into the simulation by the process engineer is the geometry of the furnace and the tubes, the catalyst type and the process gas feed conditions and fuel and air supply conditions. During the solution the simulation breaks the furnace down into a number of elements, and allocates heat flows within each element as shown in Figure 2 for a side-fired reformer. Figure 2: Reformer Model Heat Flows Therefore, the reformer simulation program is entirely free to determine all reformer operational parameters such as tube wall temperature profile, process gas exit temperature, methane slip and fluegas exit temperature. This is not the case with basic furnace modelling programs, these have simplifications such as forced tube wall temperature profiles or limited heat input methods. Such programs are only of use as preliminary assessment tools. Over many years Synetix has licensed this steam reformer simulation software to many reformer designers, resulting in many reformers world-wide having been built on the basis of this program. This ultimately provides a substantial validation base for the computational methods used. Figures 3 and 4 show the predicted tube wall temperature profiles for a top-fired reformer and a side-fired reformer against measured temperatures. 5

8 Figure 3: Top-Fired Reformer Tube Wall Temperatures Figure 4: Side-Fired Reformer Tube Wall Temperatures By taking the data from the heat and mass balance analysis a reformer simulation can be carried out using the reconciled process conditions along with the fuel and combustion air rates. The key target data from the plant are the tube wall temperatures, the process gas composition, process gas exit temperature and the fluegas exit temperature. The variables available to achieve a fit to these target data in the reformer simulation program are limited to the catalyst activity and the heat input profile from the burners. The fuel flow is not a variable as this has been fixed by the heat and mass balance reconciliation and the reformer simulation must agree with this. Therefore, there are fewer variables (2) than the number of independent targets (4) which forces the reformer simulation to achieve a single unique solution. 6

9 By using the combined analysis of the plant data with the heat and mass balance programs and the reformer simulation the most realistic assessment of the actual operating conditions within the steam reformer is obtained. By the use of all available plant data, including accurate tube wall temperatures, the final simulations are so constrained that only one solution is feasible. From this assessment the performance of the reformer can be understood, and the underlying problem, if any can be identified. As the assessment of the primary reformer operation involves the production of a fitted heat and mass balance and a reformer simulation, these models can then be used to assess the potential for optimization of the reformer operating parameters. The assessment can also be used to evaluate the performance of the primary reformer in terms of the balance of heat input within the reformer, and the potential to increase reformer throughput within the tube wall design temperature. ASE STUDY 1 7 C The operator of a world scale methanol plant based around a Foster Wheeler terraced-wall reformer brought to our attention his concern over the low packed density obtained on a recent reformer change out. During start-up of the charge of catalyst it was felt that the tube wall temperatures were hotter and the approach to equilibrium higher than expected. In order to assess this, a full reformer survey was carried out to determine the actual catalyst performance. The data collected from the plant were extensive, including all reformer measurements as listed above, the whole of the plant steam system and methanol production rates. Full chemical analysis of all reformer process streams, fuels and fluegas was obtained. The tube wall temperature measurement was carried out using both a gold cup pyrometer and an infrared pyrometer. The measured temperatures are given below in Table 1, which illustrates the need to measure the tube temperatures with an emissivity of 1.00 and the efficacy of the correction techniques for the infrared pyrometer. Table 1: Case Study 1 Measured Tube Wall Temperatures. Gold Cup IR pyrometer IR pyrometer IR pyrometer Corrected raw data raw data Pyrometer - n/a emissivity setting Upper peephole C F Middle peephole C F Lower peephole C F

10 The FLOWPACK flowsheeting package was used to reconcile the plant data by closing the thermal balance around the reforming section of the plant. In order to achieve this the flowsheet model included not only the process streams, but also the fluegas side of the reformer and convection section, the plant steam system and the heat recovery from the synthesis gas cooling. The only doubtful measurement on the plant was the purge fuel consumption on the reformer as this was measured by two independent flow meters and these gave different flows. The heat and mass balance was able to determine the actual fuel consumption on the reformer as the model included all relevant heat sources and heat sinks. The key plant data are listed below in Table 2: Table 2: Case Study 1 Data Reconciliation Results. Plant Data Reconciled Data Natural gas feed Kmol/hr lbmol/hr Hydrogen recycle Kmol/hr lbmol/hr Reaction steam Te/hr m lb/hr Reformer exit temperature C F Methane slip mol%dry Approach to equilibrium C n/a 10.5 F 19 Natural gas fuel Kmol/hr lbmol/hr Purge gas fuel Kmol/hr 1800 or lbmol/hr 3969 or Fluegas temperature C F Fluegas temperature at stack C F Steam raised Te/hr m lb/hr

