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Simulation of the Scrubber Section of a Fluid Coker Jankovic, Jasna 2005

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SIMULATION OF THE SCRUBBER SECTION OF A FLUID COKER  by Jasna Jankovic B.A.Sc, University of Belgrade, Yugoslavia, 1996  A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE  in  THE FACULTY OF GRADUATE STUDIES CHEMICAL AND BIOLOGICAL ENGINEERING  THE UNIVERSITY OF BRITISH COLUMBIA April, 2005 © lasna Jankovic, 2005  Abstract HYSYS.Plant Version 3.0.1 in steady-state mode was used to simulate the Scrubber Section of a Syncrude Canada Ltd.'s Fluid Coker, a plant for oil sand bitumen upgrading. In this scrubber, hot vapours from the Fluid Coker are contacted counter-currently with cooler oils to remove heavy components. The objective was to develop a reliable simulation model, which would describe the plant operation as closely as possible, and to use this model to investigate possible process improvements, by changing process and design parameters. Plant data was used to define the composition, flow rate, temperature and pressure of all inlet streams, as well as parameters for all unit operation blocks. Additional data was provided to evaluate the reliability of the simulation model. The Scrubber Section was simulated using a number of unit operation blocks and process streams. The HYSYS Peng-Robinson property package was utilized. Heavy hydrocarbon mixtures were defined using pseudo-components derived from input laboratory assays data: boiling curves, density and viscosity. An investigation on presence of liquid phase in the vapour streams and heavy components in the Scrubber Overhead was undertaken, and its suggestions taken into account during the simulation. When the whole flowsheet was set up and a converged solution obtained, the HYSYS optimizer tool was used to determine unknown parameters in the system, such as tray and section efficiencies in the sheds and the packed section of the Scrubber, respectively, and fractions of vapour and liquid that reach equilibrium above the scrubber pool. An objective function was defined to quantify the extent of matching of model predictions with the plant data. The unknown parameters were varied to minimize the objective function. The set of parameters that resulted in the smallest deviation from the plant data was chosen and fixed as the "Base Case". Results of the simulation match the plant data very well (within 3.2% of the plant data). Eleven case studies were carried out in which different operating parameters and design changes were simulated to study their effects on predicted process performance: ATB Flow Rate, HGO Wash Flow Rate, HGO Underwash Flow Rate, HGO Wash Temperature, HGO Underwash In and Out of Service, Number of Trays in the Sheds, Number of Grid Sections, Simulation of the Conditions from Start of Run to End of Run, Water Instead of HGO Underwash, Saturated Steam Instead of HGO Underwash and Overhead Recycle Cut Point Changes. Based on the results of the case studies the suggestions for further process improvements were made, as well as recommendations for additional investigations. ii  Table of Contents  Abstract  ii  Table of Contents  iii  List of Tables  vi  List of Figures  ix  ACKNOWLEDGMENTS  xiii  Chapter 1 - Introduction  1  1.1. Oil Sand Processing Background  2  1.2. Fluid Coker  3  1.3. Scrubber Section  5  1.4. Project Objective  7  Chapter 2 - Process Simulator HYSYS Plant  8  2.1. Introduction to HYSYS - Literature Review  8  2.2. HYSYS Simulation Basis  10  2.3. Property Package and Flash Calculation  11  2.4. Operation Units and Logical Operations  14  Chapter 3 - Scrubber Section Simulation Model  16  3.1. Introduction  16  3.2. Simulation Structure Set Up  17  3.2.1. Property Package  17  3.2.2. Oil Characterization  17  3.2.3. Core Blocks and Simulation Components  17  3.2.4. Simulation Flowsheet  20  3.2.5. Input Plant Data  23  3.3. Optimizer Tool and the Base Case  26  Chapter 4 - Presence of Liquid Phase in the Vapour Streams  30  4.1. Introduction  30  4.2. Droplet Size Estimation  31 iii  4.3. Trajectory of the Liquid Droplets  32  Chapter 5 - Presence of Heavy Components in the Scrubber Overhead  37  5.1. Introduction  37  5.2. Liquid Entrainment in the Shed Section  38  5.3. Packed Section  40  5.4. Conclusion  46  Chapter 6 - Case Studies: Results and Discussion  47  6.1. Introduction  47  6.2. Case Studies  49  I. ATB Flow Rate  49  II. HGO Wash Flow Rate  57  III. HGO Underwash Flow Rate  65  IV. HGO Wash Temperature  73  V. HGO Underwash In and Out of Service  81  VI. Number of Trays in the Sheds  90  VII. Number of Grid Sections  98  VIII. Simulation of the Conditions from Start of Run to End of Run  106  IX. Water Instead of HGO Underwash  114  X. Saturated Steam Instead of HGO Underwash...  123  XI. Overhead Recycle Cut Point Changes  132  Chapter 7- Summary of Proposed Process Performance Improvements.. 140 7.1. Overhead Product Quality  140  7.2. Overhead Production Rate  142  7.3. Fouling in the Koch Grid  143  Chapter 8 - Conclusions and Recommendations  145  8.1. Conclusions  145  8.2. Recommendations  148  Glossary of Terms  150  References  152  Appendix I - Peng-Robinson Equation of State  156  Appendix II - Flash Block Calculation  159 iv  Appendix III - Scrubber Section Streams Data  170  Cyclone Product  170  ATB Assay  175  HGO Assay  177  Scrubber Overhead  179  Appendix IV - Cyclone Liquid Droplets Trajectory  184  V  List of Tables Table 3.1 Stream input data - information obtained from Syncrude Canada Ltd  24  Table 3.2 Input data and information for operation units obtained from Syncrude Canada Ltd..25 Table 3.3 Base Case parameter values and deviation from the plant data  28  Table 3.4 Determined unknown parameters (primary variables)  29  Table 4.1 Parameter values used in Equation (4.1)  32  Table 5.1 Parameter values for calculation the flow and capacity parameter for Figure 5.1  40  Table 5.2 Packed tower rating data calculated by Koch-Glitsch KG-Tower software.  43  Table 5.3 Parameter values for calculation the flow and capacity parameter for Figure 5.3  45  Table 1-1 Effect of ATB flow rate on Scrubber parameters  52  Table 1-2 Effect of ATB flow rate on Scrubber Overhead properties  53  Table 1-3 Effect of ATB flow rate on Scrubber Bottom properties  54  Table II-l Effect of HGO Wash flow rate on Scrubber parameters  60  Table II-2 Effect of HGO Wash flow rate on Scrubber Overhead properties  61  Table II-3 Effect of HGO Wash flow rate on Scrubber Bottom properties  62  Table III-l Effect of HGO Underwash flow rate on Scrubber parameters  68  Table III-2 Effect of HGO Underwash flow rate on Scrubber Overhead properties  69  Table III-5 Effect of HGO Underwash flow rate on Scrubber Bottom properties  70  Table IV-1 Effect of HGO Wash temperature rate on Scrubber parameters  76  Table IV-2 Effect of HGO Wash temperature on Scrubber Overhead properties  77  Table IV-3 Effect of HGO Wash temperature on Scrubber Bottom properties  78  Table V-l Effect of HGO Underwash service rate on Scrubber parameters  85  Table V-2 Effect of HGO Underwash service on Scrubber Overhead properties  86  Table V-3 Effect of HGO Underwash service on Scrubber Bottom properties  87  Table VI-1 Effect of number of Sheds trays on Scrubber parameters  93  Table VI-2 Effect of number of Sheds trays on Scrubber Overhead properties  94  Table VI-3 Effect of number of Sheds trays on Scrubber Bottom properties  95  Table VII-1 Effect of number of Grid sections on Scrubber parameters  101  Table VII-2 Effect of number of Grid sections on Scrubber Overhead properties...._  102  Table VII-3 Effect of number of Grid sections on Scrubber Bottom properties  103  Table VIII-1 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber parameters  109 vi  Table VIII-2 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead properties  110  Table VIII-3 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom properties  Ill  Table IX-1 Effect of water instead of HGO Underwash on Scrubber parameters  118  Table IX-2 Effect of water instead of HGO Underwash on Scrubber Overhead properties  119  Table IX-3 Effect of water instead of HGO Underwash on Scrubber Bottom properties  120  Table X-l Effect of saturated steam instead of HGO Underwash on Scrubber parameters  127  Table X-2 Effect of saturated steam instead of HGO Underwash on Scrubber Overhead properties  128  Table X-3 Effect of sat. steam instead of HGO Underwash on Scrubber Bottom properties.. ..129 Table XI-1 ATB flow rate effect on Overhead TBP distillation curve  134  Table XI-2 Effect of ATB flow rate on Scrubber parameters  135  Table XI-3 HGO Wash flow rate effect on Overhead TBP distillation curve  136  Table XI-4 Effect of HGO Wash flow rate on Scrubber parameters  137  Table XI-5 HGO Underwash flow rate effect on Overhead TBP distillation curve  138  Table XI-6 Effect of HGO Underwash flow rate on Scrubber parameters  139  Table AH.l Parameters for the flash block system components  163  Table AII.2 PR EOS parameters for pure substances  164  Table AII.3 Interaction parameters for Hydrogen-Methane-Ethane system  164  Table AIII.l Composition of hypothetical cyclone stream  171  Table AIII.2 Composition of Light Ends fraction of cyclone stream  171  Table AIII.3 CGO assay; Method: ASTM 2887 with HTSB enhancement  172  Table AIII.4 CGO TBP data; Method: TBP calculated by HYSYS  172  Table AIII.5 OTSB Assay; Method: ASTM 2887 & SCFE-composite data  173  Table AIII.6 OTSB TBP data; Method: TBP calculated by HYSYS  173  Table AIII.7 Cyclone Product TBP data; Method: TBP calculated by HYSYS  174  Table AIII.8 Cyclone Product composition  174  Table AIII.9 ATB assay; Method: ASTM 2887 with HTSD enhancement  176  Table AIII.l0 ATB TBP data; Method: TBP data calculated by HYSYS  176  Table AIII.l 1 ATB composition calculated by HYSYS  176  Table AIII.12 HGO assay; Method: ASTM 2887 with HTSB enhancement  178  Table AIII.13 HGO TBP data; Method: TBP calculated by HYSYS  178 vii  Table AIII.14 HGO composition (HYSYS)  178  Table AIIL15 Scrubber Overhead fractions  180  Table AIII.16 Sour Gas composition  180  Table AIIL17 CGO Assay; Method: SIM Dist  181  Table AIIL18 CGO TBP data; Method: TBP calculated by HYSYS  181  Table AIII.19 Naphtha Assay; Method: SIM Dist  182  Table AIII.20 Naphtha TBP data; Method: TBP calculated by HYSYS  182  Table AIII.21 "Plant" Scrubber Overhead TBP data; Method: TBP calculated by HYSYS.. ..183 Table AIII.22 "Plant" Scrubber Overhead composition and fraction distribution  183  List of Figures Figure 1.1 Schematic of oil sand processing  3  Figure 1.2 Schematic of a Fluid Coker  4  Figure 1.3 Schematic of the Scrubber Section of the Fluid Coker  5  Figure 3.1 Core blocks chosen to represent the Scrubber Section of the Fluid Coker  19  Figure 3.2 Simulation flowsheet of the Scrubber Section  22  Figure 4.1 Trajectory of a liquid droplet carried with Cyclone Product jet  35  Figure 5.1 Flooding correlation for columns with cross-flow plates  39  Figure 5.2 Design pressure drop chart for Koch Flexigrid Type 2 structured packing  42  Figure 5.3 Generalized flooding-pressure drop correlation of Eckert and Leva, modified by Strigle  44  Figure 1-1 Effect of ATB flow rate on temperatures along the Scrubber  50  Figure 1-2 Effect of ATB flow rate on temperature profile along the Scrubber  50  Figure 1-3 Effect of ATB flow rate on mass flow rate of Scrubber Overhead and Bottom  51  Figure 1-4 Effect of ATB flow rate on mass flow rate of other streams  51  Figure 1-5 Effect of ATB flow rate on Scrubber Overhead TBP curve  53  Figure 1-6 Effect of ATB flow rate on Scrubber Bottom TBP curve  54  Figure 1-7 Effect of ATB flow rate on Scrubber Overhead composition  55  Figure 1-8 Effect of ATB flow rate on Scrubber Bottom composition  55  Figure II-l Effect of HGO Wash flow rate on temperatures along the Scrubber  57  Figure II-2 Effect of HGO Wash flow rate on temperature profile along the Scrubber  57  Figure II-3 Effect of HGO Wash flow rate on mass flow rate of Scrubber Overhead and Bottom  59  Figure II-4 Effect of HGO Wash flow rate on mass flow rate of other streams  59  Figure II-5 Effect of HGO Wash flow rate on Scrubber Overhead TBP curve  61  Figure II-6 Effect of HGO Wash flow rate on Scrubber Bottom TBP curve  62  Figure II-7 Effect of HGO Wash flow rate on Scrubber Overhead composition  63  Figure II-8 Effect of HGO Wash flow rate on Scrubber Bottom composition  63  Figure III-l Effect of HGO Underwash flow rate on temperatures along the Scrubber  65  Figure III-2 Effect of HGO Underwash flow rate on temperature profile along the Scrubber....65  Figure III-3 Effect of HGO Underwash flow rate on mass flow rate of Scrubber Overhead and Bottom  67  Figure III-4 Effect of HGO Underwash flow rate on mass flow rate of other streams  67  Figure III-5 Effect of HGO Underwash flow rate on Scrubber Overhead TBP curve  69  Figure III-6 Effect of HGO Underwash flow rate on Scrubber Bottom TBP curve  70  Figure III-7 Effect of HGO Underwash flow rate on Scrubber Overhead composition  71  Figure III-8 Effect of HGO Underwash flow rate on Scrubber Bottom composition  71  Figure TV-1 Effect of HGO Wash temperature on temperatures along the Scrubber  73  Figure IV-2 Effect of HGO Wash temperature on temperature profile along the Scrubber  73  Figure TV-3 Effect of HGO Wash temperature on mass flow rate of Scrubber Overhead and Bottom  75  Figure IV-4 Effect of HGO Wash temperature on mass flow rate of other streams  75  Figure IV-5 Effect of HGO Wash temperature on Scrubber Overhead TBP curve  77  Figure IV-6 Effect of HGO Wash temperature on Scrubber Bottom TBP curve  78  Figure IV-7 Effect of HGO Wash temperature on Scrubber Overhead composition  79  Figure IV-8 Effect of HGO Wash temperature on Scrubber Bottom composition  79  Figure V-l Effect of HGO Underwash service on temperatures along the Scrubber  82  Figure V-2 Effect of HGO Underwash service on temperature profile along the Scrubber  82  Figure V-3 Effect of HGO Underwash service on mass flow rate of Scrubber Overhead and Bottom  84  Figure V-4 Effect of HGO Underwash service on mass flow rate of other streams  84  Figure V-5 Effect of HGO Underwash service on Scrubber Overhead TBP curve  86  Figure V-6 Effect of HGO Underwash service on Scrubber Bottom TBP curve  87  Figure V-7 Effect of HGO Underwash service on Scrubber Overhead composition  88  Figure V-8 Effect of HGO Underwash service on Scrubber Bottom composition  88  Figure VI-1 Effect of number of Sheds trays on temperatures along the Scrubber Figure VI-2 Effect of number of Sheds trays on temperature profile along the Scrubber  91 91  Figure VI-3 Effect of number of Sheds trays on mass flow rate of Scrubber Overhead and Bottom  92  Figure VI-4 Effect of number of Sheds trays on mass flow rate of other streams  93  Figure VI-5 Effect of number of Sheds trays on Scrubber Overhead TBP curve  94  Figure VI-6 Effect of number of Sheds trays on Scrubber Bottom TBP curve  95  Figure VI-7 Effect of number of Sheds trays on Scrubber Overhead composition  96  Figure VI-8 Effect of number of Sheds trays on Scrubber Bottom composition  96  Figure VII-1 Effect of number of Grid sections on temperatures along the Scrubber  98  Figure VII-2 Effect of number of Grid sections on temperature profile along the Scrubber  98  Figure VII-3 Effect of number of Grid sections on mass flow rate of Scrubber Overhead and Bottom  100  Figure VII-4 Effect of number of Grid sections on mass flow rate of other streams  100  Figure VII-5 Effect of number of Grid sections on Scrubber Overhead TBP curve  102  Figure VII-6 Effect of number of Grid sections on Scrubber Bottom TBP curve  103  Figure VII-7 Effect of number of Grid sections on Scrubber Overhead composition  104  Figure VII-8 Effect of number of Grid sections on Scrubber Bottom composition  104  Figure VIII-1 Effect of pressure drop in Grid and absolute pressure in the Scrubber on temperatures along the Scrubber  107  Figure VIII-2 Effect of pressure drop in Grid and absolute pressure in the Scrubber on temperature profile along the Scrubber  107  Figure VIII-3 Effect of pressure drop in Grid and absolute pressure in the Scrubber on mass flow rate of Scrubber Overhead and Bottom  108  Figure VIII-4 Effect of pressure drop in Grid and absolute pressure in the Scrubber on mass flow rate of other streams  108  Figure VIII-5 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead TBP curve  110  Figure VIII-6 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom TBP curve  Ill  Figure VIII-7 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead composition  112  Figure VIII-8 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom composition  112  Figure IX-1 Effect of water instead of HGO Underwash on temperatures along the Scrubber  115  Figure IX-2 Effect of water instead of HGO Underwash on temperature profile along the Scrubber  115  Figure IX-3 Effect of water instead of HGO Underwash on mass flow rate of Scrubber Overhead and Bottom  117  Figure IX-4 Effect of water instead of HGO Underwash on mass flow rate of other streams...! 17 xi  Figure IX-5 Effect of water instead of HGO Underwash on Scrubber Overhead TBP curve. ..119 Figure IX-6 Effect of water instead of HGO Underwash on Scrubber Bottom TBP curve  120  Figure IX-7 Effect of water instead of HGO Underwash on Scrubber Overhead composition. 121 Figure IX-8 Effect of water instead of HGO Underwash on Scrubber Bottom composition... 121 Figure X-l Effect of saturated steam instead of HGO Underwash on temperatures along the Scrubber  124  Figure X-2 Effect of saturated steam instead of HGO Underwash on temperature profile along the Scrubber  124  Figure X-3 Effect of saturated steam instead of HGO Underwash on mass flow rate of Scrubber Overhead and Bottom  126  Figure X-4 Effect of saturated steam instead of HGO Underwash on mass flow rate of other streams  126  Figure X-5 Effect of sat. steam instead of HGO Underwash on Scrubber Overhead TBP curve  128  Figure X-6 Effect of sat. steam instead of HGO Underwash on Scrubber Bottom TBP curve  129  Figure X-7 Effect of saturated steam instead of HGO Underwash on Scrubber Overhead composition  130  Figure X-8 Effect of saturated steam instead of HGO Underwash on Scrubber Bottom composition  130  Figure XI-1 ATB flow rate effect on Overhead TBP distillation curve  134  Figure XI-2 HGO Wash flow rate effect on Overhead TBP distillation curve  136  Figure XI-3 HGO Underwash flow rate effect on Overhead TBP distillation curve  138  Figure AII.l Schematic of the flash block  159  Figure AIII.l Cyclone Product TBP curve  174  Figure AIII.2 Cyclone Product molecular weight distribution curve  174  Figure AIII.3 Cyclone Product density distribution curve  174  Figure AIII.4 ATB TBP curve  176  Figure AIII.5 HGO TBP curve  178  Figure AIII.6 "Plant" Scrubber Overhead TBP curve  183  Figure AIII.7 "Plant" Scrubber Overhead molecular weight distribution curve  183  Figure AIII.8 "Plant" Scrubber Overhead density distribution curve  183  Figure AIV.l Trajectory of a liquid droplet carried with the Cyclone Product jet  188 xii  ACKNOWLEDGMENTS  I would like to express my sincere thanks to Dr. Paul Watkinson and Dr. Dusko Posarac, my supervisors, for their support and guidance throughout the duration of my work. Special thanks to Dr. Iftikhar Huq from Syncrude Canada Ltd. for his help and valuable suggestions during this project. Financial support provided by Syncrude Canada Ltd. and NSERC is gratefully acknowledged. I would like to dedicate this thesis to my family, my husband Bosko and my children, for their patience, great support and encouragement, which gave me the strength over these years.  Xlll  Chapter I - Introduction  Chapter 1 - Introduction Today's world industry, economy and politics are greatly dependent on the fossil fuel energy availability. In 2004, 40% of world energy consumed is oil, 20% gas, 20% coal and the remaining 20% is hydro-electric power, biomass and renewable energy [1]. Energy demand is expected to increase over the next period until 2025 at an average of 2% per year. Fossil fuels dominate the global energy demand, with up to 90% of the total demand. OPEC Oil Outlook to 2025 reports that the volume of oil demand will increase from 77 million barrels per day in 2002 to 115 million barrels per day in 2025 [2]. Some sources estimate that at this rate of consumption the current recoverable reserves of oil will be spent in about 50 years [3,4]. The estimates of the world ultimately recoverable reserves (URR) given by the US Geological Survey (USGS) are about 3.3 trillion barrels [5]. These reserves include a huge amount of Canadian oil sands as well, making Canada the second-largest holder of reserves after Saudi Arabia. Total recoverable oil reserves in Alberta are estimated at over 334 billion barrels, with the oil sand production of 964,000 barrels per day and conventional crude oil production of 629,000 barrels per day in 2003, [5]. Syncrude Canada Ltd. and Suncor Inc., located in the Northern Alberta, produce crude oil from oil sand, which is about 18% of total crude oil production in Canada. Cost of the oil production from oil sand is still high, comparing to the conventional crude oil production. Oil sand recovering and processing improvements lead to the decrease of the cost per barrel of oil, as well as increase in the ability to recover and process more of oil sand. Since the reserves of conventional fossil fuels are in decline, and having in mind huge reserves of oil sand, this could have a significant positive impact on current fossil fuel energy situation. Syncrude Canada Ltd., as one of the largest producers of sweet crude oil and other products recovered from oil sand, has been improving the processes for recovering and upgrading oil sand bitumen over many years. Continuous research and plant development led by Syncrude Canada Ltd. include also use of modern means of computer process simulation in business planning, plant design and process optimization. The majority of their bitumen upgrading stages has been simulated so far. In this project, HYSYS.Plant Version 3.0.1 process simulator was used to 1  Chapter I — Introduction  simulate the Scrubber Section of a Syncrude Canada Ltd.'s Fluid Coker, a plant for upgrading the bitumen that originates from the oil sand. General oil sand processing, as well as Fluid Coker and detailed Scrubber Section operation are described in Sections 1.1-1.3.  1.1. Oil Sand Processing Background Oil sands are deposits composed of sand, bitumen, mineral rich clays and water. Bitumen is a very thick, viscous product of the oil sand. In order to be transportable by pipeline and usable by conventional refineries it must be upgraded to synthetic crude oil or diluted with lighter hydrocarbons [6, 7]. Oil sand processing starts with digging the oil sand by mining shovels and transporting by trucks to crushing stations, where it is broken down to chunks about 45 cm. After that, the ore is fed to rotating drums for further reducing the size to 5 cm. At this point, warm water is added to the oil sand to create slurry. The slurry is pumped through a pipeline to the extraction unit. The mixing during the slurry transport from the mine to the plant begins the separation process and recovers over 90% of the bitumen. The resulting bitumen froth is separated from the water and sand in froth settlers, where a hydrocarbon solvent is added to separate the remaining solids, water and heavy asphaltenes. The clean, diluted bitumen is low in contaminants and with relatively low viscosity is easily transported by pipeline to upgrading process. The upgrading process of the diluted bitumen starts with Diluent Recovery Unit. This is an atmospheric distillation column, which serves to separate diluent naphtha (used as a solvent in bitumen cleaning process), to remove light components and to produce Atmospheric Topped Bitumen (ATB) as feedstock for the Fluid Cokers, LC-Finer and Vacuum Distillation Unit. The Vacuum Distillation Unit processes about 55% of ATB. It removes light and heavy gas oils which are then sent directly to hydro treat ers. The residual - Vacuum Topped Bitumen (VTB) is blended with the other 45% of ATB and then sent to the LC-Finer and Fluid Cokers for further processing. Bitumens have low H/C ratios, which can be raised by either adding hydrogen or removing carbon. LC-Fining is a catalytic process in which hydrogen is added to increase the hydrogen to carbon ratio in the feed hydrocarbon material, and light gas oil (LGO) is produced. The unreacted residue from the LC-Finer is sent to a Fluid Coker for further cracking. ATB, VTB and 2  Chapter I - Introduction  LC-Finer residue are fed to the Fluid Coker. The coking process removes part of the carbon content of the feedstock by thermal cracking of long hydrocarbon chains in bitumen. The product vapours from the Coker and LC-finer are combined together and fractionated into Naphtha, Light and Heavy Gas Oil (Combined Gas Oil, CGO). Further treatment (hydrotreatment to remove heavy metals, sulphur and nitrogen) and blending of different products result in Sweet Blend crude oil, a 100% sweet, light, low-sulphur crude that is shipped by pipeline to refineries and mostly used for production of gasoline and diesel fuel [7, 8, 9,10]. A partial schematic of oil sand processing is shown in Figure 1.1, where the bitumen feed is taken to include ATB, VTB and LC-Finer residuum.  Figure 1.1 Schematic of oil sand processing [11]  1.2. Fluid Coker Hot ATB, VTB and LC-Finer residuum are fed continuously to the Fluid Coker unit where the feed is thermally cracked or broken down into lighter products (Figure 1.2). VTB and residuum feed are sprayed into a fluidized bed of coke particles positioned in the middle part of the reactor. Coking reactions occur on the surface of the particles at temperature of 510-530°C. Liquid that remains on the coke after the coking reactions is stripped off by steam in the Fluid Coker Stripper Section, located in the bottom part of the reactor. The coke is sent to the Burner, 3  Chapter I - Introduction where the coke is partially burned and recycled to the Coker to supply heat needed for the coking reaction. Excess coke is removed and stored for potential future use. In the Coker, the lighter products of cracking reactions (vapour) rise from fluidized zone through cyclones where coke particulates and most of the liquid droplets are removed. Product from the cyclone then enters the upper part of the Fluid Coker - the Scrubber Section [12]. In this project the Scrubber Section of a Syncrude Canada Ltd.'s Fluid Coker has been simulated. Therefore, this section will be described in more detail in the next section.  Figure 1.2 Schematic of a Fluid Coker  4  Chapter I - Introduction  1.3. Scrubber Section The Scrubber Section of a Syncrude Canada Ltd.'s Fluid Coker is approximately 17 m high and 9 m diameter section positioned at the top part of the Fluid Coker. The Scrubber Section itself consists of three main parts: the Scrubber Pool at the bottom, six sets of Sheds in the middle part and the Koch Grid - ten layers of Koch Flexigrid Type 2 structured packing at the top [13]. The purpose of the Scrubber Section is to remove ("scrub") heavy components from the hot rising vapour from the Coker cyclones, by contacting the lower temperature falling hydrocarbon liquids. The main product of the Fluid Coker is Scrubber Overhead, a mainly vapour product with the boiling range between -250 and 690°C. Its characteristics are given in Appendix III. This product exits from the top of the Scrubber Section and enters the Fractionator where four fractions are separated: Sour Gas, Butane, Naphtha and a Combined Gas Oil (CGO), consisting of Light Gas Oil (LGO) and Heavy Gas Oil (HGO). As mentioned in Section 1.1, after the hydrotreatment, Naphtha and CGO are used for blending into Sweet Blend crude. A schematic of the Scrubber Section of the Coker is shown in Figure 1.3.  Figure 1.3 Schematic of the Scrubber Section of the Fluid Coker 5  Chapter I - Introduction  The primary feed to the Scrubber Section, named Cyclone Product, comes out of the six cyclone snouts, positioned at the bottom part of the Scrubber, at a velocity of 76 m/s and a temperature of 540°C. The Cyclone Product is mainly vapour, but it is suspected that it contains some liquid and even solid particles of heavy hydrocarbons with boiling temperatures of over 1000°C. Due to the cyclone nozzles position and orientation, the vapour is expected to cause a swirling effect in the 1.5 m high section between the Scrubber Pool and the first row of Sheds. It exchanges heat and mass with the down-flowing liquid. This still hot rising vapour passes through six trays of Sheds, being contacted by colder liquids from the upper part of the Scrubber and ATB feed. ATB enters the Scrubber above the Sheds at 325°C and serves to scrub the heavy fractions and particulates from the rising vapour. Vapour further rises through the Koch Grid. Both below and above the Koch Grid, Heavy Gas Oil (HGO) enters the Scrubber also at 325°C. This HGO stream is one part of the Scrubber Overhead product, which is recycled from the downstream Fractionator, to help scrub heavy components from the vapour. It keeps the grid wet and controls the temperature in order to reduce fouling of the grid. Fouling can occur in processing equipment, particularly at temperatures above 400°C, and where liquids are stagnant. Heavy components partially volatilize, crack and "coke", building layers of deposits from both liquid and gas phases [14]. These deposits affect cyclone snouts and the Koch Grid the most, causing increases in pressure drop and decreases in process performance. For that reason, it is very important to reduce fouling, either by keeping the temperature low enough or by reducing stagnant zones which contain heavy liquid fractions. Scrubber Overhead vapour from the top of the Koch Grid exits the Scrubber at 390-400°C. As already mentioned in this section, this product is sent to the Fractionator and separated into Sour Gas, Butane, Naphtha and CGO (LGO and HGO) used for further treatment and blending, while one part of the HGO is recycled to the Scrubber Section. Liquid containing heavy fractions from HGO, ATB and Cyclone Product passes downward through the Koch Grid and the Sheds, scrubbing the rising vapour, and collects in the Scrubber Pool. Mixing of the Scrubber Pool Liquid by high pressure saturated steam -Agitation Steam keeps all particulates suspended. This liquid, which is pumped from the pool, is split in two streams: one that joins the VTB feed for the Coker and the other that is cooled by the Scrubber  6  Chapter I — Introduction  Pool Liquid Cooler (SPL Cooler) and recycled to the Scrubber Pool in order to keep its temperature below 400°C and reduce fouling.  1.4. Project Objective The Scrubber Section of the Coker involves complex mixtures of hydrocarbons with a wide boiling range; three phases - vapour, liquid and even solid; possible liquid entrainment in the vapour phase; multistage processes; fouling reactions etc. The whole process is not fully understood. The product quality and the system performance depend on process parameters, choice and properties of inlet streams and design of the units. A HYSYS process simulation of the Scrubber Section can help increase understanding, leading to process improvements. An attempt to simulate the Scrubber Section of the Syncrude Canada Ltd.'s Fluid Coker was by M . Williston as a Bachelor's Thesis project at UBC in 2002 [8]. This work, although successful in matching some plant data, showed some uncertainties. Not too much attention was paid to composition of the product stream, which is a crucial parameter for successful plant simulation. Also, some sections of the Scrubber were not represented in enough detail, which caused relatively high deviations from the plant data ( within 10%). In this project, a more detailed and realistic model of the Scrubber Section of Syncrude Canada Ltd.'s Fluid Coker was developed. The objective of the project was to develop a reliable simulation model for the Scrubber Section and to use this model to investigate possible process improvements, by changing process and design parameters. The model was utilized for different case studies with the goal to investigate the effects of parameter and design changes on process performance and gain better understanding of process behavior.  7  Chapter 2 - Process Simulator HYSYS Plant  Chapter 2 - Process Simulator H Y S Y S Plant 2.1. Introduction to HYSYS - Literature Review In order to remain competitive in the market and to meet government regulations, the process industries must improve and optimize their operations, making them more efficient, profitable, safe and reliable. Improvements to the process have to be undertaken throughout the plant lifecycle, quickly and without risky and costly on-site design changes. Process simulators are very efficient tools in improving design, evaluation of different operation changes, monitoring of equipment performance, optimizing the process and production planning. Process simulators have been widely used in the oil and gas, and petroleum refining industries for more than 30 years. Refinery unit operations are very specific, and most of the commercial process simulators are not efficient enough to model the whole process as an integrated system. However, some of the process simulators, such as Aspen Plus and Aspen RefSYS by Aspen Technology, Inc., HYSYS by Hyprotech, Ltd., and Pro/II by Simulation Sciences, Inc., are improved and adapted for use in petroleum process simulations [15]. In this project, the HYSYS process simulator was used. HYSYS is powerful engineering simulation software. It contains a variety of built-in property packages, a data base with experimental data for more than 1500 components and 16000 fitted binaries, a wide range of estimation methods for components not included in the data-base, and a regression package [16]. It also offers the ability for the user to include a specific property calculation, set of experimental data or coefficients, in order to improve accuracy for a specific simulation system. HYSYS has built-in routines to solve a wide range of specialized unit operations: separation operations, columns, heat transfer equipment, reactors, piping equipment (tees, mixers, valves), rotating equipment, solid separation operations, electrolyte operations, logical operations (adjuster, recycle, controller) [17]. HYSYS can be used in both steady state and dynamic modeling environment. Steady state simulations can be switched to dynamic mode by specifying additional engineering details, including pressure-flow relationships and equipment dimensions. Aspects of the HYSYS process simulator application in industry and research are various: process design (synthesis of new designs, analysis of current designs, process optimization), process operation (monitoring, control, data collection, operator training) and process 8  Chapter 2 -Process Simulator HYSYS Plant  management (production planning and scheduling, quality control), as well as application in order to obtain more data on a process and understand the process behaviour. Following are some examples of HYSYS applications: At a Chevron Canada gas plant both steady-state and dynamic HYSYS simulation were applied to investigate a modified Claus sulphur recovery plant. The aim of the study was to determine the effect of three different control schemes on the efficiency of the plant [18]. At a HOVENSA LLC refinery, a model for the optimization of the deisopentanizer tower was developed with the HYSYS process simulator, using averaged process and lab data. The average deviation from main plant parameters (temperature profile, compositions) was around 7% [19].  Lars et al. [20] report application of the HYSYS simulator to model the glycol regeneration processes after natural gas dehydration by absorption in triethylene glycol. Soave et al. [21] investigated the options for saving energy in industrial distillation towers by preheating the feed (or one part of the feed) with the heat recovered from the bottom product. The HYSYS process simulator is used to determine the optimum split ratio of the feed and feed tray, showing the economical impact of the proposed solution. In steady state and dynamic modeling of the xylene distillation column from the Mizushima Oil Refinery [22], temperature profile, flow rates and other parameters showed average deviation from the plant data of less than 10%. Process simulators used for petroleum process simulation (Aspen Plus, HYSYS, Pro/II), commonly use pseudo-components for petroleum mixture characterization. However, highly predictive and reliable models require accurate presentation of the phase-equilibrium behavior and hence more detailed defining of the streams composition. Analytical techniques such as chromatography, mass spectrometry and nuclear magnetic resonance spectroscopy give information that could be used in calculation of fluid properties. The application of these techniques leads to more detailed, but much larger process models. There are still not available algorithms for these kinds of models. Briesen et al. [23] have tried to apply this new approach to a refinery process simulation using continuous mixture representation instead of the commonly used pseudo-component approach. This continuous mixture approach assumes that the number of chemical species present in a petroleum mixture is so large that it can be considered a continuous 9  Chapter 2 -Process  Simulator HYSYS Plant  rather than a discrete distribution. The authors developed a new solution strategy for this problem and applied it to a 9-stage distillation column and tested for different feed mixtures. All tests showed better accuracy and efficiency for the proposed method compared to the conventional pseudo-component approach. Today, many companies are reported [24] to implement process simulation to improve process efficiency, optimize existing operation or to assist in business planning. Petro-Canada, Lurgi Oel-Gas-Chemie, Syncrude Canada Ltd, NOVA Chemicals, Akzo-Nobel and Alkon are some of the names mentioned [24].  