delayed coker in hysys 1

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


4.2. Droplet Size Estimation 31

iii

4.3. Trajectory of the Liquid Droplets 32

Chapter 5 - Presence of Heavy Components in the Scrubber Overhead 37


5.2. Liquid Entrainment in the Shed Section 38

5.3. Packed Section 40

5.4. Conclusion 46

Chapter 6 - Case Studies: Results and Discussion 47


6.2. 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


Scrubber Overhead properties 110


Scrubber Bottom properties I l l

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 91

Figure VI-2 Effect of number of Sheds trays on temperature profile along the Scrubber 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


Scrubber Bottom TBP curve I l l


Scrubber Overhead composition 112


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


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


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


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


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


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 pseudo-

components if complex mixtures are involved;

• Installing the reaction components and formulating reactions, if they occur;

10


• 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 hydrocarbon-

hydrocarbon and hydrocarbon - nonhydrocarbon pairs available in the HYSYS library [16]. The

Peng-Robinson EOS is presented below:

P = - a- (2.1) V-b V(V + b) + 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


molecules. These parameters can be determined from critical values P c and T c, and the acentric

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 3 - ( \ -B)Z 2 +Z(A-3B2-2B)-(AB-B2-B3) = 0

R 2 T 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 c, pressure P 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 c, P c, acentric factor and molecular weight are used [16, 26]. Twu's model for viscosity

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


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. Vapour-

liquid 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 rp /?-r-ln<D, = [(\>i-R—)-dp

0 P

v . = ( — ) T (2.5)

dV where the molar volume v and the derivative can be calculated using EOS.

dnt

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


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 up-

going 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


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


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. Al 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


Core operation blocks:

1. The Koch Grid is modeled as a packed absorption column, with the same height (1.8 m),

diameter (9 m) and the type of packing (Koch Flexigrid Type 2) as in the plant. The

number of sections in the 10-layer grid was set to be 2. This is the minimum number of

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 of

liquid that decreases available space for the gas flow. A t the start of run (SOR) of the

Syncrude Canada Ltd. 's Fluid Coker the pressure in the Scrubber was 117.21 kPa and

pressure at the top of the Koch Gr id was 117.14 kPa. During operation, fouling of the

grid and the cyclone exit nozzles occurs due to the coke formation, causing an increase in

pressure drop. In order to maintain sufficient production of the Overhead, pressure in the

Scrubber was raised gradually by the operators. A t the end of run (EOR), it was typically

186.16 kPa, and the pressure at the top of the K o c h Gr id was 185.53 kPa. For the Base

Case, start-of-run (SOR) conditions were used, and pressure effects from S O R to E O R

were simulated in Case Study VIII, Chapter 6.

2. The Shed section consists of six sets of sheds, which are about 1-m wide and with 1.2-m

horizontal spacing between them. Sloped from the both sides in the direction of liquid

flow, and with serrated weirs, the sheds improve the distribution of 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 of which reflect the Shed section

geometry exactly. The Shed section was therefore simulated as an absorption column of

six trays (one for each set of sheds). Although there was a concern that a tray column can

simulate the Sheds sufficiently well (in the Sheds, there is no bubbling of 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 of trays specified in the tray

column corresponds to the number of shed sets in the Shed section. The dimensions of

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


4. 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 by

passes 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


External streams, such as Cyclone Product, A T B feed, HGO 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.

An 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


Sheds through the six trays, contacting the falling liquids from the top of the Sheds - ATB 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. An 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

S c r u b b e r O v e r h e a d

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 m V h

MotocUarWaigM 70.77

Mass Density 2.834 kgAn3

GrtdUquM

Temperature 395.0 C

Pressure 17.00 p**g

Mass Flow • 2.179e+O05 kgm

Actual Volume Flow - 304.7 nvVTi

H G O ( u n d e r w a s h )

ShedVapour

Temperature 404.9 C

Pressure 17 00 psig

Mass Ftow 8 705o'005 kg/h

Actual Votume Flow 2.773e«00S m3m

HGO (ondorwash)

Temperature 325.0 C

Pressure 200.0 psig

Mass Flow S.247e*O04 kgm

Actual Volume Ftow 66 26 m3m

Shed Liquid

Temperature 473.4 c -

Pressure 17.00 ps<g

Mass Flow 3.613o*OQS k g *

Actual Voluma Ftow 490.5

L l q l ids to So jbber '

"A

Agnation Steam

AgfUfon Steam

Temperature 185.0 C

Pressure . 150.0 psig

Mass Ftow 2,04U*004 kgm

Actual Voluma Ftow 23.47 m3fh

ATB

Temperature 325.0 C

Pressure 17.00

Actual Voluma Ftow 363.5

Mass Ftow 3.019tt*O05 k g *

Vapour to Sheds

Temperature 514.0 C

Pressure 17.00 psig

Mass Ftow 6.S96**005 kgm

Actual Volume Ftow 3.1SOe*«M rrvVh

Cyctorw Product

Temperature 539.9 C

Pressure 17.00 P»«g

Mass Ftow S.470e*OOS kg/ti

Actual Votume Ftow 2.846e*O05 raJ/H

Ligh t E n d s

• Scrubb.Pooi liquid'

Temperature 375.0 C

Pressure 17.00 psig

Mass Ftow 6 4370*005 kgm

Actual Volume Ftow 789.0 HVWl

Molecular Weight 6370

Mass Den iffy 815 9 kgAn3

Viscosay 0 5665 CP

• Hot Scr.Recyde

Actual Volume Flow 459.2 mim

Mass Flow' 3.746fl*O05 kgm

Actual Volume Ftow

Figure 3.2 Simulation flowsheet of the Scrubber Section

22


3.2.5. Input Plant Data

Plant data obtained from Syncrude Canada Ltd. was used as inputs in the simulation system.

Flow rates, temperatures, densities, viscosities and distillation data (or composition) of the

following streams were provided:

-Cyc lone Product - Cyclone Product is the product o f the coking zone of 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 of liquid and

even solids are also present. The composition of this stream was defined in the 1980's

when the Coker was run in "once through" mode (no recycles or additional input

streams were used during the operation). It contained water, light ends, C G O (Coker

Gas Oil) fraction and O T S B fraction (OTSB-Once Through Scrubber Bottom, a

mixture of heavy fractions, some of which boil above 1000°C). The characterization of

the Cyclone Product is explained in detail in [31]. The weight percents of these four

fractions, as wel l as composition of 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 High Temperature Simulated Distillation (HTSD)

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 (SCFE) was used for fractions above 524°C [32].

- A T B - Atmospheric Topped Bitumen is a product o f atmospheric distillation of

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 of 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 kg/m3

Viscosity, cP at 20°C/at 30(fC

Temp. °C

Pressure psig (kPa)

HGO Wash 24 kbpdb ASTM D2887/HTSD 987 1737/ 15 325 200(1380) HGO Underwash lOkbpd A S T M D2887/HTSD 987 1737/ 15 325 200(1380) ATB 55 kbpd A S T M D2887/HTSD 1024 214016/3 325 17(117.21) Overhead/ATB ratio 2.51 kg/kg Cyclone Product Water Light Ends CGO OTSB

150 kg/s 10wt% 12wt% 61 wt% 17 wt%

Known composition A S T M D2887/HTSD ASTM D2887/SCFE

540 17(117.21)

Agitation Steam 6 kg/s 185 150(1035) Split SPL Rec.rTo Coker0 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 m3/day

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


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. a , b

Unit Parameter Value

Koch Grid

Type of packing Number of layers Height, ft Diameter, ft Top pressure, psig Bottom pressure, psiq

Koch Flexigrid 2 10 6 30

16.99 (SOR)-26.91 (EOR) 17 (SOR)-27 (EOR)

Sheds Number of trays Pressure, psig Delta P, psi

6 17 (SOR)-27 (EOR)

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

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


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


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


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"10) to match the Scrubber Overhead content of heavy fractions (524°C+) to

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.(%) T(Ovhd) (°C) 390 393 0.8 T(Grid Bottom) (°C) 395 395 0.0 T(Shed Top) (°C) 407 405 0.5 T(Shed Bottom) (°C) 470 473 0.6 T(Scrubber Pool) (°C) 375 375 0.0 Ratio Overhead/ATB Mass Flows 2.50 2.57 2.8 Act. Vol. Flow(To Coker) (kbarrel/day) 50 49 2.0 Act. Vol. Flow( SPL Rec) (kbarrel/day) 74 73 1.4 Overhead 524°C- mole fraction 0.94 0.91 3.2 S u m of deviations: 11.3

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


Table 3.4 Determined unknown parameters (primary variables)

Primary Variables Value Shed Trays Efficiency 0.53a

Koch Grid Overall Efficiency 0.78a

Splitter TEE-102 (Upgoing Vapours): Vol. flow fraction to Flash Block 0.41 Splitter TEE-103 (Scrubber Liquid): Vol. flow fraction to Flash Block 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]:

where Dpth is theoretical particle size removed by the cyclone, p g is viscosity of the gas

(Cyclone Product vapour phase), Bc= Dc/4 where D c is the cyclone diameter. N s is effective

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. Vjn=74 m/s is the velocity at the inlet of the cyclone,

calculated based on the volume flow rate (80.6 m3/s) and the cyclone inlet cross-section (d=0.48

m), pp is the density of the Cyclone Product liquid phase and pg is the density of the Cyclone

Product gas phase.

All values used in Equation (4.1) are listed in Table 4.1.