11 The heat and mass balance study confirmed that the plant data were consistent and provided an accurate set of data on which a reformer simulation could be based. The data reconciliation confirmed which of the purge fuel flowmeters was correct. The only significant difference between the plant data and the heat and mass balance was the fluegas exit temperature from the radiant section of the reformer, which was given as 83 C (149 F) higher than the measured value. The measurement of the fluegas temperature is difficult due to the large amount of radiation within the reformer, and a large error is normally encountered in this measurement. The above data were then used as the inputs to the steam reformer simulation program. The reformer simulation was carried out with the catalyst packing density reduced from the standard value to match the value achieved on the plant. As stated in the introductory sections, the reformer simulation is constrained to achieving a fit with the plant temperatures, gas compositions, tube wall temperatures and the heat and mass balance. The simulation achieved this fit without any additional variation of catalyst parameters, and the key data are reported below in Table 3. Table 3: Case Study 1 Reformer Simulation Results. Reconciled Data Reformer Simulation Reformer exit temperature C F Methane slip mol%dry Approach to equilibrium C F Natural gas fuel Kmol/hr lbmol/hr Purge gas fuel Kmol/hr lbmol/hr Fluegas temperature C F Maximum tube temperature C F Figure 5 shows the tube wall temperature profile predicted by the reformer simulation compared to the gold cup and corrected infrared temperatures. 9

12 Figure 5: Case 1 Tube Wall Temperatures The reformer simulation program was then used to determine the effects of the low density charge on the reformer operating parameters. The differences are given below in Table 4: Table 4: Case Study 1 Effects of Catalyst Charged Density. Low density Normal density Methane slip mol%dry Approach to equilibrium C F Natural gas fuel Kmol/hr lbmol/hr Fluegas temperature C F Maximum tube temperature C F From the results of the assessment, the catalyst was in fact performing slightly worse than originally expected, but only marginally so, and had the catalyst been charged to the expected density then the performance would have been exactly as expected. During the survey it was established that the plant was operating to a perceived tube wall temperature limitation and this set the plant operational philosophy. The reformer firing was increased until the tube wall temperatures approached the design temperature. This then fixed the maximum throughput of natural gas and also fixed the production of steam within the convection section. This resulted in no flexibility in the operation of the plant steam system. 10

13 The tube design temperature was 970 C (1780 F) and the tubes were considered to be operating at approximately 950 C (1740 F). The tube wall temperature measurements were taken with an infrared pyrometer with the emissivity set below 1.00 and no correction was made for background radiation. As discussed above, the emissivity should always be set to 1.00 on an infrared pyrometer to minimize the temperature errors. The experience with the tubes confirmed the operating temperature. The tubes were almost 20 years old, had never suffered a failure, and according to metallurgical inspection there was no measurable creep. Whilst this is a conservative and extremely safe manner in which to operate a steam reformer, it does limit the operational flexibility to a greater extent than is realistically required. The operator was interested in achieving fuel savings by reducing the amount of steam raised on the plant. In fact the operator was considering the installation of a pre-reformer to achieve this goal. The survey and subsequent data analysis resulted in the production of a highly accurate plant model which could be used to explore variations in plant operation. The model was used to investigate how the reformer could be operated to achieve the requirements of the operator, and improve the efficiency and operational flexibility of the plant. As the tube wall temperatures were established beyond any doubt to be substantially below the design temperature, the firing could in fact be increased on the bottom terrace of the furnace without jeopardizing tube life. Fuel burnt at the bottom of a terraced-wall reformer is more efficient since it has to flow upwards past the whole tube length, which allows it to give up more heat than fuel fired on the upper terrace. Therefore, the optimization of the reformer altered the fuel firing pattern from 45% of the fuel fired on the bottom terrace to 55% fired on the bottom terrace, with no change in the process side conditions within the reformer, or methanol production rate from the plant. The key figures are produced below in Table 5: Table 5: Case Study 1 Reformer Optimization Results. Original operation Optimized operation Natural gas feed Kmol/hr lbmol/hr Reformer exit temperature C F Methane slip mol%dry Approach to equilibrium C F Natural gas fuel Kmol/hr lbmol/hr Purge gas fuel Kmol/hr lbmol/hr Fuel fired at base % Fluegas temperature C F Tube wall temperature C F

14 This analysis shows that the reformer can be operated in a more flexible manner and that there is an improved operating method which results in a reduction in the fuel consumption on the plant for the same methanol production rate. Furthermore, operation in this manner involves no capital investment as is required by a pre-reformer. ASE STUDY 2 8 C The operator of a large steam reformer was concerned that the catalyst in the reformer was significantly under performing, with an approach to equilibrium of 19 C (35 F), more than twice the figure expected. A full reformer survey was carried out to determine the true performance of the catalyst and to determine the true approach to equilibrium. The data collected from the plant were extensive, including all reformer measurements as listed above and the whole of the plant steam system. Full chemical analysis of all reformer process streams, fuels and fluegas was obtained. The tube wall temperature measurement was carried out using both a gold cup pyrometer and an infrared pyrometer. The measured temperatures are given below in Table 6: Table 6: Case Study 2 Measured Tube Wall Temperatures. Gold Cup IR pyrometer IR pyrometer IR pyrometer Corrected raw data raw data Pyrometer - n/a emissivity setting Upper peephole C F Middle peephole C F Lower peephole C F The exit composition was determined to have an equilibrium temperature of 853 C (1567 F) based on a process mass balance across the reformer, and the analysis of the process gas exit the reformer. The operator was adding 34 C (61 F) onto the measured temperature inlet the downstream equipment to produce an estimated temperature at the catalyst exit. This gave a catalyst exit temperature of 872 C (1602 F), and resulted in the calculated approach to equilibrium being 19 C (35 F). However, from the tube wall temperature survey the reformer tubes were no hotter than normal, which suggested that the catalyst had not been deactivated. The temperatures within the furnace were even throughout the furnace, confirming that uneven flows through the tubes or uneven firing were not responsible for the higher than expected approach to equilibrium. In this case the correction to the reformer exit temperature of 34 C (61 F) was called into doubt, as there was no measured value for this. It was possible that as the reformer had just undergone an overhaul, the heat lost from the reformer exit system may have been reduced due to maintenance work. Therefore, 34 C (61 F) may have been acceptable for the reformer in the past, but at the time of the survey it was too high. 12