2.2. HYSYS Simulation Basis In order to solve equations representing material and energy balances, the stream connections and the relations representing the equipment functions within a simulation flowsheet, HYSYS performs sequential modular process simulation [24]. In the sequential modular method, the process is represented by a collection of modules. A module is a model of an individual element in a flowsheet that can be isolated from the flowsheet and interpreted separately. Unit sequences (modules) are solved sequentially, iteratively, one by one until the convergence is met. HYSYS uses subroutines to model these process units, but in contrast to other simulators, it has the ability to perform calculation in both directions (forward and reverse). Also, reported by [25], HYSYS immediately interpret the commands, as they are entered, which makes the response of the program fast. The following steps are used to set up a new simulation model: •  Selecting a component list from HYSYS data base for known components included in the modelled system;  •  Defining an appropriate property package (Equation of State (EOS) or Activity model);  •  Supplying data (laboratory assays and bulk properties) for defining the pseudocomponents if complex mixtures are involved;  •  Installing the reaction components and formulating reactions, if they occur;  10  Chapter 2 -Process  •  Simulator HYSYS Plant  Defining the streams by providing their compositions, flow rates and two property variables (usually temperature and pressure). Automatic Flash calculation for all other properties of the streams, physical and transport (liquid density, vapour density, viscosity, thermal conductivity, surface tension, etc.), is done by HYSYS using property package with its physical and transport functions;  •  Installing the operation units and defining needed parameters;  •  Connecting the elements (streams and operation units);  Based on Vapour-Liquid Equilibrium (VLE), mass and energy balance and relations representing equipment operations, HYSYS performs calculations needed for model solution and convergence.  2.3. Property Package and Flash Calculation In the simulation process, one of the most important steps is the choice of the thermodynamic property package. It enables calculation of many stream properties: physical and transport properties, PVT relationships, VLE calculations, number of phases, phase composition, and hence affects the accuracy of material and energy balances. The choice of the property package depends on the chemical nature of the system (hydrocarbons, electrolytes, sour water, etc.), conditions (T, P), and parameter availability. For oil, gas and petrochemical systems, the Peng-Robinson EOS is one of recommended property packages. It contains enhanced binary interaction parameters for all hydrocarbonhydrocarbon and hydrocarbon - nonhydrocarbon pairs available in the HYSYS library [16]. The Peng-Robinson EOS is presented below:  P =^ V-b  V(V + b) +  (2.1)  a  b(V-b)  Here a and b represent deviation from ideal behaviour. Term a represents the strength of attraction between two molecules (interaction force), and b is proportional to the size of the  11  Chapter 2 -Process Simulator HYSYS Plant  molecules. These parameters can be determined from critical values P and T , and the acentric c  c  factor co for pure substances. Peng-Robinson EOS is presented in more detail in Appendix I. Based on the Peng-Robinson EOS the following properties can be calculated: the compressibility factor Z, molar volume, enthalpy, entropy, heat capacity, fugacity coefficient, fugacity of a phase, etc. In HYSYS, the compressibility factor is calculated as the root of the following equation, where the smallest root corresponds to the liquid phase and the largest for the vapour phase. Z -(\-B)Z 3  R  2  T  2  +Z(A-3B -2B)-(AB-B -B ) 2  2  3  =0  2  RT  Molar volume for the liquid or vapour phase can be calculated from:  ZRT v = ——  (2.3)  Phase equilibrium computations for heavy hydrocarbon mixtures are difficult because of the complexity of the mixtures and lack of experimental data. Critical temperature T , pressure P c  c  and acentric factor co of each component, needed for EOS calculations, are not available for all components present in complex hydrocarbon mixtures. They have to be estimated from measured properties for boiling point fractions: specific gravity, viscosity, molecular weight and distillation curve. Numerous relationships can be used for these purposes [26]. These correlations for critical properties and acentric factor and correlations for physical and transport properties viscosity, density, thermal conductivity, surface tension, etc. - are automatically selected by HYSYS based on the system under study. In the present simulation, the Lee-Kesler correlations for T , P , acentric factor and molecular weight are used [16, 26]. Twu's model for viscosity c  c  determination is chosen for heavy hydrocarbon mixture [16, 26]. Katz-Firoozabadi correlations were used for density and boiling points calculation, because they are accurate for hydrocarbons 12  Chapter 2 -Process Simulator HYSYS Plant  up to C45 [16, 26]. The Missenar and Reidel method is used for thermal conductivity calculation [16, 27] and for surface tension a modified equation of Brock and Bird is used [16, 28]. The equations of the selected property package, and the physical and transport property functions are used for the flash calculations to determine all thermodynamic, physical and transport properties of a stream. Based on degrees of freedom concepts HYSYS determines when and what type of flash calculation on stream it can perform. If stream composition and two property variables are known (temperature and/or pressure, and vapour fraction, enthalpy or entropy) the stream is thermodynamically defined. These properties are either specified by the user or calculated by an operation. Depending on known stream property variables, HYSYS automatically performs the flash calculations: T-P, T-VF, T-S, T-H, P-VF, P-S or P-H. [16]. Flash calculation is based on system tendency to reach thermodynamic equilibrium. Vapourliquid equilibrium ratio for a component i is given by the following equation:  K. = A = — i -  (2.4)  where y>\ and x, are mole fractions of component i in vapour and liquid phases, and O", and O', are the fugacity coefficients for the component i in the vapour and liquid phases. Fugacity coefficients can be calculated from a general thermodynamic equation: P  /?-r-ln<D, = 0  v.=(—)  rp [(\>i-R—)-dp  P  (2.5)  T  dV  where the molar volume v and the derivative  can be calculated using EOS. dn  t  As a starting point, the composition and molar volume of each phase must be estimated. EOS equation is used to improve the values during iteration. The equilibrium ratio, K for each 13  Chapter 2 -Process Simulator HYSYS Plant  component is calculated and compared to assumed phase composition. Further iteration leads to the convergence of the solution [29].  2.4. Operation Units and Logical Operations Unit operations are represented by models, or sets of equations, which include the mass and energy balance, equilibrium and kinetic relations, and specific unit operation functions. The interconnections between the units are represented by material and energy flows. The model equations require physical property data, e.g., density, enthalpy or volume. These properties are calculated by the property package. The property equations are solved iteratively each time a unit operation model is evaluated. This approach is used in almost all steady state and dynamic simulation systems [30]. Appendix II shows the procedure of manually solving of a simple flash block containing a ternary mixture and the comparison with the HYSYS solution for the same problem. Even such a simple system of equations, with only three pure components, takes a long time to solve manually, while HYSYS needs less than a second to obtain the solution, which is in good agreement with the manually calculated one (the average difference in the vapour and liquid composition is about 4.5%). The most complex operation units that HYSYS simulates are multi-stage mass transfer towers (columns) [17]. Columns consist of a series of equilibrium or non-equilibrium flash stages. For each feed stream, location, composition, flow rate and two property variables (T, P, S, H or vapour fraction) have to be known. To determine pressure and temperature drop along the column HYSYS uses simple linear interpolation between specified bottom and top values. The driving force for any distillation is a favorable vapour-liquid equilibrium. Reliable VLE relationships are essential for distillation column design and for most other operations involving liquid-vapour phase contacting. The Flash calculation within the column follows several steps: •  For a first stage, the entire component flow (liquid and vapour) and the enthalpy of the external feed are added to the components flows and the enthalpy of the internal streams entering the stage;  14  Chapter 2 — Process Simulator HYSYS Plant  •  HYSYS performs the Flash calculation of the combined mixtures based on the total enthalpy at the pressure of the stage. This calculation gives the conditions and composition of the vapour and liquid phases leaving the stage.  •  Further, the next stage is solved in the same way, until all stages are solved.  Trays of the column are considered ideal, if efficiency is not specified by the user. If specified by the user, even if the efficiency is one hundred percent, the trays are considered to be real [17]. Fractional efficiency less than unity is equivalent to by-passing of a part of the upgoing stream around the stage or the whole column. Calculations for other equipment, such as mixers, tees, coolers, heaters, etc. are based mostly on mass and energy balances, and are much easier to solve. In addition to the above mentioned units, HYSYS uses sets of several logical operations that enable better control and functioning of the whole flowsheet. In this project one adjuster and several recycles were used. An adjuster varies the value of one independent variable in a stream or operation, to meet the required value (specified by the user) in another stream or operation. Trial-and-error technique is used. Recycles are used whenever downstream material mixes with upstream material. The calculation around the recycle starts with the assumption of the unknown parameter. HYSYS then compares the assumed value in the stream to the calculated value of the opposite stream. If different, HYSYS generates a new assumption and repeats calculations until assumed and calculated values are close within the specified tolerance [17].  15  Chapter  3 - Scrubber  Section Simulation  Model  Chapter 3 - Scrubber Section Simulation Model 3.1. Introduction A steady-state computer simulation of the Scrubber Section of Syncrude Canada Ltd.'s Fluid Coker was developed, in order to predict effects of process and design variables changes on the Scrubber Section performance. The HYSYS.Plant Version 3.0.1 process simulator was used. Data from Syncrude Canada Ltd. was used to define the composition, flow rate, temperature and pressure of all inlet streams, as well as to provide parameters for all unit operation blocks. The Scrubber Section was simulated using a number of unit operation blocks and process streams. The operation blocks used to model the Scrubber were selected through consideration of the actual process, which is described in Section 1.3. The Koch Grid, which consisted often layers of structured packing, is modeled as a packed absorption column; the Shed section with six sets of sheds is modeled as a six-tray absorption column; the Scrubber Pool is modeled as a stirred tank; the space between the Shed section and the Scrubber Pool, where down flowing liquids from the Sheds and the rising product from the cyclones get in contact and are assumed to partially exchange mass and energy, is modeled as a flash block with two by-passes for the liquid and vapour fractions that do not reach the equilibrium. Mixers, splitters, coolers, pumps and adjusters are added to represent all stream and mass and heat transfer connections. When the whole system was set up, the HYSYS optimizer tool was used to determine unknown parameters in the system: the Koch Grid overall efficiency, the Sheds tray efficiency, and split ratios in the splitters around the flash block. These parameters were varied to minimize a suitable objective function, defined to quantify the extent of matching of model predictions with plant data. The set of parameters that minimized the deviation of predicted values from plant data was designated the "Base Case". Based on this "Base Case" different case studies were performed with the goal to investigate the effects of parameter and design changes on process performance.  16  Chapter 3 - Scrubber Section Simulation Model  3.2. Simulation Structure Set Up 3.2.1. Property Package As a starting point for the simulation, a property package was chosen. The Peng-Robinson equation of state, which is one of the usual choices for vapour-liquid equilibrium calculations for hydrocarbon systems, was adopted in this work.  3.2.2. Oil Characterization Hydrocarbon streams associated with the Scrubber are complex mixtures of huge numbers of components. The composition of these mixtures, especially the heavy fractions, is impossible to know since not all compounds are identified. Molecules can contain from 1 to more that 130 carbon atoms. Although HYSYS has a database for more than 1500 pure component properties, only hydrocarbons up to C30 are available [16]. In this simulation, only the light components, Ci to C 4 are characterized individually. A l l heavier fractions are characterized based on laboratory assays (boiling curves, densities and viscosities). Based on this input HYSYS forms "working curves" for TBP, molecular weight, density and viscosity. In  order to obtain discrete components  these fractions were divided into 20  pseudocomponents by "cutting" the assay distillation curve into 20 cuts (a higher number means higher accuracy, but also longer calculation time during the simulation runs). HYSYS automatically calculates NBP, molecular weight, density and viscosity of these components based on the correlations mentioned in Chapter 2.  3.2.3. Core Blocks and Simulation Components As described above, in order to simulate the Scrubber Section, it is broken down into four core operation blocks, a set of external and internal streams, and additional auxiliary units (splitters, mixers, adjusters, recycle streams). The core blocks are presented in Figure 3.1.  17  Chapter 3 - Scrubber Section Simulation Model Core operation blocks: 1. The K o c h Grid is modeled as a packed absorption column, with the same height (1.8 m), diameter (9 m) and the type o f packing (Koch Flexigrid Type 2) as in the plant. The number o f sections i n the 10-layer grid was set to be 2. This is the minimum number o f sections that allows for different pressures at the bottom and the top stage (to account for the pressure drop present in the plant). In counterflow packed columns, the vapour phase experiences a pressure drop due to the small free cross-section space and the presence o f liquid that decreases available space for the gas flow. A t the start o f run (SOR) o f the Syncrude Canada Ltd.'s Fluid Coker the pressure in the Scrubber was 117.21 kPa and pressure at the top o f the K o c h G r i d was 117.14 kPa. During operation, fouling o f the grid and the cyclone exit nozzles occurs due to the coke formation, causing an increase i n pressure drop. In order to maintain sufficient production o f the Overhead, pressure in the Scrubber was raised gradually by the operators. A t the end o f run ( E O R ) , it was typically 186.16 kPa, and the pressure at the top o f the K o c h G r i d was 185.53 kPa. For the Base Case, start-of-run ( S O R ) conditions were used, and pressure effects from S O R to E O R were simulated i n Case Study VIII, Chapter 6. 2. The Shed section consists o f six sets o f sheds, which are about 1-m wide and with 1.2-m horizontal spacing between them. Sloped from the both sides i n the direction o f liquid flow, and with serrated weirs, the sheds improve the distribution o f liquid that showers downwards. Gas passes through the same openings, contacting the liquid. H Y S Y S has several basic column templates which can be used, none o f which reflect the Shed section geometry exactly. The Shed section was therefore simulated as an absorption column o f six trays (one for each set o f sheds). Although there was a concern that a tray column can simulate the Sheds sufficiently well (in the Sheds, there is no bubbling o f gas through the liquid layer as in the tray column, and the contact efficiency is much smaller), this appeared to be the closest representation.  Additional parameters were specified to  represent the real column as close as possible. The number o f trays specified in the tray column corresponds to the number o f shed sets i n the Shed section. The dimensions o f the column and weirs were specified, as well. 3. The  Scrubber  Pool,  where  heavy  liquids from the  Scrubber  are  collected and  continuously mixed by Agitation steam, is modeled as a mixing tank. 18  Chapter 3 - Scrubber Section Simulation  4.  Model  The space between the Sheds and the Scrubber Pool is assumed to have significant exchange of mass and heat between the rising vapour and the down-flowing liquid. The vapour exits from the cyclone nozzles at a high velocity (76 m/s) and at a small angle to the horizontal. The tangential direction of snouts of the nozzles causes swirling of the exiting vapour, allowing it to spend enough time within this space to get in contact with the liquid. This space is represented as a flash block. In a flash block, H Y S Y S performs flash calculation, assuming vapour-liquid equilibrium. Since the vapour and the liquid may not reach equilibrium between the Sheds and the Scrubber Pool, by-passes for both streams are included in the model for this section, to account for the part of the streams that do not reach equilibrium. Two splitters around the flash block are used to divide the main vapour (rising vapour from the cyclones and the Scrubber Pool) and liquid stream (Shed Liquid ) into the fraction that goes directly to the flash block, and the part that bypasses it. In this way, a non-equilibrium stage of the process was accounted for.  Figure 3.1 Core blocks chosen to represent the Scrubber Section of the Fluid Coker 19  Chapter 3 - Scrubber Section Simulation Model  External streams, such as Cyclone Product, A T B feed, H G O Wash and Underwash, and Agitation Steam are input streams, while stream that goes back to the Coker and the Scrubber Overhead, as the main product, are the outlet streams. Internal streams leading to and from each unit represent rising vapours and falling liquids. Three splitters were incorporated into model. Beside two splitters around the flash block, mentioned above, an additional splitter was included into the flowsheet. The Scrubber Pool Liquid (Scrubber Bottom) splits to a Scrubber Pool Recycle (that is cooled and recycled back to the Scrubber Pool to keep the pool temperature around 375°C) and a stream that goes directly back to the Coker. Mixers are included when combining streams from different sources. A n adjuster is used to control the Scrubber Pool temperature (keep it constant at 375°C) by changing the SPL cooler duty. Recycles are included whenever downstream material mixes with upstream material.  3.2.4. Simulation Flowsheet The simulation flowsheet of the Scrubber Section is shown in Figure 3.2. As a primary feed, Cyclone Product from the top part of the Coker enters the Scrubber Section above the Scrubber Pool. H Y S Y S calculation suggests that this stream contains a small amount of liquid (3 wt.%). This is explained in Chapter 4. Cyclone Product is mixed with the vapour from the Scrubber Pool (Tank Vapour) in the space between the Scrubber Pool and the Sheds. This mixture is named Upgoing Stream. As mentioned in point 4 in Section 3.2.3, one part of this mixture (Upgoing Stream (to flash)), which accounts for the part that reach equilibrium with the falling liquid from the Sheds (Shed Liquid), enters the Flash Block. The small amount of liquid present in the Cyclone Product fraction of this mixture is removed and one hundred percent vapour mixture leaves the Flash Block and enters the bottom of the Sheds. The other part of the Upgoing Stream by-passes the Flash Block going directly to the bottom of the Sheds. This stream still contains liquid. It is mixed with the vapour product from the Flash Block and enters the bottom of the Sheds, as a stream named Vapour to Sheds. This stream, although called vapour stream, contains about 2 wt.% of liquid phase. This stream rises from the bottom of the 20  Chapter 3 - Scrubber Section Simulation Model Sheds through the six trays, contacting the falling liquids from the top of the Sheds - A T B feed (enters above the very top Shed set) and mixture of HGO Underwash (enters under the Grid) and Grid Liquid. The resulting Shed Vapour, leaves the top of the Sheds and enters the bottom of the Grid. The contact along the Koch Grid between the rising Shed Vapour and falling liquid HGO Wash that enters above the Koch Grid (top section of the Koch Grid in the simulation), results in the final vapour product of the Scrubber and the Fluid Coker, the Scrubber Overhead, which leaves the top of the Koch Grid. All liquids containing heavy components are collected in the Scrubber Pool. Agitation steam enters as a side stream to the Scrubber Pool. The Scrubber Pool Liquid exits the pool and splits into two fractions, one that goes back to the Fluid Coker, and the other, Scrubber Pool Liquid Recycle (SPL Recycle), which is cooled by Scrubber Pool Liquid Cooler (SPL Cooler) and recycled to the Scrubber Pool. A n adjuster is used to control the temperature of the SPL Recycle (by adjusting the cooling duty of the SPL Cooler) in order to keep the temperature of the Scrubber Pool at 375°C.  21  Chapter 3 —Scrubber Section Simulation Model  Scrubber Overhead  Scrubber Overhead  Temperature  383 2  C  Pressure  16.95  P«9  M a i s Flow  7.787«»O0S  kg/h  Actual Votoroe Flow  2.74S«*0Q5  mVh  MotocUarWaigM  70.77  Mass Density  2.834  kgAn3  GrtdUquM Temperature  395.0  C  Pressure  17.00  p**g  • 2.179e+O05  kgm  Mass Flow Actual Volume Flow  304.7  -  nvVTi  HGO (underwash) H G O (ondorwash) ShedVapour  Temperature  325.0  C  200.0  psig  S.247e*O04  kgm  66 26  m3m  Temperature  404.9  C  Pressure  Pressure  17 00  psig  Mass Flow  Mass Ftow  8 705o'005  kg/h  Actual Volume Ftow  Actual Votume Flow  2.773e«00S  m3m  ATB Temperature  325.0  Pressure  17.00  Actual Voluma Ftow  363.5  C  3.019tt*O05  Mass Ftow  kg*  Shed Liquid Temperature  473.4  c-  Pressure  17.00  ps<g  3.613o*OQS  kg*  Mass Flow Actual Voluma Ftow  Vapour to Sheds  490.5  Temperature  514.0  C  Pressure  17.00  psig  Mass Ftow  6.S96**005  kgm  Actual Volume Ftow  3.1SOe*«M  rrvVh  Cyctorw Product  L l q lids to  Temperature  539.9  C  Pressure  17.00  P»«g  Mass Ftow  S.470e*OOS  kg/ti  Actual Votume Ftow  2.846e*O05  raJ/H  So jbber '  "A  Light Ends  Agnation Steam AgfUfon Steam Temperature  185.0  C  Pressure .  150.0  psig  Mass Ftow  2,04U*004  kgm  Actual Voluma Ftow  23.47  •  m3fh  Scrubb.Pooi liquid'  Temperature  375.0  C  Pressure  17.00  psig  6 4370*005  kgm  Mass Ftow Actual Volume Ftow  789.0  Molecular Weight  6370 815 9  Mass Den iffy  0 5665  Viscosay  HVWl  kgAn3 CP  • Hot Scr.Recyde Actual Volume Flow Mass Flow'  459.2  mim  3.746fl*O05  kgm  Actual Volume Ftow  Figure 3.2 Simulationflowsheetof the Scrubber Section 22  Chapter 3 - Scrubber Section Simulation Model 3.2.5.  Input Plant Data  Plant data obtained from Syncrude Canada Ltd. was used as inputs i n the simulation system. Flow rates, temperatures, densities, viscosities and distillation data (or composition) o f the following streams were provided: - C y c l o n e Product - Cyclone Product is the product o f the coking zone o f the Fluid Coker and the vapour feed to the Scrubber Section. Mostly, this stream contains vapour phase when it exits from the cyclone nozzles, but small amounts o f liquid and even solids are also present. The composition o f this stream was defined i n the 1980's when the Coker was run i n "once through" mode (no recycles or additional input streams were used during the operation). It contained water, light ends, C G O (Coker Gas  O i l ) fraction and O T S B fraction (OTSB-Once Through Scrubber Bottom, a  mixture o f heavy fractions, some o f which boil above 1000°C). The characterization o f the Cyclone Product is explained in detail in [31]. The weight percents o f these four fractions, as w e l l as composition o f light ends, laboratory assays for C G O and O T S B , and the Cyclone Product T B P data, composition by boiling fractions, molecular weight and density distribution generated by H Y S Y S are shown in Appendix III. C G O is characterized using A S T M D2887 method, applicable for fractions up to 538°C, and for higher boiling components the H i g h Temperature Simulated Distillation ( H T S D ) enhancement, a method that extends A S T M D2887 to 760°C is used. For the O T S B assay, as an enhancement for A S T M D2887 method, the Supercritical Fluid Extraction method ( S C F E ) was used for fractions above 524°C [32]. - A T B - Atmospheric Topped Bitumen is a product o f atmospheric distillation o f bitumen, with 50 wt.% that boils above 560°C. The experimental assay is collected by Syncrude Canada Ltd. using A S T M D2887 method with H T S D enhancement for high boiling components. Distillation assay used as input and the T B P data, composition by boiling fractions, molecular weight and density distribution calculated by H Y S Y S are shown in Appendix III. - H G O - Heavy Gas O i l , which is one part o f the Overhead product after fractionation (343-524°C fraction), is recycled and serves to scrub heavy fractions and particulates from rising vapour in the Scrubber. It is injected both above ( H G O Wash) and below 23  Chapter 3 -Scrubber Section Simulation Model  (HGO Underwash) the Koch Grid. Experimental assay is also collected using ASTM D2887 method with HTSD enhancement. This assay and the TBP data, composition by boiling fractions, molecular weight and density distribution given by HYSYS are shown in Appendix III. -Agitation steam - 185°C saturated steam that serves to mix the Scrubber Bottom Liquid in the Scrubber Pool and keep particulates suspended. Note that flow rate and composition of the Scrubber Overhead, as a product stream, are not part of the input plant data; these are to be calculated in the simulation and values compared to the plant data. Input data is shown in Table 3.1. Table 3.1 Stream input data - information obtained from Syncrude Canada Ltd: Name  Flow  Boiling curve  Density at 15°C  Viscosity, cP  Temp.  kg/m  at 20°C/at 30(fC  °C  Pressure psig (kPa)  987 987 1024  1737/ 15 1737/ 15 214016/3  325 325 325  200(1380) 17(117.21)  540  17(117.21)  185  150(1035)  3  24 kbpd lOkbpd 55 kbpd  HGO Wash HGO Underwash ATB Overhead/ATB ratio Cyclone Product Water Light Ends CGO OTSB  b  2.51 kg/kg 150 kg/s 10wt% 12wt% 61 wt% 17 wt%  Agitation Steam Split SPL Rec.rTo Coker  0  A S T M D2887/HTSD A S T M D2887/HTSD A S T M D2887/HTSD  200(1380)  Known composition A S T M D2887/HTSD A S T M D2887/SCFE  6 kg/s 60%:40%  a) A S T M 2887-simulated distillation method applicable to all petroleum products boiling below 538°C. HTSD-High Temperature Simulated Distillation extends A S T M D2887 to 760°C boiling points. SCFE-Supercritical Fluid Extraction method, new method capable of analyzing high molecular weight residue fractions. b) 1 kbpd = 158.9 m /day 3  c) Volume flow ratio  HYSYS does not have an option for input HTSD or SCFE assay data, and the fractions above 538°C were inserted as ASTM D2778 data. Since ASTM D2778 method is applicable to petroleum fractions boiling below 538°C, HYSYS extrapolates the boiling point curve beyond 538°C. This extrapolation may cause inaccurate stream property estimation. An investigation on accuracy of the Scrubber Section model using this extrapolation method, given in [8], reports that the model matches the plant data within 10%, and the method was accepted. 24  Chapter 3 - Scrubber Section Simulation Model  Beside the stream characteristics, dimensions, position, types, temperatures and pressures were available for the following operation units: Koch Grid, Sheds, Cyclone exit tube snouts, SPL cooler. Data is presented in Table 3.2. Table 3.2 Input data and information for operation units obtained from Syncrude Canada Ltd. Unit  Parameter  Koch Grid  Sheds  Type of packing Number of layers Height, ft Diameter, ft Top pressure, psig Bottom pressure, psiq  Value Koch Flexigrid 2 10 6 30 16.99 (SOR)-26.91 (EOR)  Number of trays  17 (SOR)-27 (EOR) 6  Pressure, psig  17 (SOR)-27 (EOR)  Delta P, psi  0  Scmfober Pool  Delta P, psi Temperature, °C  0 375  Cvclone snout  Exit velocity of aas. ft/s  250  SPR Pumn  Delta P. psi  333  a)  SOR-start of ran; EOR-end of run  b)  1 psig = 6.8948 kPa; 1 ft = 0.3048 m  a,b  Based on these parameters the streams and operation units within the simulation flowsheet were defined. Several unknown parameters remained to be determined: •  Tray efficiencies in the Sheds  •  Section efficiencies in the Koch Grid  •  Split ratio in two splitters (for vapour and liquid) that surround the flash block. In order to determine the unknown parameters, the HYSYS optimizer tool was used.  Changing the values for these parameters, the best fit of the simulation model to the real plant data was obtained. This is explained in Section 3.3. The presence of liquid phase in vapours along the Scrubber Section and heavy fractions (above 524°C) in the Scrubber Overhead product are additional issues that had to be addressed. They are discussed in Chapters 4 and 5.  25  Chapter 3 - Scrubber Section Simulation Model  3.3. Optimizer Tool and the Base Case When the flowsheet was set up and a converged solution was obtained, the HYSYS optimizer tool was used to find the operating parameters that best match the plant data. HYSYS has several modes of Optimizer. The "Original Optimizer" was used in this simulation. The procedure is based on the "Complex" method of Box [33], the Downhill Simplex algorithm of Press et al. [34] and Box algorithm of Kuester and Mize [35]. The procedure can be found in [17]. In order to use the optimizer tool, primary variables (varied variables) and objective function had to be defined. Primary variables are values manipulated in order to minimize or maximize the objective function. The definition of the objective function is very important for obtaining a reliable simulation model. In this work, the objective function was defined based on the purpose of the simulation - to match the following important plant parameters: -Temperatures (Overhead temperature was taken to be the most important) -Flow rates (Scrubber Overhead and To Coker) -Scrubber Overhead composition (especially 524°C+ fractions). All these parameters were included in the objective function. Normalized values (normalized deviation of the parameters from the plant data), based on the following equation, were used:  X=ABS (1-Xmodel/Xplant)  (3.1)  where X represents the temperature, flow rate or mole fraction of 524°C+ components. Weight factors in the objective function were chosen based on the estimate of the importance of each parameter. Temperature of the Scrubber Overhead, as the main product of the Fluid Coker, and its composition (especially presence of heavy fractions) had to be matched the best. While for other parameters included in the objective function the weight factors were chosen to be 1, for these two parameters several options were tried (100, 50, 10 and 1). The optimizer tool was run for each case in order to find the optimal solution. Deviations of the 26  Chapter 3 - Scrubber Section Simulation Model  parameters listed in Table 3.3 were summed for each case, and the sums were compared. Weight factors 100 and 50 when applied to temperature and fraction boiling over 524°C of the Scrubber Overhead gave too high deviations for all other parameters, while a weight factor of unity for all parameters could not match any of the plant data very well. The sums of deviations for all these three options (weight factors 100, 50 and 1) were much higher compared to the case where weight factors were chosen to be 10 for the above two key parameters. As a result, the following objective function was finally defined:  OF = lO-T(Ovhd) + T(Grid Bottom) + T(Shed Top) + T(Shed Bottom) + (3.2)  +Flow Ratio(OvhoVATB) + SPL Flow(To Coker) + 10-Mol. Fraction 524°C (Ovhd) +  where all parameters are normalized functions as shown by Equation (3.1). In the present study, the primary variables were unknown process parameters mentioned in Section 3.2.4: Shed trays efficiency, Grid sections efficiency and split ratio in the two splitters around the flash block. Values for the above parameters were to be determined in order to obtain the process model that matches the plant data as closely as possible. Different efficiencies for each of six trays in the Sheds and two sections in the Koch Grid, together with split ratios in two splitters resulted in ten primary variables. This large number of varied variables resulted in excessive running time of the optimizer without obtaining the optimum solution. The number of varied variables was decreased to four by assuming that all six trays in the Sheds have the same efficiency, and both sections in the Koch Grid have the same efficiency. The remaining four parameter values were changed simultaneously, from 0 to 1. When the optimum set of parameters and good matching with the plant data was obtained, values for split ratios in the two splitters and the Sheds trays efficiency were fixed, and efficiencies for the two Koch Grid sections were assumed different and changed simultaneously from 0 to 1. With efficiency of 0.55 for the top section and 1 for the bottom section, the optimizer tool yielded an even smaller value for the objective function, and a better match with the plant data. Hence, the overall efficiency of the Koch Grid was taken to be the average of the two, 0.78. Further, obtained efficiencies for the Koch Grid sections (0.55 and 1) were fixed along with the two split ratios, and efficiencies for the six trays of the Sheds were 27  Chapter 3 - Scrubber Section Simulation  Model  varied simultaneously. Since the optimizer tool could not find a lower value for the objective function, the previous solution with the same efficiencies for the six trays was accepted. Additionally, as will be explained in Chapter 4, the efficiencies for the heaviest components had to be decreased (to 10" ) to match the Scrubber Overhead content of heavy fractions (524°C+) to 10  plant data. In this way the Base Case was designated. The parameter values and deviations from the plant data are shown in Table 3.3, while values for primary variables (unknown parameters) determined by using the optimizer tool are in Table 3.4. The values for primary variables were fixed for further simulation. The Base Case model matches the plant data well, with the average parameter deviation from the plant data of 1.4% and the highest deviation in Overhead 524°C+ fraction of 3.2%. This Base Case was further used as a starting point for all case studies. Table 3.3 Base Case parameter values and deviation from the plant data Type of Parameter  Plant  Model  Dev.(%)  390  393  395 407 470  395 405 473  0.8 0.0 0.5 0.6  375 2.50  375 2.57  0.0 2.8  Act. Vol. Flow(To Coker) (kbarrel/day)  50  49  2.0  Act. Vol. Flow( SPL Rec) (kbarrel/day) Overhead 524°C- mole fraction S u m of deviations:  74 0.94  73 0.91  1.4 3.2 11.3  T(Ovhd) (°C) T(Grid Bottom) (°C) T(Shed Top) (°C) T(Shed Bottom) (°C) T(Scrubber Pool) (°C) Ratio Overhead/ATB Mass Flows  In Table 3.3 all model values are results of HYSYS calculation, except the Scrubber Pool temperature, which is adjusted to be 375°C. It should be noted that the objective function is not unique - other terms could be included and different weight factors incorporated. The justification of this objective function is that it captures key features of the Scrubber system operation, and with the chosen weight factors gives acceptable deviations from the plant data. Such a low deviation from the plant data gives more confidence in using the defined process model for different case studies, and higher reliability of obtained results of the case studies.  