(4.1)

31


Table 4.1 Parameter values used in Equation (4.1)

Parameter Value Parameter Value a

3.1 MO' 5 kg/ms N s for vo=76 m/s 5.8

D c 1.702 m PP 3 759.32 kg/m3

Bc= Dc/4 0.43 P g a 1.8298 kg/m3

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

24-105. 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, 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,=vn-K — for 1 < — < 100 or 4.06 <X< 58m

(4.2) K = 6.2 for v0 =10 to 50 mis

32


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 " 2 1 4 ' ) for distance 0<X<4m (4.3)

X(t) = y/16.5 + 547.3(f - 0.0665) 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-FD=m— (4.5) dt

In this equation g is gravitational acceleration, m is mass of the droplet, and F D is drag force:

FD=CD^pgv2X^dp) (4.6)

Velocity v=vp - vg is the slip velocity between the droplet (particle) and surrounding gas, dp

is the droplet (particle) diameter, and Co is drag coefficient. For the spherical particles it can be

calculated from:

33


24 24« CD = — = ^ — (4.7)

Re pg-v-dp

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 ( 0 = v r o . f + ^

g exp(—^--0-1

v. ro

(4.8)

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 S (4.9)

where vT is the terminal velocity in the flowing surrounding gas, vg = 9.8 m/s is the velocity of

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) = Y0(t)-vg-t (4.10)

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"6 m), assuming that all others would be carried even further. The

trajectory is presented in Figure 4.1.

0 1 2 3 4 5 6 7 8 9 o 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


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 by

pass. 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"10). For all these cases the sum of deviations from

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


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 m3/h, and liquid

volume flow rates in hundreds of m3/h), suggesting that extensive liquid entrainment may be 38


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(Pg/P[)05 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


Table 5.1 Parameter values for calculation the flow and capacity parameter for Figure 5.1

Parameter Value Parameter Value

U N F 1.35 m/s a 13.895 dyn/cm

PL 706.1 kg/m3 L 158 kg/s

P G 2.64 kg/m3 G 183 kg/s

The calculated value for the flow parameter F L G is 0.052, and for the capacity parameter, C S B

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 C S B 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


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 s , and liquid loading, a design pressure drop can be

determined [40]:

41


F s , m/s • (kg/m3)1

0.5 0.6 0,8 I

1 O.S

0:6 0;5 0.4

03

0.2

<1

FLEXIORID 2

0.1 ; —^ 0.08 < 0.06 0.05 0.04 0.03 •

0.02

_ liquid U M J I J JbpCwieBosae gpm,«J m V h

171 14?

sa 122 & sa

73 m 10 24

s 12 0 tt

S (S«m Aif-Walaf, AiflUwil Tower: M B a m a W

0.4 0 5 0 A 0.8 I — 2

F s . ft/s • (lb/ft3)'*

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/m3 and gas velocity through the net area (Shed Vapour), UG= 1.17 m/s the

superficial factor is calculated to be Fs=2.07 m/s-(kg/m3)0,5( vertical dashed line on the chart).

This value along with the liquid loading:

150m3//? m3/h L P = Vol. Flow rate/Cross sec. area = — - = 2.28 — (steep dashed line on the chart) 65.58m' 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 1 DESCRIPTION Packed Colum BED NUMBER

LOADINGS Vapor Rate Scale Factor 1.00 Vapor Rate kg/hr 870500 Vapor Density kg/m3 3.1400 Vapor Volume m3/s 77.01 Vapor Viscosity CP 0.0010

Liquid Rate Scale Factor 1.00 Liquid Rate kg/hr 126100 Liquid Density kg/m3 840.00 Liquid Volume m3/h 150.12 Surface Tension dyne/cm 14.69 Liquid Viscosity CP 0.565

System Factor 1.00

Packing Type FLEX IGR ID® 2 SS

METAL

Tower Diameter mm 9141.00 Tower Area m2 65.63

Fs rn/s(kg/m3)/U5 2.08 Cv m/s 0.07 Liquid Loading m3/h/m2 2.29

Pressure Drop mbar/m <0.5 O.fiJ-

Calculated Capacity % 34 Constant L/V

WARNINGS: 1,

WARNINGS: 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 0 0 5

PG (5.1)

44


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 p = 4 ft"1, the calculated value for

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 t 1.17 m/s V 16.53 cS

P L 840.6 kg/m3 L 35 kg/s

PG 3.14 kg/m3 G 241 kg/s

In Table 5.3 U t is Shed Vapour superficial velocity through the net area, po is Shed Vapour

density, p L is liquid density (HGO Wash), v is kinematic viscosity of the liquid, L is liquid

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 20/ft

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


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"10), which resulted in the presence of high boiling fractions in the

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


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


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


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


- Grid Top

- Grid Bottom

- Shed Top

- Shed Bottom

- Scrubber Fbol

50 60 70 80 ATB Flow Fate, kbarrel/day

(1 barrel=0.0049684 m3/h)

Figure 1-1 Effect of ATB flow rate

on temperatures along the Scrubber

90

• — 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

10 20 30 40 50 Position, feet from bottom of scr. pool

Figure 1-2 Effect of ATB flow rate on

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 I-

7).

• Scrubber Bottom properties:

-Actual volume and mass flow rate increase by 75% and 71% (Table 1-1).

- Density drops from 816 to 796 kg/m3 (Table 1-1).

-Average molecular weight changes from 637 to 594 (Table 1-3).

50


-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).


mass flow rate of Scrubber Overhead

and Bottom


mass flow rate of other streams

51

Chapter 6 - Case Studies: Results and Discussion Table 1-1 Effect of ATB flow rate on Scrubber parameters

AJBRowRie ATB\fclirre RowRie Posftkn ft rrffh 273 298 323 348 373 397

fronthe kbarrel/da/ 55 eo 65 70 75 80 pod bcttcm (Base Case) % % % % %

KrhQidTcp 43 Tcp Sags TerrpHft) 393 390 -0.7 388 -1.4 385 -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 -0.8 398 -1.6 395 -24 392 ^ 2 383 4.8 S-BdsBA 22 BatariSagsTa-rpHft) 473 467 -1.4 461 -26 457 ^ 5 453 4.3 447 -56 Sditrja-Rrl 0 BJk Liqjd Terrperalre f t ) 375 382 1.8 387 32 391 4.3 394 51 398 62

Row Rates& Densities SmiJberOatHad AdLEl VJirre Rcw(rrf/h) 274,787 273272 -0.6 271,620 -12 269,930 -1.8 268209 -24 266,026 -32

MassRcwfkgti) 778,651 770,080 -1.1 760233 -24 749,638 •3.7 738449 -52 723279 -7.1 MassDaBty^rrf) 2S 282 -06 280 -12 278 -20 275 -28 272 4.1

Soii t iRxlUojd Adual\«LnBHcw(rrfyh) 78 90EJ 14.7 1,024 298 1,147 453 1271 61.1 1,386 755 MassRow(l<glri) 64372c 729293 133 817,363 27.0 907,871 41.0 999,796 553 1,103,082 71.4 l\tettnsity(l^rr?) 816 806 -12 798 -21 792 -29 786 -36 796 -24

OsfEad/AlB MassRowFaio 25E 234 -93 21; -172 1.SE -242 1.8C -30.3 1.65 -331 QidUc)jd A i d Mtiurre Rcw(rr?/h) 30: 296 -30 287 -58 279 -8.5 270 -11.2 26C -14.5

MassRowfkJh) 217,831 211,767 -28 206219 -53 200,755 -7.9 195,376 -103 183,811 -133 MBsDarstyfkg'rri) 714.97 71644 02 71824 Q5 7202C Q7 72229 1.0 724.99 1.4

3T6d\^xr AdLEl Vtiure Row(rr?/ri) 277,32 275,780 -0.6 274,07E -1.2 272,336 -1.8 270,551 -2 263281 -33 MassRcw(l^h) 870,45c 855787 -1.7 840,392 -35 824,337 -53 807,76E -72 783,02 -97 ^[irsty(te|'rrr) an 310 -1.1 307 -23 306 <16 299 4.9 293 •6.7

SBdu'cud AdLel\/tiuTeRcw(rr?/ri) 497 532 132 627.66 263 694 397 760 529 819 64.9 IVteRcw(kglh) 361,32 407,740 128 454,777 259 502,535 391 549994 522 598,614 64.3 M3EsD3Taty(ko/rr?) 724.7E 724.94 Q0 724.56 QO 724.31 -Q1 724.06 -Q1 724.53 QO

ToCtter Aiel UiLrre Rcw(rr?/h) 33C 378 14.7 42c 298 479 453 531 61.1 579 755 MassRcw(t0h) 269,07? 304,846 133 341,656 27.0 379490 41.0 417,915 553 431,03E 71.4 ^ D r a t y f l ^ r r r ) 8158: 803.17 -12 79334 -21 791.82 -29 78642 ^ 6 79542 -24

AUIiud irtcrrrdjcn NpCLTtDSBJS Ta-rperakreft) 514 511 -0.6 506 -12 505 -1.7 503 -22 501 -26 LpgdngSreari Terrpadueft) 534 534 -01 533 -Q1 533 -02 533 -03 533 -Q3 SBCfe SagsBlidaTy Q53 Q53 Q0 Q53 QO Q53 QO Q53 QO Q53 QO HfacfiQid SageBlidercy Q75 Q75 00 075 QO Q75 QO 075 QO Q75 QO SR-Cder DJy(MvBuri) 44.32 44.32 QO 44.32 QO 44.32 QO 44.32 QO 7543 702

52


Table 1-2 Effect of ATB flow rate on Scrubber Overhead properties ATB Flow Rate (kbarrel/d) 55 (Basic) 80 Cut Point [%] 55 80