15 In order to determine the catalyst exit temperature iteration was required between both the heat and mass balance program and the reformer simulation program. On their own neither could independently resolve this problem. Both methods reached a simultaneous solution at a reformer exit temperature of 862 C (1584 F) with an approach to equilibrium of 9 C (17 F). The key results of both simulations are given in Table 7 below: Table 7: Case Study 2 Data Reconciliation Results. Plant Data Reconciled Reformer Data simulation Natural gas feed Kmol/hr lbmol/hr Hydrogen recycle Kmol/hr lbmol/hr Reaction steam Te/hr m lb/hr Reformer exit temperature C n/a F Methane slip mol%dry Approach to equilibrium C n/a 9 9 F Natural gas fuel Kmol/hr lbmol/hr Purge gas fuel Kmol/hr lbmol/hr Fluegas temperature C F Tube wall temperature C 897 n/a 896 F Fluegas stack temperature C n/a F Steam raised Te/hr n/a m lb/hr Figure 6 shows the predicted tube wall temperature profile for an exit temperature of 862 C (1584 F) compared to the gold cup and corrected infrared temperatures. 13

16 Figure 6: Case 2 Tube Temperatures at 862 C (1584 F) The heat and mass balance study confirmed that the plant data were consistent and provided an accurate set of data on which a reformer simulation could be based. The reformer simulation also fitted the plant data with a high degree of accuracy, with the simulation being carried out on the basis of standard catalyst properties. The fact that both the heat and mass balance and the reformer simulation are consistent together and with the plant data proves that this was the correct analysis of the plant operation. From this it was concluded that the catalyst was performing as expected and the source of the error was in the correction for heat losses in the reformer exit. In order to prove that the catalyst was not in fact deactivated and causing a high approach with an exit temperature of 872 C (1602 F), the simulations were run on the basis of this exit temperature. The results from these simulations are given below in Table 8: 14

17 Table 8: Case Study 2 Data at Alternate Exit Conditions. Plant Data Reconciled Reformer Data simulation Natural gas feed Kmol/hr lbmol/hr Hydrogen recycle Kmol/hr lbmol/hr Reaction steam Te/hr m lb/hr Reformer exit temperature C n/a F Methane slip mol%dry Approach to equilibrium C n/a F Natural gas fuel Kmol/hr lbmol/hr Purge gas fuel Kmol/hr lbmol/hr Fluegas temperature C F Tube wall temperature C 897 n/a 913 F Fluegas stack temperature C n/a F Steam raised Te/hr n/a m lb/hr From the results in the above table it can be seen that the assumption of 872 C (1602 F) exit temperature and a high approach to equilibrium are not consistent as the reformer simulation can not match the tube wall temperatures. Figure 7 shows the predicted tube wall temperature profile compared to the gold cup and corrected infrared temperatures for an exit temperature of 872 C (1602 F), which confirms that the tube wall temperatures are not consistent. 15

18 Figure 7: Case 2 Tube Temperatures at 872 C (1602 F) The operator accepted the findings of the survey and associated data analysis and operated the catalyst for several additional years in the knowledge that it was performing perfectly satisfactorily. 9 C ONCLUSIONS The case studies contained within the paper serve to illustrate how a combination of reformer assessment techniques can be used together to determine the true operation of a primary steam reformer. The techniques used include accurate tube wall temperature measurement, heat and mass balance modelling of the process and furnace sides of the reformer and reformer simulation. Each of the techniques used in isolation, or to a lesser degree of rigour would not in these or similar cases have resulted in a true assessment of the reformers actual performance. The cases also validate the approach taken for reformer tube wall temperature measurement and correction methods, and for the reformer thermal modelling techniques as used within Synetix. 10 R EFERENCES 1. Tube Wall Temperature Measurement in Steam Reformers. B J Cromarty & S C Beedle AIChE Ammonia Safety Symposium. 2. Theory of Radiative Heat Transfer in Co-Current Tube Furnaces. F C Roesler. Chemical Engineering Science, 1967, volume

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