28  Chapter 3 - Scrubber Section Simulation Model  Table 3.4 Determined unknown parameters (primary variables) Primary Variables Shed Trays Efficiency Koch Grid Overall Efficiency Splitter TEE-102 (Upgoing Vapours): Vol. flow fraction to Flash Block Splitter TEE-103 (Scrubber Liquid): Vol. flow fraction to Flash Block  Value 0.53 0.78 a  a  0.41 0.58  a) In addition to these efficiencies, very low efficiencies for heavy components (524°C+ fraction) in the Sheds and the Grid were applied.  29  Chapter 4 — Presence of Liquid Phase in the Vapour Streams  Chapter 4 - Presence of Liquid Phase in the Vapour Streams 4.1. Introduction When the Cyclone Product assay (as given in Appendix III) was used for the Cyclone Product stream definition for the simulation, previous HYSYS calculations [8] suggested that under the given conditions (pressure, temperature and stream composition) this vapour stream contained a small amount of liquid phase. This was confirmed also in this work. As a consequence, almost all "vapour" streams in the current simulation contain some amount of liquid phase: Cyclone Product (3 wt.%), vapour that goes to the Sheds (2 wt.%), vapour from the Sheds to the Grid (19 wt.%) and even Scrubber Overhead (8 wt.%). The reason for this is the presence of heavy components in the Cyclone Product. This stream which is generated in the Fluid Coker contains very heavy fractions, some of which boil above 1000°C. It contains some liquids and even solids. The Cyclone Product passes through six cyclones in parallel where most of these liquids and solids are removed. The Cyclone Product is expected to be vapour under the given conditions, but as HYSYS simulation and Syncrude Canada Ltd.'s sources suggest [13, 14], there is some liquid phase present. Note that HYSYS calculations are based only on the provided assay data and equilibrium calculations. HYSYS has no ability to recognize liquid entrainment. An investigation of the fouling process within the cyclones in Syncrude Canada Ltd.'s Fluid Coker (Nelms [14]) suggested that approximately 40% of the source of the foulant is from the entrainment of the liquid (other 60% is from condensation of vapour). This entrainment occurs because some liquid droplets of the feed sprayed into the fluidized bed of coke particles do not come in contact with coke particles and are carried upwards with the vapour. Some of the droplets fall back to the bed, but the rest are carried all the way to the cyclones. Although most of the liquid and solid particles should be removed by the cyclones, some remain within the cyclone as foulant, and some are carried out of the cyclone with the vapour jet. This stream is the Cyclone Product. For the HYSYS simulation it was necessary to determine whether droplets that are carried with the Cyclone Product vapour jet reach the bottom of the Sheds and enter the Sheds with the rising vapour, or fell down into the Scrubber Pool. Both cases depend on the liquid droplets 30  Chapter 4 -Presence  of Liquid Phase in the Vapour Streams  trajectory when exit the cyclone nozzle, as well as possibility that these droplets are being washed by the falling liquid from the Sheds. In the case when all liquid droplets reach the Sheds, the model would allow the original Cyclone Product with the fraction of liquid phase to reach the Sheds and join the up-going stream along the Scrubber. Since the liquid phase contains heavy components, this would affect the composition and properties of the Overhead product. In the case when all liquid droplets end up in the Scrubber Pool, a flash block would be required to separate heavy liquid components from the rising Cyclone Product vapour and send them to the Scrubber Pool. The third option is that only one part of the Cyclone Product containing liquid phase enters the Shed section, and the other part enters the flash block.  4.2. Droplet Size Estimation In the plant cyclones, coke solids from the bed and extra scouring coke enter the cyclones along with droplets. The discussion below ignores these solids and considers only vapour and liquid. In order to determine the behavior of the liquid droplets, the size of the droplets that cannot be removed by the cyclones and are carried upwards was calculated first, based on the equation for cyclone cut point given in [36]:  (4.1)  where D h is theoretical particle size removed by the cyclone, p is viscosity of the gas pt  g  (Cyclone Product vapour phase), B = D /4 where D is the cyclone diameter. N is effective c  c  c  s  number of spiral paths taken by the gas within the cyclone, determined graphically based on the exit velocity from the cyclone, vo=76 m/s. Vj=74 m/s is the velocity at the inlet of the cyclone, n  calculated based on the volume flow rate (80.6 m /s) and the cyclone inlet cross-section (d=0.48 3  m), p is the density of the Cyclone Product liquid phase and p is the density of the Cyclone p  g  Product gas phase. All values used in Equation (4.1) are listed in Table 4.1.  31  Chapter 4 -Presence  of Liquid Phase in the Vapour Streams  Table 4.1 Parameter values used in Equation (4.1) Value  Parameter  3.1 M O ' kg/ms  N for vo=76 m/s  5.8  1.702 m  PP  759.32 kg/m  Parameter a  5  D  c  B = D /4 c  0.43  c  Value  s  Pg  3  3  1.8298 kg/m  a  3  a) Note: These values are calculated by HYSYS  As stated above, Equation (4.1) ignores the effects of bed coke and of scouring coke which is added to the vapours upstream of the cyclone entrance. Based on this calculation appears that liquid droplets that are smaller than 11 um are carried with the vapour jet, if they have not been impacted by scouring coke. Larger droplets are removed by the cyclone.  4.3. Trajectory of the Liquid Droplets The trajectory of the liquid droplets carried with the Cyclone Product jet was then considered. Although the cyclone snouts are positioned at a small angle to the horizontal, the stream can be considered as a nearly horizontal jet above the Scrubber Pool. The surrounding vapour velocity (9.8 m) was estimated based on the total volume flow rate of the Scrubber Pool vapour and the Cyclone Product, and cross-section area of the column. Longitudinal distribution of velocity, v/, for the droplets was calculated based on equations for a turbulent free jet, given in [37]. A turbulent jet is a free jet with the Reynolds number greater than 2000. In the case of the Cyclone Product, the Reynolds number is calculated to be 24-10 . The equation is applicable for the air jet into the surrounding air. Density gradient 5  between the jet fluid and surrounding fluid has effect on the spread of the jet. Since Cyclone Product vapour, as jet fluid, has similar density as the surrounding vapour, as in the original case with  air, this  v,=v -K n  —  equation for  was used without change  1 < — < 100  or  4.06 <X<  for  the  system under  study:  58m  (4.2) K  = 6.2  for  v =10 to 50 mis 0  32  Chapter 4 -Presence  of Liquid Phase in the Vapour Streams  In this equation vo is the exit velocity of the jet, 76 m/s, X is the horizontal distance from the exit of the nozzle and Do is the nozzle diameter, 0.58 m. Equation (4.2) is applicable for the distances from the nozzle 7 < X/Do < 100, which is in the present case 4.06 < X < 58 m from the nozzle. However, the Scrubber has a diameter of 9 m and the nozzle snout is close to the wall. From the exit of the nozzle up to 4 m, a linear change of velocity was assumed. After equation derivation, integration in time and values input, the following two equations for the horizontal distance change in time were obtained:  X(0 = ^ - ( l - e "  X(t)  2 1 4  ')  = y/16.5 + 547.3(f  - 0.0665)  for distance 0 < X < 4 m  (4.3)  for distance 4 <X < 58m  (4.4)  Detailed equation derivation is shown in Appendix IV. Vertical distribution of velocity can be calculated from the vertical force balance (weight of the droplet against drag force) and the terminal velocity of the droplet:  mg-F =m— D  (4.5)  dt  In this equation g is gravitational acceleration, m is mass of the droplet, and F is drag force: D  (4.6)  F =C ^p v X^d ) 2  D  D  g  p  Velocity v=v - v is the slip velocity between the droplet (particle) and surrounding gas, d p  g  p  is the droplet (particle) diameter, and Co is drag coefficient. For the spherical particles it can be calculated from: 33  Chapter 4 -Presence  C  D  =  24  =  —  Re  of Liquid Phase in the Vapour Streams  24« ^ —  (4.7)  p -v-d g  p  Terminal velocity can be calculated from Equations (4.5), (4.6) and (4.7) when— = 0. dt  In this investigation, the surrounding gas was considered to be the vapour that originates from the Scrubber Pool in combination with the Cyclone Product vapour. Properties of this combined vapour were used for the calculations. If the droplet of 11 um diameter was moving through a stagnant gas, the terminal velocity would be VTO=0.0016 m/s for the system under the study. From Equations (4.5), (4.6) and (4.7), integrating the velocity in time, the vertical distance from the nozzle would be (Y is set up to be directed downwards):  y (0  = v  0  r  o  . f +  ^  g  exp(—^--0-1  (4.8)  v.ro  However, droplets are not moving through a stagnant gas. The surrounding gas, as mentioned above, a combination of the Cyclone Product vapour and the Scrubber Pool vapour, is moving upward. The velocity of these vapours is calculated based on the total volume flow and the scrubber cross-section, and its value is 9.8 m/s. This velocity is included in Equations (4.5), (4.6) and (4.7) through the slip velocity. The terminal velocity is also affected by the velocity of the surrounding gas, and is not same as the terminal velocity in the stagnant gas:  V  T =  V  T O - V  (4.9)  S  where v is the terminal velocity in the flowing surrounding gas, v = 9.8 m/s is the velocity of T  g  the surrounding gas and VTO is the terminal velocity in the stagnant gas.  34  Chapter 4 - Presence of Liquid Phase in the Vapour Streams  From Equations (4.5), (4.6) and (4.7), and taking slip velocity into account, the vertical distance from the nozzle as a function of time would be:  Y(t) = Y (t)-v -t 0  (4.10)  g  Detailed derivation is presented in Appendix IV. Integrating both horizontal and vertical velocity in time, the horizontal and vertical distances were obtained. Simply inserting the time in Equations (4.3), (4.4) and (4.10), the trajectory of the droplets within the space above the Scrubber Pool has been estimated. The calculation has been done for the largest droplet diameter present in the jet (11-10" m), assuming that all others would be carried even further. The 6  trajectory is presented in Figure 4.1.  0  1  2  3  4  5  6  7  8  9o  10  CO  Horizontal distance from the nozzle- x, m  Figure 4.1 Trajectory of a liquid droplet carried with the Cyclone Product jet Maximum possible distance of the nozzle snout from the Scrubber wall is presented. The actual distance is much smaller (data was not available)  As mentioned above, in this calculation horizontal injection has been assumed, although the snouts are positioned at an angle. Also, because of the tangential direction, the six snouts affect  35  Chapter 4 -Presence  of Liquid Phase in the Vapour Streams  each other's jet trajectory, resulting in probably more spiral and upward directed moving of the vapour jet. This calculation suggests that droplets not removed in the cyclone are carried up with the vapour and reach the Sheds. Some portion is probably washed down by the liquid falling from the Sheds or hit the wall and condenses. Based on these considerations, the simulation structure of the Scrubber Section was set up to allow one part of the liquid phase of the Cyclone Product to reach the Sheds directly via a bypass. The remainder is passed through a flash block to account for the fraction that is being washed down by the Shed liquid or hit the wall and falls back to the Scrubber Pool.  36  Chapter 5 -Presence of Heavy Components in the Scrubber Overhead  Chapter 5 - Presence of Heavy Components in the Scrubber Overhead 5.1. Introduction Syncrude Canada Ltd.'s data (Appendix III) shows that Scrubber Overhead contains some fractions boiling over 700°C. In the initial version of the simulation model, such heavy fractions did not appear in the calculated Scrubber Overhead. The maximum NBP was around 540°C, because of equilibrium conditions and tray efficiency. The presence of heavy components in the plant scrubber overhead was therefore attributed to non-equilibrium conditions. Several options were considered for change to the simulation model in order to simulate non-equilibrium conditions and account for high boiling fractions in the Overhead: 1. By-passing liquid from the Sheds directly to the Overhead; 2. By-passing liquid from the Sheds to the Grid bottom; 3. Decreasing the Shed tray efficiency of the 524°C+ components; 4. Decreasing the Koch Grid section efficiency of the 524°C+ components; 5. Decreasing both the Shed tray and Koch Grid sections efficiency of 524°C+ components. The optimizer tool was used with the first two options to attempt to match the plant data. In addition to the varied parameters used previously, the by-pass fraction of the Shed Liquid was used. For the last three options, all parameters were left unchanged, except that the 524°C+ components efficiency was decreased from the average tray efficiency of 0.53 in the Sheds and/or 0.78 in the Grid to essentially zero (10" ). For all these cases the sum of deviations from 10  plant parameters were compared. Options 1 and 5 showed the lowest deviations from the plant data, while all others could not match the plant data, especially the Overhead composition, well enough. Between the two successful options, Option 5 with the low component efficiency in both the Sheds and the Koch Grid was chosen for implementation in the simulation, because it showed better match to the plant data. 37  Chapter 5 -Presence  of Heavy Components in the Scrubber  Overhead  An investigation of the possible real physical cause for the presence of heavy components in the overhead was undertaken to justify the hypothesis on low heavy components efficiency. The methods for estimation of conditions in the Sheds and the Koch Grid are shown in the following sections. The investigation suggests that possible causes for appearance of heavy components in the Overhead product is liquid entrainment that occurs in the Sheds and the very high gas flow rate in the packed section (Koch Grid), which also causes some liquid entrainment and excessive pressure drop. Both conditions lead to decreased column efficiency. The most affected are the heavy boiling fractions because of their low volatility under the given conditions. Remaining in the form of liquid, they are the species that can be entrained in the vapour phase.  5.2. Liquid Entrainment in the Shed Section The Shed section and the liquid and gas distribution are described briefly in Section 3.2.3. The description suggests that the Shed section may be considered as a counterflow plate column (in counterflow plate columns, liquid and gas utilize the same openings for flow), similar to a baffle plate column [38]. In plate columns, when a gas passes through a liquid, it generates fine liquid droplets. In cross-flow plate columns (sieve, valve or bubble plates) that happens even at very low gas loads (bubble regime) due to the bursting of bubbles [29]. In counterflow columns (shed or baffle column), high gas flow rate can also generate liquid droplets. If the terminal velocity of the droplets is lower than the gas velocity, they will be entrained in the gas stream. Under very high gas load and velocity, even large droplets can be thrown upwards. Some of the droplets (larger ones) fall back to the liquid stream, but the smaller ones can reach the upper tray. In this way liquid of lower volatility may reach a tray with higher volatility liquid and even some very heavy components could possibly reach the overhead products. When a column runs under a very high gas loading and low liquid loading, the conditions could result in excessive liquid entrainment in the vapour and the downward flow of liquid is destroyed. This is referred to as "entrainment flooding". Entrainment can cause lowered tray efficiency and even different component efficiency [38]. In the Shed column, the difference in the gas and liquid volume flow rate is of several orders of magnitude (vapour volume flow rates measure in hundred thousands of m /h, and liquid 3  volume flow rates in hundreds of m /h), suggesting that extensive liquid entrainment may be 3  38  Chapter 5 -Presence  of Heavy Components in the Scrubber  Overhead  present. To assess this, an estimation of entrainment conditions was done based on the chart given in Figure 5.1 [38]. The abscissa term L/G(P /P[)  05  g  is called the flow parameter and the  ordinate term CSB is called the capacity parameter.  Figure 5.1 Flooding correlation for columns with cross-flow plates [38] All the values for the liquid and the vapour entering the Shed column are calculated by HYSYS and given below, in Table 5.1. In this table U N F is the gas velocity through the net area (for counterflow plates, net area is the same as column area). This velocity is calculated based on actual volumetric flow of the gas (3.110 m/h) and cross section area of the column (63.58 m ). PL is the liquid density, pG is the gas density, cr is the liquid surface tension, L is the liquid loading (mass flow sum of liquid that comes from the Koch Grid, HGO Underwash and ATB feed) and G is the gas loading (Vapour to Sheds).  39  Chapter 5 -Presence  of Heavy Components in the Scrubber  Overhead  Table 5.1 Parameter values for calculation the flow and capacity parameter for Figure 5.1 Parameter  Parameter  Value  UNF  1.35 m/s  PL  706.1 kg/m  PG  2.64 kg/m  3  3  Value  a  13.895 dyn/cm  L  158 kg/s  G  183 kg/s  The calculated value for the flow parameter F L G is 0.052, and for the capacity parameter, CSB is 0.087. Figure 5.1 shows that for the flow parameter 0.052, and tray spacing 762 mm, which is the case in the Shed column, entrainment flooding would occur at the CSB value of-0.12. The calculated value of 0.087 shows that the Shed column is not within the flooding regime, but at 72% of flooding conditions. This suggests that at the given conditions within the Shed column, a significant entrainment could be present. This entrainment contains not only the liquid carried by Cyclone Product vapour, but also some additional liquid from the Shed column itself. Column efficiency is decreased, especially for the heavy fractions because of their low volatility, possibly allowing them to reach the top of the column.  5.3. Packed Section In packed columns, the vapour-liquid contacting takes place in continuous beds of solid packing elements rather than on discrete individual plates. The vapour enters the column below the bottom bed and flows upward through the column. The liquid enters at the top through the liquid distributor and flows downward through the packing counter-currently to the rising vapour. Packed beds may be divided into two categories: Those containing packing elements that are placed in the column in a random arrangement, usually by dumping; and those containing carefully installed elements designed specifically to fit the column dimensions. The former elements are called random, or dumped, packing. The latter are called ordered, or structured, packing. In Syncrude Canada Ltd.'s Scrubber, Koch Flexigrid Type 2 structured packing is used. It is designed to have maximum capacity, limited liquid holdup and minimum pressure drop.  40  Chapter 5 -Presence  of Heavy Components in the Scrubber  Overhead  A countercurrent flow of the gas and liquid phases over a high surface area packing should provide highly efficient mass transfer and separation. However, the efficiency decreases if the liquid flow is not uniform through the bed. Ideally, all the surfaces should be wetted by the liquid and liquid flow through the bed should be uniform [38]. Within a packed column, if a gas flows counter-currently to liquid flow, pressure drop is a consequence of flow through the series of small openings in the bed. For low liquid flows pressure drop is proportional approximately to the square of the gas velocity [38]. As the gas flow rate increases, the liquid is only partially enabled to flow downwards, and tends to remain trapped in the void space of the packing. Consequently, space available for the gas flow is reduced, causing increase in pressure drop. Further increase in gas flow rate may lead to a point when the liquid cannot flow any more. This situation is called flooding and is analogous to entrainment flooding in a plate column. At this point pressure drop radically increases with a small change in gas flow rate. Flooding conditions affect the mass-transfer efficiency of the column [39]. In order to justify the hypothesis of low Koch Grid efficiency for heavy components, given in Section 5.1, conditions within the Koch Grid were determined. One of the indicators for the flooding conditions, or near-flooding conditions, is increased pressure drop within the column. In the present study, two methods were used to calculate the design pressure drop for the type of packing used in the Koch Grid at the present conditions. The results were compared with the pressure drop obtained by using generalized flooding-pressure drop correlation of Eckert and Leva, modified by Strigle [38] and the pressure drop in the real plant at the start of run (before any fouling has occurred ), which was 0.99 mbar. Two design methods for calculating the pressure drop within the Koch Flexigrid Type 2 structured packing are, as follows: 1)  Koch Glitsch chart for the Koch Flexigrid Type 2 structured packing, shown in Figure  5.2: Based on superficial factor for gas, F , and liquid loading, a design pressure drop can be s  determined [40]:  41  Chapter 5 -Presence of Heavy Components in the Scrubber Overhead  F , m/s • (kg/m ) 3  1  s  0.5 0.6  I  FLEXIORID 2  _  1  0,8  liquid U M J I J  JbpCwieBosae gpm,« m V h 171 14? sa 122 & sa 73 J  O.S  0:6 0;5 0.4  m 24  10 s 0  03  12  tt  S ( S « m Aif-Walaf, AiflUwil Tower: M B a m a W  0.2 <1  0.1 — ^ ;  0.08 < 0.06 0.05 0.04 0.03 • 0.02 0.4 0 5 0  A  I  0.8  —2  F . ft/s • (lb/ft )'* 3  s  Figure 5.2 Design pressure drop chart for Koch Flexigrid Type 2 structured packing [40] For the given parameters for gas and liquid phase entering the Koch Grid: gas density (Shed Vapour), po = 3.14 kg/m and gas velocity through the net area (Shed Vapour), UG= 1.17 m/s the 3  superficial factor is calculated to be F =2.07 m/s-(kg/m ) ( vertical dashed line on the chart). 3  0,5  s  This value along with the liquid loading: 150m //? — - = 2.28 3  L = Vol. Flow rate/Cross sec. area = P  65.58m'  m /h — (steep dashed line on the chart) 3  m  gives on the chart a very low value for the pressure drop of around 0.12 mbar/m ( arrow on the chart) what is 0.22 mbar for the total height of the column (1.8 m). This value is, actually, out of 42  Chapter 5 - Presence of Heavy Components in the Scrubber Overhead  the range of the chart. For the given conditions the design pressure drop is estimated to be 0.22 mbar, what is much smaller than the plant value of 0.99 mbar. 2)  A similar result is obtained using KG-Tower, software offered by Koch Glitsch [40],  which can calculate the design pressure drop for the given conditions and specific packing type. Pressure drop calculated with this software was 0.157 mbar/m, which is 0.286 mbar along the column. The result is shown in Table 5.2. There is a warning from the software that given parameters are out of the range of applicability of the method. Table 5.2 Packed tower rating data calculated by Koch-Glitsch KG-Tower software [40]  ZONE DESCRIPTION BED NUMBER  1 Packed Colum  LOADINGS Vapor Rate Scale Factor Vapor Rate Vapor Density Vapor Volume Vapor Viscosity  kg/hr kg/m3 m3/s CP  1.00 870500 3.1400 77.01 0.0010  Liquid Rate Scale Factor Liquid Rate Liquid Density Liquid Volume Surface Tension Liquid Viscosity  kg/hr kg/m3 m3/h dyne/cm CP  1.00 126100 840.00 150.12 14.69 0.565 1.00  System Factor Packing Type  FLEXIGRID® 2 SS METAL mm m2  9141.00 65.63  Liquid Loading  rn/s(kg/m3) U5 m/s m3/h/m2  2.08 0.07 2.29  Pressure Drop  mbar/m  <0.5  %  34  Tower Diameter Tower Area Fs Cv  /  Calculated Capacity Constant L/V WARNINGS: WARNINGS:  O.fiJ-  1, 1. The capacity rating is extrapolated for given physical properties. Please contact KOCH-GLITSCH.,  43  Chapter 5 - Presence of Heavy Components in the Scrubber Overhead  As was mentioned above, flooding conditions will affect the pressure drop in the column. The generalized flooding-pressure drop correlation of Eckert and Leva, modified by Strigle [38], shown in Figure 5.3, enables prediction of pressure drop and flooding conditions in packed columns based on liquid and gas .loading and characteristics, as well as packing parameters.  Figure 5.3 Generalized flooding-pressure drop correlation of Eckert and Leva, modified by Strigle [38] The ordinate term is the capacity parameter:  i0.50  PG  /r°V  005  (5.1)  PG  44  Chapter 5 -Presence  of Heavy Components in the Scrubber Overhead  For the gas and liquid entering the Koch Grid, with the parameters given in Table 5.3 and packing factor for Koch Flexigrid Type 2 structured packing F = 4 ft" , the calculated value for 1  p  the capacity parameter is Cs= 0.52.  Table 5.3 Parameter values for calculation the flow and capacity parameter for Figure 5.3 Parameter  Value  Parameter  Value  u  1.17 m/s  V  16.53 cS  L  35 kg/s  G  241 kg/s  t  PL  840.6 kg/m  PG  3.14 kg/m  3  3  In Table 5.3 U is Shed Vapour superficial velocity through the net area, po is Shed Vapour t  density, p is liquid density (HGO Wash), v is kinematic viscosity of the liquid, L is liquid L  loading (mass flow sum of liquid that comes from the Koch Grid and ATB feed) and G is gas loading. The abscissa term is the same flow parameter used for plate columns, but applied for liquid and gas entering the Koch Grid. Using liquid and gas loading from Table 5.3, the calculated value for the flow parameter is  FLG=  0.0088. From the values for capacity and flow parameter  and Figure 5.3, the pressure drop in the packed column is determined to be around 0.06 in H 0/ft 2  or 0.4 in H 2 O (or 0.99 mbar) for the total length of the column. This pressure drop corresponds to the pressure drop in the plant (0.99 mbar). Since, based on the generalized flooding-pressure drop correlation of Eckert and Leva, pressure drop of 1.5 in H20/ft represents the flooding condition, from Figure 5.3 it can be concluded that the Koch Grid column is not close to this range. However, the difference between first two calculations (Koch Glitsch for the design purpose) and the last one shows that very high flow rate of the gas definitely has an effect on the pressure drop along the Koch Grid and must be taken into account. Consideration at the beginning of the Section 5.2 suggests that high volume flow rate of the gas compared to the liquid, causes high liquid holdup and pressure drop, 45  Chapter 5 -Presence  of Heavy Components in the Scrubber  Overhead  and the efficiency of the column may be lowered. The above presented calculations for the pressure drop showed that the real pressure drop along the Koch Grid is higher than the design value, suggesting that liquid holdup and hence partial entrainment may be present in the Koch Grid (bubbling of the gas through the liquid may cause some liquid entrainment, lowering the efficiency for heavy fractions). Based on this conclusion, the hypothesis for low efficiency for heavy fractions was accepted with more confidence.  5.4. Conclusion Calculations based on conditions in both columns indicate that column efficiencies may be lowered, either due to the liquid entrainment within the Shed column, or due to the very high gas loading and increased liquid holdup within the Koch Grid. Since simply decreasing overall column efficiencies could not provide satisfactory matching with the plant data, the option with decreased heavy component efficiency in both Shed and Koch Grid column was applied. As mentioned before, the efficiency for heavy fractions (524°C+, 40 out of 120 components) was radically decreased (to 10" ), which resulted in the presence of high boiling fractions in the 10  Overhead. The Base Case matched the plant data sufficiently well to perform different case studies, which are presented in the following chapter.  46  Chapter 6 - Case Studies: Results and  Discussion  Chapter 6 - Case Studies: Results and Discussion 6.1. Introduction Once the Base Case is set up, different design and parameter changes can be applied and their effect on process performance investigated. Eleven case studies have been performed where following parameters have been changed:  I.  ATB Flow Rate - gradual change of ATB actual volume flow rate from original 55 kbarrel/day to 80 kbarrel/day.  II.  HGO Wash Flow Rate - change of HGO Wash flow rate from 24 kbarrel/day in the Base Case to 30 and 40 kbarrel/day.  III.  HGO Underwash Flow Rate - change of HGO Underwash flow rate from 0 to 10 (Base Case) and 20 kbarrel/day.  IV.  HGO Wash Temperature - change of HGO Wash temperature from 250 to 350°C (Base Case 325°C).  V.  HGO Underwash In and Out of Service - investigates the effect of Grid Underwash function on Scrubber parameters. Considers four options: 1. HGO Underwash is in service - flow rate of HGO Underwash is 10 kbarrel/day; 2. HGO Underwash is out of service and Overhead temperature is not controlled. 3. HGO Underwash is out of service and Overhead temperature is controlled by HGO Wash flow rate. 4. HGO Underwash is out of service and Overhead temperature is controlled by ATB feed flow rate.  VI.  Number of Trays in the Sheds - change the number from 2 to 10 (Base Case - 6 trays).  47  Chapter 6 - Case Studies: Results and Discussion  VII.  Number of Grid Sections - change the number from 2 to 10 (Base Case - 2  sections). VIII. Simulation of Conditions from Start of Run to End of Run - investigates effect of condition changes from Start of Run to End of Run of the Fluid Coker. IX.  Water Instead of HGO Underwash (T=30-40°C) - how much flow is required to decrease Grid entrance temperature by 10-20°C (keep the temperature of the Overhead and the Scrubber Pool the same).  X.  Saturated Steam Instead of HGO Underwash) - how much flow is required to drop Grid entrance temperature by 10-20°C (keep the temperature of the Overhead and the Scrubber Pool the same).  XI.  Recycle Cut Point Changes - drop Recycle cut point (RCP) by 15, 30 and 45°C on the CGO (Overhead) SimDist: Related to the 95% cut point. Options to drop RCP: • Increase fresh ATB flow rate; • Increase top of Grid Wash flow rate; • Increase bottom of Grid Wash flow rate.  Simulation Output For each case, actual volume and mass flow rates, and densities of all vapour and liquid streams in the Scrubber, the temperature profile up the Scrubber from the Scrubber Pool to the Overhead stream have been determined. Also, Scrubber Overhead and Bottom properties, including composition, average molecular weight, fraction distribution and SimDist curve were calculated. Some additional information is also included in the results. For each case study, all results are organized in tables and charts in the same way. Numbering of the tables and charts is adapted to this organization. Roman numerals represent the numeral of the case study, and Arabic numerals the number of the table or figures. Following is a list of tables and figures that can be found for every case:  48  Chapter 6 - Case Studies: Results and Discussion  Table 1: Process parameters changes-effect of particular parameter change (ATB flow rate, HGO Wash flow rate...); Tables 2 and 3: Scrubber Overhead and Scrubber Bottom properties, TBP distillation temperatures and fraction distribution data for some specific cases; Figures 1 and 2: Temperature change and profile along the Scrubber-effect of parameters change; Figures 3 and 4: Mass flow rate changes for process streams; Figures 5 and 6 Scrubber Overhead and Scrubber Pool Liquid distillation curves; Figures 7 and 8: Scrubber Overhead and Scrubber Bottom composition comparison (mole fractions). Every case consists of main observations of the changes as a consequence of parameter change, followed by discussion.  6.2. Case Studies I. ATB Flow Rate Atmospheric Topped Bitumen (ATB) flow rate has been changed from original 55 kbarrel/day to 80 kbarrel/day, with a step of 5 kbarrel/day, and effect on Scrubber parameters, Scrubber Overhead and Scrubber Pool Liquid flow rates, composition and properties have been investigated. In this case, the Scrubber Pool temperature was not kept constant, in order to see the effect of ATB flow rate change. Instead, the cooling duty of SPL cooler was kept the same. Observations: By increasing ATB flow rate from 55 to 80 kbarrel/day: • Temperature profile:  - A l l temperatures along the Scrubber decrease by 17-26°C. Only Scrubber Pool temperature increases by 23°C (Table 1-1, Figures 1-1 and 1-2)  49  Chapter 6 - Case Studies: Results and  Discussion  • — 55 kbarrel/day ATB • — 60 kbarrel/day ATB 4—  65 kbarrel/day ATB 70 kbarrel/day ATB 75 kbarrel/day ATB 80 kbarrel/day ATB  - Grid Top - Grid Bottom - Shed Top - Shed Bottom - Scrubber Fbol  50  60  70  80  10 20 30 40 Position, feet from bottom of scr. pool  90  ATB Flow Fate, kbarrel/day (1 barrel=0.0049684 m3/h)  50  Figure 1-1 Effect of ATB flow rate  Figure 1-2 Effect of ATB flow rate on  on temperatures along the Scrubber  temperature profile along the Scrubber  • Overhead  properties:  -Actual volume and mass flow rate of the Scrubber Overhead drop by 3% and 7%, respectively, based on Base Case (Table 1-1, Figure 1-3). -Density decreases from 2.83 to 2.72 kg/m (Table 1-1). -Average molecular weight drops from 71 to 62 (Table 1-2). -Composition shows lower presence of 400°C+ fractions (Table 1-2, Figures 1-5 and I7). • Scrubber Bottom  properties:  -Actual volume and mass flow rate increase by 75% and 71% (Table 1-1). - Density drops from 816 to 796 kg/m (Table 1-1). 3  -Average molecular weight changes from 637 to 594 (Table 1-3).  50  Chapter 6 - Case Studies: Results and  Discussion  -Composition shows much higher presence of middle fractions (400-500°C), while heavy fractions are diluted (Table 1-3, Figures 1-6 and 1-8). • Other.  -Sheds Vapour and Grid Liquid volume and mass flow decrease, while Liquid from the Sheds actual volume and mass flow increase (Figure 1-4).  Figure 1-3 Effect of ATB flow rate on  Figure 1-4 Effect of ATB flow rate on  mass flow rate of Scrubber Overhead  mass flow rate of other streams  and Bottom  51  Chapter 6 - Case Studies: Results and Discussion  Table 1-1 Effect of ATB flow rate on Scrubber parameters Posftkn  ft fronthe pod bcttcm  KrhQidTcp 43  AJBRowRie rrffh kbarrel/da/ Tcp Sags T e r r p H f t )  ATB\fclirre RowRie 348 70 % %  273 55 (Base Case)  298 eo  393  390  -0.7  388  -1.4  385  323 65 %  373 75  397 80 %  %  -22  382  -29  376  4.3  KxhGidEtt 38  ECBonSags Terrp B f t )  395  392  -0.8  389  -1.6  385  -24  382  ^32  376  4.9  SnedsTcp  34  TcpSagsTerrpBtft)  405  402  398  -1.6  395  -24  392  ^2  383  S-BdsBA  22  4.8  BatariSagsTa-rpHft) BJk Liqjd Terrperalre f t )  473  467  -0.8 -1.4  461  -26  457  ^5  453  4.3  447  -56  375  382  1.8  387  32  391  4.3  394  51  398  62  Sditrja-Rrl  0  Row Rates& Densities SmiJberOatHad  AdLEl VJirre Rcw(rrf/h) MassRcwfkgti)  SoiitiRxlUojd  Adual\«LnBHcw(rrfyh)  274,787 778,651  271,620 -12 760233 -24  2S  273272 -0.6 770,080 -1.1 282 -06  280 -12  278 -20  275 -28  272 4.1  78 64372c  90EJ 14.7 729293 133  1,024 298 817,363 27.0  1,147 453 907,871 41.0  1271 61.1 999,796 553  1,386 755 1,103,082 71.4 796 -24  MassDaBty^rrf) MassRow(l<gri) l\tettnsity(l^rr?) l  806  -12  798 -21  792 -29  786 -36  25E  234 -93  21; -172  1.SE -242  30: 217,831  287 -58 206219 -53  279 -8.5 200,755 -7.9  71824 Q5 274,07E -1.2 840,392 -35  7202C Q7 272,336 -1.8 824,337 -53  1.8C -30.3 270 -11.2 195,376 -103 72229 1.0  306 <16 694 397 502,535 391  MassRowFaio A i d Mtiurre Rcw(rr?/h) MassRowfkJh)  714.97  296 -30 211,767 -28 71644 02  3T6d\^xr  AdLEl Vtiure Row(rr?/ri) MassRcw(l^h)  277,32 870,45c  275,780 -0.6 855787 -1.7  ^[irsty(te|'rrr) AdLel\/tiuTeRcw(rr?/ri)  an  310 -1.1  307 -23  497 361,32  532 132 407,740 128  627.66 263 454,777 259  Q0  724.56 QO  378 14.7 304,846 133 803.17 -12  42c 298 341,656 27.0 79334 -21  MBsDarstyfkg'rri)  IVteRcw(kgh) M3EsD3Taty(ko/rr?) l  ToCtter  Aiel UiLrre Rcw(rr?/h) MassRcw(t0h) ^Dratyfl^rrr)  268209 -24 738449 -52  816  OsfEad/AlB QidUc)jd  SBdu'cud  269,930 -1.8 749,638 •3.7  724.7E 33C 269,07? 