Temperature [°C] 393 376 TBP [°C] TBP [°C] Pressure [psig] 16.99 16.99 0 -253 -253 Molecular Weight 70.77 61.50 1 -237 -240 Mass Density [kg/m3] 2.83 2.49 2 -207 -211 Act. Volume Flow [m3/h] 282,483 273,389 3.5 -167 -174 Mass Enthalpy [kj/kg] -2924 -3279 5 -136 -143 Mass Entropy [kJ/kg-C] 5.30 5.48 7.5 -102 -107 Mass Heat Capacity [kJ/kg-C] 2.74 2.66 10 -85 -92 Vapor Phase Fraction ( Mass Basis) 0.93 0.92 12.5 -51 -62 Specific Heat [kJ/kgmole-C] 193.80 163.43 15 -34 -45 Std. Gas Flow [STD_m3/h] 260,141 255,072 17.5 -3 -8 Watson K 11.38 11.43 20 266 206 Liq. Mass Density (Std. Cond) [kg/m3] 930.01 921.05 25 310 298 Molar Volume [m3/kgmole] 24.98 24.65 30 336 324 Mass Heat of Vap. [kj/kg] 2825 2843 35 354 341

40 373 363 Fraction Distribution Data 45 391 377

Volume fraction 50 405 391 C4-(<177°C) 0.060 0.066 55 420 405 LGO (177-343°C) 0.259 0.287 60 439 419 HGO (343-524°C) 0.580 0.553 65 442 436 524+ (>524°C) 0.101 0.094 70 459 441

75 471 448 80 486 466 85 496 484 90 525 523

92.5 537 532 95 550 546

96.5 556 556 98 616 617 99 744 753

100 871 888 100

10 -

0 J , , , , , , , , , 0 100 200 300 400 500 600 700 800 900 1000

Temperature, °C

Figure 1-5 Effect of ATB flow rate on Scrubber Overhead TBP curve

5 3


Table 1-3 Effect of ATB flow rate on Scrubber Bottom properties

ATB Flow Rate (kbld) 55 (Basic) 80 Cut Point [%] 55 80

Temperature [°C] 375 398 TBP [UC] TBP [°C] Pressure [psig] 17.00 17.00 0 412 387 Molecular Weight 637.04 594.06 1 441 421 Mass Density [kg/m3] 815.85 796.42 2 466 440 Act. Volume Flow [m3/h] 812 1,426 3.5 490 463 Mass Enthalpy [kJ/kg] -1329 -1325 5 504 466 Mass Entropy [kJ/kg-C] 3.39 3.38 7.5 514 483 Mass Heat Capacity [kJ/kg-C] 2.89 2.92 10 517 487 Vapor Phase Fraction ( Mass Basis) 0.00 0.00 12.5 520 491 Specific Heat [kJ/kgmole-C] 1843.61 1731.85 15 525 496 Std. Gas Flow [STD_m3/h] 23,893 43,904 17.5 548 512 Watson K 11.41 11.44 20 556 514 Kinematic Viscosity [cSt] 0.69 0.74 25 564 519 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1023.66 30 592 525 Molar Volume [m3/kgmole] 0.78 0.75 35 605 556 Mass Heat of Vap. [kJ/kg] 1309 1532 40 630 590 Surface Tension [dyne/cm] 15.47 14.74 45 635 597 Thermal Conductivity [W/m-K] 0.13 0.13 50 682 628 Viscosity [cP] 0.57 0.59 55 684 635

0.78 0.75 60 693 682 Fraction Distribution Data 65 706 687


92.5 892 854 95 917 897

96.5 964 940 98 1031 965 99 1047 1002

100 1055 1041 100

1100 Temperature, °C

Figure 1-6 Effect of ATB flow rate on Scrubber Bottom TBP curve 3±


0.50 1 m 55 kbarrel/day ATB : Light Ends:20%; Water: 65% ; 100> fraction: 15%

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

0.10

0.00

• 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%

200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>


Figure 1-8 Effect of ATB flow rate on Scrubber Bottom composition

55


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


I I . HGO Wash F l o w 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:


- 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)

500

480

460

of 440

| 420

400

380

360

- Grid Top - Grid Bottom - Shed Top • Shed Bottom - Scrubber Bool

— A —

=8=

20 30 HGO Flow Rate, kbarrel/day

(1 bar re 1=0.0049684 m3/h)

480

460

440

420

400

380

360 40 0

- • — 24 kbarrel/day HGO

- • - - 30 kbarrel/day HGO

- 40 kbarrel/day HGO

10 20 30 40 Position, feet from bottom of scr . pool

50

Figure II-l Effect of HGO Wash flow

rate on temperatures along the Scrubber

Figure II-2 Effect of HGO Wash flow

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/m3 (Table II-l).

-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/m3 (Table II-l).

-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


800

700

600

20 30 40 HGO Flow Rate, kbarrel/day

(1 barrel=0.0049684m3/h)

Figure II-3 Effect of HGO Wash flow

rate on mass flow rate of Scrubber

Overhead and Bottom

900

850

800

750

700

650

600

550

500

450

400

350

300 -I 250

200

150

100

- Grid Liquid

- Shed Vapor

- Shed Liquid


(1 barrel=0.0049684m3/h)

40

Figure II-4 Effect of HGO Wash flow rate

on mass flow rate of other streams

59


Table II-l Effect of HGO Wash flow rate on Scrubber parameters

HGO Wash Flow Rate HGO Wash Volume Flow Rate Position ' ft m3/h 119 149 199

from the kbarrel/day 24 30 40 pool bottom (Base Case) % %

Koch Grid Top 43 Top Stage Temp Est (°C) 393 390 -0.7 386 -1.9 Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 395 393 -0.5 389 -1.6 Sheds Top 34 Top Stage Temp Est (°C) 405 404 -0.3 400 -1.1 Sheds Bot. 22 Bottom Stage Temp Est (°C) 473 466 -1.5 456 -3.7 Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0

Flow Rates& Densities Scrubber Overhead Actual Volume Flow (m3/h) 274,787 274,291 -0.2 273,178 -0.6

Mass Flow (kg/h) 778,651 788,004 1.2 797,485 2.4 Mass Density (kg/m3) 2.83 2.87 1.4 2.92 3.0

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 859 8.9 1,001 26.8 Mass Flow (kg/h) 643,728 695,136 8.0 797,839 23.9 Mass Density (kg/m3) 816 809 -0.9 797 -2.3

Overhead / ATB Mass Flow Ratio 2.58 2.61 1.2 2.64 2.4 Grid Liquid Actual Volume Flow (m3/h) 305 347 13.9 425 39.5

Mass Flow (kg/h) 217,861 247,811 13.7 303,893 39.5 Mass Density (kg/m3) 714.97 713.83 -0.2 715.03 0.0

Shed Vapor Actual Volume Flow (m3/h) 277,329 277,648 0.1 277,899 0.2 Mass Flow (kg/h) 870,450 878,893 1.0 892,151 2.5 Mass Density (kg/m3) 3.14 3.17 0.9 3.21 2.3

Shed Liquid Actual Volume Flow (m3/h) 497 546 9.9 639.68 28.8 Mass Flow (kg/h) 361,323 394,390 9.2 459,545 27.2 Mass Density (kg/m3) 724.76 722.61 -0.3 718.40 -0.9

To Coker Actual Volume Flow (m3/h) 330 359 8.9 418 26.8 Mass Flow (kg/h) 269,078 290,567 8.0 333,497 23.9 Mass Density (kg/m3) 815.85 808.87 -0.9 797.16 -2.3

Additional information Vapour to Sheds Temperature (°C) 514 511 -0.7 505 -1.8 Upgoing Stream Temperature (°C) 534 534 -0.1 533 -0.2 Sheds Stage Efficiency 0.53 0.53 0.0 0.53 0.0 Koch Grid Stage Efficiency 0.75 0.75 0.0 0.75 0.0 SPL Coler Duty (MMBtu/h) 44.72 45.75 2.3 48.25 7.9

6 0


Table II-2 Effect of HGO Wash flow rate on Scrubber Overhead properties

HGO Wash Flow Rate (kbarrel/day) 24 (Basic) 40 Cut Point [%] 24 40

Temperature [ °C] 393 386 TBP[°C] TBP[°C] Pressure [psig] 16.99 16.99 0 -253 -253 Molecular Weight 70.77 71.83 1 -237 -237 Mass Density [kg/m3] 2.83 2.92 2 -207 -206 Act. Volume Row [m3/h] 282,860 281,203 3.5 -167 -166 Mass Enthalpy [kJ/kg] -2924 -2909 5 -136 -134 Mass Entropy [kJ/kg-C] 5.30 5.17 7.5 -102 -101 Mass Heat Capacity [kJ/kg-C] 2.74 2.71 10 -85 -81 Vapor Phase Fraction (Mass Basis) 0.93 0.91 12.5 -51 -49 Specific Heat [kJ/kgmole-C] 193.80 194.57 15 -34 -22 Std. Gas Flow [STD_m3/h] 260,141 262,526 17.5 -3 2 Watson K 11.38 11.33 20 266 274 Liq. Mass Density (Std. Cond) [kg/m3] 930.01 940.10 25 310 311 Molar Volume [m3/kgmole] 24.98 24.60 30 336 336 Mass Heat of Vap. [kJ/kg] 2825 2800 35 354 352


Volume fraction 50 405 397 C4-(<177 °C) 0.060 0.052 55 420 407 LGO (177-343 °C) 0.259 0.269 60 439 420 HGO (343-524 °C) 0.580 0.584 65 442 440 524+ (>524 °C) 0.101 0.096 70 459 444