8158:  724.94  724.31 -Q1 479 453 379490 41.0 791.82 -29  266,026 -32 723279 -7.1  1.65 -331 26C -14.5 183,811 -133  270,551 -2 807,76E -72  724.99 1.4 263281 -33 783,02 -97  299 4.9  293 •6.7  760 529 549994 522  819 64.9 598,614 64.3  724.06 -Q1 531 61.1 417,915 553 78642 ^ 6  724.53 579 431,03E 79542  755 71.4  QO  -24  AUIiud irtcrrrdjcn  N^pCLTtDSBJS  Ta-rperakreft)  514  511  -0.6  506  -12  505  -1.7  503  -22  501  -26  LpgdngSreari SBCfe HfacfiQid SR-Cder  Terrpadueft) SagsBlidaTy SageBlidercy DJy(MvBuri)  534 Q53 Q75 44.32  534 Q53 Q75 44.32  -01  533 Q53 075 44.32  -Q1 QO QO QO  533 Q53 Q75 44.32  -02 QO QO QO  533 Q53 075 44.32  -03 QO QO QO  533 Q53 Q75 7543  -Q3 QO QO 702  Q0 00 QO  52  Chapter 6 - Case Studies: Results and  Discussion  Table 1-2 Effect of ATB flow rate on Scrubber Overhead properties ATB Flow Rate (kbarrel/d)  55 (Basic)  80  Cut Point [%]  393 16.99 70.77 2.83 282,483 -2924 5.30 2.74 0.93 193.80 260,141 11.38 930.01 24.98 2825  376 16.99 61.50 2.49 273,389 -3279 5.48 2.66 0.92 163.43 255,072 11.43 921.05 24.65 2843  Volume fraction 0.060 0.259 0.580 0.101  0.066 0.287 0.553 0.094  Temperature [°C] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kj/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kj/kg] Fraction Distribution Data C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+ (>524°C)  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  55  80  TBP [°C] -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 550 556 616 744 871  TBP [°C] -253 -240 -211 -174 -143 -107 -92 -62 -45 -8 206 298 324 341 363 377 391 405 419 436 441 448 466 484 523 532 546 556 617 753 888  100  10 0  J  ,  ,  ,  0  100  200  300  ,  ,  ,  400 500 600 Temperature, °C  ,  ,  ,  700  800  900  1000  Figure 1-5 Effect of ATB flow rate on Scrubber Overhead TBP curve  53  Chapter 6 - Case Studies: Results and  Discussion  Table 1-3 Effect of ATB flow rate on Scrubber Bottom properties ATB Flow Rate (kbld) Temperature [°C] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  55 (Basic)  80  375 17.00 637.04 815.85 812 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57 0.78  398 17.00 594.06 796.42 1,426 -1325 3.38 2.92 0.00 1731.85 43,904 11.44 0.74 1023.66 0.75 1532 14.74 0.13 0.59 0.75  Volume fraction 0.000 0.000 0.089 0.911  0.000 0.000 0.196 0.804  Fraction Distribution Data C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+ (>524°C)  Cut Point [%]  55  80  T B P [ C] 412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706 743 750 760 807 852 892 917 964 1031 1047 1055 U  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  T B P [°C] 387 421 440 463 466 483 487 491 496 512 514 519 525 556 590 597 628 635 682 687 699 742 750 761 816 854 897 940 965 1002 1041  100  1100 Temperature, °C  Figure 1-6 Effect of ATB flow rate on Scrubber Bottom TBP curve 3±  Chapter 6 - Case Studies: Results and 0.50  Discussion  m 55 kbarrel/day ATB : Light Ends:20%; Water: 65% ; 100> fraction: 15%  1  B 80 kbarrel/day ATB : Light ends: 21%; Water: 65% ; 100> fraction: 14%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  Components' Boiling Temperatures Range, °C  Figure 1-7 Effect of ATB flow rate on Scrubber Overhead composition  • 55 kbarrel/day ATB : Light Ends: 0% ; Water: 1%; 100> fraction: 99% ru 80 kbarrel/day ATB : Light ends: 0% ; Water: 1% ; 100> fraction: 99%  0.10  0.00 200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  Components' Boiling Temperatures Range, °C  Figure 1-8 Effect of ATB flow rate on Scrubber Bottom composition  55  Chapter 6 - Case Studies: Results and Discussion  Discussion: ATB is a low temperature stream (325°C) and its higher flow rate into the system lowers all temperatures along the Scrubber (except Scrubber Pool temperature, which will be explained later). ATB flow rate increase causes less evaporation and consequently smaller amount of vapours, and more liquids. Only lighter fractions are able to evaporate, decreasing the density and average molecular weight of the Overhead. Fractions below 400°C show higher presence in the Overhead, which can be seen in fraction distribution data, TBP curve and composition. Presence of the LGO fractions is higher, of the HGO is lower, but in total, CGO (LGO plus HGO) fraction does not change. Middle fractions (400-500°C) end up in the Scrubber Bottom diluting heavy fractions and also decreasing its density and molecular weight. These fractions are present in higher amount in the Scrubber Bottom (see fraction distribution data, TBP curve and composition). Scrubber Pool temperature rises because in this case SPL Cooler duty is kept constant. More and more liquid passes through the cooler, recycling to the Scrubber Pool. With the constant cooler duty, cooler is not able to sufficiently cool down all this liquid, causing the rise in the Scrubber Pool temperature.  56  Chapter 6 - Case Studies: Results and I I .  HGO Wash  Flow  Discussion  Rate  Flow rate of Heavy Gas Oil (HGO) Wash, a stream that enters at the top of the Grid, has been changed from 24 kbarrel/day in the Base Case to 30 and 40 kbarrel/day. Effect on Scrubber parameters - temperatures and streams flow rates and densities, as well as Scrubber Overhead and Scrubber Pool Liquid composition and properties have been studied. Observations: By increasing HGO Wash flow rate from 24 to 40 kbarrel/day: •  Temperature  profile:  - A l l temperatures along the Scrubber drop by 5-7°C. Shed Bottom temperature decreases by 18°C (Table II-1, Figures II-1 and II-2)  480  500  - • — 24 kbarrel/day HGO - • - - 30 kbarrel/day HGO  460  480  - 40 kbarrel/day HGO  460  440 - Grid Top - Grid Bottom - Shed Top • Shed Bottom - Scrubber Bool  of 440 | 420  420  400  —A—  400  =8=  380  380 360 20  30  40  HGO Flow Rate, kbarrel/day  360  0  10  20  30  40  50  Position, feet from bottom of s c r . pool  (1 bar re 1=0.0049684 m3/h)  Figure II-l Effect of HGO Wash flow  Figure II-2 Effect of HGO Wash flow  rate on temperatures along the Scrubber  rate on temperature profile along the Scrubber  57  Chapter 6 - Case Studies: Results and Discussion Overhead properties:  -Actual volume of the Scrubber Overhead drop by 0.6% and mass flow rate increases by 2.4%. (Table II-l, Figure II-3). -Density increases from 2.83 to 2.92 kg/m (Table II-l). 3  -Average molecular weight increases from 71 to 72 (Table II-2). -Composition shows higher presence of 300-500°C fractions that originate from HGO (Table II-2, Figures II-5 and 11-7). Scrubber Bottom properties:  -Actual volume and mass flow rate increase by 27% and 24% (Table II-l). - Density drops from 816 to 797 kg/m (Table II-l). 3  -Average molecular weight changes from 637 to 596 (Table II-3). -Composition shows higher presence of 400-600°C fractions that also originate from HGO (Table II-3, Figures II-6 and II-8). Other.  -Sheds Vapour volume and mass flow slightly increase, while Grid Liquid and Liquid from the Sheds actual volume and mass flow significantly increase (Figure II-4).  58  Chapter 6 - Case Studies: Results and Discussion 900  800  850 800 750 700 650  - Grid Liquid  600  - Shed Vapor  550  - S h e d Liquid  500  700  450 400 350 300  -I  250 200 150 100  600 20  30 HGO Flow Rate, kbarrel/day (1 barrel=0.0049684m3/h)  40  20  30 HGO Flow Rate, kbarrel/day  40  (1 barrel=0.0049684m3/h)  Figure II-3 Effect of HGO Wash flow  Figure II-4 Effect of HGO Wash flow rate  rate on mass flow rate of Scrubber  on mass flow rate of other streams  Overhead and Bottom  59  Chapter 6 - Case Studies: Results and  Discussion  Table II-l Effect of HGO Wash flow rate on Scrubber parameters Position  ' ft from the pool bottom Koch Grid Top 43 Koch Grid Bot. 38 Sheds Top 34 Sheds Bot. 22 Scrubber Pool 0  Flow Rates& Densities Scrubber Overhead  HGO Wash Flow Rate m3/h kbarrel/day Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)  2.58 305 217,861 714.97 277,329 870,450 3.14 497 361,323 724.76 330 269,078 815.85  2.61 347 247,811 713.83 277,648 878,893 3.17 546 394,390 722.61 359 290,567 808.87  2.64 425 303,893 715.03 277,899 892,151 3.21 639.68 459,545 718.40 418 333,497 797.16  3  3  3  3  3  Additional information Vapour to Sheds Upgoing Stream Sheds Koch Grid SPL Coler  % -1.9 -1.6 -1.1 -3.7 0.0  Mass Flow Ratio Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  3  3  To Coker  386 389 400 456 375  273,178 -0.6 797,485 2.4 2.92 3.0 1,001 26.8 797,839 23.9 797 -2.3  3  Shed Liquid  % -0.7 -0.5 -0.3 -1.5 0.0  274,291 -0.2 788,004 1.2 2.87 1.4 859 8.9 695,136 8.0 809 -0.9  3  Shed Vapor  199 40  274,787 778,651 2.83 789 643,728 816  3  3  Overhead / ATB Grid Liquid  Flow Rate  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) 3  Scrubb.Pool Liquid  HGO Wash Volume 119 149 24 30 (Base Case) 393 390 395 393 405 404 473 466 375 375  Temperature (°C) Temperature (°C) Stage Efficiency Stage Efficiency Duty (MMBtu/h)  514 534 0.53 0.75 44.72  511 534 0.53 0.75 45.75  1.2 13.9 13.7 -0.2 0.1 1.0 0.9 9.9 9.2 -0.3 8.9 8.0 -0.9 -0.7 -0.1 0.0 0.0 2.3  505 533 0.53 0.75 48.25  2.4 39.5 39.5 0.0 0.2 2.5 2.3 28.8 27.2 -0.9 26.8 23.9 -2.3 -1.8 -0.2 0.0 0.0 7.9  60  Chapter 6 - Case Studies: Results and Discussion  Table II-2 Effect of HGO Wash flow rate on Scrubber Overhead properties HGO Wash Flow Rate (kbarrel/day)  24 (Basic)  Temperature [ °C]  40  Cut Point [%]  393 16.99 70.77 2.83 282,860 -2924 5.30 2.74 0.93 193.80 260,141 11.38 930.01 24.98 2825  386 16.99 71.83 2.92 281,203 -2909 5.17 2.71 0.91 194.57 262,526 11.33 940.10 24.60 2800  Volume fraction 0.060 0.259 0.580 0.101  0.052 0.269 0.584 0.096  Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Row [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction (Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg]  40  TBP[°C] 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  Fraction Distribution Data C4-(<177 °C) LGO (177-343 °C) HGO (343-524 °C) 524+ (>524 °C)  24  TBP[°C]  -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 550 556 616 744 871  -253 -237 -206 -166 -134 -101 -81 -49 -22 2 274 311 336 352 366 382 397 407 420 440 444 465 477 491 524 532 546 555 600 733 863  100 90 80  70-I 24 kbarrel/day HGO 60  40 kbarrel/day HGO  a 50 £  40 30 20 10  100  200  300  400  500  600  700  800  900  1000  Temperature, °C  Figure II-5 Effect of HGO Wash flow rate on Scrubber Overhead TBP curve 61  Chapter 6 - Case Studies: Results and  Discussion  Table II-3 Effect of HGO Wash flow rate on Scrubber Bottom properties HGO Wash Flow Rate (kbarrel/day) Temperature [°C] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  24 (Basic)  40  375 17.00 637.04 815.85 812 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57  375 17.00 596.68 797.16 1,030 -1322 3.39 2.92 0.00 1739.52 31,616 11.43 0.71 1024.29 0.75 1535 14.81 0.13 0.57  Volume fraction 0.000 0.000 0.089 0.911  0.000 0.000 0.193 0.807  Fraction Distribution Data C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+(>524°C)  Cut Point [%]  24  40  T B P [ C] 412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706 743 750 760 807 852 892 917 964 1031 1047 1055 U  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  T B P [ C] 343 398 438 460 467 485 490 495 504 513 515 519 525 555 565 593 625 634 681 685 697 741 750 761 817 861 898 945 993 1021 1048 U  1100 Temperature, °C  Figure II-6 Effect of HGO Wash flow rate on Scrubber Bottom TBP curve  62  Chapter 6 - Case Studies: Results and Discussion 0.50  • 24 kbarrel/day HGO : Light Ends: 20%; Water: 65% ; 100> fraction: 15% rj 40 kbarrel/day HGO : Light Ends: 20%; Water: 64% ; 100> fraction: 16%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  C o m p o n e n t s ' Boiling T e m p e r a t u r e s Range, °C  Figure H-7 Effect of HGO Wash flow rate on Scrubber Overhead composition  • 24 kbarrel/day HGO : Light Ends: 0% ; Water: 1%; 100> fraction: 99% B 40 kbarrel/day HGO : Light Ends: 0% ; Water: 1%; 100> fraction: 99%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  C o m p o n e n t s ' Boiling T e m p e r a t u r e s Range, °C  Figure II-8 Effect of HGO Wash flow rate on Scrubber Bottom composition  63  Chapter 6 - Case Studies: Results and Discussion  Discussion: HGO Wash is also a low temperature stream (325°C) and with its higher flow rate, all temperatures in the Scrubber decrease. The drop is less significant than in the case with ATB because total flow rate change is lower. As HGO Wash increases, less evaporation occurs in the system, less vapours and more liquids are produced. The reason that the flow rate of the vapours is slightly higher is that some middle fractions from HGO end up in the vapour. The major increase in the liquid flow rate is in Shed Bottom (Figure II-4), which also causes better cooling of rising vapour, and therefore much lower temperature of the Shed Bottom than in the Base Case. Overhead mass production rate is slightly higher. Opposite to the Case Study I, Overhead density and average molecular weight increase. Also, composition shows higher presence of middle fractions (300-500°C) (CGO). Although most of the middle fractions should end up in the Scrubber Bottom, increased amount appears in the Overhead as well. These fractions originate mostly from HGO and since higher amount of HGO is present in the system, they show up both in the Overhead and Bottom. Scrubber Bottom also has lower density and average molecular weight, and higher presence of these middle fractions, which originate from HGO.  64  Chapter 6 - Case Studies: Results and Discussion  III. HGO Underwash Flow Rate Heavy Gas Oil (HGO) Underwash enters the Scrubber under the Grid with the flow rate of 10 kbarrel/day. Change in flow rate from 0 to 10 and 20 kbarrel/day has been simulated and effect on Scrubber parameters and stream properties have been followed. Observations: By increasing HGO Underwash flow rate from 0 to 20 kbarrel/day: • Temperature  profile:  - A l l temperatures along the Scrubber drop between 8-12°C. Shed Bottom temperature decreases by 35°C (Table III-1, Figures III-l and III-2) 500  500  480  480  460  - 0 kbarrel/day HGO • 10 kbarrel/day HGO  j  - 20 kbarrel/day HGO  - Grid Top - Grid Bottom  - 440  - Shed Top - Shed Bottom  =• 420  - Scrubber Pool  400  380  360 10  20  HGO Flow Rate, kbarrel/day  360 0  10  20  30  40  50  Position, feet from bottom of scr. pool  (1 barrel=0.0049684m3/h)  Figure I I M Effect of HGO Underwash  Figure III-2 Effect of HGO Underwash  flow rate on temperatures along the  flow rate on temperature profile along  Scrubber  the Scrubber  Overhead  properties:  -Volume flow rate of the Scrubber Overhead drops by 0.3% and mass flow rate increases by 4.4%. (Table III-1, Figure III-3). 65  Chapter 6 - Case Studies: Results and Discussion  -Density increases from 2.76 to 2.89 kg/m (Table III-l). 3  -Average molecular weight increases from 69 to 71 (Table III-2). -Composition shows higher presence of 300-500°C fractions that originatefromHGO (Table III-2, Figures III-5 and III-7). Scrubber Bottom properties:  -Actual volume and mass flow rate increase by 30% and 26%, respectively (Table III1). - Density drops from 828 to 803 kg/m (Table III-l). 3  -Average molecular weight changes from 664 to 610 (Table III-3). -Composition shows higher presence of 400-600°C fractions that also originate from HGO (Table III-3, Figures III-6 and III-8). Other.  -Grid Liquid volume and mass flow slightly decrease, while Shed Liquid flow rates significantly increase. Sheds Vapour actual volume flow slightly decreases, and mass flow increases (Figure III-4).  66  Chapter 6 — Case Studies: Results and Discussion  800 •  900 850 800 750 -) 700  700  .C  - Overhead  600  - Scrubber Bottom  aT ra  i  o 600  650  4  550  -Grid Liquid -Shed Vapor -Shed Liquid  500 450 400  1  350 300 250 200 150  500 • 10  20  HGO Flow Rate, kbarrel/day  100 0  10  20  HGO Flow Rate, kbarrel/day (1 barrel=0.0049684 m3/h)  (1 barrel=0.0049684 m3/h)  Figure III-3 Effect of HGO Underwash  Figure III-4 Effect of HGO Underwash  flow rate on mass flow rate of Scrubber  flow rate on mass flow rate of other  Overhead and Bottom  streams  67  Chapter 6 - Case Studies: Results and  Discussion  Table III-l Effect of HGO Underwash flow rate on Scrubber parameters Position  ft from the pool bottom  HGO Uderw. Flow Rate m3/h kbarrel/day  0 0  393  397  0.9  393  0.0  389  -1.2  %  99 20  %  %  Koch Grid Top  43  Koch Grid Bot.  38  Bottom Stage Temp Est (°C)  395  400  1.3  395  0.0  389  -1.6  Sheds Top  34  Top Stage Temp Est (°C)  405  410  1.4  405  0.0  398  -1.6  Sheds Bot.  22  Bottom Stage Temp Est (°C)  473  495  4.6  473  0.0  460  -2.8  0  Bulk Liquid Temperature (°C)  375  375  0.0  375  0.0  375  0.0  Scrubber Pool  Top Stage Temp Est (°C)  HGO Underwash Flow Rate 50 10  50 10 (Base Case)  Flow Rates& Densities Scrubber Overhead  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  274,787 778,651  274,718 0.0 757,412 -2.7  274,787 778,651  0.0 0.0  2.83  2.76 -2.7  2.83  0.0  789 643,728  688 -12.9 569,357 -11.6  789 643,728  0.0 0.0  1.9 921 16.8 739,959 14.9  1.5  816  0.0  803 -1.6  2.51 -2.7  2.58  0.0  305 217,861  0.0 0.0  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  Mass Density (kg/m )  816  3  Overhead / ATB  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  2.58 3  305 217,861  Mass Density (kg/m )  714.97  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  3  714.97  0.0  717.71 0.4  0.0 0.0  276,366 -0.3 877,301 0.8  3.14  3.09 -1.6  3.14  0.0  497 361,323  440 -11.4 316,491 -12.4  496.80 361,323  0.0 0.0  587 18.1 423,013 17.1  724.76  718.94 -0.8  724.76  0.0  720.99 -0.5  330 269,078  287 -12.9 237,991 -11.6  330 269,078  0.0 0.0  385 16.8 309,303 14.9  815.85  0.0  803.04 -1.6  3  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m ) 3  1.5  296 -2.7 212,704 -2.4  277,329 870,450  Mass Density (kg/m ) To Coker  715.50  2.62  0.1  3  Actual Volume Flow (m /h) Mass Flow (kg/h)  3.2 3.3  277,339 0.0 856,313 -1.6  Mass Density (kg/m ) Shed Liquid  314 224,963  2.89  277,329 870,450  3  Shed Vapor  828  273,909 -0.3 790,657 1.5  3  815.85  828.00  1.5  3.17  1.1  Additional information Vapour to Sheds  Temperature (°C)  514  521  1.3  514  0.0  508  -1.2  Upgoing Stream Sheds Koch Grid SPL Coler  Temperature (°C) Stage Efficiency Stage Efficiency Duty (MMBtu/h)  534 0.53 0.75 44.72  535 0.53 0.75 47.35  0.1 0.0 0.0 5.9  534 0.53 0.75 44.72  0.0 0.0 0.0 0.0  534 0.53 0.75 46.13  -0.1 0.0 0.0 3.2  68  Chapter 6 - Case Studies: Results and Discussion  Table III-2 Effect of HGO Underwash flow rate on Scrubber Overhead properties HGO Underw. Flow Rate (kbarrel/day) Temperature [ C]  10  0 397 16.99 69.39 2.76 274,718 -2961 5.40 2.75 0.94 191.12 258,103 11.42 922.13 25.17 2848  U  Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg]  20  Cut Point [%] 389 16.99 0 71.44 1 2.89 2 281,956 3.5 -2913 5 5.21 7.5 2.72 10 0.92 12.5 194.29 15 261,669 17.5 11.34 20 936.54 25 24.75 30 2809 35 40 45 50 0.055 55 60 0.266 0.581 65 0.099 70 75 80 85 90 92.5 95 96.5 98 99 100  393 16.99 70.77 2.83 274,787 -2924 5.30 2.74 0.93 193.80 260,141 11.38 930.01 24.98 2825  Fraction Distribution Data Volume fraction 0.060 0.260 0.571 0.109  C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+ (>524°C)  Cut Point [%]  0.060 0.259 0.580 0.101  0  10  TBP [°C]  TBP [ C]  -253 -238 -208 -170 -138 -103 -88 -55 -43 -4 255 309 336 353 374 392 406 420 440 446 465 482 491 514 526 539 552 556 632 754 879  20 U  -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 550 556 616 744 871  TBP ["C] -253 -237 -206 -166 -135 -101 -82 -50 -26 -1 271 310 336 352 367 385 402 417 423 441 448 466 482 493 524 535 548 556 608 737 866  10 \ 0 -I 0  r  ,  ,  100  200  300  ,  ,  .  400 500 600 Temperature, °C  1  1  1  1  700  800  900  1000  Figure III-5 Effect of HGO Underwash flow rate on Scrubber Overhead TBP curve  69  Chapter 6 - Case Studies: Results and Discussion  Table III-3 Effect of HGO Underwash flow rate on Scrubber Bottom properties HGO Underw. Flow Rate (kbarrel/day) Temperature [°C] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  10  20  375 17.00 663.75 828.00 708 -1333 3.39 2.88 0.00 1911.25 20,282 11.39 0.69 1048.30 0.80 1704 15.92 0.14 0.58  375 17.00 637.04 815.85 119,107 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57  375 17.00 609.86 803.04 139,097 -1324 3.39 2.91 0.00 1774.01 28,688 11.43 0.70 1028.93 0.76 1553 15.02 0.13 0.56  Volume fraction 0.000 0.000 0.045 0.955  0.000 0.000 0.089 0.911  0.000  0  Cut Point [%]  Fraction Distribution Data C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+ (>524°C)  0.000 0.153 0.847  0 TBP [°C]  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  427 464 492 514 520 527 555 556 560 577 592 599 627 633 680 683 685 693 704 742 748 755 764 815 884 898 949 965 1037 1049 1057  10  20  TBP [°C]  TBP [°C]  412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706 743 750 760 807 852 892 917 964 1031 1047 1055  402 435 459 482 486 493 504 513 515 517 520 526 556 575 596 625 634 681 684 692 708 747 754 772 820 884 903 957 1023 1037 1052  100  1100  Temperature, °C  Figure III-6 Effect of HGO Underwash flow rate on Scrubber Bottom TBP curve  70  Chapter 6 - Case Studies: Results and  Discussion  • 0 kbarrel/day HGO Underwash: Light Ends: 20% ; Water: 65%; 100> fraction: 15% • 10 kbarrel/day HGO Underwash: Light Ends: 20% ; Water: 65%; 100> fraction: 15% • 20 kbarrel/day HGO Underwash: Light Ends: 20% ; Water: 65%; 100> fraction: 15%  200-300  300-400  400-500  500-600  600-700  700-800  Components' Boiling Temperatures Range,  800-900  900-1000  1000>  °C  Figure III-7 Effect of HGO Underwash flow rate on Scrubber Overhead composition  • 0 kbarrel/day HGO Underw ash: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 10 kbarrel/day HGO Underwash: Light Ends: 0% Water: 1%; 100> fraction: 99% • 20 kbarrel/day HGO Underwash: Light Ends: 0% Water: 1%; 100> fraction: 99%  200-300  300-400  400-500 500-600 600-700 700-800 800-900 C o m p o n e n t s ' Boiling Temperatures Range, °C  900-1000  1000>  Figure III-8 Effect of HGO Underwash flow rate on Scrubber Bottom composition  71  Chapter 6 — Case Studies: Results and  Discussion  Discussion: HGO Underwash has the same composition and temperature as HGO Wash; the only difference is the amount and the position where the stream enters the Scrubber. Hence, in this case study, all effect and trends are similar to the Case Study II. Again, with higher flow rate of this low temperature stream, less evaporation occurs in the system, less vapour and more liquid are produced. Increase of Shed Liquid flow rate is the same as for the previous case, but cooling ability is higher, because this liquid has lower temperature (HGO Underwash enters the Scrubber at lower point and it is still cold enough when reach the Scrubber Bottom). That is the reason why Shed Bottom temperature is decreased much more than in the previous case (35°C, comparing to 18°C). Overhead mass production rate slightly increase. Its density and average molecular weight increase, and composition shows higher presence of middle fractions (300-500°C). The reason is explained in the previous case. Scrubber Bottom also has lower density and average molecular weight, and higher presence of these middle fractions, which originate from HGO.  72  Chapter 6 - Case Studies: Results and Discussion  IV. HGO Wash Temperature In this case study the effect of HGO Wash temperature on Scrubber parameters has been studied. Temperature has been gradually changed from 250°C to 350°C (in the Base Case HGO Wash temperature is 325°C). Observations: By changing HGO Wash temperature from 250 to 350°C: • Temperature  profile:  - A l l temperatures along the Scrubber increase. Shed Top temperature does not change very much, while most significant change of about 13°C is in the case of Shed Bottom. By using colder HGO Wash (250°C) the Overhead temperature is lowered by 7°C, comparing to the Base Case, while Grid Bottom temperature is still high (Table IV-1, Figures IV-1 and IV-2).  Figure IV-1 Effect  of HGO Wash  Figure IV-2 Effect of HGO Wash  temperature on temperatures along the  temperature on temperature profile along  Scrubber  the Scrubber 73  Chapter 6 - Case Studies: Results and  Overhead  Discussion  properties:  - B y increasing HGO Wash temperature actual volume of the Scrubber Overhead rises by 2.5% and mass flow rate increases by 4.2%. (Table IV-1, Figure IV-3). -Density increases from 2.80 to 2.84 kg/m (Table IV-1). 3  -Average molecular weight increases from 69 to 71 (Table IV-2). -Composition shows higher presence of 500-600°C fractions (Table IV-2, Figures IV-5 andIV-7). Scrubber Bottom  properties:  -Actual volume and mass flow rate drop by 12% and 11% (Table IV-1). - Density changes from 807 to 817 kg/m (Table IV-1). 3  -Average molecular weight changes from 619 to 642 (Table IV-3). -Composition shows lower presence of fractions up to 600°C, while heavier fractions are more concentrated (Table IV-3, Figures IV-6 and IV-8). Other:  - A l l three streams (Grid and Shed Liquid, and Shed Vapour) volume and mass flow decrease, but in the case of liquids the change is more radical (Figure IV-4).  74  Chapter 6 - Case Studies: Results and 800  Discussion  1,000 950  2 780  900 850  760  800 750  740  - Grid Liquid  700 • 720  650 -  - Shed Vapor - S h e d Liquid  600  700  550 500  680  450 660  400 ± 350  640 A  300 250  620 250  275  300  325  350  HGO Wash T e m p e r a t u r e , °C  0-  200 250  275  300  325  350  HGO Wash T e m p e r a t u r e , °C  Figure IV-3 Effect  of HGO Wash  Figure IV-4 Effect of HGO Wash  temperature on mass flow rate of Scrubber  temperature on mass flow rate of other  Overhead and Bottom  streams  75  Chapter 6 - Case Studies: Results and  Discussion  Table IV-1 Effect of HGO Wash temperature rate on Scrubber parameters Position  ft from the pool bottom  HGO Wash Temperaure Temperatures °C  300 %  325 (Base Case)  %  1.3  393  1.9  0.7  395  0.9  396  1.1  0.3  405  0.4  405  0.5  350 %  Koch Grid Top  43  Koch Grid Bot.  38  Bottom Stage Temp Est (°C)  392  394  Sheds Top  34  Top Stage Temp Est (°C)  403  405  Sheds Bot.  22  Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)  464  470  1.4  473  2.1  477  2.9  375  375  0.0  375  0.0  375  0.0  269,827 754,323  273,227 772,077  1.3 2.4  274,787 778,651  1.8 3.2  276,497 786,167  2.5 4.2  2.80  2.83  1.1  2.83  1.4  2.84  1.7  Scrubber Pool  0  Top Stage Temp Est (°C)  250 386  391 .  396  2.6  Flow RatesS Densities Scrubber Overhead  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  870 701,660  Mass Density (kg/m ) 3  Overhead / ATB  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  Shed Vapor  Actual Volume Flow (m /h) Mass Flow (kg/h)  Shed Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  3  0.7  816  1.2  817  1.3  2.50  2.56  2.4  2.58  3.2  2.60  4.2  322 -10.6 229,561 -10.3 0.4 -0.1 -1.0  277,329 -0.2 870,450 -1.5  277,138 -0.3 867,023 -1.9  3.15  -0.9  3.14 -1.4  3.13 -1.7  513.94 -7.1 372,207 -6.9  497 -10.2 361,323 -9.6  484 -12.5 350,719 -12.2  553 399,580  Mass Density (kg/m )  722.23  724.22  0.3  364 293,294  339 275,723  -6.6 -6.0  806.53  812.15  0.7  3  3  289 -19.9 206,915 -19.1  712.98  3.18  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  305 -15.4 217,861 -14.8  277,633 875,577  3  To Coker  812  710.21  Mass Density (kg/m ) 3  767 -11.9 626,246 -10.7  277,834 884,067  3  3  789 -9.3 643,728 -8.3  807  360 255,809  Mass Density (kg/m )  812 -6.6 659,624 -6.0  3  714.97  724.76  0.7  0.3  381 4.8 269,078 -8.3 815.85  1.2  717.02  724.78  1.0  0.4  320 -11.9 261,771 -10.7 816.90  1.3  Additional information Vapour to Sheds  Temperature (°C)  510  513  0.6  514  0.9  515  1.1  Upgoing Stream Sheds  Temperature (°C)  534  534  0.1  534  0.53  0.53  0.53  Koch Grid SPL Coler  Stage Efficiency Duty (MMBtu/h)  0.75 44.73  0.75 43.35  0.0 0.0 -3.1  534 0.53  0.1  Stage Efficiency  0.1 0.0  0.75 44.72  0.0 0.0  0.75 42.60  0.0 0.0 -4.8  76  Chapter 6 - Case Studies: Results and  Discussion  Table IV-2 Effect of HGO Wash temperature on Scrubber Overhead properties 250  HGO Wash Temperature (°C) Temperature [°C]  325 (Basic)  Cut Point [%]  350  250  325  350  T B P [°C]  TBP fC]  T B P [°C]  386  393  396  Pressure [psig]  16.99  16.99  16.99  0  -253  -253  -253  Molecular Weight  68.88  70.77  71.35  1  -239  -237  -237  2.80  2.83  2.84  2  -209  -207  -206  269,827  282,860  276,497  3.5  -170  -167  -167  -3009  -2924  -2898  5  -139  -136  -135  Mass Entropy [kJ/kg-C]  5.31  5.30  5.30  7.5  -103  -102  -102  Mass Heat Capacity [kJ/kg-C]  2.72  2.74  2.74  10  -89  -85  -83  Vapor Phase Fraction ( Mass Basis)  0.91  0.93  0.93  12.5  -55  -51  -50  187.49  193.80  195.76  15  -43  -34  -28  258,922  260,141  260,513  17.5  -5  -3  -1  11.39  11.38  11.37  20  252  266  270  Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg]  Specific Heat [kJ/kgmole-C] Std. G a s Flow [STD_m3/h] Watson K  927.59  930.01  930.77  25  309  310  311  Molar Volume [m3/kgmole]  24.64  24.98  25.10  30  331  336  337  Mass Heat of Vap. [kJ/kg]  2836  2825  2820  35  350  354  354  40  365  373  376  45  381  391  391  50  397  405  406  Liq. Mass Density (Std. Cond) [kg/m3]  Fraction Distribution Data Volume fraction C4-(<177°C)  0.054  0.060  0.057  55  409  420  420  L G O (177-343°C)  0.277  0.259  0.261  60  423  439  440  H G O (343-524°C)  0.569  0.580  0.581  65  440  442  443  524+(>524°C)  0.101  0.101  0.101  70  448  459  463  75  466  471  474  80  483  486  488  85  494  496  496  90  525  525  525  92.5  536  537  538  95  549  550  551  96.5  556  556  556  98  618  616  617  99  730  744  743  100  878  871  869  100  90  -I  80  HGO T=250°C HGO T=325°C  70  HGO T=350°C 60 <jj 50 >  40 30 20 \ 10  0  100  200  300  400 500 600 Temperature, °C  700  800  900  1000  Figure IV-5 Effect of HGO Wash temperature on Scrubber Overhead TBP curve 77  Chapter 6 - Case Studies: Results and  Discussion  Table IV-3 Effect of HGO Wash temperature on Scrubber Bottom properties HGO Wash Temperature (°C) Temperature [°C]  250  325 (Basic)  350  Cut Point [%]  250  325  T B P [°C]  350  T B P ["C]  T B P ["C]  375  375  375  17.00  17.00  17.00  0  406  412  415  Molecular Weight  619.51  637.04  642.88  1  439  441  451  Mass Density [kg/m3]  806.53  815.85  816.90  2  464  466  482  870  812  767  3.5  483  490  494  -1325  -1329  -1321  5  489  504  512  Mass Entropy [kJ/kg-C]  3.40  3.39  3.40  7.5  496  514  515  Mass Heat Capacity [kJ/kg-C]  2.91  2.89  2.90  10  513  517  519  Vapor P h a s e Fraction ( M a s s Basis)  0.00  0.00  0.00  12.5  515  520  524  Specific Heat [kJ/kgmole-C]  1800.34  1843.61  1863.12  15  517  525  544  Std. G a s Flow [STD_m3/h]  26,780  23,893  23,033  17.5  519  548  555  11.43  11.41  11.40  20  522  556  556  0.70  0.69  0.68  25  555  564  590  1031.82  1038.96  1041.23  30  561  592  597  Molar Volume [m3/kgmole]  0.77  0.78  0.79  35  592  605  625  M a s s Heat of Vap. [kJ/kg]  1425  1309  1667  40  603  630  632  Surface Tension [dyne/cm]  15.13  15.47  15.45  45  630  635  680  Thermal Conductivity [W/m-K]  0.13  0.13  0.13  50  635  682  683  Viscosity [cP]  0.56  0.57  0.55  55  682  684  686  1425  1309  1667  60  686  693  695  65  697  706  737  70  739  743  745  749  750  751  Pressure [psig]  Act. Volume Flow [m3/h] M a s s Enthalpy [kJ/kg]  Watson K Kinematic Viscosity [cSt] Liq. M a s s Density (Std. Cond) [kg/m3]  Fraction Distribution Data Volume fraction Volume fraction C4-(<177°C)  0.000  0.000  0.000  75  LGO(177-343°C)  0.000  0.000  0.000  80  756  760  761  H G O (343-524°C)  0.129  0.089  0.076  85  800  807  809  524+ (>524°C)  0.871  0.911  0.924  90  821  852  856  92.5  887  892  893  95  905  917  937  96.5  963  964  964  98  1026  1031  1032  99  1046  1047  1047  100  1054  1055  1055  100  400  500  600  700  800  900  1000  1100  Temperature, °C  Figure IV-6 Effect of HGO Wash temperature on Scrubber Bottom TBP curve 78  Chapter 6 - Case Studies: Results and Discussion 0.50  • HGO T=250°C: Light Ends: 20%; Water: 65% ; 100> fraction: 15% • HGO T=325°C: Light Ends: 20%; Water: 65% 100> fraction: 15%  0.40  • HGOT=350°C: Light Ends: 20%; Water: 65% 100> fraction: 15% (A  0.30  c o  o  s 4-  a> o  S  0.20 0.10  0.00  11 200-300  300-400  400-500  500-600  Components'  600-700  700-800  800-900  900-1000  1000>  Boiling T e m p e r a t u r e s R a n g e , °C  Figure IV-7 Effect o f H G O Wash temperature on Scrubber Overhead composition  0.50  • HGO T= 250°C: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • HGO T=325°C: Light Ends: 0% ; Water: 1 %; 100> fraction: 99%  0.40  • HGO T=350°C: Light Ends: 0% ; Water: 1%; 100> fraction: 99% c o  0.30  u  2 <+ffi  o  S  0.20 0.10  0.00 200-300  _ tri  300-400  400-500  500-600  Components'  600-700  700-800  800-900  900-1000  1000>  Boiling T e m p e r a t u r e s R a n g e , °C  Figure IV-8 Effect o f H G O Wash temperature on Scrubber Bottom composition  79  Chapter 6 - Case Studies: Results and Discussion  Discussion: By increasing HGO Wash temperature from 250 to 350°C, all the temperatures in the Scrubber are higher. The major changes are at Shed Bottom and Grid Top position. Higher temperature causes more evaporation and higher amount of vapour, and less liquid. Smaller amount of Shed Liquid gets in contact with hot vapour from the cyclones, causing significant increase in Shed Bottom temperature. At the Grid Top position, higher flow rate of hot vapour get in contact with unchanged flow rate but hotter HGO Wash, resulting in higher temperature of the Grid Top. Although higher temperature produces more vapour, it could be noticed that Shed Vapour flow rate drops a little bit. The explanation could be that less liquid comes to the Sheds from the Grid, resulting in smaller total amount of evaporated liquid. Density and average molecular weight of the Overhead are higher, while composition, fraction distribution data and TBP curves show lower presence of heavier fractions (500°C+) in the Overhead. Most heavy fractions end up in the Scrubber Bottom, which can be seen from the increased density, molecular weight and presence of heavy fractions in the Bottom.  80  Chapter 6 - Case Studies: Results and Discussion V. HGO Underwash In and Out of Service This study investigates the effect o f Grid Underwash function on Scrubber parameters and properties o f Scrubber Overhead and Scrubber Bottom (Scrubber Pool Liquid). Four options have been investigated: 1. H G O Underwash is i n service - flow rate o f H G O Underwash is 10 kbarrel/day; 2. H G O Underwash is out o f service and Overhead temperature is not controlled. 3. H G O Underwash is out o f service and Overhead temperature is controlled by H G O Wash flow rate - i n order to maintain the overhead temperature constant (around 393°C), H G O Wash actual volume flow rate should be increased by 37.5% (from 24 kb/day to 33 kb/day). 4. H G O Underwash is out o f service and Overhead temperature is controlled by A T B feed flow rate - increase o f A T B actual volume flow b y 14.5% (from 55 kb/day to 63 kb/day) is able to keep the overhead temperature constant.  Observations: • Temperature profile - H G O Underwash off, uncontrolled: A l l temperatures increase by 4-5°C, only Shed Bottom temperature is 22°C higher. - H G O Underwash off, controlled by H G O Wash: Overhead temperature is kept the same (393°C), while all others are increased b y 4-5°C, compared to Base Case; - H G O Underwash o f f controlled by A T B : This case has almost the same effect as when H G O Underwash is i n service. There is only a small change in all observed temperatures (Table V - l , Figures V - l and V - 2 ) .  81  Chapter 6 — Case Studies: Results and  Discussion  Figure V - l Effect of H G O Underwash  Figure V - 2 Effect of H G O Underwash  service on temperatures along the Scrubber  service on temperature profile along the Scrubber  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases.  • Overhead  •  properties:  HGO Underwash off, uncontrolled:  -Actual volume flow of the Scrubber Overhead changes insignificantly and mass flow rate drops by 3%, comparing to the case when H G O Underwash is in service (Table V - l , Figure V-3). -Density changes from 2.83 to 2.76 kg/m (Table V - l ) . 3  -Average molecular weight drops from 71 to 69 (Table V-2). -Composition shows lower presence of 300-500°C fractions that originate from H G O (Table V-2, Figures V-5 and V-7). 82  Chapter 6 - Case Studies: Results and Discussion  •  H G O Underwash off, controlled by H G O Wash:  -There is no significant change in any of Scrubber Overhead properties, compared to Base Case. •  H G O Underwash off, controlled by A T B :  -Actual volume and mass flow rate drops by 0.7 and 3.9%, respectively. -Density changes from 2.83 to 2.74 kg/m . 3  -Average molecular weight drops from 71 to 69 (Table V-2). -Composition shows slightly lower presence of 300-500°C fractions that originate from HGO (Table V-2, Figures V-5 and V-7).  Scrubber Bottom properties: •  H G O Underwash off, uncontrolled:  -Actual volume and mass flow rate drop by 13% and 12% (Table V - l ) . - Density increases from 816 to 828 kg/m (Table V - l ) . 3  -Average molecular weight changes from 637 to 663 (Table V-3). -Composition shows lower presence of 400-600°C fraction that originates from H G O (Table V-3, Figures V-6 and V-8). •  H G O Underwash off, controlled by H G O Wash:  -Again, for all Scrubber Bottom properties there are just minor changes, comparing to the case when H G O Underwash is in service. •  H G O Underwash off, controlled by A T B :  -Actual volume and mass flow rate increase by 7.5% and 8% (Table V - l ) . - Density increases from 816 to 821 kg/m (Table V - l ) . 3  -Average molecular weight changes from 637 to 645 (Table V-3). -Composition shows lower presence of 500-600°C fraction that originates from HGO, but slightly higher presence of 6 0 0 ° C fractions (Table V-3, Figures V-6 and V-8). +  83  Chapter 6 - Case Studies: Results and Discussion  • Other. •  HGO Underwash off, uncontrolled: -Grid Liquid volume and mass flow slightly increase; Shed Vapour mass flow decreases a little, and Shed Liquid volume and mass flow significantly decrease.  •  H G O Underwash off, controlled by H G O Wash: - G r i d Liquid volume and mass flow rate radically increase, Shed Vapour does not change much, and Shed Liquid flow rate decreases a little.  •  HGO Underwash off, controlled by A T B : -Grid Liquid and Shed vapour flow rate show a minor drop, while Shed Liquid flow rate increases by 5% (Figure V-4).  Figure V-3 Effect of H G O Underwash  Figure V-4 Effect of H G O Underwash  service on mass flow rate of Scrubber  service on mass flow rate of other  Overhead and Bottom  streams  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases. 84  Chapter 6 - Case Studies: Results and Discussion  Table V-l Effect of H G O Underwash service on Scrubber parameters  Position  HGO Underwash ON (Base Case)  ft from the pool bottom  1  2  %  Top Stage Temp Est (°C)  393  397  0.9  38  Bottom Stage Temp Est (°C)  395  400  34  Top Stage Temp Est (°C)  405  410  22  Bottom Stage Temp Est (°C)  473  0  Bulk Liquid Temperature (°C)  375  Koch Grid Top  43  Koch Grid Bot. Sheds Top Sheds Bot. Scrubber Pool  OFF Uncontrolled  HGO Underwash OFF Contr. by HGO 3 %  OFF Contr. by ATB  4  %  393  0.0  0.9  396  0.3  1.2  406  0.3  477  0.8  476  0.6  375  0.0  372  -0.7  393  0.0  1.3  399  1.4  410  495  4.6  375  0.0  Note 1: When in service H G O Underwash flow rate is 10,000 barrel/day Note 2: To keep the Overhead T constant, H G O Wash should be increased from 24 to 33 kbpd, and A T B from 55 to 63 kbpd.  Flow RatesS Densities Scrubber Overhead  Actual Volume Flow (m /h)  274,787  274,718  0.0  274,692  0.0  272,863 -0.7  Mass Flow (kg/h)  778,651  757,412  -2.7  778,344  0.0  748,181  -3.9  2.83  2.76  -2.7  2.83  0.0  2.74  -3.2  789  688 -12.9  771  -2.3  848  7.5  643,728  569,357 -11.6  631,095  -2.0  696,159  8.1  0.3  821  0.6  2.16 -16.2  3  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  816  828  1.5  818  Overhead / A T B  Mass Flow Ratio  2.58  2.51  -2.7  2.58  0.0  Grid Liquid  Actual Volume Flow (m Ih)  305  314  3.2  378  24.1  297 -2.4  217,861  224,963  3.3  268,867  23.4  212,905 -2.3  3  Mass Flow (kg/h)  714.97  715.50  0.1  710.86  -0.6  Actual Volume Flow (m Ih)  Mass Density (kg/m )  277,329  277,339  0.0  278,645  0.5  275,476 -0.7  Mass Flow (kg/h)  870,450  856,313  -1.6  873,944  0.4  835,027  3.14  3.09  -1.6  3.14  -0.1  440 -11.4 316,491 -12.4  488.15  -1.7  520  4.7  353,507  -2.2  378,953  4.9  724.18  -0.1  728.33  0.5  3  Shed Vapor  Mass Density (kg/m ) 3  Shed Liquid  Actual Volume Flow (m Ih) Mass Flow (kg/h)  497 361,323  Mass Density (kg/m ) To Coker  -0.8  0.1 -4.1  3.03 -3.4  330  287 -12.9  322  -2.3  354  7.5  269,078  237,991 -11.6  263,798  -2.0  290,994  8.1  818.40  0.3  821.10  0.6  Actual Volume Flow (m /h) 3  Mass Flow (kg/h)  718.94  724.76  3  715.94  Mass Density (kg/m ) J  815.85  828.00  1.5  Additional information Vapour to Sheds  Temperature (°C)  514  521  1.3  515  0.3  515  0.2  Upgoing Stream  Temperature (°C)  534  535  0.1  534  0.0  534  0.0  Sheds  Stage Efficiency  0.53  0.53  0.0  0.53  0.0  0.53  0.0  Koch Grid  Stage Efficiency  0.75  0.75  0.0  0.75  0.0  0.75  0.0  S P L Coler  Duty (MMBtu/h)  44.72  47.35  5.9  45.64  2.0  56.04  25.3  85  Chapter 6 - Case Studies: Results and  Discussion  Table V-2 Effect of HGO Underwash service on Scrubber Overhead properties HGO OFF  HGO ON Temperature f t ]  HGO OFF  (uncontr.)  HGO OFF  (contr. by HGO) (contr. by ATB)  Cut Point [%]  HGO OFF  HGO ON  HGO OFF  (uncontr.)  TBP f C]  HGO OFF  (contr. by HGO) (contr. by ATB)  TBP fC]  TBP f t ]  TBP [°C]  393  397  393  393  Pressure [psig]  16.99  16.99  16.99  16.99  0  -253  -253  -253  Molecular Weight  70.77  69.39  70.78  68.61  1  -237  -238  -237  -239  2.83  2.76  2.83  2.74  2  -207  -208  -207  -209  Mass Density [kg/m3]  -253  282,860  282,789  282,762  280,880  3.5  -167  -170  -167  -171  Mass Enthalpy [kj/kg]  -2924  -2961  -2924  -2996  5  -136  -138  -136  -140  Mass Entropy [kJ/kg-C]  5.30  5.40  5.30  5.40  7.5  -102  -103  -102  -104  Mass Heat Capacity [kJ/kg-C]  2.74  2.75  2.74  2.74  10  -85  -88  -85  -89  Vapor Phase Fraction (Mass Basis)  0.93  0.94  0.92  0.93  12.5  -51  -55  -51  -56  Specific Heat [kJ/kgmole-C]  193.80  191.12  193.98  188.17  15  -34  -43  -34  -43  Std. Gas Flow [STDjn3/h]  260,141  258,103  260,017  257,826  17.5  -3  -4  -3  -5  11.38  11.42  11.38  11.42  20  266  255  266  248  Act. Volume Flow [m3/h]  Walson K Liq. Mass Density (Std. Cond) [kg/m3]  930.01  922.13  929.28  922.13  25  310  309  310  309  Molar Volume [m3/kgmole]  24.98  25.17  24.98  25.02  30  336  336  336  335  Mass Heal ofVap. [kJ/kg]  2825  2848  2826  2748  35  354  353  354  351  40  373  374  374  367  45  391  392  391  389  50  405  406  406  405  Fraction Distribution Data Volume fraction C4-(<177°C)  0.060  0.060  0.060  0.055  55  420  420  420  420  LGO (177-343°C)  0.259  0.260  0.259  0.268  60  439  440  439  439  HGO(343-524°C)  0.580  0.571  0.580  0.575  65  442  446  443  443  524+(>524°C)  0.101  0.109  0.101  0.101  70  459  465  461  459  75  471  482  473  474  80  486  491  487  487  85  496  514  496  496  90  525  526  525  525  92.5  537  539  537  538  95  550  552  550  551  96.5  556  556  556  556  98  616  632  617  623  99  744  754  784  752  100  871  879  871  880  10 O  I O  ,  100  ,  200  ,  300  ,  400  ,  500  .  600  1  700  1  800  1  900  1000  T e m p e r a t u r e , °C  Figure V-5 Effect of HGO Underwash service on Scrubber Overhead TBP curve 86  Chapter 6 - Case Studies: Results and Discussion  Table V-3 Effect of HGO Underwash service on Scrubber Bottom properties HGO ON Temperature fC]  HGO OFF HGO OFF HGO OFF Cut Point [%] (uncontr.) (contr. by HGO) (contr. by ATB)  HGO ON  HGO OFF HGO OFF HGO OFF (uncontr.) (contr. by HGO) (contr. by ATB)  TBP [°C]  TBP ["C]  TBP ["C]  TBP f C]  375  375  375  375  17.00  17.00  17.00  17.00  0  412  427  414  412  Molecular Weight  637.04  663.75  641.74  645.71  1  441  464  450  441  Mass Density [kg/m3]  815.85  828.00  818.40  821.10  2  466  492  482  469  812  708  794  873  3.5  490  514  494  493  Mass Enthalpy [kJ/kg]  •1329  •1333  •1330  -1337  5  504  520  512  512  Mass Entropy [kJ/kg-C]  3.39  3.39  3.39  3.38  7.5  514  527  515  515  Mass Heat Capacity [W/kg-C]  2.89  2.88  2.89  2.88  10  517  555  519  519  Vapor Phase Fraction (Mass Basis)  0.00  0.00  0.00  0.00  12.5  520  556  524  524  Specific Heat [kJ/kgmole-C]  1843.61  1911.25  1854.96  1861.72  15  525  560  543  552  Std. Gas Flow[STD_m3/h]  23,893  20,282  23,252  25,492  17.5  548  577  555  556  11.41  11.39  11.40  11.41  20  556  592  556  558  0.69  0.69  0.69  0.71  25  564  599  588  591  1038.96  1048.30  1040.88  1041.87  30  592  627  595  597  Molar Volume [m3/kgmole]  0.78  0.80  0.78  0.79  35  605  633  624  626  Mass Heat of Vap. [kJ/kg]  1309  1704  1657  1647  40  630  680  632  633  Surface Tension [dyne/cm]  Pressure [psig]  Act. Volume Flow [m3/h]  Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3]  15.47  15.92  15.57  15.71  45  635  683  671  680  Thermal Conductivity [W/m-K]  0.13  0.14  0.14  0.14  50  682  685  683  683  Viscosity [cP]  0.57  0.58  0.57  0.58  55  684  693  685  687  60  693  704  695  696  65  706  742  708  738  70  743  748  744  745 752  Fraction Distribution Data Volume fraction C4-(<177°C)  0.000  0.000  0.000  0.000  75  750  755  751  LGO(177-343°C)  0.000  0.000  0.000  0.000  80  760  764  760  761  HGO (343-524°C)  0.089  0.045  0.077  0.075  85  807  815  809  809  524+(>524°C)  0.911  0.955  0.923  0.925  400  500  600  700  800  90  852  884  854  853  92.5  892  898  893  893 926  95  917  949  932  96.5  964  965  964  964  98  1031  1037  1032  1031  99  1047  1049  1047  1046  100  1055  1057  1055  1054  900  1000  1100  Temperature, °C  Figure V-6 Effect of HGO Underwash service on Scrubber Bottom TBP curve  87  Chapter 6 - Case Studies: Results and  Discussion  • HGO ON: Light Ends: 20%; Water: 65% ; 100> fraction: 15%  1 •  200-300  300-400  400-500  • HGO OFF-uncontrolled: Light Ends: 20%; Water: 65% ; 100> fraction: 15% O HGO OFF-controlled by HGO Wash: Light Ends: 20%; Water 65% ; 100> fraction: 15% 0 HGO OFF-controlled by ATB: Light Ends: 21%; Water: 65% ; 100> fraction: 14%  1  500-600  600-700  700-800  800-900  900-1000  1000>  C o m p o n e n t s ' Boiling T e m p e r a t u r e s Range, °C  Figure V-7 Effect o f H G O Underwash service on Scrubber Overhead composition  • HGO ON: Light Ends: 0% ; Water; 1%; 100> fraction: 99% • HGO OFF-uncontrolled: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • HGO OFF-controlled by HGO Wash: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • HGO OFF-controlled by ATB: Light Ends: 0% Water: 1%; 100> fraction: 99%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  C o m p o n e n t s ' Boiling T e m p e r a t u r e s Range, °C  Figure V-8 Effect o f H G O Underwash service on Scrubber Bottom composition  8S  Chapter 6 - Case Studies: Results and Discussion  Discussion: As was mentioned before, HGO Underwash is a cold stream that enters the Scrubber at 325°C, cooling down hot vapour from the Sheds. If Underwash were out of service, this vapour would still be hot, causing higher temperature in the whole system, more evaporation and consequently lower flow rate of the Shed Liquid and Scrubber Bottom. Lower flow rate of Shed Liquid causes such high jump in temperature for the Shed Bottom, because it is not enough to cool down the hot vapour from cyclones. The significant drop in Bottom flow rate is also due to the overall mass balance (less mass "in" since HGO Underwash is out of service). Although more evaporation occurs, vapour in the Scrubber show slightly lower mass flow rates. The reason is lower total mass "in". In this case, Scrubber Overhead contains less middle fractions that originate from HGO, lower density and molecular weight. Scrubber Bottom also has less middle fractions, heavy fractions are concentrated, and density and molecular weight are higher. If Overhead temperature were controlled by HGO Wash flow rate (same composition and almost same amount as HGO Underwash), all other temperatures would be increased by several degrees. Total flow rate of Overhead and Scrubber Bottom would be almost the same as for the Base Case. The same is true for Overhead and Bottom properties and composition. The difference is in distribution of internal vapour and liquid streams, caused by different entrance position of HGO Wash and Underwash. ATB flow rate seems to provide better control over the whole temperature profile. ATB has higher cooling (heating) capacity than HGO. It also doesn't have any significant effect on vapour and liquid flow rates and properties, since the temperature and position of this stream is similar to HGO Underwash. Both Scrubber Overhead and Bottom contain a little bit more heavy fractions in this case, because ATB is a heavier feedstock than HGO.  89  Chapter 6 — Case Studies: Results and Discussion  VI. Number of Trays in the Sheds Originally, number of trays in the Shed column was six. In this case study this number has been changed from 2 to 10, with a step of 2. Trays efficiency has been kept the same - 53%. The parameters are compared to the original case where the Sheds has 6 rows. Observations: By changing the number of Sheds' trays from 2 to 10: • Temperature profile:  -Grid Top temperature drops from 400°C to 392°C; -Grid Bottom temperature decreases up to 6 trays in Sheds, but after that increases to 401°C and remains the same; -Shed Top temperature decreases also up to 6 trays in Sheds, and then remains the same; -Shed Bottom temperature increases gradually (Table VI-1, Figures VI-1 and VI-2).  90  Chapter 6 - Case Studies: Results and Discussion  Figure VI-1 Effect of number of Sheds  Figure VI-2 Effect of number of Sheds  trays on temperatures along the Scrubber  trays on temperature profile along the Scrubber  Overhead properties:  - A s the number of trays increases, actual volume flow of the Scrubber Overhead first rises (up to 6 trays in the Sheds) and then remains almost the same, while mass flow rate decreases gradually by 7%. (Table VI-1, Figure VI-3). -Density changes from 2.84 to 2.70 kg/m (Table VI-1). 3  -Average molecular weight drops from 72 to 67 (Table VI-2). -Composition shows higher and higher presence of fractions up to 500°C fractions, and lower presence of heavier fractions (Table VI-2, Figures VI-5 and VI-7). Scrubber Bottom properties:  -Both actual volume and mass flow rate increase approximately by 21-22% (Table VI-  91  Chapter 6 - Case Studies: Results and Discussion  - Density drops from 822 to 802 kg/m (Table VI-1). 3  -Average molecular weight changes from 649 to 634 (Table VI-3). -Composition shows higher presence of 500-600°C fraction, while both lighter and heavier fractions are less included (Table VI-3, Figures VI-6 and VI-8). • Other.  -Both volume and mass flow rate of Grid Liquid increase up to 6 trays in Sheds, and start to drop with higher number of trays. -Shed Vapour volume flow decreases up to 6 trays, then increases and stays more or less the same. Mass flow rate increases up to 6 trays, after what stars to drop. -Both volume and mass flow rate of Shed Liquid increase (Figure VI-4).  800 •  750 700 650 • Overhead - Scrubber Bottom  - Grid Liquid  600  - Shed Vapor  550  - Shed Liquid  500 450 400 350 300 250 200 150 100 4  6  l  10  Number of Sheds' trays  4  6  8  Number of Sheds' trays  Figure VI-3 Effect of number of  Figure VI-4 Effect of number of Sheds  Sheds trays on mass flow rate of  trays on mass flow rate of other streams  Scrubber Overhead and Bottom  92  Chapter 6 - Case Studies: Results and  Discussion  Table VI-1 Effect of number of Sheds trays on Scrubber parameters Position  ft  Number of Sheds' trays  from the  6  pool bottom  Base Case  Koch Grid Top  43  Top Stage Temp Est  2  6  4  %  393  400  1.8  %  8  10  %  %  394  0.3  393  0.0  392  %  -0.2  392  -0.3  Koch Grid Bot.  38  Bottom Stage Temp Est  395  409  3.6  397  0.5  395  0.0  401  1.5  401  1.6  Sheds Top  34  Top Stage Temp Est  405  417  3.0  407  0.4  405  0.0  406  0.2  405  0.1  Sheds Bot.  22  Bottom Stage Temp Est  473  466  -1.5  472  -0.4  473  0.0  492  4.0  494  4.3  0  Bulk Liquid Temperature  375  375  0.0  375  0.0  375  0.0  375  0.0  375  0.0  Scrubber Pool  Flow RatesS Densities Scrubber Overhead  Actual Volume Flow (m /h)  274,787  279,507  1.7  275,568  0.3  274,787 0.0  275,284 0.2  275,149 0.1  Mass Flow (kg/h)  778,651  794,406 2.0  782,235  0.5  778,651 0.0  746,571 -4.1  741,836 -4.7  2.83  2.84 0.3  2.84  0.2  2.83 0.0  2.71 -4.3  2.70 -4.9  789  738 -6.5  778 -1.3  789 0.0  895 13.5  912 15.6  643,728  606,279 -5.8  635,952 -1.2  643,728 0.0  720,241 11.9  731,830 13.7  816  822 0.7  817  0.1  816 0.0  804 -1.4  802 -1.6  2.58  2.63 2.0  2.59  2.46 -4.7  3  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m /h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  Overhead/ATB  Mass Flow Ratio  Grid Liquid  Actual Volume Flow(m /h)  0.5  2.58 0.0  2.47 -4.1  305  193 -36.7  293 -3.8  305 0.0  217 -28.7  211 -30.7  217,861  139,382 -36.0  209,570 -3.8  217,861 0.0  153,354 -29.6  149,067 -31.6  3  Mass Flow (kg/h) Mass Density (kg/m )  714.97  714.94  0.0  714.97 0.0  705.88 -1.3  278,195  0.3  277,329 0.0  279,033 0.6  278,954  865,746 -0.5  870,450 0.0  773,865 -11.1  764,843 -12.1  2.86 -8.9  3.11 -0.8  3.14 0.0  2.77 •11.6  2.74 -12.6  445 -10.4  487.08 -2.0  497 0.0  601 21.0  612 23.2  361,323  329,675 -8.8  354,643 -1.8  361,323 0.0  411,200 13.8  420,275 16.3  724.76  740.33 2.1  728.11 0.5  724.76 0.0  683.68 -5.7  677.97 -6.5  330  308 -6.5  325 -1.3  381 15.6  374 13.5  378 14.6  269,078  253,425 -5.8  265,828 -1.2  269,078 0.0  301,061 11.9  305,905 13.7  815.85  821.88 0.7  817.03 0.1  815.85 0.0  804.45 -1.4  802.40 -1.6  3  Shed Vapor  277,329  282,557  Mass Flow (kg/h)  870,450  807,726 -7.2  3.14 497  3  Mass Density (kg/m ) 3  Shed Liquid  Actual Volume Flow (m /h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  To Coker  722.93 1.1  Actual Volume Flow (m /h)  Actual Volume Flow(m /h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  1.9  705.56 -1.3 0.6  Additional information Vapour to Sheds  Temperature (°C)  514  514  0.1  514  0.0  514  0.0  517  0.6  517  0.5  Upgoing Stream  Temperature (°C)  534  534  0.0  534  0.0  534  0.0  534  0.0  534  0.0  Sheds  Stage Efficiency  0.53  0.53  0.0  0.53  0.0  0.53  0.0  0.53  0.0  0.53  0.0  Koch Grid  Stage Efficiency  0.75  0.75  0.0  0.75  0.0  0.75  0.0  0.75  0.0  0.75  0.0  SPLColer  Duty (MMBtu/h)  44.72  36.69  -18.0  42.74  -4.4  44.72  0.0  68.10  52.3  70.85  58.4  93  Chapter 6 - Case Studies: Results and Discussion  Table VI-2 Effect of number of Sheds trays on Scrubber Overhead properties Number of Sheds' trays  2  Temperature fC] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg]  6  10  Cut Point [%]  400  393  392  16.99 72.00 2.84 279,507 -2863 5.30 2.75 0.95 198.04 260,892 11.37 931.79 25.33 2814  16.99 70.77 2.83 274,787 -2924 5.30 2.74 0.93 193.80 260,141 11.38 930.01 24.98 2825  16.99 67.71 2.70 275,149 -3012 5.35 2.71 0.99 183.21 259,068 11.38 930.82 25.11 1751  2  6  10  TBP f C]  TBP fC]  TBP [°C]  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45  -253 -237 -206 -166 -134 -101 -82 -49 -23 2 274 314 337 356 377 392  -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391  50  406  405  -253 -238 -208 -169 -137 -103 -87 -53 -42 -3 260 309 335 352 367 388 404  C4-(<177°C)  0.052  0.060  0.058  55  420  420  419  LGO(177-343°C)  0.264  0.259  0.264  60  0.576  0.580 0.101  0.591 0.087  65  439 442  437  HGO (343-524°C) 524+(>524°C)  440 445  70  465  75 80 85 90 92.5 95 96.5 98 99 100  482 490 513 526 539 552 556 615 741 866  459 471 486 496 525 537 550 556 616 744 871  453 467 484 494 520 527 548 556 599 735 872  Fraction Distribution Data Volume fraction  0.108  441  10 \  Ot 0  ,  ,  ,  100  200  300  ,  ,  ,  400 500 600 Temperature, °C  r  ,  ,  1  700  800  900  1000  Figure VI-5 Effect of number of Sheds trays on Scrubber Overhead TBP curve 94  Chapter 6 — Case Studies: Results and  Discussion  Table VI-3 Effect of number of Sheds trays on Scrubber Bottom properties Number of Sheds' trays Temperature fC] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kj/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kj/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  2  6  Cut Point [%]  375 17 649 822 738 -1331 3.39 2.89 0.00 1872.29 22,098 11.39 0.69 1043.55 0.79 1656 15.68 0.14 0.57  375 17 637 816 789 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57  17 634 811 827 -1326 3.41 2.90 0.00 1842.15 24,988 11.43 0.66 1035.18 0.78 1654 15.34 0.13 0.53  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70  411 440 466 489 512 517 523 542 555 556 563 592 602 629 634 681 684 690 699 740 749  412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706  412 441 480 495 512 515 517 522 526 541 547 558 590 597 626 633 681 684 691 701  10  Volume fraction  6  TBP [°C]  375  Fraction Distribution Data  2  10  TBP [°C]  TBP [°C]  743  742  C4-(<177°C)  0.000  0.000  0.000  75  753  750  749  LGO(177-343°C)  0.000  0.000  0.000  80  762  760  758  HGO (343-524°C) 524+(>524°C)  0.064  0.089  0.091  85  812  807  804  0.936  0.911  0.909  90  880  852  822  92.5 95 96.5 98 99 100  895 941 964 1034 1048 1056  892 917 964 1031 1047 1055  889 907 963 1028 1046 1055  100  400  500  600  700 800 Temperature, °C  900  1000  1100  Figure VI-6 Effect of number of Sheds trays on Scrubber Bottom TBP curve 95  Chapter 6 - Case Studies: Results and Discussion • 2 Sheds' trays: Light Ends: 20%; Water: 64% ; 100> fraction: 16% • 6 Sheds' trays: Light Ends: 20%; Water: 65% ; 100> fraction: 15% • 10 Sheds' trays: Light Ends: 20%; Water: 65% ; 100> fraction: 15%  200-300  300-400  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  1000>  Figure VI-7 Effect of number of Sheds trays on Scrubber Overhead composition  • 2 Sheds' trays: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 6 Sheds' trays: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 10 Sheds' trays: Light Ends: 0% ; Water: 1%; 100> fraction: 99%  0.00 -i 200-300  300-400  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  1000>  Figure VI-8 Effect of number of Sheds trays on Scrubber Bottom composition  96  Chapter 6 - Case Studies: Results and Discussion  Discussion: Increasing the number of trays up to 6 in the Sheds column means better separation ability better contact between vapour and liquid, removal of heavy fractions from the vapour and lighter fractions from the liquid. That is why Shed Vapour mass first increases - it is losing small amounts of heavy fractions (condensation), but is probably getting higher amounts of lighter fractions (evaporation). This vapour is being quenched with HGO Wash stream, producing more liquids, and less Overhead product. But the composition of Shed Vapour definitely affects the properties of the Overhead: its density and average molecular weight decrease, and composition shows higher presence of middle fractions, but lower of heavier fractions. Grid Liquid amount is increased as a consequence of higher Shed Vapour amount. Shed Liquid is losing lighter fractions and getting heavier, and its mass flow rate increases as well. Since in this stage, the evaporation of lighter fractions is dominant, due to the heat of evaporation, the temperatures along the Scrubber drop. Only Shed Bottom temperature increases radically. Liquid in the Sheds has to pass more contacting stages to reach the bottom, contacting hot vapours. When it reaches the bottom, its temperature is raised significantly. Better separation removes heavy fractions from the Overhead and directs them to the Bottom. That can be seen from the properties and compositions of Scrubber Overhead and Bottom. Further increasing the number of trays, from 6 to 8 or 10, changes the situation. The results suggest that most of the light fractions have already been evaporated from the liquids, but not all heavier fractions have been condensed from the vapours. Vapours lose heavy components, losing mass flow rate (Shed Vapour), while liquids gain mass. Since the evaporation is decreased, temperature in the system does not drop any more - at some positions temperature even starts to rise.  97  Chapter 6 - Case Studies: Results and Discussion  VII. Number of Grid Sections As was mentioned in the introductory description of Core operation blocks, the Koch Grid is simulated as a packed column with 2 packing sections, with the same diameter and height as the plant packed section. The overall efficiency is determined to be 78%. In this case study the number of section has been changed from 2 to 5 and 10, with no change of efficiency and the effect on Scrubber parameters and stream properties have been studied. Observations: • Temperature profile:  - B y changing the number of Grid sections from 2 to 5, only Shed Bottom temperature increases by 13°C, while all other temperatures remain almost the same. Further increasing the number of sections has no significant effect on temperature change along the reactor. (Table VII-1, Figures VII-1 and VII-2). 500  500  - 2 Grid sections  480  -o - - 5 Grid sections -10 Grid sections  460 - Grid Top 440  - Grid Bottom - Shed Top - Shed Bottom  420 A  - Scrubber Pool  400  380  360  360 4  6  8  Number of Grid sections  10  10  20  30  40  50  Position, feet from bottom of scr. pool  Figure VII-1 Effect of number of Grid  Figure VII-2 Effect of number of Grid  sections on temperatures along the  sections on temperature profile along the  Scrubber  Scrubber  98  Chapter 6 - Case Studies: Results and Overhead  Discussion  properties:  - B y increasing the number of Grid sections from 2 to 5 actual volume flow of the Scrubber Overhead drops by 0.2% and mass flow rate increases by 2%. (Table VII-1, Figure VII-3). -Density increases from 2.83 to 2.89 kg/m (Table VII-1). 3  -Average molecular weight increases from 71 to 72 (Table VII-2). -Composition shows slightly lower presence of fractions up to 500°C, and higher presence of heavier fractions (Table VII-2, Figures VII-5 and VII-7). -Further increasing the number of sections from 5 to 10 has no significant effect on any of the properties. Scrubber Bottom  properties:  - B y changing the number of Grid sections from 2 to 5 actual volume and mass flow rate drop by 8% and 7%, respectively (Table VII-1). - Density drops from 816 to 822 kg/m (Table VII-1). 3  -Average molecular weight changes from 637 to 650 (Table VII-3). -Composition shows lower presence of components boiling up to 600°C, while heavier fractions are more concentrated (Table VII-3, Figures VII-6 and VII-8). -Change in number of Grid sections from 5 to 10 has no effect on Bottom properties. Other.  -While changing Grid section number from 2 to 5, volume and mass flow of all three streams (Grid and Shed Liquid, and Shed Vapour) decrease by 15%, 1% and 8%, respectively. Further change in number of sections has no effect. (Figure VII-4).  99  Chapter 6 - Case Studies: Results and Discussion  Figure VII-3 Effect of number of Grid  Figure VII-4 Effect of number of Grid  sections on mass flow rate of Scrubber  sections on mass flow rate of other  Overhead and Bottom  streams  100  Chapter 6 - Case Studies: Results and Discussion  Table VII-1 Effect of number of Grid sections on Scrubber parameters Position  ft from the pool bottom  Koch Grid Top Koch Grid Bot. Sheds Top Sheds Bot. Scrubber Pool  Base Case 2  43 38 34 22 0  Top Stage Temp Est Bottom Stage Temp Est Top Stage Temp Est Bottom Stage Temp Est Bulk Liquid Temperature Note: Overall Grid section efficiency is 0.75  Number of Grid sections 5 10  393 395 405 473 375  393 397 406 486 375  % -0.1 0.6 0.2 2.7 0.0  393 397 406 486 375  % -0.1 0.6 0.2 2.6 0.0  -0.2 2.0  274,318 793,567  -0.2 1.9  Flow Rates& Densities Scrubber Overhead  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  274,787 778,651  274,318 793,996  2.83  2.89  789 643,728  725 595,823  -8.2 -7.4  726 596,652  -8.0 -7.3  816  822  0.8  822  0.8  2.58  2.63  2.0  2.63  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  Mass Density (kg/m ) 3  Overhead / A T B  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  Shed Vapor  Actual Volume Flow (m /h) Mass Flow (kg/h)  Shed Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  To Coker  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  3  305 217,861  Mass Density (kg/m )  1.9 259 -14.8 185,481 -14.9  714.77  0.0  714.82  0.0  276,845 860,927  -0.2 -1.1  276,852 860,709  -0.2 -1.1  3.14  3.11  -0.9  3.11  -0.9  497 361,323  461 332,768  -7.1 -7.9  461.72 333,193  -7.1 -7.8  724.76  718.81  -0.8  719.11  -0.8  330 269,078  303 249,054  -8.2 -7.4  303 249,400  -8.0 -7.3  815.85  822.17  0.8  822.05  0.8  3  Mass Density (kg/m ) 3  3  2.1  714.97  Mass Density (kg/m ) 3  259 -15.0 185,180 -15.0  2.89  277,329 870,450  3  3  2.1  J  Additional information Vapour to Sheds  Temperature (°C)  514  517  0.7  517  0.6  Upgoing Stream Sheds Koch Grid SPL Coler  Temperature (°C) Stage Efficiency Stage Efficiency Duty (MMBtu/h)  534 0.53 0.75 44.72  534 0.53 0.75 45.46  0.0 0.0 0.0 1.7  534 0.53 0.75 45.38  0.0 0.0 0.0 1.5  101  Chapter 6 — Case Studies: Results and  Discussion  Table VII-2 Effect of number of Grid sections on Scrubber Overhead properties Number of Grid sections  2  Temperature [°C]  5  10  Cut Point [%]  2  5  393  393  393  Pressure [psig]  16.99  16.99  16.99  0  -253  -252  -252  Molecular Weight  70.77  71.96  71.93  1  -237  -226  -226  2.83  2.89  2.89  2  -206  -200  -200  274,787  274,318  274,310  3.5  -166  -164  -163  -2924  -2895  -2896  5  -134  -133  -132  Mass Entropy [kJ/kg-C]  5.30  5.26  5.26  7.5  -101  -100  -100  Mass Heat Capacity [kJ/kg-C]  2.74  2.74  2.74  10  -82  -78  -78  Phase Fraction ( Mass Basis)  0.93  0.90  0.90  12.5  -49  -48  -47  193.80  197.47  197.37  15  -23  -16  -15  260,141  260,885  260,861  17.5  2  2  2  11.38  11.37  11.37  20  274  279  277 315  Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kj/kg]  Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3]  T B P [°C]  10  T B P [°C]  TBP fC]  930.01  932.00  931.96  25  314  315  Molar Volume [m3/kgmole]  24.98  24.86  24.86  30  337  345  348  Mass Heat of Vap. [kJ/kg]  2825  2815  2815  35  356  357  358  Fraction Distribution Data Volume fraction  40  377  380  382  45  392  399  400  50  406  409  410  C4-(<177°C)  0.060  0.050  0.050  55  420  422  423  LGO(177-343°C)  0.259  0.250  0.250  60  440  442  444  H G O (343-524°C)  0.580  0.579  0.579  65  445  454  454  524+ (>524°C)  0.101  0.121  0.121  70  465  475  474  75  482  484  484  80  490  506  508  85  513  515  518  90  526  532  534  92.5  539  542  545  10 0  95  552  558  558  96.5  556  571  571  98  615  623  623  99  741  746  746  100  866  866  867  1 1  ,  ,  ,  0  100  200  300  ,  ,  ,  400 500 600 Temperature, °C  , 700  ^_ 800  ,  1  900  1000  Figure VII-5 Effect of number of Grid sections on Scrubber Overhead TBP curve 102  Chapter 6 - Case Studies: Results and Discussion  Table VII-3 Effect of number of Grid sections on Scrubber Bottom properties Number of Sheds' trays  Temperature ["C] Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kJ/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  2  5  Cut Point [%]  10  375  375  375  17.00 637.04 815.85 812 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57 0.78  17.00 650.69 822.17 746 -1331 3.39 2.89 0.00 1877.95 21,651 11.40 0.69 1043.70 0.79 1332 15.70 0.14 0.57 0.79  17.00 650.44 822.05 747 -1331 3.39 2.89 0.00 1877.33 21,689 11.40 0.69 1043.61 0.79 1331 15.69 0.14 0.57 0.79  Fraction Distribution Data  2  5  TBP [°C] 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60  10  TBP [°C]  TBP  pC]  419 446 474 494 511 520 531 532 543 564 583 592 598 632 638 642 687 692 696 707  419 446 474 493 510 520 530 532 542 562 581 590 598 631 636 642 686 690 696  65  412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706  70  743  747  746 753  705  C4-(<177°C)  0.000  0.000  0.000  75  750  754  LGO(177-343°C)  0.000  0.000  0.000  80  760  765  765  HGO (343-524°C) 524+(>524°C)  0.089  0.056  0.056  85  807  809  809  0.911  0.944  0.944  90 92.5 95 96.5 98 99 100  852  855  854  892 917 964 1031 1047 1055  895 919 989 1032 1048 1056  893 917 989 1032 1048 1056  100  1100 Temperature, °C  Figure VII-6 Effect of number of Grid sections on Scrubber Bottom TBP curve  103  Chapter 6 - Case Studies: Results and Discussion  0.50  • 2 Grid Sections: Light Ends: 20%; Water: 65% ; 100> fraction: 15% • 5 Grid Sections: Light Ends:20%; Water: 65% ; 100> fraction: 15%  0.40  • 10 Grid Sections: Light Ends:20%; Water: 65% ; 100> fraction: 15%  V)  c o 0.30 o 2 •50.20  2  0.10  0.00  TO 200-300  300-400  ID  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  1000=  Figure VII-7 Effect of number of Grid sections on Scrubber Overhead composition  • 2 Grid Sections: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 5 Grid Sections: Light Ends: 0% ; Water: 1%; 100> fraction: 99% O 10 Grid Sections: Light Ends: 0% ; Water: 1% 100> fraction: 99%  200-300  300-400  In  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  1000>  Figure VII-8 Effect of number of Grid sections on Scrubber Bottom composition 104  Chapter 6 - Case Studies: Results and  Discussion  Discussion: When the number of Grid sections increases from 2 to 5 that means better contact between rising vapour through the Grid and falling liquid, mainly HGO Wash. Fraction 500-600°C from HGO evaporates and ends up in vapour. Mass flow rate, density and average molecular weight of Overhead rises, and composition shows higher presence of these heavier fractions. Liquids in the Scrubber contain less middle fractions; Heavy fractions are more concentrated, causing higher density and average molecular weight of the Scrubber Bottom. Mass flow rate of all liquids along the Scrubber decreases. Lower mass flow rate of the Grid Liquid causes lower flow rate of the Shed Vapour. Also, less liquid contacts the hot vapours from the cyclones, and the Shed bottom temperature rises. Except Shed bottom temperature, all temperatures along the Scrubber do not change significantly. Further increasing the number of Grid sections beyond five has no effect on Scrubber parameters or stream properties.  105  Chapter 6 - Case Studies: Results and Discussion  VIII. Simulation of the Conditions from Start of Run to End of Run In this case study, the effect of the conditions (absolute pressure and the pressure drop) at the start of run (SOR) and the end of run (EOR) of the Fluid Coker, explained in Section 3.2.3 have been investigated. At the SOR the pressure in the Scrubber is set to be 17 psig (117.21 kPa), and pressure drop along the grid 0.4 in of water (0.07 kPa). As the fouling of the grid and the cyclone nozzles due to the coke formation occurs, the pressure drop increases. In order to maintain sufficient production of the Overhead, pressure in the Scrubber is raised gradually from 17 psig (117.21 kPa) at the SOR to 27 psig (186.16 kPa ) at the EOR. Pressure drop increases from 0.4 in of water to 2.5 in of water (0.36 kPa). This change in pressure drop results from changes in hydrodynamics at the two pressures considered. Fouling itself is not directly accounted for. Accordingly, the pressure at the top of the grid changes from 16.6 psig (117.14 kPa) to 26.9 psig (185.53 kPa) . Note that Overhead production rate is not constant. a  The effect of change in absolute pressure in the Scrubber and pressure drop in the Grid has been simulated and effects on Scrubber temperature profile, Scrubber Overhead and Scrubber Pool Liquid flow rates, composition and properties have been investigated. Observations: • Temperature profde:  - A s the pressure drop increases, all temperatures along the Scrubber slightly increase (a few degrees). Only Sheds bottom temperature drops for 6°C (Table VIII-1, Figures VIII-1 andVIII-2).  a  ) In this case study, units for the pressures and pressure drops are adapted to the commonly  used units in the plant. For conversion: 1 psig = 6.8948 kPa 106  Chapter  6 - Case Studies:  Results  and  -•- -Pdrop0.4in; Abs. P=17psig *—Pdrop=0.81 in; Abs. P=19psig • --Pdrop=1.23in; Abs P=21 psig Pdrop=1.65in; Abs P=23 psig Pdrop2.07in; Abs.P=25 Pdrop2.5in; Abs.P=27  480  460 - Grid Top - Grid Bottom 440  e  - Shed Top  3  - Shed Bottom  1  - Scrubber Pool  8. £  Discussion  420  380  0.00  0.50 17  1.00 19  2.00 21  23  2.50  25  27  10  Top scale: P r e s s u r e d r o p in Grid, in. of water Bottom s c a l e : Absolute p r e s s u r e in the S c r u b b e r , psig  20  30  40  50  Position, feet from bottom of scr. pool  Figure VIII-1 Effect of pressure drop in  Figure VIII-2 Effect of pressure drop in  the Grid and absolute pressure in the  Grid  Scrubber  Scrubber on  on  temperatures  along  the  Scrubber  •  and  absolute  pressure in the  temperature profde along  the Scrubber  Overhead  properties:  -Actual volume and mass flow rate of the Scrubber Overhead decrease gradually by 24% and 4.4%, respectively, based on the SOR case (Table VIII-1, Figure VIII-3). -Density increases from 2.83 to 3.58 kg/m (Table VIII-1). 3  -Average molecular weight drops from 71 to 63 (Table VIII-2). -Composition shows lower presence of 400°C + fractions (Table VIII-2, Figures VIII-5 and VIII-7). •  Scrubber  Bottom  properties:  -Actual volume and mass flow rate increase by 15% and 13% (Table VIII-1).  107  Chapter 6 - Case Studies: Results and Discussion  - Density drops from 816 to 802 kg/m (Table VIII-1). 3  -Average molecular weight changes from 637 to 610 (Table VIII-3). -Composition shows higher presence of middle fractions (400-600°C) (Table VIII-3, Figures VIII-6 and VIII-8). •  Other.  -Sheds Vapour and Grid Liquid mass flow decrease, while Liquid from the Sheds actual volume and mass flow increase (Figure VIII-4).  Figure  V I I I - 3  Effect of pressure drop  Figure  V I I I - 4  Effect of pressure drop  in the Grid and absolute pressure in the  in Grid and absolute pressure in the  Scrubber on mass flow rate of the  Scrubber on mass flow rate of other  Scrubber Overhead and Bottom  streams  108  Chapter 6 - Case Studies: Results and Discussion  v  Table VIII-1 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber parameters KxhGid Top Fressuie (psig) BaoTiFtBssue (psig) ftfiftkn ft FtessureDop( in of water) pod uottcrri FreasureDcpfpa) KxtiGidTcp 43 Tqp Sags Tsip Est (t)  BED Case 1699 17 040 0.014  1857 19 081 0029  %  2096 21 1.23 0.044  393  394  02  FtesureorcpintreKxhGrid 2294 2493 23 25 1.65 207 % Q06 % Q075  394  02  395  395  0.5  %  26.91 27 250 Q09  %  Q5  395  Q5  Hb+iQidEtt. 38  BrJt0Ti9ar^Termaft)  395  395  Q1  396  Q2  396  Q3  397  Q4  04  34  Tqp Stags Torp Est CQ  397  STedsTqp  405  405  QO  403  02  406  Q3  406  Q3  407  Q4  STSOSBI  22  EdtDrnSageTerrp Est CQ BJk Liqud Terperatije ft)  473  471  •0.4  470  -0.7  459  -1.0  468  -12  467  -1.4  375  375  QO  375  QO  375  0.0  375  Q0  375  QO  Soxtba-Pod  0  HojvFaes& Densities  SoUtja-aetead  A i d VflirreRa/vtrrfTh)  Uqj'd  Ad^VfluTCRcwfrrf/h) IVassHcw(kgfri)  255259 -60 771,718 -0.9  213,570 -11.4 764,769 -1.8  23Q4C -161 757,6« -27  218600 -2Q4 750,931 -36  28!  