75 471 465 80 486 477 85 496 491 90 525 524

92.5 537 532 95 550 546

96.5 556 555 98 616 600 99 744 733

100 871 863

100

90

80

70-I

60

a 50

£ 40

30

20

10

24 kbarrel/day HGO

40 kbarrel/day HGO

100 200 300 400 500 600 Temperature, °C

700 800 900 1000

Figure II-5 Effect of HGO Wash flow rate on Scrubber Overhead TBP curve

61


Table II-3 Effect of HGO Wash flow rate on Scrubber Bottom properties

HGO Wash Flow Rate (kbarrel/day) 24 (Basic) 40 Cut Point [%] 24 40

Temperature [°C] 375 375 TBP [UC] TBP [UC] Pressure [psig] 17.00 17.00 0 412 343 Molecular Weight 637.04 596.68 1 441 398 Mass Density [kg/m3] 815.85 797.16 2 466 438 Act. Volume Flow [m3/h] 812 1,030 3.5 490 460 Mass Enthalpy [kJ/kg] -1329 -1322 5 504 467 Mass Entropy [kJ/kg-C] 3.39 3.39 7.5 514 485 Mass Heat Capacity [kJ/kg-C] 2.89 2.92 10 517 490 Vapor Phase Fraction ( Mass Basis) 0.00 0.00 12.5 520 495 Specific Heat [kJ/kgmole-C] 1843.61 1739.52 15 525 504 Std. Gas Flow [STD_m3/h] 23,893 31,616 17.5 548 513 Watson K 11.41 11.43 20 556 515 Kinematic Viscosity [cSt] 0.69 0.71 25 564 519 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1024.29 30 592 525 Molar Volume [m3/kgmole] 0.78 0.75 35 605 555 Mass Heat of Vap. [kJ/kg] 1309 1535 40 630 565 Surface Tension [dyne/cm] 15.47 14.81 45 635 593 Thermal Conductivity [W/m-K] 0.13 0.13 50 682 625 Viscosity [cP] 0.57 0.57 55 684 634


Volume fraction 70 743 697 C4-(<177°C) 0.000 0.000 75 750 741 LGO (177-343°C) 0.000 0.000 80 760 750 HGO (343-524°C) 0.089 0.193 85 807 761 524+(>524°C) 0.911 0.807 90 852 817

92.5 892 861 95 917 898

96.5 964 945 98 1031 993 99 1047 1021

100 1055 1048


Figure II-6 Effect of HGO Wash flow rate on Scrubber Bottom TBP curve

62


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> Components ' Boiling Temperatures 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>

Components ' Boiling Temperatures Range, °C

Figure II-8 Effect of HGO Wash flow rate on Scrubber Bottom composition

63


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


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:


- 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

480

460

- 440

=• 420

400

380

360

- Grid Top

- Grid Bottom

- Shed Top

- Shed Bottom

- Scrubber Pool

500

10 20 HGO Flow Rate, kbarrel/day (1 barrel=0.0049684m3/h)

Figure I I M Effect of HGO Underwash

flow rate on temperatures along the

Scrubber

480 j

360

- 0 kbarrel/day HGO

• 10 kbarrel/day HGO

- 20 kbarrel/day HGO

0 10 20 30 40 50

Position, feet from bottom of scr. pool

Figure III-2 Effect of HGO Underwash

flow rate on temperature profile along

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


-Density increases from 2.76 to 2.89 kg/m3 (Table III-l).

-Average molecular weight increases from 69 to 71 (Table III-2).

-Composition shows higher presence of 300-500°C fractions that originate from HGO

(Table III-2, Figures III-5 and III-7).


-Actual volume and mass flow rate increase by 30% and 26%, respectively (Table III-

1).

- Density drops from 828 to 803 kg/m3 (Table III-l).

-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 •

700 - Overhead

- Scrubber Bottom

600 1

500 • 10

HGO Flow Rate, kbarrel/day (1 barrel=0.0049684 m3/h)

20

.C

aT ra i o

900 850 800 750 -) 700 650 600 4 550 500 450 400 350 300 250 200 150 100

-Grid Liquid -Shed Vapor -Shed Liquid


(1 barrel=0.0049684 m3/h)

20


flow rate on mass flow rate of Scrubber

Overhead and Bottom


flow rate on mass flow rate of other

streams

67


Table III-l Effect of HGO Underwash flow rate on Scrubber parameters

HGO Uderw. Flow Rate HGO Underwash Flow Rate Position ft m3/h 50 0 50 99

from the kbarrel/day 10 0 10 20 pool bottom (Base Case) % % %

Koch Grid Top 43 Top Stage Temp Est (°C) 393 397 0.9 393 0.0 389 -1.2 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 Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 375 0.0

Flow Rates& Densities Scrubber Overhead Actual Volume Flow (m3/h) 274,787 274,718 0.0 274,787 0.0 273,909 -0.3

Mass Flow (kg/h) 778,651 757,412 -2.7 778,651 0.0 790,657 1.5 Mass Density (kg/m3) 2.83 2.76 -2.7 2.83 0.0 2.89 1.9

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 688 -12.9 789 0.0 921 16.8 Mass Flow (kg/h) 643,728 569,357 -11.6 643,728 0.0 739,959 14.9 Mass Density (kg/m3) 816 828 1.5 816 0.0 803 -1.6

Overhead / ATB Mass Flow Ratio 2.58 2.51 -2.7 2.58 0.0 2.62 1.5 Grid Liquid Actual Volume Flow (m3/h) 305 314 3.2 305 0.0 296 -2.7

Mass Flow (kg/h) 217,861 224,963 3.3 217,861 0.0 212,704 -2.4 Mass Density (kg/m3) 714.97 715.50 0.1 714.97 0.0 717.71 0.4

Shed Vapor Actual Volume Flow (m3/h) 277,329 277,339 0.0 277,329 0.0 276,366 -0.3 Mass Flow (kg/h) 870,450 856,313 -1.6 870,450 0.0 877,301 0.8 Mass Density (kg/m3) 3.14 3.09 -1.6 3.14 0.0 3.17 1.1

Shed Liquid Actual Volume Flow (m3/h) 497 440 -11.4 496.80 0.0 587 18.1 Mass Flow (kg/h) 361,323 316,491 -12.4 361,323 0.0 423,013 17.1 Mass Density (kg/m3) 724.76 718.94 -0.8 724.76 0.0 720.99 -0.5

To Coker Actual Volume Flow (m3/h) 330 287 -12.9 330 0.0 385 16.8 Mass Flow (kg/h) 269,078 237,991 -11.6 269,078 0.0 309,303 14.9 Mass Density (kg/m3) 815.85 828.00 1.5 815.85 0.0 803.04 -1.6

Additional information Vapour to Sheds Temperature (°C) 514 521 1.3 514 0.0 508 -1.2 Upgoing Stream Temperature (°C) 534 535 0.1 534 0.0 534 -0.1 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 SPL Coler Duty (MMBtu/h) 44.72 47.35 5.9 44.72 0.0 46.13 3.2

68


Table III-2 Effect of HGO Underwash flow rate on Scrubber Overhead properties

HGO Underw. Flow Rate (kbarrel/day) 0 10 20 Cut Point [%] 0 10 20

Temperature [UC] 397 393 389 Cut Point [%] TBP [°C] TBP [UC] TBP ["C] Pressure [psig] 16.99 16.99 16.99 0 -253 -253 -253 Molecular Weight 69.39 70.77 71.44 1 -238 -237 -237 Mass Density [kg/m3] 2.76 2.83 2.89 2 -208 -207 -206 Act. Volume Flow [m3/h] 274,718 274,787 281,956 3.5 -170 -167 -166 Mass Enthalpy [kJ/kg] -2961 -2924 -2913 5 -138 -136 -135 Mass Entropy [kJ/kg-C] 5.40 5.30 5.21 7.5 -103 -102 -101 Mass Heat Capacity [kJ/kg-C] 2.75 2.74 2.72 10 -88 -85 -82 Vapor Phase Fraction ( Mass Basis) 0.94 0.93 0.92 12.5 -55 -51 -50 Specific Heat [kJ/kgmole-C] 191.12 193.80 194.29 15 -43 -34 -26 Std. Gas Flow [STD_m3/h] 258,103 260,141 261,669 17.5 -4 -3 -1 Watson K 11.42 11.38 11.34 20 255 266 271 Liq. Mass Density (Std. Cond) [kg/m3] 922.13 930.01 936.54 25 309 310 310 Molar Volume [m3/kgmole] 25.17 24.98 24.75 30 336 336 336 Mass Heat of Vap. [kJ/kg] 2848 2825 2809 35 353 354 352

40 374 373 367 Fraction Distribution Data 45 392 391 385

Volume fraction 50 406 405 402 C4-(<177°C) 0.060 0.060 0.055 55 420 420 417 LGO (177-343°C) 0.260 0.259 0.266 60 440 439 423 HGO (343-524°C) 0.571 0.580 0.581 65 446 442 441 524+ (>524°C) 0.109 0.101 0.099 70 465 459 448