293 55  314 10.8  207,931 -24.3 744,287 AA  3Z 160  344 212  353 263  78E 64372c  813 30 66Q619 26  836 59 676,903 52  85E 89 698,319 7.7  882 11.8 703,53: 102  906 14.6 725,577 127  IvassfinatyWrrT)  816  813 -0.4  810 -0.7  807 -1.1  80E -1.4  802 -1.7  CvBtead/ATB  fVtesflcwFaio  25  255 -0.9  25; -1.8  251 -27  249 -36  247 4.4  QidUqJd  AdijalVtiLrreRo«(rrf/h) lvB5sRcw(kgh)  303 -04 216233 -Q7  302 -0.9 214,747 -1.4  301 -1.1 213606 -20  300 -1.4 212,529 -24  300 -1.7 211,580, -29  M3ssRow(kgh) ^te^tyfkc/rr )  274,787 778,651  3  SQttbPod  Nte irBity(kg'rr ) r  SnedVfepcr  3  AJi^\<tiLneRcw(rr /h) 3  M3ssRcw(kgti) 3-BrJUQjd  Adu=l\fluTBRo/v(rr?/h) M3ssRow(kg4i)  ToOter  A u^\^LrreRcw(rr 7h) fVB3sRcw(kgh)  KteD&^flg'rr?) +  3  3D: 217,831 714.97  71282 -Q3  710.83 -Q6  7031C -1.2  26Q753 -60 851,£S -1.0  243,023 -11.3 863,457 -20  70915 -08 232,791 -161 84519C -29  707.54 -1.0  277,32 87Q45C  22Q921 837,399 -38  314  331 53  347 10.5  363 157  21Q16E -242 829,807 A.7  379 20.8  497 361,323  51C 27 33904 22  521.83 50 376,88: 4.3  533 7.4 384,44E 64  54E 97 391,949 85  567 120 399316 1Q5  724.7E  72363 -02  72224  -Q3  72Q6E -06  71907 -08  717.46 -1.0  33C 26907t  34C 30 276,139 26  3*  59 282,946 52  359 89 28Q820 7.7  369 11.8 293,607 1Q2  37E 14.6 308291 127  8158E  81279 -Q4  80998 -Q7  807.2c -1.1  804.63 -1.4  80211 -1.7  39E 258  AcUtiorol irfarnation \*por to Shads  Ta-rp=ratLre(C)  514  513  -02  512  -Q3  512  -05  511  -Q6  510  -Q7  IpgaraSjBam Sheds KrfiGid SFLOda-  TerparalrefC) SageEffidary Stage Efiidaxy rOfy(MvButi)  534 Q53 Q75 44.72  534 Q53 Q75 4565  -Q1 QO QO 21  533 Q53 Q75 4681  •02 QO QO 4.7  533 Q53 Q75 4812  -02 0.0 QO 7.6  533 Q53 Q75 4941  -03 QO QO 1Q5  532 Q53 Q75 5072  -Q3 QO QO 134  109  Chapter 6 - Case Studies: Results and Discussion  Table VIII-2 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead properties Pressure drop in the Grid, in. water Pressure in the Scrubber, psig Temperature [ C] Pressure [psig] Molecular Weight Mass Density [kg/m 3] Act. Volume Flow [m3/h] Mass Enthalpy [kj/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapour Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kj/kg]  17 393 16.99 70.77 2.83 274,787 -2924 5.30 2.74 0.93 193.80 260,141 11.38 930.01 24.98 2825  395 26.91 62.93 3.27 207,839 -3179 5.49 2.71 0.91 170.42 255,492 11.43 921.04 19.23 2853  Volume fraction 0.060 0.259 0.580 0.101  0.067 0.268 0.564 0.100  U  Fraction Distribution Data C4-(<177 °C) LGO (173-743 °C) HGO (343-524 °C) 524+ (>524 °C)  Gut Point [%]  2.5 27  0.4  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  2.5 27 TBP["C] -253 -239 -209 -171 -140 -104 -90 -57 -42 -6 244 305 325 350 365 379 395 407 421 440 447 466 482 493 525 536 549 555 615 743 869  0.4 17 TBP["C] -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 550 556 616 744 871  100 90 80 70  •Pdrop 0.4 in; Abs. P=17 psig  60  Pdrop 2.5 in; Abs. P=27 psig  » 50  5  40 30 20 10 0  100  200  300  400 500 600 Temperature, °C  700  800  900  1000  Figure VIII-5 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead TBP curve  110  Chapter 6 - Case Studies: Results and  Discussion  Table VIII-3 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom properties P r e s s u r e d r o p in t h e G r i d , in. w a t e r  0.4  2.5  P r e s s u r e in t h e S c r u b b e r , p s i g  17  27  Temperature [ C] U  Cut Point [%]  375 27.00 610.37 802.11 905 -1324 3.40 2.91 0.00 1777.47 28,107 11.44 0.69 1028.41 0.76 1654 14.91 0.13 0.55 0.76  Volume fraction 0.000 0.000 0.089 0.911  0.000 0.000 0.152 0.848  Fraction Distribution Data  C4-(<177°C) LGO (177-343°C) HGO (343-524°C) 524+ (>524°C)  2.5  17  375 17.00 637.04 815.85 789 -1329 3.39 2.89 0.00 1843.61 23,893 11.41 0.69 1038.96 0.78 1309 15.47 0.13 0.57 0.78  Pressure [psig] Molecular Weight Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kj/kg] Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Phase Fraction ( Mass Basis) Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass Heat of Vap. [kj/kg] Surface Tension [dyne/cm] Thermal Conductivity [W/m-K] Viscosity [cP]  0.4  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  27  TBP [°C] 412 441 466 490 504 514 517 520 525 548 556 564 592 605 630 635 682 684 693 706 743 750 760 807 852 892 917 964 1031 1047 1055  TBP [ C] 389 423 452 466 485 492 505 513 515 517 520 527 556 584 597 626 634 681 685 693 724 745 754 787 821 885 904 959 1023 1038 1053 U  100  400  500  600  700  800  900  1000  1100  T e r n p e r a t u r e , °C  Figure VIII-6 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom TBP curve 111  Chapter 6 - Case Studies: Results and Discussion  • Pdrop 0.4 in; Abs. P=17 psig: Light Ends: 20%; Water: 65% ; 100> fraction: 15%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  Components' Boiling Temperatures Range, °C  Figure VIII-7 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead composition  • Pdrop 0.4 in; Abs. P=17 psig: Light Ends: 0% Water: 1%; 100> fraction: 99% U Pdrop 2.5 in; Abd. P=27 psig: Light Ends: 0% ; Water: 1%; 100> fraction: 99%  200-300  300-400  400-500  500-600  600-700  700-800  800-900  900-1000  1000>  Components' Boiling Temperatures Range, °C  Figure VIII-8 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom composition 112  Chapter 6 — Case Studies: Results and Discussion  Discussion: In this case study, all changes are caused by increase in absolute pressure in the Scrubber, not by pressure drop along the Grid. As the higher pressure is applied in the whole Scrubber, less evaporation occurs both in the tray column (Sheds) and the Koch Grid. Due to the less heat of evaporation spent, the temperatures in the whole system slightly increase. Less vapour and more liquid are produced. Accumulation of the liquid occurs mostly in the bottom of the Sheds, cooling the rising vapour more effectively. That is the reason why only Shed bottom temperature drops by 6°C. Gas volume is very responsive to pressure and the raise in Scrubber Overhead density is due to the increase in pressure, not to the change in composition. Table VIII-2 shows that the average molecular weight of the Overhead is lower at the EOR than at the SOR, what would lead to drop of density. In this case the pressure effect on density is dominant. Figure VIII-7 shows that at higher pressures in the Scrubber, less heavy components (above 400°C NBP) are present in the Overhead. Their ability to evaporate is reduced. Decrease in liquid's densities is due to the fact that as the pressure increases, less heavy (above 400°C NBP) components evaporate and go to the Overhead, and more fractions under 500°C are present in the liquid. Heavy fractions are still present in the liquid in the same amount, but their percentage in the total liquid is lower, because of the presence of these lighter fractions. This is shown in Figure VIII-8. Also, the average molecular weight of the Scrubber Pool Liquid is lower, due to the same reason (Table VIII-3).  113  Chapter 6 - Case Studies: Results and  Discussion  IX. Water Instead of HGO Underwash HGO Underwash stream enters the Scrubber under the Grid and serves to help HGO Wash stream to scrub particulates and heavy components from rising vapour as well as to cool it down in order to prevent fouling of the Grid. In this case study water at 40°C is used instead of HGO Underwash. The goal is to decrease Grid entrance (Grid Bottom) temperature by 10-20°C, in order to reduce fouling in the Grid. If Scrubber Overhead temperature was not controlled while water is applied, it would be possible to achieve a significant drop in Grid entrance temperature, but in that case the Overhead temperature gets too low. For that reason Overhead temperature has to be controlled either by HGO Wash flow rate, or ATB flow rate. Both cases are considered. If water flow rate is 700 barrel/day (~4600kg/h), the temperature profile remains similar to the original case with the HGO Underwash, without control. Higher water flow rate (2.5 kbarrel/day = -17000 kg/h), controlled by HGO or ATB flow rate, decreases Grid entrance temperature by 5 and 7°C, respectively, keeping the Overhead temperature constant. In order to control the Overhead temperature, ATB actual flow rate should be deceased by 2.6% (from 55 kbarrel/day to 52.5 kbarrel/day) and HGO Wash actual flow rate by 54% (from 24 kbarrel/day to 11 kbarrel/day). These three cases are summarized below. Observations: • Temperature  profile:  -With 700 barrel/day of water instead of HGO Underwash, without any control of Overhead temperature, Scrubber temperature profile remains almost the same as in the original case. Shed Bottom temperature increases by 10°C. -2.5 kbarrel/day of water is able to decrease Grid Bottom temperature by 5°C, if the Overhead temperature is controlled by ATB flow rate. Shed Top temperature is also several degrees lower and Shed Bottom temperature is 11 degrees higher. - 2.5 kbarrel/day of water, with the HGO control, shows better ability to decrease Grid entrance temperature, while keeping the Overhead temperature constant (Table IX-1, Figures IX-1 and IX-2).  114  Chapter 6 - Case Studies: Results and Discussion  500 •  500  480  480  460  460  - Grid Top  440 A  0 kbarrel/day water, 10 kbarrel/day HGO 0.7 kbarrel/day w ater, uncontrolled 2.5 kbarrel/day w ater, controlled! by ATB 2.5 kbarrel/day w ater, controlled by HGO wash  440  - Grid Bottom - Shed Top  420 •  - Shed Bottom  420 4  - Scrubber Pool 400  400  380  380  360  360  2 3 Cases: 1: Base case- 0 kbarrel/day water (10 kbarrel/day HGO); 2: 0.07 kbarrel/day water unontr.; 3:2.5 kbarrel/day water contr. by ATB; 4: 2.5 kbarrel/day water contr. by HGO wash  10  20  30  40  Position, feet from bottom of scr. pool  Figure IX-1 Effect of water instead of HGO  Figure IX-2 Effect of water instead of HGO  Underwash on temperatures along the  Underwash on temperature profile along the  Scrubber  Scrubber  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases.  • Overhead properties: -In three cases (700 barrel/day of water, uncontrolled, 2.5 kbarrel/day of water, controlled by HGO and 2.5 kbarrel/day of water, controlled by ATB, actual volume of the Scrubber Overhead increases by 1.6, 6.5 and 5.6% and mass flow rate decreases by 3.5, 3.9 and 10.4%, respectively. (Table IX-1, Figure IX-3). -Density changes from 2.83 to 2.69, 2.56 and 2.40 kg/m (Table IX-1). 3  115  Chapter 6 - Case Studies: Results and  Discussion  -Average molecular weight decreases from 71 to 67, 64 and 60 (Table IX-2). - A l l three cases show lower presence of middle fractions (300-500°C), but higher content of both lighter and heavier fractions (Table IX-2, Figures IX-5 and IX-7). Scrubber Bottom  properties:  -700 barrel/day of water, uncontrolled, 2.5 kbarrel/day of water, controlled by HGO and 2.5 kbarrel/day of water, controlled by ATB cases make actual volume of Scrubber Bottom to drop by 8.6, 26.5 and 13.9%, respectively. Mass flow rate drops by 7.8, 26 and 13% (Table IX-1). - Density changes from 816 to 823, 822 and 824 kg/m (Table IX-1). 3  -Average molecular weight changes from 637 to 654, 655, 664 (Table LX-3). -Composition shows lower presence of fractions up to 600°C, while heavier fractions are more concentrated (Table LX-3, Figures IX-6 and IX-8). Other.  -Grid Liquid volume and mass flow rate slightly increase for first two cases, while for the last one both radically drop. -Shed Vapour and Shed Liquid mass flow rate drop for all three cases, compared to original one (Figure IX-4).  116  Chapter 6 - Case Studies: Results and Discussion  780  900  760  850  740  800  720 \  750  700 680 -I  - Overhead  700  660  - Scrubber Bottom  650  -I  -I  640  600 -  620  550 -  - Shed Vapor  500  - S h e d Liquid  600  -I  580  450  560  400  - Grid Liquid  540 520 500 480 460 440 420 1 C a s e s : 1: Base c a s e - 0 kbarrel/day water (10 kbarrel/day HGO); 2: 0.07 kbarrel/day water unontr.; 3: 2.5 kbarrel/day water contr. by ATB; 4: 2.5 kbarrel/day water contr. by HGO w a s h  C a s e s : 1: Base c a s e - 0 kbarrel/day water (10 kbarrel/day HGO); 2: 0.07 kbarrel/day water unontr.; 3: 2.5 kbarrel/day water contr. by A T B ; 4: 2.5 kbarrel/day water contr. by HGO w a s h  Figure IX-3 Effect of water instead of HGO  Figure IX-4 Effect of water instead of  Underwash on mass flow rate of Scrubber  HGO Underwash on mass flow rate of  Overhead and Bottom  other streams  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases.  117  Chapter 6 - Case Studies: Results and Discussion  Table IX-1 Effect of water instead of HGO Underwash on Scrubber parameters  Position  ft from the pool bottom  Koch Grid Top  43  Water Flow Rate m3/h kbarrel/day Case:  Water flow rate (instead of HGO Underwash) (barrel/day) 0 3 12 12 0 (10 kbl/day HGO) 0.70 2.50 2.50 (Base Case) Uncontrolled Controlled by ATB Controlled by HGO 1 2 % 3 % 4  Top Stage Temp Est (°C)  393  393  0.0  Koch Grid Bot. 38  Bottom Stage Temp Est (°C)  395  396  Sheds Top  Top Stage Temp Est (°C)  405  406  Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)  473  483 375  34  Sheds Bot. 22 Scrubber Pool 0  375  %  389  -1.1  390  -0.8  0.3  390  -1.2  387  -2.1  0.3  401  -1.0  395  -2.4  2.1 0.0  484  2.3 0.0  483 376  0.3  375  2.0  Flow RatesS Densities Scrubber Overhead  Actual Volume Flow (rrfVh) Mass Flow (kg/h)  274,787 778,651  279,116 1.6 751,648 -3.5  292,649 6.5 748,329 -3.9  290,197 5.6 697,520 -10.4  2.83  2.69 -5.0  2.56 -9.8  2.40 -15.2  789 643,728  721 -8.6 593,311 -7.8  580 -26.5 476,648 -26.0  679 -13.9 559,999 -13.0  816  823 0.9  822 0.8  824 1.0  Mass Density (kg/m) 3  Scrubb.Pool Liquid  Actual Volume Flow (m/h) Mass Flow (kg/h) 3  Mass Density (kg/m) 3  Overhead / ATB  Mass Flow Ratio  2.58  2.49 -3.5  3.15 22.2  2.31 -10.4  Grid Liquid  Actual Volume Flow (m/h) Mass Flow (kg/h) Mass Density (kg/m)  305 217,861 714.97  309 1.3 220,857 1.4 715.70 0.1  318 4.2 227,254 4.3 715.74 0.1  202 -33.7 146,778 -32.6 726.36 1.6  Actual Volume Flow (m/h) Mass Flow (kg/h)  277,329 870,450  281,729 1.6 846,437 -2.8  295,180 6.4 849,522 -2.4  290,634 4.8 798,526 -8.3  3.14  3.00 -4.3  2.88 -8.3  2.75 -12.5  497 361,323  455 -8.4 329,985 -8.7  384.31 -22.6 277,213 -23.3  423 -14.9 308,849 -14.5  724.76  725.14 0.1  721.32 -0.5  330 269,078 815.85  301 -8.6 248,004 -7.8 822.97 0.9  242 -26.5 199,239 -26.0  730.15 0.7 284 -13.9 234,080 -13.0 824.18 1.0  3  3  Shed Vapor  3  Mass Density (kg/m) 3  Shed Liquid  Actual Volume Flow (m/h) Mass Flow (kg/h) 3  Mass Density (kg/m) 3  To Coker  Actual Volume Flow (m/h) Mass Flow (kg/h) Mass Density (kg/m) 3  822.45 0.8  Additional information Vapour to Sheds  Temperature (°C)  514  518  0.7  514  0.0  518  0.9  Upgoing Stream Sheds Koch Grid SPL Coler  Temperature (°C) Stage Efficiency Stage Efficiency Duty (MMBtu/h)  534 0.53 0.75 44.72  535 0.53 0.75 44.26  0.1 0.0 0.0 -1.0  534 0.53 1.00 42.74  0.0 0.0 33.3 -4.4  535 0.53 1.00 39.06  0.1 0.0 33.3 -12.7  118  Chapter 6 - Case Studies: Results and Discussion Table IX-2 Effect of water instead of HGO Underwash on Scrubber Overhead properties Water flow rate  0.7 kbpd H 0  Basic  2  (10 kbpd HGO Wash) Uncontrolled Temperature fC]  393  2.5 kbpd H 0  2.5 kbpd H 0  2  2  (10 kbpd HGO Wash)  Controlled by ATBControlled by HGO 393  389  0.7 kbpd HjO  Basic  Cut Point [%]  Uncontrolled  TBP fC]  390  Pressure [psig]  16.99  16.99  16.99  16.99  0  -253  -253  Molecular Weight  70.77  67.47  63.58  60.07  1  -237  •239  2.83  2.70  2.56  2.40  2  •207  -209  282,860  287,383  301,246  298,723  3.5  -167  •171  Mass Enthalpy [kJ/kg]  •2924  •3056  -3265  -3414  5  •136  •140  Mass Entropy [kJ/kg-C]  5.30  5.44  5.56  5.75  7.5  -102  -104  Mass Heat Capacity [kJ/kg-C]  2.74  2.74  272  2.73  10  •85  •90  Vapor Phase Fraction (Mass Basis)  0.93  0.93  0.92  0.94  12.5  -56  193.80 260,141  184.96 263,719  173.16 278,271  164.05 274,537  15 17.5  •51 •34 -3  Mass Density [kg/m3] Act. Volume Flow [m3/h]  Specific Heat [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K  •43 -5  11.38  11.42  11.44  11.51  20  266  248  930.01  922.31  921.94  907.94  25  310  309  Molar Volume [m3/kgmole]  24.98  25.03  24.87  24.99  30  336  335  Mass Heat of Vap. [kJ/kg]  2825 1.05  2862 1.05  2899  2939 1.05  35 40  354 373  351 367  45  391  391  50  405  406  Liq. Mass Density (Std. Cond) [kg/m3]  1.05  Fraction Distribution Data Volume fraction C4-(<177°C)  0.060  0.055  0.073  0.074  55  420  420  LGO (177-343°C)  0.259  0.268  0.263  0.264  60  439  439  HGO(343-524°C)  0.580  0.569  0.558  0.554  65  442  444  524+(>524°C)  0.101  0.107  0.107  0.108  70  459  465  75 80  471 486  481 489  85  496  512  90  525  526  92.5  537  539  95  550  552  96.5  556  556  98  616  630  99  744  755  100  871  882  100 90 80  • O kbarrel/day water, 10 kbarrel/day HGO  70 • 0.7 kbarrel/day w ater, uncontrolled  60 | O  >  -2.5 kbarrel/day water, controlled by ATB  50  •2.5 kbarrel/day w ater, controlled by HGO wash  40 30 \ 20 10 \ 0  -I O  , 100  , 200  , 300  , 400  , 500  • 600 '  Temperature, °C  , 700  • 800  , 900  1 1000  Figure IX-5 Effect of water instead of HGO Underwash on Scrubber Overhead TBP curve 119  Chapter 6 - Case Studies: Results and Discussion Table IX-3 Effect of water instead of HGO Underwash on Scrubber Bottom properties Water flow rate  Basic  0.7 kbpd H 0  2.5 kbpd H 0  (10kbpd  Uncontrolled  Controlled  Controlled  by ATB  by HGO  2  HGO Wash) Temperature f C]  375  Pressure [psig]  2  375  375  2.5 kbpd HjO Cut Point  Basic  0.7 kbpd H 0 2500 kbpd H 0 2  (10 kbpd  [%]  Uncontrolled  HGO Wash) TBPfC]  375  TBP[°C]  2500 kbpd H 0  2  2  Controlled  Controlled  by ATB  by HGO  TBPfC]  TBPfC]  17.00  17.00  17.00  17.00  0  412  420  425  Molecular Weight  423  637.04  654.11  654.85  664.51  1  441  456  460  Mass Density [kg/m3]  458  815.85  822.97  822.45  824.18  2  466  485  487  486  812  741  597  699  3.5  490  509  512  512  -1329  -1330  -1330  -1326  5  504  514  514  514  Mass Entropy [kJ/kg-C]  3.39  3.39  3.40  3.41  7.5  514  519  519  519  Mass Heat Capacity [kJ/kg-C]  2.89  2.89  2.89  2.89  10  517  525  524  542 556  Act. Volume Flow[m3/h] Mass Enthalpy [kJ/kg]  Vapor Phase Fraction (Mass Basis) Specific Heal [kJ/kgmole-C] Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt]  0.00  0.00  0.00  0.00  12.5  520  1843.61  1887.95  1890.50  1920.13  15  525  555 556  551 556  23,893  21,417  17,210  19,926  17.5  548  559  558  563  11.41  11.40  11.40  11.41  20  556  564  564  591  558  0.69  0.69  0.71  0.69  25  564  592  592  597  1038.96  1044.56  1043.80  1046.30  30  592  605  604  626  Molar Volume [m3/kgmole]  0.78  • 0.79  35  605  630  630  632  1309  1685  0.80 1703  0.81  Mass Heat of Vap. [kJ/kg]  1707  40  630  635  635  Surface Tension [dyne/cm]  680  15.47  15.72  15.72  1572  45  635  681  681  683  0.13  0.14  0.14  0.14  50  682  684  684  685  0.57  0.57  0.58  0.57  55  684  691  691  693  60  693  700  701  704  65  706  740  740  742  Liq. Mass Density (Std. Cond) [kg/m3]  Thermal Conductivity [W/m-K] Viscosity [cP]  ,.  Fraction Distribution Data Volume fractior C4-(<177°C)  0.000  0.000  0.000  0.000  70  743  746  749  748  LGO(177-343°C)  0.000  0.000  0.000  0.000  75  750  753  754  755  HGO(343-524°C)  0.089  0.049  0.057  0.048  80  760  763  764  764  524+(>524°C)  0.911  0.951  0.943  0.952  85  807  813  815  815  90  852  881  883  884  92.5  892  896  898  898  95  917  944  953  950  96.5  964  964  965  965  98  1031  1035  1040  1037  99  1047  1048  1050  100  1055  1056  1060  1049 1057  100  400  500  600  700  800  900  1000  1100  T e m p e r a t u r e , °C  Figure IX-6 Effect of water instead of HGO Underwash on Scrubber Bottom TBP curve  120  Chapter 6 - Case Studies: Results and Discussion  • 0 kbarrel/day water, 10 kbarrel/day HGO: Light Ends: 20% ; Water: 65%; 100> fraction: 15% • 0.7 kbarrel/day water, uncontrolled: Light Ends: 19% ; Water: 68%; 100> fraction: 14% • 2.5 kbarrel/day water, controlled by ATB: Light Ends: 19% ; Water: 68%; 100> fraction: 13% 0 2.5 kbarrel/day water, controlled by HGO: Light Ends: 19% Water: 69%; 100> fraction: 11%  200-300  300-400  400-500 500-600 600-700 700-800 800-900 C o m p o n e n t s ' Boiling Temperatures Range, °C  900-1000  1000>  Figure IX-7 Effect of water instead of HGO Underwash on Scrubber Overhead composition  • 0 kbarrel//day water, 10 kbarrel/day HGO: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 0.7 kbarrel/day water, uncontrolled: Light Ends: 0% ; Water: 1%; 100> fraction: 99% • 2.5 kbarrel/day water, controlled by ATB: Light Ends: 0% ; Water: 1%; 100> fraction: 99% 0 2.5 kbarrel/day water, controlled by HGO wash: Light Ends: 0% ; Water: 1%; 100> fraction: 99%  •rrrg] 200-300  300-400  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  m 1000>  Figure IX-8 Effect of water instead of HGO Underwash on Scrubber Bottom composition 121  Chapter 6 - Case Studies: Results and  Discussion  Discussion: When water at 40°C is applied instead of HGO Underwash, without any control of the Overhead temperature, 700 barrel/day is enough to achieve the same temperature profile as 10,000 barrel/day HGO. The temperature of the water is much lower than 325°C for HGO. Overhead mass flow is lower due to the lower total mass "in". Its density and average molecular weight is decreased because most of the water ends up in the Overhead, and some middle fractions that originate from HGO are not present there any more. Scrubber Bottom mass flow rate significantly decreases, because of lower total mass flow "in", and missing HGO fraction that would be present in the Bottom if HGO Underwash was in service. This also causes more concentrated heavier fractions, and consequently higher density and average molecular weight of the Bottom. With controlled Overhead temperature, either by HGO or ATB flow rate, effects are similar with the case without control. The difference is that, actually, total mass flow "in" is much lower than in the first case, because along with applying 2.5 kbarrel/day of water, ATB has to be decreased by 2.5 kbarrel/day, and HGO by 13 kbarrel/day, in order to keep the Overhead temperature the same. That affects the flow rates of Overhead and Bottom, decreasing both radically. However, the volumetric production rate of LGO and HGO is improved. When ATB flow rate is decreased for the control, it mostly affects Scrubber Bottom flow rate (decreases by 26%), because ATB is a heavier stream and most of its components end up in the Bottom. HGO is lighter, and its flow rate affects both Overhead (decreases by 10%) and Bottom (decreases by 13%).  122  Chapter 6 - Case Studies: Results and Discussion  X. Saturated Steam Instead of HGO Underwash In this case study saturated steam at 150 psig (1035 kPa) is used instead of HGO Underwash. The goal is the same as in the previous case study - to decrease Grid entrance temperature, while keeping the Overhead temperature the same. Mass flow of the steam has been set from 0 kg/h (when 10000 barrel/day HGO was used-approximately 49000 kg/h) to 22000 kg/h. An additional case is considered - where saturated steam mass flow rate is the same as in the previous case study, where only 700 barrel/day (4627 kg/h) of water is used. If saturated steam flow rate was 7000 kg/h, the temperature profile would remain similar to the original case with HGO Underwash, without control. Higher flow rate of 22000 kg/h, controlled by HGO or ATB flow rate, decreases Grid entrance temperature by 4°C and 9°C, respectively, keeping the Overhead temperature constant. In order to control the Overhead temperature ATB actual flow rate should be decreased by 4.5% (from 55 to 52 kbarrel/day), and HGO Wash by 65% (from 24 to 8.3 kbarrel/day). All effects are similar to the previous case, with water used instead of HGO Underwash. In the case where same mass flow rate of saturated steam is used as water in the previous case, all effects are very similar. Temperatures along the Scrubber are slightly higher (1-2°C), Overhead volume and mass flow rate a little bit higher, and Scrubber Bottom volume and mass flow rate slightly lower. Properties of both Overhead and Scrubber Bottom are almost the same for two cases. Observations: • Temperature profile:  o  7000 kg/h of saturated steam instead of HGO Underwash, without any control of Overhead temperature, makes the Scrubber temperature profile to remain almost the same as in the original case. Only Shed Bottom temperature increases by 9°C.  o  22000 kg/h of saturated steam decreases Grid Bottom temperature by 4°C, if the Overhead temperature is controlled by ATB flow rate. Shed Top temperature is also several degrees lower and Shed Bottom temperature is 15°C higher.  123  Chapter 6 - Case Studies: Results and Discussion  o  22000 kg/h of saturated steam, with the HGO control, is able to decrease Grid entrance temperature by 9°C, while keeping the Overhead temperature at 390°C (Table X - l , Figures X - l and X-2).  Figure X-l Effect of saturated steam instead  Figure X-2 Effect of saturated steam  of HGO Underwash on temperatures along  instead  the Scrubber  temperature profile along the Scrubber  of  HGO  Underwash  on  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases.  • Overhead properties:  -In the first three cases (7000 kg/h of saturated steam, uncontrolled, 22000 kg/h of saturated steam, controlled by HGO and 22000 kg/h saturated steam, controlled by 124  Chapter 6 - Case Studies: Results and  Discussion  ATB) actual volume of the Scrubber Overhead increases by 2.6, 9.3 and 8.2 and mass flow rate decreases by 3.3, 2.9 and 10.7%, respectively (Table X - l , Figure X-3). -Density changes from 2.83 to 2.67, 2.52 and 2.34 kg/m (Table X-l). 3  -Average molecular weight decreases from 71 to 67, 64 and 60 (Table X-2). -Composition shows lower presence of middle fractions (300-500°C), but higher content of both lighter and heavier fractions (Table X-2, Figures X-5 and X-7). Scrubber Bottom  properties:  -In 7000 kg/h of saturated steam, uncontrolled, 22000 kg/h of saturated steam, controlled by HGO, 22000 kg/h saturated steam, controlled by ATB, both actual volume and mass flow rate of Scrubber Bottom drops by around 7, 35 and 16%, respectively(Table X-l). - Density changes from 816 to 822, 827 and 827 kg/m (Table X-l). 3  -Average molecular weight changes from 637 to 652, 664, 670 (Table X-3). -Composition shows lower presence of fractions up to 600°C, while heavier fractions are more concentrated (Table X-3, Figures X-6 and X-8). Other:  -Grid Liquid volume and mass flow rate slightly increase for first two cases, while for the third one both radically drop and for the last one slightly drop. -Shed Vapour volume flow rate increases, while mass flow rate decreases. -Both Shed Liquid volume and mass flow rate significantly drop (Figure X-4).  125  Chapter 6 - Case Studies: Results and Discussion  2  3  4  C a s e s : 1: 0 kg/h s t e a m , 10 kbarrel/day HGO; 2: 7000 kg/h s t e a m , uncontrolled; 3:22000 kg/h s t e a m , controlled by A T B ; 4:22000 kg/h s t e a m , controlled by HGO w a s h ; 5:4627 kg/h s t e a m , uncontrolled P r e s s u r e drop in Grid, in. of water  2  3  4  C a s e s : 1:0 kg/h s t e a m , 10 kbarrel/day HGO; 2: 7000 kg/h s t e a m , uncontrolled; 3: 22000 kg/h s t e a m , controlled by A T B ; 4:22000 kg/h s t e a m , controlled by HGO w a s h ; 5: 4627 kg/h s t e a m , uncontrolled  Figure X-3 Effect of saturated steam instead  Figure X-4 Effect of saturated steam  of HGO Underwash on mass flow rate of  instead of HGO Underwash on mass  Scrubber Overhead and Bottom  flow rate of other streams  Note: Lines that connect data points do not present trend lines. They are shown to help comparison between different cases.  126  Chapter 6 - Case Studies: Results and Discussion  Table X-l Effect of saturated steam instead of HGO Underwash on Scrubber parameters Saturated Steam Flow Rate Position  ft  kg/h  from the  Saturated steam flow rate instead of HGO Underwash 0 (10 kbl/day HGO)  7000  22000  Base Case  Uncontrolled  Controlled by ATB  22000  4627  Controlled by HGO  Same as water  1  2  %  3  %  4  %  Koch Grid Top  43  Top Stage Temp Est (°C)  393  392  -0.2  389  -1.0  390  Koch Grid Bot.  38  Bottom Stage Temp Est (°C)  395  395  0.0  391  -1.1  386  Sheds Top  34  Top Stage Temp Est (°C)  405  405  0.1  401  -0.9  Sheds Bot.  22  Bottom Stage Temp Est (°C)  473  482  1.7  497  0  Bulk Liquid Temperature (°C)  375  375  0.0  375  pool bottom  Scrubber Pool  Case:  5  %  -0.7  394  0.2  -2.4  397  0.5  394  -2.7  407  0.5  5.0  486  2.7  486  2.6  0.0  375  0.0  375  0.0  Flow Rates& Densities Scrubber Overhead  Actual Volume Flow (m/h)  274,787  281,987  2.6  300,449  297,419 8.2  279,618 1.8  Mass Flow (kg/h)  778,651  752,644  •3.3  755,865 -2.9  695,494 -10.7  754,755 -3.1  2.83  2.67  -5.8  2.52 -11.2  2.34 -17.5  2.70 -4.7  789  727  -7.9  510 -35.3  661 -16.2  713 -9.6  643,728  597,497  -7.2  422,224 -34.4  546,779 -15.1  587,558 -8.7  816  822  0.8  827  2.58  2.49  -3.3  3.48 35.0  305  307  0.6  324  6.3  189 -38.1  310  1.8  217,861  219,461  0.7  232,176 6.6  137,495 -36.9  221,960  1.9  714.97  715.98  0.1  716.77  0.3  729.26 2.0  Actual Volume Flow (m/h)  277,329  284,611  2.6  303,004  9.3  297,520 7.3  282,240  Mass Flow (kg/h)  870,450  846,044  •2.8  861,981 -1.0  800,297 -8.1  850,652 -2.3  3.14  2.97  -5.3  2.84 -9.4  2.69 -14.3  497  458  -7.8  349.92 -29.6  411 -17.3  450 -9.3  361,323  332,365  -8.0  249,529 -30.9  300,237 -16.9  326,195 -9.7  724.76  725.57  0.1  713.11 -1.6  730.43 0.8  724.23 -0.1  330  304  -7.9  213 -35.3  276 -16.2  298 -9.6  269,078  249,754  -7.2  176,490 -34.4  228,554 -15.1  245,599 -8.7  815.85  822.01  0.8  3  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  Actual Volume Flow (m/h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  Overhead / ATB  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m/h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  Shed Vapor  3  Mass Density (kg/m ) 3  Shed Liquid  Actual Volume Flow (m/h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  To Coker  Actual Volume Flow (m/h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  827.32  9.3  1.4  827 1.4  824  2.30 -10.7  1.4  826.99 1.4  1.0  2.50 -3.1  715.73 0.1 1.8  3.01 -4.0  824.18  1.0  Additional information Vapour to Sheds  Temperature (°C)  514  517  0.6  522  1.6  520  1.1  518  0.9  Upgoing Stream  Temperature (°C)  534  534  0.1  535  0.1  535  0.1  535  0.1  Sheds  Stage Efficiency  0.53  0.53  0.0  0.53  0.0  0.53  0.0  0.53  0.0  Koch Grid  Stage Efficiency  0.75  0.75  0.0  0.75  0.0  0.75  0.0  0.75  0.0  SPLColer  Duly (MMBtu/h)  4472  43.82  -2.0  26.03  -41.8  39.74  -11.1  44.81  0.2  127  Chapter 6 - Case Studies: Results and Discussion Table X-2 Effect of saturated steam instead of HGO Underwash on Scrubber Overhead properties Water flow rate  Basic  7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam Cut Point  (10 kbpd  Uncontrolled  HGO Wash) Temperature ft]  393  392  Pressure [psig]  16.99  Molecular Weight  70.77  16.99 66.72  Mass Density [kg/m3] Act. Volume Flow [m3/h] Mass Enthalpy [kJ/kg]  Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kgmole] Mass HeatofVap. [kJ/kg]  Controlled  by ATB  by HGO 389  Uncontrolled  Controlled  HGO Wash)  16.99  16.99  62.60 2.52  58.50 2.34 306,156  Controlled  by ATB  by HGO  TBP [°C]  390 0 1  TBP ft]  -253  -253  -253  -237 -207 •167  -239  -136  -215 •179 -149  -102  •140 •104  -239 -210 -172 •141  -85 -51 •34  •90 •57 •42  -105 •91  •112 •96  -59 •44  •73 •48 •29  •253 •242  2.67 290,271 •3095  309,276 -3326  5.30 2.74  5.46 2.74  5.60 2.72  0.93 182.60 266,718  0.93 170.32 285,477  0.92 159.56  0.95 164.05  7.5 10 12.5 15 17.5  -3  -5  -6  11.42  281,095 11.44  274,537  11.38  11.53  20  266  246  237  •3  930.01  922.64  922.97  906.82  25  310  303  289  24.98 2825 1.05  25.00  24.88  25.02  2868 1.05  2950 1.05  2960 1.06  30 35 40  336 354 373  308 334  325 351 366  340 364  45  391  390  390  380  50  405  406  406  399  420  420  420 438  -3510 5.82 2.73  Fraction Distribution Data Volume fraction C4-(<177°C)  7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam  (10 kbpd  [%]  2.83 282,860 -2924  Mass Entropy [kJ/kg-C] Mass Heat Capacity [kJ/kg-C] Vapor Phase Fraction (Mass Basis) Specific Heat [kJ/kgmole-CJ Std. Gas Flow[STD_m3/h] Watson K  Controlled  Basic  0.060  0.059  2 3.5 5  0.075  0.093  55  420  -209 -171  351 367  322  LGO(177-343°C)  0.259  0.266  0.259  0.262  60  439  439  439  HGO (343-524°C)  0.580  0.568  0.557  0.529  65  442  443  444  524+(>524°C)  444  0.101  0.107  0.110  0.116  70  459  464  465  464  75 80  471  480  482  482  486  85  496  489 512  491 513  490 515  90 92.5  525 537  526 539  527 541  527 542  95 96.5 98  550 556 616  552 556 630  553 557 640  555 563 655  99 100  744 871  755 882  767 889  782 909  HGO —  —  7000 kg/h water,uncontrolled  22000 kg/h water, controlled by A T B 22000 kg/h water, controlled by HGO  // •  .  '  -  O  100  200  300  400  500  600  700  800  900  1000  Temperature, °C  Figure X-5 Effect of sat.steam instead of HGO Underwash on Scrubber Overhead TBP curve 128  Chapter 6 - Case Studies: Results and Discussion Table X-3 Effect of sat. steam instead of HGO Underwash on Scrubber Bottom properties Water flow rate  Basic  7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam Cut Point  (10 kbpd  Uncontrolled  HGO Wash) Temperature fC] Pressure [psig]  375  375  Controlled  Controlled  by ATB  by HGO  Basic  7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam  (10 kbpd  [%]  Uncontrolled  HGO Wash)  375  TBPfC]  375  TBP fC]  Controlled  Controlled  by ATB  by HGO  TBPfC]  TBPfC)  17.00  17.00  17.00  17.00  0  412  419  433  425  Molecular Weight  637.04  652.28  664.75  670.10  • 1  441  455  465  461  Mass Density [kg/m3]  815.85  822.01  827.32  826.99  2  466  484  495  488  812  748  525  681  3.5  490  501  515  512  Mass Enthalpy [kJ/kg]  •1329  •1330  •1332  •1330  5  504  513  520  515  Mass Entropy [kJ/kg-C]  3.39  3.39  3.40  3.41  7.5  514  517  526  523  Mass Heat Capacity [kJ/kg-C)  2.89  2.89  2.88  2.88  10  517  523  555  555  Vapor Phase Fraction (Mass Basis)  0.00  0.00  0.00  0.00  12.5  520  547  556  556  23,893  21,659  15,018  19,293  15  525  556  560  561  11.41  11.40  11.39  11.41  17.5  548  557  569  590  0.69  0.69  0.72  0.70  20  556  562  591  592  1038.96  1043.84  1047.30  1047.97  25  564  593  598  602  0.78  0.79  0.80  0.81  30  592  603  626  628  Act. Volume Flow [m3/h]  Std. Gas Flow [STD_m3/h] Watson K Kinematic Viscosity [cSt] Liq. Mass Density (Std. Cond) [kg/m3] Molar Volume [m3/kg mole) Mass Heat of Vap. [kJ/kg]  1309  1681  1728  1715  35  605  629  632  634  Surface Tension [dyne/cm]  15.47  15.69  15.91  15.85  40  630  634  680  681  Thermal Conductivity [W/m-K]  0.13  0.14  0.14  0.14  45  635  681  683  683  Viscosity [cP]  0.57  0.57  0.59  0.58  50 55  682 684  684 690  685 694  687 695  60  693  699  706  706  65  706  740  743  742  Fraction Distribution Data Volume fraction C4-(<177°C)  0.000  0.000  0.000  0.000  70  743  746  749  749  LGO(177-343°C)  0.000  0.000  0.000  0.000  75  750  753  756  756  HGO (343-524°C)  0.089  0.064  0.037  0.045  80  760  762  790  765  524+(>524°C)  0.911  0.936  0.963  0.955  85  807  812  818  816  90  852  881  887  886  92.5  892  895  902  899 953  95  917  943  962  96.5  964  964  987  965  98  1031  1035  1043  1039  99 100  1047  1048  1055  1056  1053 1062  1049 1057  1100 T e m p e r a t u r e , °C  Figure X-6 Effect of sat. steam instead of HGO Underwash on Scrubber Bottom TBP curve 129  Chapter 6 - Case Studies: Results and Discussion  • 0 kg/h water, 10 kbarrel/day HGO: Light Ends: 20%; Water: 65% ; 100> fraction: 15% • 7000 kg/h water, uncontrolled: Light Ends: 20%; Water: 67% ; 100> fraction: 14% • 22000 kg/h water, controlled by ATB: Light Ends: 19%; Water 69% ; 100> fraction: 12% 0 22000 kg/h water, controlled by HGO w a s h : Light Ends: 19%; Water: 70% ; 100> fraction: 11%  200-300  300-400  1  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  1000>  Figure X-7 Effect of saturated steam instead of HGO Underwash on Scrubber Overhead composition • 0 kg/h water, 10 kbarrel/day HGO: Light Ends: 0%; Water: 1% ; 100> fraction: 99% • 7000 kg/h water, uncontrolled: Light Ends: 0%; Water: 1% ; 100> fraction: 99% • 22000 kg/h water, controlled by ATB: Light Ends: 0%; Water: 1% ; 100> fraction: 99% 0 22000 kg/h water, controlled by HGO wash: Light Ends: 0%; Water: 1% ; 100> fraction: 99%  1 , wrrm , 200-300  300-400  400-500 500-600 600-700 700-800 800-900 Components' Boiling Temperatures Range, °C  900-1000  mrm 1000>  Figure X-8 Effect of saturated steam instead of HGO Underwash on Scrubber Bottom composition  130  Chapter 6 - Case Studies: Results and Discussion Discussion: Since all the effects are very similar to the Case Study IX where water was used instead of HGO Underwash, also all explanations are very similar. When saturated steam at 185°C is applied instead of HGO Underwash, without any control of the Overhead temperature, 7000 kg/h is enough to achieve the same temperature profile as 10,000 barrel/day HGO (approximately 49000 kg/h). The temperature of the steam is low comparing to 325°C for HGO, and it has higher cooling (heating) capacity. Again, Overhead mass flow is lower due to the lower total mass "in". Its density and average molecular weight drop because Overhead contains most of the water from the steam, and doesn't contain middlefractionsthat originate from HGO any more. Scrubber Bottom mass flow rate significantly decreases, because of lower total mass flow "in", and missing HGO fraction that would be present in the Bottom if HGO Underwash was in service. This also causes more concentrated heavier fractions, and consequently higher density and average molecular weight of the Bottom. With controlled Overhead temperature, either by HGO or ATB flow rate, total mass flow "in" is much lower than in the first case, because in order to keep the Overhead temperature constant, ATB flow rate has to be 3 kbarrel/day lower, and HGO 16 kbarrel/day lower. That again affects the flow rates of Overhead and Bottom, decreasing it radically. However, in the case with ATB control, the overall volumetric production rate of HGO is improved. When ATB flow rate is decreased for the control, it mostly affects Scrubber Bottom flow rate (decreases by 34%), because ATB is heavier stream and most of its components end up in the Bottom. HGO is lighter, and its flow rate affects both Overhead (decreases by 10%) and Bottom (decreases by 15%).  131  Chapter 6 - Case Studies: Results and Discussion  XI. Overhead Recycle Cut Point Changes Scrubber Overhead is the final product of the Scrubber (and Fluid Coker). It contains significant amount of heavy fractions, with NBP above 524°C, which are not desirable. This study investigates ways to reduce the presence of these fractions. In other words, to decrease the 95% cut point on the Overhead distillation curve. The specific objective was to investigate required increase in: •  ATB feed,  •  HGO Wash and  •  HGO Underwash  flow rates sufficient to drop the recycle cut point on the Overhead product distillation curve by 15°C, 30°C and 45°C (related to the 95% cut point). Observations and discussion: It was found in this case study that even with radical increases in ATB, HGO Wash or HGO Underwash flow rate, the cut point was not decreased more than 10°C. The reason for that is very low efficiency (10" ) for high boiling components (524°C+) that 10  was used in the Base Case in order to match the Overhead composition. This low efficiency means that these components, actually, by-pass directly to the Scrubber Overhead, without getting in contact with down-flowing liquids (ATB or HGO). Hence, ATB or HGO flow rate does not have any effect on high boiling end of the distillation curve. Theoretically, this was related to the liquid entrainment in the vapour. As already mentioned in Chapter 4 and 5, the vapour has very high volume flow rate comparing to liquid (hundred thousand's comparing to hundred's), and its velocity is very high (-10 m/s). Some liquid droplets are being carried up with vapour and finally end up in the Overhead. Tray and component efficiency is decreased. Even with increase of liquid streams (HGO or ATB) flow rate, more vapour is produced, and the ratio vapour/liquid doesn't change much. Entrainment is still present, resulting in still very low efficiency for heavy fractions and their presence in the Overhead. That reflects on the Overhead distillation curve, especially on the high temperature end. 132  Chapter 6 - Case Studies: Results and Discussion  The results (process parameters and Overhead TBP distillation curves) for all three cases are shown below. Temperature profiles, flow rate and composition charts are not presented, because they all correspond to ATB, HGO Wash and HGO Underwash Flow Rate studies. ATB feed flow rate: By increasing ATB volume flow rate by 82%, the recycle cut point on Overhead distillation curve is lowered just by 5°C (Table XI-1 and Figure XI-1). It has effect on lower cut points (middle fractions), but not so much on this higher boiling fractions. With ATB higher flow rates, temperatures along the Scrubber get too low, what could affect separation and other process parameters (Table XI-2). HGO Wash flow rate: By increasing HGO Wash flow rate by 320%, 95% recycle cut point is decreased by 8°C (Table XI-3, Figure XI-2). Scrubber temperatures are again very low (Table XI-4). HGO Wash seems to have better ability to decrease distillation cut point than ATB, since increase of 15 kbarrel/day in flow rate can decrease cut point in the same extent as 25 kbarrel/day increase of ATB flow rate. The reason is probably that ATB contains more heavy components, and by increasing their content in the Scrubber, it is harder to reduce the cut point. HGO Underwash flow rate: If HGO Underwash volume flow rate is increased ten times from original case (10 kbarrel/day to 100 kbarrel/day), 95% distillation cut point is reduced only by 10°C (Table XI-5, Figure XI-3). Temperatures and all other properties are affected by this increase (Table XI-2). Increase of 20 kbarrel/day in HGO Underwash flow rate has similar effect as 15 kbarrel/day of HGO Wash and 25 kbarrel/day of ATB. HGO Underwash has the same composition as HGO Wash, and its smaller influence on cut point is due to the position of its inlet (smaller cooling effect on the Overhead).  133  Chapter 6 - Case Studies: Results and Discussion Table XI-1 ATB flow rate effect on Overhead TBP distillation curve ATB Flow Rate (kbarrel/day) Cut Point [%]  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5  55  80  TBP fC]  TBP ["C]  -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 5M 556 616 744 871 1  "96.5 98 99 100  100 -253 -240 -211 -174 -143 -107 -92 -62 -45 -8 206 298 324 341 363 377 391 405 419 436 441 448 466 489 523 532 i4(. 556 617 753 888 r  TBP fC] -253 -243 -216 -181 -151 -114 -97 -76 -50 -42 -5 282 310 335 349 364 377 391 403 407 420 440 445 480 517 531 '545 555 648 780 911  134  Chapter 6 - Case Studies: Results and Discussion  Table XI-2 Effect of ATB flow rate on Scrubber parameters Position  ft from the pool bottom Koch Grid Top 43 Koch Grid Bot. 38 Sheds Top 34 Sheds Bot. 22 Scrubber Pool 0  ATB Flow Rate m3/h kbarrel/day Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)  273 55 (Base Case) 393 395 405 473 375  397 80 376 376 386 447 398  % -4.3 -4.9 -4.8 -5.6 6.2  497 100 364 362 372 436 371  % -7.4 -8.3 -8.1 -7.9 -0.9  257,200 663,909 2.58 1,997 1,575,214 789  -6.4 -14.7 -8.9 153.1 144.7 -3.3  1.15 227 167,283 736.24 258,944 705,130 2.72 1,115 810,640 727.01 835 658,440 788.78  -55.4 -25.4 -23.2 3.0 -6.6 -19.0 -13.2 124.4 124.4 0.3 153.1 144.7 -3.3  Flow Rates& Densities Scrubber Overhead  Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kq/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kq/m )  274,787 778,651 2.83 789 643,728 816  Mass Flow Ratio Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kq/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kq/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kq/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  2.58 305 217,861 714.97 277,329 870,450 3.14 497 361,323 724.76 330 269,078 815.85  J  3  Scrubb.Pool Liquid  3  3  Overhead / A T B Grid Liquid  3  3  Shed Vapor  J  3  Shed Liquid  3  3  To Coker  J  3  Vapour to Sheds Upqoinq Stream Sheds Koch Grid S P L Coler  Temperature (°C) Temperature (°C) Staqe Efficiency Staqe Efficiency Duty (MMBtu/h)  514 534 0.53 0.75 44.32  266,026 -3.2 723,279 -7.1 2.72 -4.1 1,385 75.5 1,103,082 71.4 796 -2.4 1.65 260 188,811 724.99 268,281 786,029 2.93 819 593,614 724.56 579 461,088 796.42 501 533 0.53 0.75 75.43  -36.1 -14.5 -13.3 1.4 -3.3 -9.7 -6.7 64.9 64.3 0.0 75.5 71.4 -2.4 -2.6 -0.3 0.0 0.0 70.2  492 532 0.53 0.75 102.31  -4.3 -0.5 0.0 0.0 130.9  135  Chapter 6 - Case Studies: Results and Discussion Table XI-3 HGO Wash flow rate effect on Overhead TBP distillation curve HGO Wash Flow Rate (kbarrel/day)  Cut Point [%] 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 •  96*5 98 99 100  24  40  TBP [°C] -253 -237 -207 -167 -136 -102 -85 -51 -34 -3 266 310 336 354 373 391 405 420 439 442 459 471 486 496 525 537 550 556 616 744 871  TBP ["C] -253 -237 -206 -166 -134 -101 -81 -49 -22 2 274 311 336 352 366 382 397 407 420 440 444 465 477 491 524 532 546 555 600 733 863  60  80  TBP ["CJ -253 -237 -206 -166 -134 -101 -81 -49 -22 2 274 310 333 349 364 377 391 404 407 420 440 443 465 485 521 527 543 555 601 731 861  100  TBP [°C] -253 -237 -206 -167 -135 -102 -84 -51 -31 -1 268 308 324 342 357 366 378 392 405 407 420 440 443 480 517 527 543 "554 596 730 864  TBP ["C] -253 -238 -208 -170 -138 -103 -89 -55 -43 -4 254 300 322 336 349 363 371 378 391 400 406 420 440 475 496 526  5|2  553 598 736 874  24 kbarrel/day HGO Wash 40 kbarrel/day HGO Wash 60 kbarrel/day HGO Wash 80 kbarrel/day HGO Wash 100 kbarrel/day HGO Wash  100  200  300  400  500  600  700  800  900  1000  T e m p e r a t u r e , °C  Figure XI-2 HGO Wash flow rate effect on Overhead TBP distillation curve 136  Chapter 6 - Case Studies: Results and Discussion Table XI-4 Effect of HGO Wash flow rate on Scrubber parameters HGO Wash Flow Rate ft  Position  from the pool bottom Koch Grid Top  43  m3/h kbarrel/day  119 24  199 40  298 60 %  (Base Case) Top Stage Temp Est (°C)  393  397 80  386  497 100  %  -1.9  376  -4.3  % 368.5  -6.3  359  -8.6 -8.6  Koch Grid Bot.  38  Bottom Stage Temp Est (°C)  395  389  -1.6  380  -3.9  371.2  -6.0  361  Sheds Top  34  Top Stage Temp Est (°C)  405  400  -1.1  393  -3.0  384.5  -5.0  374  -7.6  Sheds Bot.  22  Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)  473  456  -3.7  441  -6.9  430.6  -9.0  420  -11.4  375  375  0.0  375  0.0  375.0  0.0  375  0.0  273,178 797,485  -0.6 2.4  270,319 798,630  -1.6 2.6  267410.3 787391.8  -2.7 1.1  2.92  3.0  2.95  4.3  2.9  3.9  1,001 26.8 797,839 23.9  1,352 1,053,191  71.4 63.6 -4.5  Scrubber Pool  0  Flow RatesS Densities Scrubber Overhead  Actual Volume Flow (m'Vh)  274,787 778,651  Mass Flow (kg/h) Mass Density (kg/m ) 3  Scrubb.Pool Liquid  2.83  Actual Volume Flow (m /h) Mass Flow (kg/h) 3  789 643,728  Mass Density (kg/m ) 3  Overhead / A I B  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m /h) Mass Flow (kg/h)  Shed Vapor  Actual Volume Flow (m /h) Mass Flow (kg/h)  3  3  J  -2.3  779  2.58  2.64  2.4  2.65  425 39.5 303,893 39.5  -5.9  759  -6.9  2.53  -2.0  2.6  2.6  1.1  93.0 95.1  739.7 541125.0  142.8 148.4  928 204.5 688,636 216.1  0.0  722.87  1.1  731.5  2.3  742.07  3.8  0.2 2.5  277,562 906,554  0.1 4.1  276478.3 910060.9  -0.3 4.6  273,949 895,642  -1.2 2.9  3.21  2.3  3.27  3.27  4.2  639.68 28.8 459,545 27.2  860 613,054  497 361,323  3  724.76 J  2.2  715.03  Mass Density (kg/m ) Actual Volume Flow (m /h) Mass Flow (kg/h) Mass Density (kg/m )  2.90  2,251 185.3 1,709,377 165.5  277,899 892,151  3.14 3  -4.1  714.97  3  Actual Volume Flow (m /h) Mass Flow (kg/h)  588 425,029  768.0  -2.0  277,329 870,450  Mass Density (kg/m )  To Coker  797  305 217,861  Mass Density (kg/m )  Shed Liquid  816  1723.4 118.4 1323490.9 105.6  263,442 762,774  330 269,078  3  815.85  718.40  4.1  3.3  4.9  73.1 69.7  1064.9 758496.6  114.3 109.9  1,325.36 166.8 947,845 162.3  -0.9  712.77  -1.7  712.3  -1.7  418 26.8 333,497 23.9  565 440,234  71.4 63.6  720.4 553219.2  118.4 105.6  941 714,520  185.3 165.5  778.82  -4.5  768.0  -5.9  759.29  -6.9  797.16  -2.3  715.16  -1.3  Additional information Vapour to Sheds  Temperature (°C)  514  505  -1.8  494  -4.0  485.4  -5.6  477  -7.2  Upgoing Stream Sheds Koch Grid S P L Coler  Temperature (°C) Stage Efficiency Stage Efficiency Duty (MMBtu/h)  534 0.53 0.75 44.72  533 0.53 0.75 48.25  -0.2 0.0 0.0 7.9  532 0.53 0.75 56.03  -0.5 0.0 0.0 25.3  530.2 0.5 0.8 62.2  -0.7 0.0 0.0  528 0.53 0.75 67.02  -1.1 0.0 0.0  39.0  49.9  137  Chapter 6 - Case Studies: Results and Discussion  Table XI-5 HGO Underwash flow rate effect on Overhead TBP distillation curve H G O U n d e r w a s h F l o w Rate (kbarrel/day) Cut Point [%]  0  10  20  50  100  Cut Point [%]  TBP [°C]  TBP [°C]  TBP ["C]  TBP f C ]  TBP f C ]  -253 -238 -208 -170 -138 -103 -88 -55 -43 -4 255 309 336 353 374 392 406 420 440 446 465 482 491 514 526 539 552 556 632 754 879  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  -253 -253 -237 -237 -207 -206 -167 -166 -136 -135 -102 -101 -85 -82 -51 -50 -34 -26 -3 -1 266 271 310 310 336 336 354 352 373 367 391 385 405 402 420 417 439 423 442 441 459 448 471 466 486 482 496 493 525 524 537 535 550 H I B i ^ g 8 556 556 616 608 744 737 871 866  -253 -237 -206 -166 -134 -101 -82 -50 -24 0 272 309 330 349 363 377 391 397 406 420 440 442 465 473 521 527 543 555 601 731 862  -253 -239 -209 -171 -139 -104 -89 -56 -43 -5 249 296 320 335 345 355 364 377 381 392 405 420 451 460 490 526 542 553 597 736 875  10 O-l 0  ,  ,  ,  100  200  300  ,  ,  ,  400 500 600 T e m p e r a t u r e , °C  ,  ,  ,  1  700  800  900  1000  Figure XI-3 HGO Underwash flow rate effect on Overhead TBP distillation curve  138  Chapter 6 - Case Studies: Results and Discussion Table XI-6 Effect of HGO Underwash flow rate on Scrubber parameters HGO Underwash Flow Rate  HGO Uderw. Flow Rate Position  ft from the  m3/h  50  0  50  99  248  kbarrel/day  10  0  10  20  50  Koch Grid Top  43  %  (Base Case)  pool bottom Top Stage Temp Est f C )  393  397  0.9  497 100  %  %  %  %  393  0.0  389  -1.2  375  -4.6  356  -9.6  Koch Grid Bot.  38  Bottom Stage Temp Est (°C)  395  400  1.3  395  0.0  389  -1.6  371  -6.0  349  -11.6  Sheds Top  34  Top Stage Temp Est (°C)  405  410  1.4  405  0.0  398  -1.6  381  -5.9  357  -11.8  Sheds Bot.  22  Bottom Stage Temp Est (°C)  473  495  4.6  473  0.0  460  -2.8  437  -7.6  412  -12.9  0  Bulk Liquid Temperature (°C)  375  375  0.0  375  0.0  375  0.0  375  0.0  375  0.0  0.0  Scrubber Pool  Flow Rates& Densities Scrubber Overhead  Actual Volume Flow (m /h)  274,787  274,718 0.0  274,787  Mass Flow (kg/h)  778,651  757,412 -2.7  778,651 0.0  2.83  2.76 -2.7  2.83  789  688 -12.9  643,728  3  Mass Density (kg/m ) 3  Scrubb.Pool Liquid  (nvVh)  Actual Volume Flow  262,182  -4.6  795,252  2.1  755,852  -2.9  0.0  2.89  1.9  2.95  4.0  2.88  1.7  789  0.0  921 16.8  569,357 -11.6  1,422 80.2  2,555 223.8  739,959 14.9 1,103,587 71.4  1,931,997 200.1  643,728  0.0  1.5  816  0.0  803 -1.6  776  -4.9  756  -7.3  2.58  2.51 -2.7  2.58  0.0  2.62  2.63  2.1  2.50  -2.9  305  314 3.2  305  0.0  296 -2.7  268 -12.0  232 -23.9  217,861  224,963 3.3  217,861 0.0  212,704 -2.4  195,892 -10.1  174,124 -20.1  714.97  715.50 0.1  714.97  0.0  717.71 0.4  Actual Volume Flow (m /h)  277,329  277,339 0.0  277,329  0.0  276,366 -0.3  272,185 -1.9  263,741 -4.9  Mass Flow (kg/h)  870,450  856313 -1.6  870,450  0.0  877,301 0.8  865,084  -0.6  803,916  -7.6  3.14  3.09 -1.6  3.14  0.0  3.17  3.18  1.3  3.05  -2.9  497  440 -11.4  496.80  0.0  587 18.1  898.84 80.9  1,456 193.1  361,323  316,491 -12.4  361,323 0.0  423,013 17.1  641,101 77.4  1,047,135 189.8  724.76  718.94 -0.8  724.76 0.0  720.99 -0.5  713.25 -1.6  330  287 -12.9  330  0.0  385 16.8  594 80.2  1,068 223.8  269,078  237,991 -11.6  269,078  0.0  309,303 14.9  461,299 71.4  807,575 200.1  815.85  0.0  803.04 -1.6  776.05 -4.9  756.30  -7.3  3  Overhead/ATB  Mass Flow Ratio  Grid Liquid  Actual Volume Flow (m /h) 3  Mass Flow (kg/h) Mass Density (kg/m ) 3  3  Mass Density (kg/m ) 3  Actual Volume Flow (m /h) 3  Mass Flow (kg/h) Mass Density (kq/m ) 3  To Coker  -1.8  1.5  828  Mass Density (kg/m )  Shed Liquid  269,959  790,657  816  Mass Flow (kg/h)  Shed Vapor  273,909 -0.3  Actual Volume Flow (m Ih) Mass Flow (kg/h) Mass Density (kg/m )  815.85  3  828.00  1.5  1.5  1.1  730.60  2.2  750.76  719.23  5.0  -0.8  Additional information Vapour to Sheds  Temperature (°C)  514  521  1.3  514  0.0  508  -1.2  491  -4.4  473  -8.0  Upgoing Stream  Temperature (°C)  534  535  0.1  534  0.0  534  -0.1  531  -0.5  527  -1.3  Sheds  Stage Efficiency  0.53  0.53  0.0  0.53  0.0  0.53  0.0  0.53  0.0  0.53  0.0  Koch Grid  Stage Efficiency  0.75  0.75  0.0  0.75  0.0  0.75  0.0  0.75  0.0  0.75  0.0  SPLColer  Duty(MMBtu/h)  44.72  47.35  5.9  44.72  0.0  46.13  3.2  55.62  24.4  64.12  43.4  139  Chapter 7— Summary of Proposed Process Performance  Improvements  Chapter 7 - Summary of Proposed Process Performance Improvements Results and conclusions from the case studies, as well as investigations regarding liquid entrainment and low column efficiencies, suggest several things that could be done to improve Scrubber Section performance in terms of: 1. Better Overhead Product quality 2. Higher productivity 3. Reduced fouling in the Koch Grid  7.1. Overhead Product Quality The final product of the Fluid Coker is the Overhead Product. After exiting the top of the Scrubber Section of the Coker it enters the fractionator where it is separated into four streams: Sour Gas, Butane, Naphtha and Combined Gas Oil (CGO), a mixture of LGO and HGO. The composition, TBP curve, molecular weight and density distribution for the Overhead Product is presented in Appendix III. The most valuable products are Naphtha and CGO, which undergo hydrotreating and mixing into Sweet Blend, the upgraded product. Overhead Product also contains about 10 vol.% of heavy fractions boiling above 524°C, which are not desirable. These fractions make up the dead load in the downstream equipment and cause fouling, and they are not desirable in the final product. Chapters 4 and 5 show how these fractions possibly end up in the Overhead Product. First, very high gas velocity exiting from the cyclones causes some liquid droplets (that mostly contain these heavy fractions) to be thrown upward and reach the Sheds. Very low liquid loading in the Sheds is presumably not able to wash them down, neither along the Sheds nor in the space below the Sheds. Liquid entrainment that is suspected to be present in the Sheds causes lower column efficiency, which affects heavy components the most. Even more, additional liquid is entrained in the vapour along the column. The liquid reaches the Koch Grid, where high gas velocity carries droplets upwards. The Koch Grid is rather efficient in terms of gas-liquid contact, and a 140  Chapter 7-Summary of Proposed Process Performance Improvements  great part of these droplets should be washed down. But, since the Koch Grid operates at too high a gas loading, which is even out of the design ranges (based on HYSYS calculation), part of the liquid still remains entrained and finally reaches the Overhead Products, affecting its quality. Several options could solve this problem: 1. Decrease the feeding rate into the middle part of the Fluid Coker: This would decrease the Cyclone Product flow rate, causing lower gas velocity through the cyclone nozzles and less injected liquid droplets. Also, lower gas loading along the Sheds would decrease liquid entrainment and the amount of the liquid that reach the Koch Grid. Lower gas loading within the Koch Grid would allow operating within the design range and significantly increase the efficiency of the column. The negative effect of this option is decreased total production, but the quality of the product would be improved. 2. Increase ATB, HGO Wash and Underwash flow rates: The last case study (XI) showed that even a drastic increase in ATB, HGO Wash and Underwash flow rates did not remove heavy fractions from the Overhead Product (decrease Overhead Product 95% recycle cut point). The reason that HYSYS simulation showed such result was the arbitrary decreased efficiency assumed for heavy fractions in both the Shed column and the Koch Grid in order to match the Overhead composition. However, in reality, by increasing liquid loading in the columns, efficiency of the columns should increase, increasing the ability to wash down the entrained liquid droplets. This would lead to the lower Overhead cut point, meaning improved quality. In this project, it was not possible to calculate this effect, since it was not known how much the efficiency improves with ATB or HGO flow rate. In order to do that, several plant tests should be done to check the change in Overhead composition with the ATB or HGO flow rates. In that way, it would be possible to estimate the efficiency of the columns and investigate the Overhead cut point change. Also, results in Case Studies II and III showed that higher HGO flow rate improved CGO content of the Overhead. The negative effect of higher ATB or HGO flow rate could be a decrease in the temperatures along the Scrubber, causing poorer separation within the system. Also, HGO is a valuable product and its increased consumption should be optimized.  141  Chapter 7-Summary of Proposed Process Performance Improvements 3. Improve Shed efficiency: It was mentioned in Section 5.2 that liquid entrainment in the Sheds could result in the presence of heavy fractions in the Overhead Product. Increasing the number of Shed trays improves separation, but still does not solve the entrainment issue. Efficiency could be improved only by replacing the current type of trays. This type has rather a large gap between two sheds (1.2 m), while sheds themselves are 0.9 m wide, which allows high gas flow without significant contact with the liquid. Trays that would enable better contact between the vapour and liquid could help improve the efficiency. However, this involves high investment and interruption of the process, and still would not be effective enough without decreased gas loading.  7.2. Overhead Production Rate Overhead Product is the main product of the Fluid Coker. Increasing the production rate is often in conflict with improved quality of the product. The case studies showed that increasing H G O flow rate or temperature does not have any significant advantage in terms of increasing the production rate. Similarly, there is no advantage to increasing the number o f Sheds over 6 trays, and changing the number of K o c h Grid sections. The same is with using water or saturated steam instead of H G O Undarwash. There is an option that could improve the production rate, but still not significantly affect the quality:  1.  Increase ATB flow rate: Although the Case Study I showed that A T B flow rate does not improve the Overhead production rate, it lowers the temperatures along the Scrubber, potentially reducing the fouling. To overcome the pressure drop due to the fouling within the cyclones and the K o c h Grid, higher pressures are applied in the whole system. Case Study VIII shows that higher pressure radically decreases the production rate. In that sense, by reducing the fouling, increasing A T B flow rate could help increasing the production rate. Also, as was mentioned before, increased A T B flow rate would increase liquid loading in the Shed column, improving the efficiency. Optimum should be found, as too high A T B flow rate could decrease temperatures along the Scrubber too much and affect the separation efficiency.  142  Chapter 7-Summary of Proposed Process Performance Improvements  7.3. Fouling in the Koch Grid As was already mentioned, at temperatures around 380°C and higher, coke formation occurs due to the cracking reactions of heavy hydrocarbons. Typically, for the 524°C+ fraction, coke formation starts in about four hours at 390°C, and two hours at 400°C. For the gas oil fractions coking reactions are somewhat slower [41]. These reactions occur everywhere in the system, but the cyclone nozzles and the Koch Grid are affected the most. In the former, deposits form by physical condensation. In the Koch Grid, layers of deposits which build over time, decrease the void space and cause fouling of the grid. High temperature and presence of heavy fractions enhance this process. Hence, in order to reduce fouling, the temperature along the Koch Grid should be kept below 400°C and the presence of heavy fractions should be reduced as much as possible. The following options could help with this issue: 1. Increase A T B flow rate: Higher A T B flow rate radically decrease the temperatures along the Scrubber, which should reduce the fouling within the Koch Grid. Also, higher liquid flow rate decreases the liquid entrainment, and hence appearance of heavy fractions in the grid which are known to increase the fouling. Although Case Study I shows that higher A T B flow rate decreases the Overhead production rate, it is already explained in Section 7.2 that by reducing the fouling, it could actually improve the production rate. The option investigated in Case Study V , without use of HGO Underwash and with temperature controlled by increased A T B , is also acceptable, since good temperature control can be achieved, with saving HGO product and not really affecting the production. The optimal flow rate of A T B has to be estimated, based on further investigation on fouling process and its dependence on temperature and liquid loading in the system. 2. Using water or saturated steam instead of H G O Underwash: The Koch Grid bottom  temperature can be decreased using water at 40°C or saturated steam at 150 psig (10.2 atm) in place of H G O Underwash. As was mentioned in Section 7.2, about 17000 kg/h of water at 40°C with 54% lower H G O Wash volume flow rate reduces Koch Grid bottom temperature by 8°C, while about 22000 kg/h of saturated steam at 150 psig (10.2 atm), with 65% lower HGO Wash volume flow rate reduces it by 9°C. This also improves the Overhead production rate. Since lower liquid loading is present in the system than in the case when H G O  143  Chapter 7-Summary of Proposed Process Performance Improvements  Unredwash is used, higher liquid entrainment is expected. This could possibly affect the fouling rate; however, the effect is probably weak compared to the effect of temperature.  144  Chapter 8 - Conclusions and Recommendations  Chapter 8 - Conclusions and Recommendations 8.1. Conclusions H Y S Y S process simulation o f the Scrubber Section o f the Fluid Coker gave insights into the process behaviour and improved understanding o f the whole process. Case studies showed trends and quantitative outcomes o f some process and design changes, which can suggest possible options for process improvement. Based on the results from Chapter 6 and considerations from Chapters 4, 5 and 7, some general and case-specific conclusions can be derived.  General conclusions: • H Y S Y S process simulator is able to effectively represent the Scrubber Section o f the Fluid Coker. Results o f the simulation match the plant data very well (within 3.2% o f the plant data), once separation efficiencies near zero were assigned to the heaviest fractions. • Calculations i n Chapter 4 suggest that liquid entrainment may be present i n the system. Entrainment o f heavy species into the vapours decreases the efficiency o f the Shed section radically. • Consideration i n Chapter 5 suggests that the K o c h Grid operates out o f the designed conditions. Too high gas loading and too low liquid flow rate result i n increased pressure drop and lower efficiency. Additional liquid entrainment is also possible within this section. • Changing gas (lowering) and liquid (increasing) loading so that the Shed section is further from the entrainment flooding and the K o c h G r i d is within the design range o f operating conditions could help improving the efficiency o f these two sections, decreasing liquid entrainment in the vapour, and improving Overhead product characteristics. • The developed  simulation can be used  for additional case  studies  and process  modifications.  145  Chapter 8 - Conclusions and Recommendations  Specific conclusions derived from case studies: • Increasing ATB flow rate has the positive effect of decreasing temperatures in the Grid and the Shed, which should reduce fouling due to the coke formation; it has no positive effect on the production rate of the desired product (Overhead), although the mass production rate decreases only 7%, and the desired CGO ( LGO plus HGO) fraction remains the same. Higher flow rate of ATB is not able to reduce 95% cut point on Overhead distillation curve significantly. An extreme increase in flow rate lowered the cut point only 5°C. • Increasing HGO Wash and Underwash flow rate has no significant effect on decreasing temperature of the Grid and on production rate. Increase of 16 kbarrel/day (79000 kg/h) of HGO Wash or 10 kbarrel/day (48000 kg/h) of HGO Underwash produces only additional 11000 kg/h and 7000 kg/h of HGO, respectively. That means that more HGO is spent than produced. Neither of the two streams is able to reduce 95% cut point on Overhead distillation curve significantly, but they have stronger effect than ATB on cut point. HGO Wash shows the best results in this sense. • Change in HGO Wash temperature has a mild effect on temperature profile. Increase in temperature slightly increases the Overhead production rate, and composition shows higher presence of middle fractions (LGO and HGO). • If HGO Underwash is out of service, the temperatures along the Scrubber increase too much (close to 400°C), and could increase fouling. That is why ATB or HGO control of the Overhead and overall temperatures is necessary. Increasing HGO Wash flow rate cannot decrease Grid bottom temperature enough, while keeping Overhead temperature at 393°C. All other properties of Overhead remain the same as when Underwash is in service. ATB flow rate changes provide a good mechanism to control all the temperatures along the Scrubber. As ATB rate increases Overhead production rate drops slightly and composition shows increased presence of LGO fractions, and decreased in HGO fractions. • Decreasing the number of Shed trays from 6 to 2 increases system temperature and has slightly increasing effect on Overhead mass and volume flow rate. Content of HGO fractions in the Overhead is lower. Increasing in number of trays above 6 changes the situation: the Grid Bottom temperature rises, the Overhead volume flow rate remains almost the same, and the content of LGO and HGO fractions in the Overhead is improved. 146  Chapter 8 - Conclusions and  Recommendations  Increasing the number of Grid sections from 2 to 10 does not have a significant effect on temperature profile along the Scrubber. Only Shed Bottom temperature changes noticeably. Scrubber Overhead production rate remains almost the same. Content of middle fractions (LGO and HGO fractions) in the Overhead drops, while heavy fractions are present in a higher amount, what is not desirable. Absolute pressure in the system definitely has a significant effect on Scrubber performance. As pressure increases, temperatures slightly increase, but Overhead production rate radically drops. Overhead contains more light fractions, while the HGO fraction is much lower. 4600 kg/h (0.7 kbarrel/day) of water (40°C) used instead of 49000 kg/h (10 kbarrel/day) of HGO Underwash is able to control all the temperatures along the Scrubber. Only the Shed Bottom temperature increases. Overhead volume production is increased, but mass flow rate is lower due to the lower density. Water content of Overhead is increased from 65 mole % to 68 mole %. Overhead contains more LGO fractions and less HGO fractions. If water was to be used to lower the Grid Bottom temperature, either ATB or HGO Wash had to be used to control the Overhead temperature. HGO showed better ability in temperature control, but resulted in lower content of HGO fraction in Overhead. Overhead volume production rate was increased, but mass flow rate was not. Both options with ATB or HGO control show higher overall volumetric HGO and LGO production. With 7000 kg/h of saturated steam instead of 49000 kg/h of HGO Underwash, without any control of Overhead temperature, the Scrubber temperature profile remains almost the same as in the original case. Only Shed Bottom temperature increases. The content of water in Overhead is slightly higher, as well as presence of LGO fractions. Overhead volume production is increased slightly, but mass flow is not. To decrease Grid Bottom temperature using saturated steam, ATB or HGO flow rate, the temperature of the Overhead must be controlled. Again HGO showed better ability for control. Overhead volume production rate was improved, but mass flow rate dropped. The composition showed lower presence of CGO fraction. Case with ATB control showed improved overall HGO volumetric production. The recycle cut point on Overhead distillation curve could be lowered by increasing ATB feed flow rate or HGO flow rate. But, the last case study (XI) showed that even a drastic 147  Chapter 8 - Conclusions and  Recommendations  increase in flow rates does not have any significant effect on 95% recycle cut point. It has more effect on lower cut points. The reason for such simulation result is explained in point 2, in Section 7.2. Too high flow rates of these three streams could decrease the temperatures along the Scrubber, which could affect separation and other process parameters. Among three options, HGO Wash stream showed the best ability to decrease the Overhead 95 % cut point. 8.2. Recommendations The conclusions and recommendations in this chapter are based on the results of the Case Studies performed within this project, as well as investigation on liquid entrainment and low column efficiency issues. Although some of the trends and process behaviour considered in this project are confirmed by plant tests, one must be careful in applying these changes. Process simulation is often able to satisfactory model real processes, but it usually includes some approximations, estimations and user's judgement. Also, it is usually not possible to include all aspects of the problem. Therefore, additional investigations, especially on fouling and liquid entrainment issue, should be undertaken, as well as plant tests to confirm the results of the studies. The fouling process definitely affects the performance of the Fluid Coker. In order to investigate fouling within the Koch Grid section it is necessary to determine all the parameters that have effect on the fouling and how they influence the fouling process. The HYSYS process simulator can help in this investigation to obtain better understanding of some parameters and outcomes correlations. Assuming that data for plant parameters change over time (pressure drop within the Koch Grid, streams' flow rates, pressures and temperatures, and Scrubber Overhead composition) are available, simulating the process changes over time and comparing with the plant results can give the insight to the fouling process. HYSYS has the option of changing characteristics of the packing type within the packed column, including the packing factor, Fp and height of packing equivalent to one theoretical plate, HETP. These characteristics would change along with the change in void fraction of the packing that accompanies fouling. If simulating the change of  148  Chapter 8 - Conclusions and Recommendations  these characteristics can produce the results that match the plant data, better perspective of fouling could be achieved. This project and another source (Nelms, [14]) suggest that liquid entrainment may be present within the Fluid Coker, and be the reason for low Sheds and Koch Grid efficiency and fouling. Based on conclusions from this project, the recommendation for further investigation would be to do several plant tests to decrease gradually gas loading within the Fluid Coker and to record the Overhead product characteristics (composition and content of liquid fraction). The same should be simulated by HYSYS and by changing the column efficiency it should be tried to match the plant results. Presumably, with the gas loading low enough no liquid entrainment should occur. In this way, some additional conclusions could be derived, and existence of liquid entrainment confirmed or denied. In general, although expensive and time consuming, some more plant tests should be done, in order to evaluate simulation results and confirm the conclusions derived from them.  149  Glossary of Terms  Glossary of Terms ASTM 2887  Simulated distillation method applicable to all petroleum products boiling below 538°C;  ^rpg  Atmospheric Topped Bitumen, a product of atmospheric distillation of bitumen, with 50 wt% that boils above 560°C;  CGO  Coker Gas Oil fraction (220-570°C fraction);  CGO  Combined Gas Oil  EOR  End of Run of the plant  EOS  Equation of State  HGO  Heavy Gas Oil, one part of the Overhead product after fractionation (343524°C fraction) that is recycled and serves to scrub heavy fractions and particulates from rising vapour in the Scrubber;  HTSD  High Temperature Simulated Distillation, which extends ASTM D2887 to 760°C boiling points  LGO  Light Gas Oil  NBP  Normal Boiling Point  OTSB  Once Through Scrubber Bottom, mixture of heavy fractions of Cyclone Product, boiling temperature up to 1090°C  PR EOS  Peng-Robinson Equation of State  PVT  Pressure-Volume-Temperature  RCP  Recycle Cut Point  SCFE  Supercritical Fluid Extraction method, new method capable of analyzing high molecular weight residue fractions  SCO  Sweet Crude Oil  SOR  Start of Run of the plant 150  Glossary of Terms  SPL  Scrubber Pool Liquid  SPR  Scrubber Pool Recycle  TBP  True Boiling Point  VLE  Vapour-Liquid Equilibrium  VTB  Vacuum Topped Bitumen, a product of vacuum distillation of bitumen, with 50 wt% that boils above 630°C  151  References  References 1. Web site: http://www.intnet.mu/iels/index.htm, date of access Feb 26, 2005. 2. "Oil outlook to 2025", OPEC Review Paper 2004. 3. Web site: http://www2.chappaqua.kl2.nv.us/hgfaculty/rooddo/chapter  21.htm, date of  access Feb 26, 2005. 4. Web site: http://www.users.on.net/~rmc/01sorry.htm, date of access Feb 26, 2005. 5. North American Oil Reserves Brochure, Alberta Energy, 2004. 6. Web site: http://www.energy.gov.ab.ca/osd/docs/osgenbrf.pdf,  date of access Feb 26,  2005.  7. Web site: http://www.syncrude.com, date of access Feb 26, 2005. 8. Williston, M., "Process Model of Scrubber Section of Syncrude's Fluid Cokers", B.A.Sc. Thesis, Dept. of Chemical and Biological Engineering, University of British Columbia, 2002.  