75 482 471 466 80 491 486 482 85 514 496 493 90 526 525 524

92.5 539 537 535 95 552 550 548

96.5 556 556 556 98 632 616 608 99 754 744 737

100 879 871 866

10 \

0 -I r , , , , . 1 1 1 1

0 100 200 300 400 500 600 700 800 900 1000 Temperature, °C

Figure III-5 Effect of HGO Underwash flow rate on Scrubber Overhead TBP curve

69


Table III-3 Effect of HGO Underwash flow rate on Scrubber Bottom properties

HGO Underw. Flow Rate (kbarrel/day) 0 10 20 Cut Point [%] 0 10 20

Temperature [°C] 375 375 375 TBP [°C] TBP [°C] TBP [°C] Pressure [psig] 17.00 17.00 17.00 0 427 412 402 Molecular Weight 663.75 637.04 609.86 1 464 441 435 Mass Density [kg/m3] 828.00 815.85 803.04 2 492 466 459 Act. Volume Flow [m3/h] 708 119,107 139,097 3.5 514 490 482 Mass Enthalpy [kJ/kg] -1333 -1329 -1324 5 520 504 486 Mass Entropy [kJ/kg-C] 3.39 3.39 3.39 7.5 527 514 493 Mass Heat Capacity [kJ/kg-C] 2.88 2.89 2.91 10 555 517 504 Vapor Phase Fraction ( Mass Basis) 0.00 0.00 0.00 12.5 556 520 513 Specific Heat [kJ/kgmole-C] 1911.25 1843.61 1774.01 15 560 525 515 Std. Gas Flow [STD_m3/h] 20,282 23,893 28,688 17.5 577 548 517 Watson K 11.39 11.41 11.43 20 592 556 520 Kinematic Viscosity [cSt] 0.69 0.69 0.70 25 599 564 526 Liq. Mass Density (Std. Cond) [kg/m3] 1048.30 1038.96 1028.93 30 627 592 556 Molar Volume [m3/kgmole] 0.80 0.78 0.76 35 633 605 575 Mass Heat of Vap. [kJ/kg] 1704 1309 1553 40 680 630 596 Surface Tension [dyne/cm] 15.92 15.47 15.02 45 683 635 625 Thermal Conductivity [W/m-K] 0.14 0.13 0.13 50 685 682 634 Viscosity [cP] 0.58 0.57 0.56 55 693 684 681


Volume fraction 70 748 743 708 C4-(<177°C) 0.000 0.000 0.000 75 755 750 747 LGO (177-343°C) 0.000 0.000 0.000 80 764 760 754 HGO (343-524°C) 0.045 0.089 0.153 85 815 807 772 524+ (>524°C) 0.955 0.911 0.847 90 884 852 820

92.5 898 892 884 95 949 917 903

96.5 965 964 957 98 1037 1031 1023 99 1049 1047 1037

100 1057 1055 1052

100


Figure III-6 Effect of HGO Underwash flow rate on Scrubber Bottom TBP curve

70


• 0 kbarrel/day HGO Underwash: 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 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%


Figure III-8 Effect of HGO Underwash flow rate on Scrubber Bottom composition

71


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


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:


- 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

temperature on temperatures along the

Scrubber


temperature on temperature profile along

the Scrubber

73



-By 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/m3 (Table IV-1).

-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).


-Actual volume and mass flow rate drop by 12% and 11% (Table IV-1).

- Density changes from 807 to 817 kg/m3 (Table IV-1).

-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


800

2 780

760

740

720

700

680

660

640 A

620

250 275 300 325

HGO Wash Temperature, °C

350


temperature on mass flow rate of Scrubber

Overhead and Bottom

1,000

950

900

850

800

750

700 •

650 -

600

550

500

450

400 ± -

350

300

250 0-200

- Grid Liquid

- Shed Vapor

- Shed Liquid

250 275 300 325

HGO Wash Temperature, °C

350


temperature on mass flow rate of other

streams

75


Table IV-1 Effect of HGO Wash temperature rate on Scrubber parameters

Position ft HGO Wash Temperaure from the Temperatures °C 250 300 325 350

pool bottom % (Base Case) % % Koch Grid Top 43 Top Stage Temp Est (°C) 386 391 1.3 393 1.9 396 2.6 Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 392 394 . 0.7 395 0.9 396 1.1 Sheds Top 34 Top Stage Temp Est (°C) 403 405 0.3 405 0.4 405 0.5 Sheds Bot. 22 Bottom Stage Temp Est (°C) 464 470 1.4 473 2.1 477 2.9 Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 375 0.0

Flow RatesS Densities

Scrubber Overhead Actual Volume Flow (m3/h) 269,827 273,227 1.3 274,787 1.8 276,497 2.5 Mass Flow (kg/h) 754,323 772,077 2.4 778,651 3.2 786,167 4.2 Mass Density (kg/m3) 2.80 2.83 1.1 2.83 1.4 2.84 1.7

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 870 812 -6.6 789 -9.3 767 -11.9 Mass Flow (kg/h) 701,660 659,624 -6.0 643,728 -8.3 626,246 -10.7 Mass Density (kg/m3) 807 812 0.7 816 1.2 817 1.3

Overhead / ATB Mass Flow Ratio 2.50 2.56 2.4 2.58 3.2 2.60 4.2 Grid Liquid Actual Volume Flow (m3/h) 360 322 -10.6 305 -15.4 289 -19.9

Mass Flow (kg/h) 255,809 229,561 -10.3 217,861 -14.8 206,915 -19.1 Mass Density (kg/m3) 710.21 712.98 0.4 714.97 0.7 717.02 1.0

Shed Vapor Actual Volume Flow (m3/h) 277,834 277,633 -0.1 277,329 -0.2 277,138 -0.3 Mass Flow (kg/h) 884,067 875,577 -1.0 870,450 -1.5 867,023 -1.9 Mass Density (kg/m3) 3.18 3.15 -0.9 3.14 -1.4 3.13 -1.7

Shed Liquid Actual Volume Flow (m3/h) 553 513.94 -7.1 497 -10.2 484 -12.5 Mass Flow (kg/h) 399,580 372,207 -6.9 361,323 -9.6 350,719 -12.2 Mass Density (kg/m3) 722.23 724.22 0.3 724.76 0.3 724.78 0.4

To Coker Actual Volume Flow (m3/h) 364 339 -6.6 381 4.8 320 -11.9 Mass Flow (kg/h) 293,294 275,723 -6.0 269,078 -8.3 261,771 -10.7 Mass Density (kg/m3) 806.53 812.15 0.7 815.85 1.2 816.90 1.3

Additional information

Vapour to Sheds Temperature (°C) 510 513 0.6 514 0.9 515 1.1 Upgoing Stream Temperature (°C) 534 534 0.1 534 0.1 534 0.1 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 SPL Coler Duty (MMBtu/h) 44.73 43.35 -3.1 44.72 0.0 42.60 -4.8

76


Table IV-2 Effect of HGO Wash temperature on Scrubber Overhead properties

HGO Wash Temperature (°C) 250 325 (Basic) 350 Cut Point [%] 250 325 350

Temperature [°C] 386 393 396 T B P f C ] T B P [°C] TBP [°C]

Pressure [psig] 16.99 16.99 16.99 0 -253 -253 -253 Molecular Weight 68.88 70.77 71.35 1 -239 -237 -237 Mass Density [kg/m3] 2.80 2.83 2.84 2 -209 -207 -206 Act. Volume Flow [m3/h] 269,827 282,860 276,497 3.5 -170 -167 -167

Mass Enthalpy [kJ/kg] -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 Specific Heat [kJ/kgmole-C] 187.49 193.80 195.76 15 -43 -34 -28 Std. Gas Flow [STD_m3/h] 258,922 260,141 260,513 17.5 -5 -3 -1 Watson K 11.39 11.38 11.37 20 252 266 270 Liq. Mass Density (Std. Cond) [kg/m3] 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

Fraction Distribution Data 45 381 391 391

Volume fraction 50 397 405 406

C4-(<177°C) 0.054 0.060 0.057 55 409 420 420

LGO (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

70

60

<jj 50

> 40

30

20 \

10

HGO T=250°C

HGO T=325°C

HGO T=350°C

0 100 200 300 400 500 600 700 800 900 1000 Temperature, °C

Figure IV-5 Effect of HGO Wash temperature on Scrubber Overhead TBP curve

77


Table IV-3 Effect of HGO Wash temperature on Scrubber Bottom properties

HGO Wash Temperature (°C) 250 325 (Basic) 350 Cut Point [%] 250 325 350

Temperature [°C] 375 375 375 T B P [°C] T B P ["C] T B P ["C]

Pressure [psig] 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 Act. Volume Flow [m3/h] 870 812 767 3.5 483 490 494 Mass Enthalpy [kJ/kg] -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 Phase Fraction ( Mass 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 Watson K 11.43 11.41 11.40 20 522 556 556 Kinematic Viscosity [cSt] 0.70 0.69 0.68 25 555 564 590 Liq. Mass Density (Std. Cond) [kg/m3] 1031.82 1038.96 1041.23 30 561 592 597 Molar Volume [m3/kgmole] 0.77 0.78 0.79 35 592 605 625 Mass 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 Fraction Distribution Data 65 697 706 737

Volume fraction Volume fraction 70 739 743 745

C4-(<177°C) 0.000 0.000 0.000 75 749 750 751

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


0.50

0.40

(A c 0.30 o

o

s 4 -a> o 0.20 S

0.10

0.00 11

• HGO T=250°C: Light Ends: 20%; Water: 65% ; 100> fraction: 15%

• HGO T=325°C: Light Ends: 20%; Water: 65% 100> fraction: 15%

• HGOT=350°C: 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> C o m p o n e n t s ' Bo i l ing T e m p e r a t u r e s Range , °C

Figure IV-7 Effect of H G O Wash temperature on Scrubber Overhead composition

0.50

0.40

c 0.30 o

u 2

<+-ffi o 0.20 S

0.10

0.00 _ tri

• HGO T= 250°C: Light Ends: 0% ; Water: 1%; 100> fraction: 99%

• HGO T=325°C: Light Ends: 0% ; Water: 1 %; 100> fraction: 99%

• HGO T=350°C: 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 ' Bo i l ing T e m p e r a t u r e s Range , °C

Figure IV-8 Effect of H G O Wash temperature on Scrubber Bottom composition

79


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 of Scrubber Overhead and Scrubber Bottom (Scrubber Pool Liquid). Four options

have been investigated:

1. H G O Underwash is in service - flow rate of H G O Underwash is 10 kbarrel/day;

2. H G O Underwash is out of service and Overhead temperature is not controlled.

3. H G O Underwash is out of service and Overhead temperature is controlled by H G O Wash

flow rate - in 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 of service and Overhead temperature is controlled by A T B feed

flow rate - increase o f A T B actual volume flow by 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 by 4-5°C, compared to Base Case;

- H G O Underwash off controlled by A T B : This case has almost the same effect as

when H G O Underwash is in service. There is only a small change in all observed

temperatures (Table V - l , Figures V - l and V-2) .