9. Web site: http://www.westernoilsands.com, date of access Feb 26, 2005. 10. Web site: http://www.mining-techiiology.com/proiects/sviicrude, date of access Feb 26, 2005.  11. Alberta oil web site: http://collections.ic.gc.ca/oil/process, date of access Feb 26, 2005. 12. Gray, M . R., "Upgrading petroleum residues and heavy oils", Marcel Dekker, Inc., New York, 1994. 13. Huq, I., Van Zanden, S., communication from Syncrude Canada Ltd., 2003-2004. 14. Nelms, C. R., "Fluid Coker Reactor Cyclone Fouling - A summary report for Syncrude Canada Ltd. and other participants", 1999. 15. Westphalen, D., Shethna, H., Aspen Technology, Inc., "Refinery wide simulation", Hydrocarbon Engineering, March 2004.  152  References  16. Brar, S., Brenner, S., Gierer, C , Strashok, C , Jamil, A., Nguyen, N . , Chau, A., Sachedina, M . , Hugo, L., Lowe, C , Hanson, K., Hyprotech, Ltd., HYSYS 3.0 Documentation, Simulation Basis, 2002. 17. Ettienne, T., A., Brenner, S., Gierer, C., Strashok, C., Jamil, A., Nguyen, N., Chau, A., Sachedina, M . , Hugo, L., Lowe, C , Hanson, K., Hyprotech, Ltd., HYSYS 3.0 Documentation, Operation Guide, 2002. 18. Brent, Y., Baker, J., Monnery, W., Svrcek, W., "Dynamic simulation for controllability of Chevron Canada Resources' Kaybob South #3 Gas Plant sulfur recovery unit", LRGCC Conference Proceedings,  2002,  52 , 239-258. nd  19. Loe, B., Pults J., "Implementing and Sustaining Process Optimization Improvements on a Deisopentanizer Tower, HOVENSA LLC", Presented at the AIChE Spring Meeting, April 25, Houston, Texas, 2001. 20. Lars, E., Tyvand Selst, E., Stavanger, K., "Process Simulation of Glycol Regeneration", for presentation at GPA Europe's meeting in Bergen, 2002. 21. Soave, G., Feliu, J.A., "Saving energy in distillation towers by feed splitting", Applied Thermal Engineering, 2002, Volume 22, Issue 8, Pages 889-896. 22. Sabharwal, A., " A Hybrid Approach Applied to an Industrial Distillation Column that Compares Physical and Neutral Network Modeling Techniques", Masters Thesis, Department of Chemical Engineering, Calgary, Alberta, 1997. 23. Briesen, H., Marquardt, W., Prozesstechnik, L., "New approach to refinery process simulation with adaptive composition representation", AIChE Journal, 2004, Volume 50, Issue 3, Pages 633-645. 24. Web site: www.aspentech.com, date of access Jan, 2005. 25. Seider, W.D., Seader, J.D., Lewin, D.R, "Process Design Principles", John Wiley & Sons, Inc., 1999. 26. Pedersen, K.S., Thomanssen, P., Fredenslund, A., "Thermodynamics of Petroleum Mixtures Containing Heavy Hydrocarbons", Ind. Eng. Chem. Process Des. Dev., 1984, Vol. 23, No. 1.  153  References  27. Reid, R.C., PrausnitzJ.M., Poling, B.E., "The Properties of Gases & Liquids", McGrawHill, Inc., 1987. 28. Reid, R.C., Prausnitz, J.M., Sherwood, T.K., "The Properties of Gases and Liquids", McGraw-Hill Book Company, 1977. 29. Stichlmair, J. G., Fair, J. R., "Distillation. Principles and Practices", Wiley-VCH, New York, 1998. 30. Eich-Soellner, E., Lory, P., Burr, P., Kroner, A., "Stationary and Dynamic Flowsheeting in the Chemical Engineering Industry", AMS Subject Classifications: 65C20, 65H10, 65L05, 65M20, 80A99, 1991. 31. Mangat, M . , "Process Model of Heavy Component Partial Condensation", B.A.Sc. Thesis, Dept. of Chemical and Biological Engineering, University of British Columbia,  2001. 32. Chung, K. H., Xu, C , Hu Y., Wang, R., "Supercritical fluid extraction reveals resid properties", Oil and Gas Journal, Jan 20, 1997. 33. Box, M.J., " A New Method of Constrained Optimization and a Comparison with Other Methods", Computer J., 1965, 8, 42-45. 34. Press, W. H., Flannery, B. P., Teukolsky, S. A., Vetterling, W. T., "Numerical Recepies in C", Cambridge University Press, 1988. 35. Kuester, J. L. and Mize, J.H., "Optimization Techniques with FORTRAN", McGraw-Hill Book Co., 1973. 36. Perry's Chemical Engineers' Handbook, Publisher New York; Montreal: McGraw-Hill, Edition 7 , Chapter 17-Gas-solids separation, 1997. th  37. Perry's Chemical Engineers' Handbook, Publisher New York; Montreal: McGraw-Hill, Edition 7th, Chapter 6- Fluid and particle dynamics, 1997. 38. Perry's Chemical Engineers' Handbook, Publisher New York; Montreal: McGraw-Hill, Edition 7th, Chapter 14-Gas-liquid contacting systems, 1997. 39. Kirk-Othmer Encyclopedia of Chemical Technology, John Wiley & Sons, Inc., 1993. 40. Web site: http://www.koch-glitsch.com , date of access Aug., 2004. 154  References  41. Yue, C , Watkinson A. P., Lucas, J. P., Chung, K. H., "Incipient coke formation during heating of heavy hydrocarbons", Fuel, 2004, 83, 1651-1658. 42. Brar, S., Channey, J., Khoshkberchi, M . , Zhao, E., Hyprotech, Ltd., HYSYS 3.0 Documentation, Reference Guide, COM Thermo, 2002.  Appendix I  Appendix I - Peng-Robinson Equation of State  The Peng-Robinson EOS is presented below:  p-^L. V-b  <,.„  2 V(V + b) +  b(V-b)  where a and b represent deviation from ideal behaviour. Term a represents the strength of attraction between two molecules (interaction force), and b is proportional to the size of the molecules. These parameters can be determined from critical values P and T , and the acentric c  c  factor co for pure substances [42]. For a pure component: a = aa c  (1-2)  a = 0.45724^--^c  b = 0.077480——^a represents temperature dependence of parameter a . ^  = l + K(\-T )  _  05  r  (I  3)  = 0.37464 +1.5422G7 - 0.26992ET  2  K  K-binary interaction coefficient co-acentric factor In this form, the Peng-Robinson EOS can be applied to a pure component. To apply it to a mixture, mixing rules are needed for a and b terms. Mixing rules states how parameters a and b for the mixture depend on composition.  156  Appendix I  a = t t (x,x a ) l  t  J  (1-4)  v  i=i y=i  where a is a measure of the strength of attraction between a molecule i and a molecule j. {]  i=i  ^ ; = ( i - ^ ) ( i - C  „ .- 0  4 5 7 2  j  )  a-7)  *'^  (1-8)  = 0.37464 +1.54226&>,. - 0.26992<y ,  2  a, < 0.49  (1-10)  Several mixing rules could be applied to the temperature dependent binary interaction parameter, £,jj. These mixing rules are mostly based on empirical equations and are suitable for some EOS and systems, but not for other. Some of them are:  ^ l - ^ + ^ r  +q r  2  and  (1-31)  157  Appendix I  ^=l-A B T ii+  where A , y  By  v  Ay, By  and  ^L  +  (I  _42)  and C are asymmetric binary interaction parameters. Values fora , a ~, u  y  C  y  can be calculated or found in tables.  H Y S Y S  b, K „ {  has a library with binary  interaction parameters for more than 16000 binaries. If the system contains pseudocomponents,  H Y S Y S  provides a wide selection of the estimation methods for all needed  parameters - T , P , co, K , etc. Methods for estimating the interaction binaries between pseudoc  c  components and library components are also available.  158  Appendix II  Appendix II - Flash Block Calculation Appendix II shows the procedure of manual solving a simple flash block containing a ternary mixture and the comparison with the HYSYS solution for the same system. The schematic of the flash block is presented in Figure AII.l. The mixture enters the flash block with the starting composition and is separated into the vapour that is richer in the low boiling, volatile, components and the liquid. As the vapour is removed continuously, the original mixture gets poorer in the light, more volatile components. The liquid that leaves the flash drum becomes richer in the heavy, less volatile components.  V (yi,y2,y3,-,y c) n  F (Zi,Z ,Z3,...,Znc) 2  L (Xi,X ,X ,...,X ) 2  3  n c  t Q  Figure AII.l Schematic of the flash block  This single-stage equilibrium separation is described by the following equations: Material balance: At a specified temperature and pressure within the flash block, one mole of starting mixture with the composition z\, z , 2  z , ( where nc is number of components) separates into L nc  moles of liquid with the composition xi, X2, composition y  u  y ,y 2  n c  z , and V moles of vapour with the nc  .  Overall mole balance on this system is:  159  Appendix II  L + V  (II-l)  =1  And the component mole balance is:  z, = x,L +  yV  =  t  L +  Xi  y.  (1 - L) = x. (l-V)  +  y  y  0 = 1,2 ,...,nc)  (II-2)  Energy balance:  Fh  F  +Q  = Vh  v  +  Lh  L  (II-3)  where h is the molar enthalpy of the feed, vapour or liquid, and Q is the heat that has to be added to the system in order to evaporate one part of the liquid. Thermodynamic requirement: Flash calculation is based on system tendency to reach thermodynamic equilibrium. Vapour-liquid equilibrium ratio for a component i is given by the following equation:  (II-4)  where O", and O', are the fugacity coefficients for the component i in the vapour and liquid phases. Combining Equations (II-2) and (II-4), two sets of equations can be obtained.  First set contains Xj :  X: =  (II-5)  160  Appendix II Since Xj must sum to unity:  nc  g  nc  ^ • - ^ L  + Kfl-L)'  1  (  "-  6 )  Another set contains y;  ' o-£'lV  y =  nc  (II  J£  nc  g  ^'-^o-^V'  1  "  (I,  7)  -  8)  To determine the composition of the vapour and the liquid leaving the flash block, K; values should be known for each component at given conditions. Kj can be calculated from fugacity coefficients <I>j and d>; using an equation of state. However, Oj and ®j depend on Xj and y;, and iterative method has to be applied. In the following example, the Peng-Robinson EOS (PR EOS) will be used. It is presented in the Appendix I, along with the mixing rules applied. Fugacity coefficients for components in the mixture can be calculated from a general thermodynamic equation: P  R• T• lncp,.  v  .  =  dV (ill.)  =  rp  f(v -R—)-dp o P  (II-9a)  ..  (H-9b)  f  161  Appendix II  dV  where the molar volume for the liquid or vapour phase, v and the derivative  can be dn  ;  calculated using EOS, actually from:  . ZRT v=—  (IMO)  Compressibility factor Z describes the real behaviour of pure fluids and can be expressed based on one of the EOS, in this example Peng-Robinson EOS and equation:  Z  =  ^  ±1  v-b  („_!!)  RT(v(v + b) + b(v-b))  Another way of presenting Equation (11-11) is in the form of a cubic equation, Equation (II13a). Equation of state for mixtures are derived from EOS for the pure components by using concentration-dependent coefficients. Substituting Equation (11-11) into Equation (II-9a) gives expression for calculating the fugacity coefficient for each component:  „ , .„ _ A ln^. = - J - ( Z - l ) - l n ( Z - 5 ) b ' ' 242B O; ,„ K  lL j ji  2  x a  b | l n [ Z + (1 + t  v  V2)7J  Z + (1 - J2)B  (11-12)  The compressibility factor of the mixture is calculated from PR EOS as the root of the following equation, where the smallest root corresponds to the liquid phase and the largest for the vapour phase. Z - ( 1 - 5 ) Z +Z(A-3B 3  2  2  -2B)-(AB-B  2  -5 ) = 0 3  (II-13a)  162  Appendix II  aP RT  (II-13b)  B=—  (II-13c)  2  2  RT  Parameters a and b are calculated from the parameters for the pure components and mixing rules presented in Appendix I by Equations (1-5) to ( 1-15). Example: A 1000 mole/h of liquid mixture containing 5 mole % Hydrogen, 70 mole % of Methane and 25 mole % of Ethane enters a flash block at the constant temperature of 200 K and pressure of 75 bar. In this example the composition of the outgoing vapour and liquid stream will be calculated, as well as flow rate of the two streams. Composition and the parameters for the three components are given in Table AII.l [38]. Table AII.l Parameters for the flash block system components [38] Hydrogen (1)  Methane (2)  Ethane (3)  Liquid  0.050  0.70  0.25  Critical temperature (K)  43.6  190.63  305.43  Critical pressure (bar)  20.47  46.17  48.84  Acentric factor co  0.00  0.01  0.099  ky factor  ki2=  -0.0222  k  2 3  =  -0.0078  ki3=  -0.1667  Parameters a, a and b for the pure substances are calculated from Equations (1-2) and (I3) in Appendix I, and presented in Table AII.2:  163  Appendix II  Table AII.2 PR EOS parameters for pure substances Substance  Pc  Tc  Hydrogen (1)  2.05E+06  Methane (2)  4.62E+06  190.63  Ethane (3)  4.88E+06  305.43  Interaction parameters  a  CO  43.6  0  a;  bi  0.327  9.61 E+03  0.01  0.981  2.44E+05  2.67E-02  0.099  1.210  7.31 E+05  4.05E-02  1.38E-02  calculated in somewhat simpler way then presented in  a-rare  Appendix I, Equation (1-6). The equation used for their calculation is Equation (11-14) and the values for the present system are given in Table AII.3.  (IM4)  a =^(l-k ) g  u  Table AII.3 Interaction parameters for Hydrogen-Methane-Ethane system Substance Hydrogen (1)  a„=  9.61 E+03 a =  4.95E+04 °13=  9.78E+04  Methane (2)  a = 2J  4.95E+04 a =  2.44E+05 Q23=  4.26E+05  Ethane (3)  031=  9.78E+04 a =  4.26E+05 Q33=  7.31 E+05  12  22  32  Based on the mixture composition, parameters a, b, A and B for the mixture can be calculated from Equations (1-4), (1-5), (II-13b) and (II-13c), respectively, and used in PR EOS.  Procedure  I iteration: 1.  As a starting point the composition of the outgoing liquid stream is assumed:  xi= 0.050; x = 0.70; x = 0.25 2  2.  3  Mixing rules ( Eq. (1-4) and (1-5)) are applied:  a=  W  l  "^3"^2^32  l  +  W  l  2  +  W  l  3  +  W  ,  +  +W  2  3  +W  3 ,  +  (  I  M  5  )  "^3"^'3^33  164  Appendix II  fl=3.2-10 Pa m /kmol 5  b=  6  2  +xb +xb 2  z  3  3  (11-16)  b=2.95-10" m /mol 2  3.  3  Calculated a and b values are used to calculate A and B parameters, as well as Z for the mixture from Equations (II-13a), (II-13b) and (II-13c). The cubic Equation (II13a) has three solutions. The lowest value corresponds to the compressibility factor of the liquid phase. Calculated value is Z= 0.28.  4.  Calculated value for Z incorporated in Equation (11-12) for the three components (hydrogen, methane and ethane) gives the fugacity coefficients of these components in the liquid phase: ti  = 6.68  ti  =0.61  ti =  5.  0.07  To estimate vapour composition Kj should be calculated from the following equation:  K =^-  (H-17)  t  Ideal behaviour of the vapour will be assumed, and hence ti -1 • From this assumption, K values for each component in the mixture are calculated: K = 6.68 y  K2  =  K3  =  6.  0.61 0.07 Vapour composition can be further estimated:  165  Appendix II  y =K,x, =6.68-0.05 = 0.334 x  y =K x  2  =0.61-0.70 = 0.427  y =K x  3  =0.07-0.25 = 0.018  2  2  3  3  7.  Since the sum of thefractionsshould be equal to unity, in this way the result can be checked:  y\  + ^ 2 + ^ 3  =0.334 + 0.427 + 0.018 = 0.779  The result is not 1, what means that the first assumption for the liquid composition is not correct. 8.  Correction of the vapour composition will be made:  y = x1  y, 2  0-334 „„ = 0.42 0.779 n  n« = °= 0.55 0.779 4 2 7  0.023  , , = * ° 1 £ 3  0.779  Based on this composition as a new guess, steps 1 to 7 is repeated, but this time for the vapour mixture. Equation (II-13a) is used to improve the values during iterations. II iteration: 1.  As a starting point the composition of the outgoing vapour stream is assumed:  x,= 0.42;x = 0.55; x = 0.023 2  2.  3  Mixing rules (Eq. (1-4) and (1-5)) are applied:  0=1.11-10 Pa m /kmol s  6  2  b=2.14-10" m /mol 2  3.  3  Calculated a and b values are used to calculate A and B parameters, as well as Z for the mixture from Equations (II-13a), (II-13b) and (II-13c). The cubic Equation (II13a) has three solutions. The highest value corresponds to the compressibility factor of the vapour phase. Calculated value is Z= 0.83.  166  Appendix II  4.  Calculated value for Z incorporated in Equation (11-12) for the three components (hydrogen, methane and ethane) gives the fugacity coefficients of these components in the vapour phase: =1.23  f = 0.62 2  <t>\ = 0.32 5.  To estimate new liquid composition Kj should be calculated from the following equation:  <f>i K values for each component in the mixture are calculated:  '  1.23  0.62  K,=™1 = 0.22 0.32 3  6.  Liquid composition can be further estimated:  JC, =y IK,  =0.42/5.43 = 0.077  x =y /K  =0.55/0.98 = 0.56  x  2  X i  2  2  = y IK, = 0.023/0.22 = 0.10 }  7. Since the sum of the fractions should be equal to unity, in this way the result can be checked: x, +x +x, =0.077 + 0.56 + 0.10 = 0.738 2  The result is not 1, what means that the assumption for the vapour composition is not correct. 8.  Correction of the liquid composition is made:  167  Appendix II  in x, = ° = n0.10 0.738 0 7 7  1  y, =  0  5  6  2  0.738  3  0.738  if, =ft 0.76  This liquid composition is used as a new guess and steps 1 to 7 are repeated. The iterations are repeated until the final solution is obtained.  Solution Equilibrium constants values: K = 6.367 K = 0.936 K, =0.158 x  2  Composition of the vapour phase: jy, =0.312 y = 0.647 y, = 0.041 2  Composition of the liquid phase: x, = 0.049 x =0.691 2  JC = 0.260 3  From Equation (II-5) the fraction of the starting mixture that leaves the flash block as liquid is: L=0.9962 The flow rate of the liquid outgoing stream is 996.2 mol/h. The vapour stream fraction is: V=0.0038  168  Appendix II  The flow rate of the vapour outgoing stream is 3.8 mol/h.  Comparison with the HYSYS calculation The same system is simulated in HYSYS process simulator. The results are as follows: Composition of the vapour phase: y = 0.294 x  y = 0.659 2  y = 0.047 3  Composition of the liquid phase: x, = 0.049 x = 0.701 x = 0.250 2  3  The flow rate of the liquid outgoing stream is 997.53 mol/h. The flow rate of the vapour outgoing stream is 2.466 mol/h. HYSYS results show the average deviation in the vapour and liquid composition of about 4.5% from the calculated ones, 0.1 % deviation in molar flow rate of the liquid phase, and 35%> deviation in molar flow rate of the vapour phase. HYSYS needs less than one second to go through the same procedure. The advantage of process simulators such as HYSYS is that they can calculate operations much complicated than a flash block, and even more they can simulate the whole processes.  169  Appendix III  Appendix III - Scrubber Section Streams Data Cyclone Product Composition of this stream was defined in the 1980's when the coker was run in "once through" mode. It contains water, light ends, CGO (Coker Gas Oil) fraction and OTSB fraction (OTSB-Once Through Scrubber Bottom, mixture of heavy fractions of cyclone vapour). Weight percents of all four fractions are shown in Table AIII. 1. Composition of Light Ends is given in Table AIII.2. CGO fraction is characterized using ASTM D2887 method with HTSD enhancement, and its assay is presented in Table AIII.3. Table AIII.4 shows TBP data calculated by HYSYS. OTSB is the heaviest fraction in Cyclone Product, which contains components boiling above 730°C. As an enhancement to ASTM D2887 method, for the fractions above 524°C SCFE technique was used. This method is capable of analyzing high molecular weight residue fractions. OTSB boiling curve was generated as a composite curve from ASTM and SCFE data. This is explained in details in M . Mangat thesis [5]. Composite data for OTSB is given in Table AIII.5. TBP data for OTSB, calculated by HYSYS are given in Table AIII.6. For the purpose of the simulation, Cyclone Product stream is formed as a mixture of above mentioned four streams: water, light ends, CGO and OTSB. Cyclone Product molar composition is given in Table AIII.8, TBP data and TBP curve calculated by HYSYS in Table AIII.7 and Figure AIII.l, molecular weight distribution in Figure AIII.2 and density distribution in Figure AIII. 3.  170  Appendix III  Table AIII.l Composition of hypothetical cyclone stream Fraction  Water Light Ends CGO OTSB  wt%  19 10 54 15  Table AIII.2 Composition of Light Ends fraction of cyclone stream Light Ends components  Hydrogen H2S Methane Ethane Ethylene Propane Propylene Butadiene Butenes i-Butane n-Butane  wt fraction (of Light Ends)  0.01 0.06 0.21 0.16 0.08 0.12 0.13 0.02 0.12 0.01 0.06  171  Appendix III  Table AIII.3 CGO assay Method: ASTM 2887 with HTSB enhancement  ASTM D2887 0 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 95 100  TBP (°C)  Vol % NBP ( C) 221 266 287 304 319 333 345 357 368 380 391 403 414 426 438 450 464 479 496 521 572 U  wt%  Table AIII.4 CGO TBP data Method: TBP calculated by HYSYS.  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  229 237 248 263 275 288 297 306 314 322 329 342 354 366 378 390 402 414 426 438 450 462 474 487 501 511 523 531 541 541 541  172  Appendix Table AIII.6 OTSB TBP data Method: TBP calculated by HYSYS  Table AIII.5 OTSB Assay Method: ASTM 2887 & SCFEcomposite data ASTM y. Boil 0 i 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  SCFE Boil Temp MT't SBotl (C) 315.7 0 318.7 1 327.4 2 3.5 343.7 5 360.7 7.5 3838 10 4007 12:5 413 6 15 423 7 17.5 432 0 20 43S0 25 45C 5 30 45S6 35 468 7 477 1 40 484 7 45 50 492 1 55 499 3 60 506:3 65 512.9 70 518.5 75 5215 80 526.0 85 529.2 90 532.4 92.5 534.0 535.7 95 96.5 536.8 98 5379 99 538 6 539 4 100  COMPOSITE 13oi Temp Boil Temp KBoil (Q (C) 3157 0 523.3 318.7 0.5 523.5 327.4 1 523.7 343.7 1.75 524.0 360.7 2.5 524.3 383.8 3.75 524.8 400.7 5 525 6 413.6 6.25 526.0 423.7 7.5 526.3 432.0 8.75 526.8 439.0 10 527.3 450.5 12.5 528.4 459.6 15 533.6 468.7 17.5 539.2 477.1 20 549.6 484.7 22.5 566.3 492.1 25 583.4 499.3 27.5 597.6 506.3 30 639.1 5119 32.5 695.7 518.5 35 759.0 522.5 37.5 825.8 526.0 40 892.9 529 2 42.5 957.1 5324 45 1015.1 534 0 46.25 1040.9 535.7 47.5 1063.9 536.8 48.25 1076.2 537.9 49 1086.1 538.6 49.5 1089.1 539.4 50 1092.2 523.5 50.5 523 7 51 524 0 51.75 524.3 52.5 524.8 53.75 525 6 55 526 0 56.25 526 3 57.5 526 8 58.75 527.3 60 528 4 62.5 65 67.5 70 72.5 75 775 80 825 85 87 5 90 925 95 96.25 97 5 98 25 99  533.6 539 2 549 6 566.3 583 4  99.5 100  1089.1  597 6 639.1 695 7 759.0 825 8 892.9 957.1 1015 1 1040S 1063 9 1076 2 1086 1 1092  2  ASTM vol % 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  TBP (°C) 341 345 355 374 394 419 436 448 458 466 472 483 492 501 509 517 525 532 540 544 547 ^52 553 555 558 558 560 561 562 562 563  SCFE TBP (°C) 523 524 524 524 524 525 526 526 526 527 527 528 534 539 550 566 583 598 639 696 759 826 893 957 1015 1041 1064 1076 1086 1089 1092  Appendix III  Table AIII.7 Cyclone Product TBP data Method: TBP calculated by HYSYS Vol % TBP [ C] -253 0 -247 1 -223 2 -192 3.5 -165 5 -128 7.5 -104 10 -93 12.5 -72 15 -50 17.5 -43 20 6 25 289 30 318 35 340 40 361 45 380 50 399 55 420 60 439 65 456 70 466 75 493 80 517 85 540 90 546 92.5 565 95 96.5 708 928 98 981 99 1028 100 U  -400 -200  0  200  400  600  800 1000 1200  Temperature, °C  Figure AIII.l Cyclone Product TBP curve  300  Figure A  I  I  600 900 1200 Molecular weight L  2  1500  1800  Cyclone Product molecular weight distribution curve  Table AIII.8 Cyclone Product composition Boiling range, "C  <=100 100-200 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>  Mole fractions  0.8921 0.0000 0.0162 0.0395 0.0353 0.0144 0.0005 0.0005 0.0004 0.0005 0.0006  Mass fractions  0.3042 0.0000 0.0599 0.2021 0.2543 0.1345 0.0057 0.0075 0.0064 0.0107 0.0148  200  400  600  1000  800  Liquid Density, kg/m  3  Figure AIII.3 Cyclone Product density distribution curve  174  1200  Appendix III  ATB Assay ATB experimental assay is collected by Syncrude Canada Ltd. using ASTM D2887 method for fractions up to 538°C, and HTSD enhancement for higher boiling components. This assay is shown in Table AIII.9. Distillation assay was used as input and TBP and composition were calculated by HYSYS. They are presented in Table AIII.10, Figure AIII.4 and Table AIII.l 1, respectively.  175  Appendix III  Table AIII.9 ATB assay Method: ASTM 2887 with HTSD enhancement ASTM D2887 wt% 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 61 62 63 64 65 66 67 68 69 70  Table AIII.10 ATB TBP data Method: TBP data calculated by HYSYS  262 277.5 296.5 308.5 318 327.5 336 343.5 350.5 357 363 369.5 375.5 381 387 392.5 397.5 403 407.5 412.5 417 421.5 426 430.5 435 439 444 448.5 453.5 458.5 463.5 468.5 473.5 478.5 483.5 488.5 494 499 504 509 514.5 519.5 524.5 530 535 540 545 550 555 560 564.5 569 573.5 578 583 587.5 591 595 599.5 604 613 617.5 622 627 631.5 636 640.5 645.5 651 655.5  TBP ("C)  Vol %  NBP (°C)  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  200  316 326 347 368 384 405 422 437 451 464 475 497 519 544 569 594 618 640 662 683 702 722 744 776 804 838 884 922 962 987 1009  400 600 800 Temperature, °C  1000  1200  Figure AIII.4 ATB TBP curve Table AIII.l 1 ATB composition calculated by HYSYS Boiling range, C <=100 100-200 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>  Mole fractions 0.000 0.000 0.000 0.109 0.258 0.238 0.212 0.105 0.061 0.010 0.007  176  Appendix III  HGO Assay Experimental assay is also collected using ASTM D2887 method with HTSD enhancement. Assay is presented in Table AIII.l2, TBP data calculated by HYSYS in Table AIII.l3 and Figure AIII.5 and the composition calculated by HYSYS in Table AIII.l 4.  177  Appendix III  Table AIII.12 HGO assay Method: ASTM 2887 with HTSB enhanc. ASTM D2887 NBP (°C)  wt% 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 61 62 63 64 65 66 67 68 69  263.5 275 289.5 298.5 305.5 310.5 315.5 320 324.5 328.5 332 336 339.5 342.5 345.5 348.5 351.5 354 357 359.9 362 364.5 367 369.5 372 374.5 377 379 381.5 384 386 388.5 390.5 393 395 397 399.5 401.5 403.5 405.5 408 410 412 414 416 418 419.5 421.5 423.5 425.5 427.5 429.5 431 433 435 437 439 441 443 445 447 449 451 453.5 455.5 457.5 460 462 464.5 467  Table AIIL13 HGO TBP data Method: TBP calculated by HYSYS  Table AIII.12 cont.  70 71 72 73 74 75 76 77 78 79 80 81 82 83 84 85 86 87 88 89 90 91 92 93 94 95 96 97 98 99 100  469 471.5 473.5 476 478.5 481 483.5 486 488.5 491.5 494.5 497 500 503 506 509.5 513 517 520.5 525 529.5 535 540.5 547 555 564.5 576 591.5 615.5 680.5 750  Vol %  TBP ( C) 281 289 302 317 327 339 348 357 365 372 378 390 402 413 424 435 445 455 465 475 486 498 510 526 544 567 603 643 719 763 797 U  0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  400 600 Temperature, °C  1000  Figure AIII.5 HGO TBP curve Table AIII.14 HGO composition (HYSYS) Boiling range, °C  <=100 100-200 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>  Mole fractions  0.000 0.000 0.027 0.334 0.493 0.113 0.022 0.008 0.004 0.000 0.000  178  Appendix III  Scrubber Overhead Scrubber Overhead is the final product of the Fluid Coker. As a vapour stream, Scrubber Overhead exits from the top of the Scrubber Section of a Fluid Coker and enters the fractionator where four fractions are separated: Sour Gas, Butane, Naphtha and Combined Gas Oil (CGO), consisted of Light Gas Oil (LGO) and Heavy Gas Oil (HGO). Composition and Simulated Distillation Data were provided by Syncrude Canada Ltd. Weight fractions of all four streams are given in Table AIII.l 5. Composition of Sour Gas is given in Table AIII.l 6. Composition and distillation data for Scrubber Overhead itself were not available. In order to define real plant Scrubber Overhead and be able to compare its characteristics with the simulated Scrubber Overhead, four fractions mentioned above were simulated as four streams and mixed together. The resulting stream was assumed to have the characteristics of the plant Scrubber Overhead. Simulated Distillation Data for Naphtha and CGO fractions, obtained from Syncrude Canada Ltd., were entered in the simulation model as assays and based on that HYSYS calculated TBP distillation curves. TBP distillation data calculated by HYSYS for Naphtha and CGO are shown in Table AIII.17 and AIII.l8, respectively. The resulting mixture, "plant" Scrubber Overhead, TBP data and TBP curve, calculated by HYSYS, are shown in Table AIII.l9 and Figure AIII.6, boiling range composition and fraction distribution in Table AIII.20, molecular weight distribution in Figure AIII.7 and density distribution in Figure AIII.8.  179  Appendix III  Table AIII.15 Scrubber Overhead fractions Fraction Sour Gas Butane Naphtha CGO  wt% 12 5 22 61  Table AIII.16 Sour Gas composition Component Hydrogen H20 H2S Methane Ethane Ethylene Propane Propene 1-Butene Biacetylene i-Butane n-Butane i-Pentane n-Pentane 13-Butadiene cis2-Butene tr2-Butene CO C02  Mole Fractions 0.113 0.029 0.138 0.317 0.133 0.062 0.086 0.061 0.001 0.001 0.004 0.010 0.000 0.001 0.002 0.002 0.011 0.007 0.021  180  Appendix III  Table AIII.17 CGO Assay Method: SIM Dist  SIM Dist  wt % 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 45 46 47 48 49 50  NBP [ C] 198 225 255 257 264 270 275 280 285 290 294 298 302 306 310 314 318 321 325 328 332 335 338 342 345 348 351 354 357 360 363 366 369 371 374 377 380 383 386 388 391 393 396 399 401 404 407 409 411 414 U  Table AIII.17 Cont  51 52 53 54 55 56 57 58 59 60 61 62 63 64 65 66 67 68 69 70 71 72 73 74 75 76 77 78 79 80 81 82 83 84 85 86 87 88 89 90 91 92 93 94 95 96 97  416 419 421 423 425 428 430 433 435 438 440 443 445 448 450 453 456 458 461 464 466 469 471 474 477 480 482 485 487 491 493 497 500 503 507 510 514 518 523 527 533 539 545 554 564 578 597  Table AIII.18 CGO TBP data Method: TBP calculated by HYSYS Vol % 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  TBP (°C) 220 241 266 276 286 298 310 321 330 339 347 362 377 391 405 417 429 441 453 466 478 492 502 520 542 564 594 628 683 716 735  181  Appendix III  Table AIII. 19 Naphtha Assay Method: SIM Dist SIM Dist wt %  NBP [°C] 1 -30 2 -7 -3 3 5 1 34 10 47 15 67 20 25 80 30 93 35 106 116 40 45 127 50 140 55 151 60 165 65 176 70 190 75 201 212 80 85 223 235 90 95 251 97 259 265 98 99 276  Table AIII.20 Naphtha TBP data Method: TBP calculated by HYSYS TBP (°C)  Vol %  0  -62  1 2  -42 -17  3.5 5  -8 0  7.5 10 12.5  15 30 40 47 53 62  15 17.5 20 25 30 35 40 45 50 55  77 88 97 107 118 131 144  60 65 70  158 171  75 80 85 90 92.5 95 96.5 98 99 100  196 207 217 229 235 243 250 258 264 264  184  Appendix III  Table AIII.21 "Plant" Scrubber Overhead TBP data Method: TBP calculated by HYSYS Vol % 0 1 2 3.5 5 7.5 10 12.5 15 17.5 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 92.5 95 96.5 98 99 100  TBP ( C) -253 -216 -178 -170 -154 -116 -94 -84 -60 -47 -41 -1 11 84 130 188 237 298 339 371 400 426 450 476 500 518 540 568 614 660 691 U  Temperature, C Figure AIII.6 "Plant" Scrubber Overhead TBP curve  100  Table AIII.22 "Plant" Scrubber Overhead composition and fraction distribution Boiling range, °C  <=100 100-200 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>  Mole fractions  0.681 0.087 0.052 0.082 0.069 0.025 0.004 0.001 0.000 0.000 0.000  200  300  400  Molecular Weight  500  600  Figure AIII.7 "Plant" Scrubber Overhead molecular weight distribution curve  Mass fractions  0.239 0.100 0.088 0.204 0.235 0.106 0.020 0.008 0.000 0.000 0.000  Vol. fraction C4-(<177°C) LGO (177-343°C) HGO(343-524°C) 524+(>524°C)  0.070 0.250 0.579 0.101  0  200  400  600  800  Liquid Density, kg/m  1000  1200  Figure AIII.8 "Plant" Scrubber Overhead density distribution curve 183  Appendix IV  Appendix IV - Cyclone Liquid Droplets Trajectory The vapour and the liquid droplets of the Cyclone Product come out of the cyclone snouts at velocity  VQ^SO  ft/s=76.1 m/s, and could be considered as a nearly horizontal jet above the  Scrubber Pool. The surrounding vapour velocity (9.8 m/s) was estimated based on the total volume flow rate for Scrubber Pool vapour and Cyclone Product and cross section area of the column. Longitudinal and vertical distribution of velocity for a droplet and distance from the nozzle can be calculated. Basic information for the calculation is given below: p = 0.11423 lb/ft = 1.7145 kg/m is density of the Cyclone Product vapour phase* 3  3  g  u.g = 3.11 • 10" cP = 3.11 • 10 kg/ms is viscosity of the Cyclone Product vapour phase* 2  5  pi = 759.32 kg/m is density of the Cyclone Product liquid phase - liquid droplet* 3  ai = 6.357-10" lbf/ft is surface tension of the Cyclone Product liquid phase - liquid droplet* 4  D =0.58 m is nozzle diameter 0  d =1.081-10" m is droplet diameter; the whole calculation is done using this diameter, calculated 5  p  based on the cyclone cut point *) Note: These values are calculated by HYSYS based on the assay data for the Cyclone Product  Longitudinal distribution of velocity (vi) for the droplets was calculated based on equations for a turbulent free jet. A turbulent jet is a free jet with Reynolds number greater than 2000. In the case of Cyclone Product, Reynolds number is:  R = °' °' D  e  V  Pg  = 24-10  5  (TV-5)  The equation is applicable for the air jet into the surrounding air. Density gradient between the jet fluid and surrounding fluid has effect on the spread of the jet. Since Cyclone Product vapour, 184  Appendix IV  as jet fluid, has similar density as the surrounding vapour, as in the original case with air, this equation was used without change for the system under study:  v, =v -K— n  A,  for 7< — <100 or 4.06<X<58m  A>  (IV-2)  AT = 6.2  for  v = 10 to 50 m/s 0  where: vj is longitudinal distribution of velocity along the center line of the jet vo = 76.1 m/s is exit velocity of the jet X is the horizontal distance from the exit of the nozzle Do = 0.58 m is the nozzle diameter This equation applies for the distance between 4.06 and 58 m from the nozzle. After inserting the values, the equation for velocity distribution within this distance is:  v , ( ^ W < 5 s = ^ | ^  far  4.06<X<58m  (IV-3)  At position X=4.06 m, the velocity would be vi = 67.4 m/s From the exit of the nozzle up to 4.06 m, linear change of velocity was assumed, changing from 76.1 m/s to 67.4 m/s. Hence, the linear equation for velocity distribution within the distance from 0 to 4.06 m would be:  ",Wo<,<,«=v„-2.14^  for  0<X<4.06m  (IV-4)  Since v, = — , after integration of both Equations (IV-3) and (IV-4) in time and values input, dt  following two equations for horizontal distance change in time were obtained:  185  Appendix IV  X(t) =  ^(l-e- ') 2U  2.14  X(t) = ^16.5 +547.3(^-0.0665)  for distance 0 < X < 4.06m  (IV-5)  for distance 4.06 <X< 58m  (IV-6)  Vertical distribution of velocity (v ) can be calculated from the vertical force balance (weight v  of the droplet against the drag force) and the terminal velocity of the droplet: mg-F =m^ D  (IV-7)  dt  In this equation g is gravitational acceleration, m is droplet mass, and F D is a drag force:  F =C {\p v ){^d )  (IV-8)  2  D  D  g  v  p  Velocity in this equation is relative velocity between droplet (particle) and surrounding gas, v =v - v , dp is droplet (particle) diameter, and Co is drag coefficient. For the spherical particles v  p  g  it can be calculated from: 24 u„  24 c  R  =  i  r  =  —  (  I  V  -  9  )  p- -p  e  v  g  d  v  dv Terminal velocity can be calculatedfromEquations (IV-7), (IV-8) and (IV-9) when—- = 0: dt dv mg-2>n/u d v = - ^ = 0 g  p  v  (IV-10)  m  Since m- p, -V - p, -(^mi ) is the mass of the spherical liquid droplet, Equation (IV-10) for 3  p  p  droplet terminal velocity becomes: Prd -g  mg  2  57tp d  p  g  p  18//,  (IV-11)  If the droplet was moving through a stagnant gas, terminal velocity would be VTO=0.0016 m/s for the system under study, and from Equations (IV-7), (IV-8) and (IV-9):  186  Appendix IV  mg-3xjU d v g  dv  v  p  v  =m  g — ^ m  dt  dy^  (IV-12)  dt  =g - ^  (IV-13)  =g  To  V  T0  y  J  Integrating velocity in time from t=0 to t: v (0 = v- l-exp(—^-0 ro  (IV-14)  v  dY  Since v = -f-, using Equation (IV-14) and integrating vertical distance Y in time from t=0 to t, v  dt  the following Equation for vertical distance change in time can be obtained: Y 't) = v -t + 0  T0  g  ^ exp(—^--0-1 T0  V  (IV-15)  J  where Yo represents vertical distance from the nozzle exit if the surrounding fluid is stagnant. Note: Y is directed downwards. However, droplets are not moving through a stagnant gas. In this investigation, the surrounding gas was considered to be the vapour that originates from Scrubber Pool in combinations with Cyclone Product vapour. Properties of this combined vapour were used for the calculations. The velocity of these vapours is calculated based on the total volume flow and Scrubber crosssection, and its value is 9.8 m/s. This velocity is included in Equations (IV-7), (IV-8) and (IV-9) through the slip velocity. The new terminal velocity also includes slip velocity: T  = T0  V  (IV-16)  ~ g  V  V  where v is the terminal velocity in the flowing surrounding gas, v = 9.8 m/s is velocity of the T  g  surrounding gas, VTO is the terminal velocity in the stagnant gas. From Equations (IV-7), (IV-8) and (IV-9), the same Equation (IV-14) is derived for change of vertical velocity in time: v (0 = v v  ro  :  l-exp(—^-0  (IV-17)  187  Appendix IV But, the real droplet velocity is influenced by the flowing velocity of the surrounding gas, and its value is: V  V«B/(0  =  (IV-18)  V (0-V^ V  and hence: vreal  8 (t) = v l - e x p ( - — t ) — v„  (IV-19)  T  dY Again, v , - — , using Equation 19 and integrating vertical distance Y in time from t=0 to t, dt vrea  the following Equation for vertical distance change in time can be obtained: (IV-20)  Y(t) = Y (t)-v-t 0  Trajectory of a liquid droplet within the space above the Scrubber Pool was calculated from Equations (IV-5, 6, 15 and 20) inserting the time. The calculation was done for the largest droplet diameter present in the jet (1.08M0" m), assuming that all others would be carried even 5  further. The trajectory is presented the Figure AIV. 1.  0) N N O  c 0)  E o L.  0) o c ro _</>  •5 « o  r >  1  2  3  4  5  6  7  8  Horizontal distance from the nozzle- x, m  Figure AIV.l Trajectory of a liquid droplet carried with the Cyclone Product jet 188  

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