81


Figure V - l Effect of HGO Underwash

service on temperatures along the Scrubber

Note: Lines that connect data points do not

present trend lines. They are shown to help

comparison between different cases.

Figure V -2 Effect of HGO Underwash

service on temperature profile along the

Scrubber


• 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 HGO Underwash is in service (Table V - l ,

Figure V-3).

-Density changes from 2.83 to 2.76 kg/m3 (Table V - l ) .

-Average molecular weight drops from 71 to 69 (Table V-2).

-Composition shows lower presence of 300-500°C fractions that originate from HGO

(Table V-2, Figures V-5 and V-7).

82


• HGO Underwash off, controlled by HGO Wash:

-There is no significant change in any of Scrubber Overhead properties, compared to Base

Case.

• HGO Underwash off, controlled by ATB:

-Actual volume and mass flow rate drops by 0.7 and 3.9%, respectively.

-Density changes from 2.83 to 2.74 kg/m3.

-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).



-Actual volume and mass flow rate drop by 13% and 12% (Table V-l) .

- Density increases from 816 to 828 kg/m3 (Table V-l) .

-Average molecular weight changes from 637 to 663 (Table V-3).

-Composition shows lower presence of 400-600°C fraction that originates from HGO

(Table V-3, Figures V-6 and V-8).


-Again, for all Scrubber Bottom properties there are just minor changes, comparing to the

case when HGO Underwash is in service.


-Actual volume and mass flow rate increase by 7.5% and 8% (Table V-l) .

- Density increases from 816 to 821 kg/m3 (Table V-l) .

-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 600°C + fractions (Table V-3, Figures V-6 and V-8).

83


• Other.


-Grid Liquid volume and mass flow slightly increase; Shed Vapour mass flow decreases a

little, and Shed Liquid volume and mass flow significantly decrease.


-Grid Liquid volume and mass flow rate radically increase, Shed Vapour does not change

much, and Shed Liquid flow rate decreases a little.


-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 HGO Underwash

service on mass flow rate of Scrubber

Overhead and Bottom

Figure V-4 Effect of HGO Underwash

service on mass flow rate of other

streams

Note: Lines that connect data points do not present trend lines. They are shown to help


84


Table V-l Effect of HGO Underwash service on Scrubber parameters

HGO Underwash HGO Underwash Position ft ON OFF OFF OFF

from the (Base Case) Uncontrolled Contr. by HGO Contr. by ATB pool bottom 1 2 % 3 % 4 %

Koch Grid Top 43 Top Stage Temp Est (°C) 393 397 0.9 393 0.0 393 0.0

Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 395 400 1.3 399 0.9 396 0.3

Sheds Top 34 Top Stage Temp Est (°C) 405 410 1.4 410 1.2 406 0.3

Sheds Bot. 22 Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)

473 495 4.6 477 0.8 476 0.6

Scrubber Pool 0

Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 372 -0.7

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 ATB from 55 to 63 kbpd. Flow RatesS Densities Scrubber Overhead Actual Volume Flow (m3/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

Mass Density (kg/m3) 2.83 2.76 -2.7 2.83 0.0 2.74 -3.2

Scrubb.Pool Liquid Actual Volume Flow (m /h) 789 688 -12.9 771 -2.3 848 7.5 Mass Flow (kg/h) 643,728 569,357 -11.6 631,095 -2.0 696,159 8.1

Mass Density (kg/m3) 816 828 1.5 818 0.3 821 0.6

Overhead / ATB Mass Flow Ratio 2.58 2.51 -2.7 2.58 0.0 2.16 -16.2

Grid Liquid Actual Volume Flow (m Ih) 305 314 3.2 378 24.1 297 -2.4 Mass Flow (kg/h) 217,861 224,963 3.3 268,867 23.4 212,905 -2.3

Mass Density (kg/m 3) 714.97 715.50 0.1 710.86 -0.6 715.94 0.1

Shed Vapor Actual Volume Flow (m Ih) 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 -4.1

Mass Density (kg/m 3) 3.14 3.09 -1.6 3.14 -0.1 3.03 -3.4

Shed Liquid Actual Volume Flow (m Ih) 497 440 -11.4 488.15 -1.7 520 4.7 Mass Flow (kg/h) 361,323 316,491 -12.4 353,507 -2.2 378,953 4.9

Mass Density (kg/m3) 724.76 718.94 -0.8 724.18 -0.1 728.33 0.5

To Coker Actual Volume Flow (m3/h) 330 287 -12.9 322 -2.3 354 7.5 Mass Flow (kg/h) 269,078 237,991 -11.6 263,798 -2.0 290,994 8.1 Mass Density (kg/m J) 815.85 828.00 1.5 818.40 0.3 821.10 0.6

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 SPL Coler Duty (MMBtu/h) 44.72 47.35 5.9 45.64 2.0 56.04 25.3

85


Table V-2 Effect of HGO Underwash service on Scrubber Overhead properties

HGO ON HGO OFF HGO OFF HGO OFF

Cut Point [%] HGO ON HGO OFF HGO OFF HGO OFF

HGO ON (uncontr.) (contr. by HGO) (contr. by ATB)

Cut Point [%] HGO ON (uncontr.) (contr. by HGO) (contr. by ATB)

Temperature f t ] 393 397 393 393 TBP f C] TBP fC] TBP [°C] TBP ft ]

Pressure [psig] 16.99 16.99 16.99 16.99 0 -253 -253 -253 -253

Molecular Weight 70.77 69.39 70.78 68.61 1 -237 -238 -237 -239

Mass Density [kg/m3] 2.83 2.76 2.83 2.74 2 -207 -208 -207 -209

Act. Volume Flow [m3/h] 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 [STD jn3/h] 260,141 258,103 260,017 257,826 17.5 -3 -4 -3 -5

Walson K 11.38 11.42 11.38 11.42 20 266 255 266 248

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

Fraction Distribution Data 45 391 392 391 389

Volume fraction 50 405 406 406 405

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 , , , , , . 1 1 1

O 1 0 0 2 0 0 3 0 0 4 0 0 5 0 0 6 0 0 7 0 0 8 0 0 9 0 0 1 0 0 0

Temperature, °C

Figure V-5 Effect of HGO Underwash service on Scrubber Overhead TBP curve

86


Table V-3 Effect of HGO Underwash service on Scrubber Bottom properties

HGO ON HGO OFF (uncontr.)

HGO OFF (contr. by HGO)

HGO OFF (contr. by ATB)

Cut Point [%] HGO ON HGO OFF (uncontr.)

HGO OFF (contr. by HGO)

HGO OFF (contr. by ATB)

Temperature fC] 375 375 375 375 TBP [°C] TBP ["C] TBP ["C] TBP f C]

Pressure [psig] 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

Act. Volume Flow [m3/h] 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

Watson K 11.41 11.39 11.40 11.41 20 556 592 556 558

Kinematic Viscosity [cSt] 0.69 0.69 0.69 0.71 25 564 599 588 591

Liq. Mass Density (Std. Cond) [kg/m3] 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] 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

60

684

693

693

704

685

695

687

696

Fraction Distribution Data 65 706 742 708 738

Volume fraction 70 743 748 744 745

C4-(<177°C) 0.000 0.000 0.000 0.000 75 750 755 751 752

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 90 852 884 854 853

92.5 892 898 893 893

95 917 949 932 926

96.5

98

964

1031

965

1037

964

1032

964

1031

99

100

1047

1055

1049

1057

1047

1055

1046

1054

400 500 600 700 800 900 1000 1100

Temperature, °C

Figure V-6 Effect of HGO Underwash service on Scrubber Bottom TBP curve

87


1 • • HGO ON: Light Ends: 20%; Water: 65% ; 100> fraction: 15%

• 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%

1 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>

Components ' Boiling Temperatures 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%


Figure V-8 Effect o f H G O Underwash service on Scrubber Bottom composition

8S


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


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:


-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


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


-As 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/m3 (Table VI-1).

-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).


-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/m3 (Table VI-1).

-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 •

• Overhead

- Scrubber Bottom

4 6 l

Number of Sheds' trays

10

750

700

650

600

550

500

450

400

350

300

250

200

150

100

- Grid Liquid

- Shed Vapor

- Shed Liquid

4 6 8

Number of Sheds' trays

Figure VI-3 Effect of number of

Sheds trays on mass flow rate of

Scrubber Overhead and Bottom

Figure VI-4 Effect of number of Sheds

trays on mass flow rate of other streams

92


Table VI-1 Effect of number of Sheds trays on Scrubber parameters

Position ft Number of Sheds' trays

from the 6 2 4 6 8 10 pool bottom Base Case % % % % %

Koch Grid Top 43 Top Stage Temp Est 393 400 1.8 394 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 Scrubber Pool 0 Bulk Liquid Temperature 375 375 0.0 375 0.0 375 0.0 375 0.0 375 0.0


Scrubber Overhead Actual Volume Flow (m3/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

Mass Density (kg/m3) 2.83 2.84 0.3 2.84 0.2 2.83 0.0 2.71 -4.3 2.70 -4.9

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 738 -6.5 778 -1.3 789 0.0 895 13.5 912 15.6

Mass Flow (kg/h) 643,728 606,279 -5.8 635,952 -1.2 643,728 0.0 720,241 11.9 731,830 13.7

Mass Density (kg/m3) 816 822 0.7 817 0.1 816 0.0 804 -1.4 802 -1.6

Overhead/ATB Mass Flow Ratio 2.58 2.63 2.0 2.59 0.5 2.58 0.0 2.47 -4.1 2.46 -4.7

Grid Liquid Actual Volume Flow(m3/h) 305 193 -36.7 293 -3.8 305 0.0 217 -28.7 211 -30.7

Mass Flow (kg/h) 217,861 139,382 -36.0 209,570 -3.8 217,861 0.0 153,354 -29.6 149,067 -31.6

Mass Density (kg/m3) 714.97 722.93 1.1 714.94 0.0 714.97 0.0 705.88 -1.3 705.56 -1.3

Shed Vapor Actual Volume Flow (m3/h) 277,329 282,557 1.9 278,195 0.3 277,329 0.0 279,033 0.6 278,954 0.6

Mass Flow (kg/h) 870,450 807,726 -7.2 865,746 -0.5 870,450 0.0 773,865 -11.1 764,843 -12.1

Mass Density (kg/m3) 3.14 2.86 -8.9 3.11 -0.8 3.14 0.0 2.77 •11.6 2.74 -12.6

Shed Liquid Actual Volume Flow (m3/h) 497 445 -10.4 487.08 -2.0 497 0.0 601 21.0 612 23.2

Mass Flow (kg/h) 361,323 329,675 -8.8 354,643 -1.8 361,323 0.0 411,200 13.8 420,275 16.3

Mass Density (kg/m3) 724.76 740.33 2.1 728.11 0.5 724.76 0.0 683.68 -5.7 677.97 -6.5

To Coker Actual Volume Flow(m3/h) 330 308 -6.5 325 -1.3 381 15.6 374 13.5 378 14.6

Mass Flow (kg/h) 269,078 253,425 -5.8 265,828 -1.2 269,078 0.0 301,061 11.9 305,905 13.7

Mass Density (kg/m3) 815.85 821.88 0.7 817.03 0.1 815.85 0.0 804.45 -1.4 802.40 -1.6


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


Table VI-2 Effect of number of Sheds trays on Scrubber Overhead properties

Number of Sheds' trays 2 6 10 Cut Point [%] 2 6 10

Temperature fC] 400 393 392 TBP fC] TBP f C] TBP [°C] Pressure [psig] 16.99 16.99 16.99 0 -253 -253 -253 Molecular Weight 72.00 70.77 67.71 1 -237 -237 -238 Mass Density [kg/m3] 2.84 2.83 2.70 2 -206 -207 -208 Act. Volume Flow [m3/h] 279,507 274,787 275,149 3.5 -166 -167 -169 Mass Enthalpy [kJ/kg] -2863 -2924 -3012 5 -134 -136 -137 Mass Entropy [kJ/kg-C] 5.30 5.30 5.35 7.5 -101 -102 -103 Mass Heat Capacity [kJ/kg-C] 2.75 2.74 2.71 10 -82 -85 -87 Phase Fraction ( Mass Basis) 0.95 0.93 0.99 12.5 -49 -51 -53 Specific Heat [kJ/kgmole-C] 198.04 193.80 183.21 15 -23 -34 -42 Std. Gas Flow [STD_m3/h] 260,892 260,141 259,068 17.5 2 -3 -3 Watson K 11.37 11.38 11.38 20 274 266 260 Liq. Mass Density (Std. Cond) [kg/m3] 931.79 930.01 930.82 25 314 310 309 Molar Volume [m3/kgmole] 25.33 24.98 25.11 30 337 336 335 Mass Heat of Vap. [kJ/kg] 2814 2825 1751 35 356 354 352


Volume fraction 50 406 405 404 C4-(<177°C) 0.052 0.060 0.058 55 420 420 419 LGO(177-343°C) 0.264 0.259 0.264 60 440 439 437 HGO (343-524°C) 0.576 0.580 0.591 65 445 442 441 524+(>524°C) 0.108 0.101 0.087 70 465 459 453

75 482 471 467 80 490 486 484 85 513 496 494 90 526 525 520

92.5 539 537 527 95 552 550 548

96.5 556 556 556 98 615 616 599 99 741 744 735

100 866 871 872

10 \

Ot , , , , , , r , , 1

0 100 200 300 400 500 600 700 800 900 1000 Temperature, °C

Figure VI-5 Effect of number of Sheds trays on Scrubber Overhead TBP curve

9 4


Table VI-3 Effect of number of Sheds trays on Scrubber Bottom properties


Temperature fC] 375 375 375 TBP [°C] TBP [°C] TBP [°C] Pressure [psig] 17 17 17 0 411 412 412 Molecular Weight 649 637 634 1 440 441 441 Mass Density [kg/m3] 822 816 811 2 466 466 480 Act. Volume Flow [m3/h] 738 789 827 3.5 489 490 495 Mass Enthalpy [kj/kg] -1331 -1329 -1326 5 512 504 512 Mass Entropy [kJ/kg-C] 3.39 3.39 3.41 7.5 517 514 515 Mass Heat Capacity [kJ/kg-C] 2.89 2.89 2.90 10 523 517 517 Phase Fraction ( Mass Basis) 0.00 0.00 0.00 12.5 542 520 522 Specific Heat [kJ/kgmole-C] 1872.29 1843.61 1842.15 15 555 525 526 Std. Gas Flow [STD_m3/h] 22,098 23,893 24,988 17.5 556 548 541 Watson K 11.39 11.41 11.43 20 563 556 547 Kinematic Viscosity [cSt] 0.69 0.69 0.66 25 592 564 558 Liq. Mass Density (Std. Cond) [kg/m3] 1043.55 1038.96 1035.18 30 602 592 590 Molar Volume [m3/kgmole] 0.79 0.78 0.78 35 629 605 597 Mass Heat of Vap. [kj/kg] 1656 1309 1654 40 634 630 626 Surface Tension [dyne/cm] 15.68 15.47 15.34 45 681 635 633 Thermal Conductivity [W/m-K] 0.14 0.13 0.13 50 684 682 681 Viscosity [cP] 0.57 0.57 0.53 55 690 684 684


Volume fraction 70 749 743 742 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) 0.064 0.089 0.091 85 812 807 804 524+(>524°C) 0.936 0.911 0.909 90 880 852 822

92.5 895 892 889 95 941 917 907

96.5 964 964 963 98 1034 1031 1028 99 1048 1047 1046

100 1056 1055 1055

100

400 500 600 700 800 900 1000 1100 Temperature, °C

Figure VI-6 Effect of number of Sheds trays on Scrubber Bottom TBP curve

95


• 2 Sheds' trays: Light Ends: 20%; Water: 64% ; 100> fraction: 16%




Figure VI-7 Effect of number of Sheds trays on Scrubber Overhead composition

0.00 -i

• 2 Sheds' trays: Light Ends: 0% ; Water: 1%; 100> fraction: 99%




Figure VI-8 Effect of number of Sheds trays on Scrubber Bottom composition

96


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

VII. Number of Grid Sections Chapter 6 - Case Studies: Results and Discussion

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:


- 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

480

460

440

420 A

400

380

360

- Grid Top

- Grid Bottom

- Shed Top

- Shed Bottom

- Scrubber Pool

500

4 6 8

Number of Grid sections

10

- 2 Grid sections

-o - - 5 Grid sections

-10 Grid sections

360 10 20 30 40 50


Figure VII-1 Effect of number of Grid

sections on temperatures along the

Scrubber


sections on temperature profile along the

Scrubber

98



- 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/m3 (Table VII-1).

-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.


- 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/m3 (Table VII-1).

-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



sections on mass flow rate of Scrubber

Overhead and Bottom


sections on mass flow rate of other

streams

100


Table VII-1 Effect of number of Grid sections on Scrubber parameters

Position ft Base Case Number of Grid sections from the 2 5 10

pool bottom % % Koch Grid Top 43 Top Stage Temp Est 393 393 -0.1 393 -0.1 Koch Grid Bot. 38 Bottom Stage Temp Est 395 397 0.6 397 0.6 Sheds Top 34 Top Stage Temp Est 405 406 0.2 406 0.2 Sheds Bot. 22 Bottom Stage Temp Est 473 486 2.7 486 2.6 Scrubber Pool 0 Bulk Liquid Temperature 375 375 0.0 375 0.0 Note: Overall Grid section efficiency is 0.75

Flow Rates& Densities Scrubber Overhead Actual Volume Flow (m3/h) 274,787 274,318 -0.2 274,318 -0.2

Mass Flow (kg/h) 778,651 793,996 2.0 793,567 1.9 Mass Density (kg/m3) 2.83 2.89 2.1 2.89 2.1

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 725 -8.2 726 -8.0 Mass Flow (kg/h) 643,728 595,823 -7.4 596,652 -7.3 Mass Density (kg/m3) 816 822 0.8 822 0.8

Overhead / ATB Mass Flow Ratio 2.58 2.63 2.0 2.63 1.9 Grid Liquid Actual Volume Flow (m3/h) 305 259 -15.0 259 -14.8

Mass Flow (kg/h) 217,861 185,180 -15.0 185,481 -14.9 Mass Density (kg/m3) 714.97 714.77 0.0 714.82 0.0

Shed Vapor Actual Volume Flow (m3/h) 277,329 276,845 -0.2 276,852 -0.2 Mass Flow (kg/h) 870,450 860,927 -1.1 860,709 -1.1 Mass Density (kg/m3) 3.14 3.11 -0.9 3.11 -0.9

Shed Liquid Actual Volume Flow (m3/h) 497 461 -7.1 461.72 -7.1 Mass Flow (kg/h) 361,323 332,768 -7.9 333,193 -7.8 Mass Density (kg/m3) 724.76 718.81 -0.8 719.11 -0.8

To Coker Actual Volume Flow (m3/h) 330 303 -8.2 303 -8.0 Mass Flow (kg/h) 269,078 249,054 -7.4 249,400 -7.3 Mass Density (kg/mJ) 815.85 822.17 0.8 822.05 0.8

Additional information Vapour to Sheds Temperature (°C) 514 517 0.7 517 0.6 Upgoing Stream Temperature (°C) 534 534 0.0 534 0.0 Sheds Stage Efficiency 0.53 0.53 0.0 0.53 0.0 Koch Grid Stage Efficiency 0.75 0.75 0.0 0.75 0.0 SPL Coler Duty (MMBtu/h) 44.72 45.46 1.7 45.38 1.5

101


Table VII-2 Effect of number of Grid sections on Scrubber Overhead properties

Number of Grid sections 2 5 10 Cut Point [%] 2 5 10

Temperature [°C] 393 393 393 T B P [°C] T B P [°C] TBP fC ] Pressure [psig] 16.99 16.99 16.99 0 -253 -252 -252 Molecular Weight 70.77 71.96 71.93 1 -237 -226 -226 Mass Density [kg/m3] 2.83 2.89 2.89 2 -206 -200 -200 Act. Volume Flow [m3/h] 274,787 274,318 274,310 3.5 -166 -164 -163 Mass Enthalpy [kj/kg] -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 Specific Heat [kJ/kgmole-C] 193.80 197.47 197.37 15 -23 -16 -15 Std. Gas Flow [STD_m3/h] 260,141 260,885 260,861 17.5 2 2 2 Watson K 11.38 11.37 11.37 20 274 279 277 Liq. Mass Density (Std. Cond) [kg/m3] 930.01 932.00 931.96 25 314 315 315 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


Volume fraction 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 HGO (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 95 552 558 558

96.5 556 571 571 98 615 623 623 99 741 746 746

100 866 866 867

10 1

0 1 , , , , , , , ^ _ , 1 0 100 200 300 400 500 600 700 800 900 1000

Temperature, °C

Figure VII-5 Effect of number of Grid sections on Scrubber Overhead TBP curve

102


Table VII-3 Effect of number of Grid sections on Scrubber Bottom properties


Temperature ["C] 375 375 375 TBP [°C] TBP [°C] TBP pC] Pressure [psig] 17.00 17.00 17.00 0 412 419 419 Molecular Weight 637.04 650.69 650.44 1 441 446 446 Mass Density [kg/m3] 815.85 822.17 822.05 2 466 474 474 Act. Volume Flow [m3/h] 812 746 747 3.5 490 494 493 Mass Enthalpy [kJ/kg] -1329 -1331 -1331 5 504 511 510 Mass Entropy [kJ/kg-C] 3.39 3.39 3.39 7.5 514 520 520 Mass Heat Capacity [kJ/kg-C] 2.89 2.89 2.89 10 517 531 530 Phase Fraction ( Mass Basis) 0.00 0.00 0.00 12.5 520 532 532 Specific Heat [kJ/kgmole-C] 1843.61 1877.95 1877.33 15 525 543 542 Std. Gas Flow [STD_m3/h] 23,893 21,651 21,689 17.5 548 564 562 Watson K 11.41 11.40 11.40 20 556 583 581 Kinematic Viscosity [cSt] 0.69 0.69 0.69 25 564 592 590 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1043.70 1043.61 30 592 598 598 Molar Volume [m3/kgmole] 0.78 0.79 0.79 35 605 632 631 Mass Heat of Vap. [kJ/kg] 1309 1332 1331 40 630 638 636 Surface Tension [dyne/cm] 15.47 15.70 15.69 45 635 642 642 Thermal Conductivity [W/m-K] 0.13 0.14 0.14 50 682 687 686 Viscosity [cP] 0.57 0.57 0.57 55 684 692 690

0.78 0.79 0.79 60 693 696 696 Fraction Distribution Data 65 706 707 705

70 743 747 746 C4-(<177°C) 0.000 0.000 0.000 75 750 754 753 LGO(177-343°C) 0.000 0.000 0.000 80 760 765 765 HGO (343-524°C) 0.089 0.056 0.056 85 807 809 809 524+(>524°C) 0.911 0.944 0.944 90 852 855 854

92.5 892 895 893 95 917 919 917

96.5 964 989 989 98 1031 1032 1032 99 1047 1048 1048

100 1055 1056 1056

100


Figure VII-6 Effect of number of Grid sections on Scrubber Bottom TBP curve

103


0.50

0.40

V) c 0.30 o o 2

•50.20 2

0.10

0.00 TO

• 2 Grid Sections: Light Ends: 20%; Water: 65% ; 100> fraction: 15%

• 5 Grid Sections: Light Ends:20%; Water: 65% ; 100> fraction: 15%

• 10 Grid Sections: Light Ends:20%; Water: 65% ; 100> fraction: 15%

ID 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000=


Figure VII-7 Effect of number of Grid sections on Scrubber Overhead composition

In

• 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%


Figure VII-8 Effect of number of Grid sections on Scrubber Bottom composition

104


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)a. Note that Overhead production rate is not constant.

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:

-As 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


480

460

440 e 3

1 8. £ 420

380

- Grid Top

- Grid Bottom

- Shed Top

- Shed Bottom

- Scrubber Pool

0.00 0.50 1.00 17 19 21 23

2.00 25

2.50 27

Top scale: Pressure drop in Grid, in. of water Bottom scale: Absolute pressure in the

Scrubber, psig

Figure VIII-1 Effect of pressure drop in

the Grid and absolute pressure in the

Scrubber on temperatures along the

Scrubber

-•- -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

10 20 30 40 50 Position, feet from bottom of scr. pool

Figure VIII-2 Effect of pressure drop in

Grid and absolute pressure in the

Scrubber on temperature profde along

the Scrubber


-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/m3 (Table VIII-1).

-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


- Density drops from 816 to 802 kg/m3 (Table VIII-1).

-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

in the Grid and absolute pressure in the

Scrubber on mass flow rate of the


Figure V I I I - 4 Effect of pressure drop

in Grid and absolute pressure in the

Scrubber on mass flow rate of other

streams

108

v Chapter 6 - Case Studies: Results and Discussion

Table VIII-1 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber parameters

KxhGid BED Case FtesureorcpintreKxhGrid Top Fressuie (psig) 1699 1857 2096 2294 2493 26.91

ftfiftkn ft BaoTiFtBssue (psig) 17 19 21 23 25 27 FtessureDop( in of water) 040 081 1.23 1.65 207 250

pod uottcrri FreasureDcpfpa) 0.014 0029 % 0.044 % Q06 % Q075 % Q09 % KxtiGidTcp 43 Tqp Sags Tsip Est (t) 393 394 02 394 02 395 0.5 395 Q5 395 Q5 Hb+iQidEtt. 38 BrJt0Ti9ar^Termaft) 395 395 Q1 396 Q2 396 Q3 397 Q4 397 04 STedsTqp 34 Tqp Stags Torp Est CQ 405 405 QO 403 02 406 Q3 406 Q3 407 Q4 STSOSBI 22 EdtDrnSageTerrp Est CQ 473 471 •0.4 470 -0.7 459 -1.0 468 -12 467 -1.4 Soxtba-Pod 0 BJk Liqud Terperatije ft) 375 375 QO 375 QO 375 0.0 375 Q0 375 QO

HojvFaes& Densities SoUtja-aetead A i d VflirreRa/vtrrfTh) 274,787 255259 -60 213,570 -11.4 23Q4C -161 218600 -2Q4 207,931 -24.3

M3ssRow(kgh) 778,651 771,718 -0.9 764,769 -1.8 757,6« -27 750,931 -36 744,287 AA ^te^tyfkc/rr3) 28! 293 55 314 10.8 3Z 160 344 212 353 263

SQttbPod Uqj'd Ad^VfluTCRcwfrrf/h) 78E 813 30 836 59 85E 89 882 11.8 906 14.6 IVassHcw(kgfri) 64372c 66Q619 26 676,903 52 698,319 7.7 703,53: 102 725,577 127 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) 3D: 303 -04 302 -0.9 301 -1.1 300 -1.4 300 -1.7

lvB5sRcw(kgh) 217,831 216233 -Q7 214,747 -1.4 213606 -20 212,529 -24 211,580, -29 NterirBity(kg'rr3) 714.97 71282 -Q3 710.83 -Q6 70915 -08 707.54 -1.0 7031C -1.2

SnedVfepcr AJi \<tiLneRcw(rr3/h) 277,32 26Q753 -60 243,023 -11.3 232,791 -161 22Q921 21Q16E -242 M3ssRcw(kgti) 87Q45C 851,£S -1.0 863,457 -20 84519C -29 837,399 -38 829,807 A.7

314 331 53 347 10.5 363 157 379 20.8 39E 258 3-BrJUQjd Adu=l\fluTBRo/v(rr?/h) 497 51C 27 521.83 50 533 7.4 54E 97 567 120

M3ssRow(kg4i) 361,323 33904 22 376,88: 4.3 384,44E 64 391,949 85 399316 1Q5 KteD&^flg'rr?) 724.7E 72363 -02 72224 -Q3 72Q6E -06 71907 -08 717.46 -1.0

ToOter A+u \ LrreRcw(rr37h) 33C 34C 30 3* 59 359 89 369 11.8 37E 14.6 fVB3sRcw(kgh) 26907t 276,139 26 282,946 52 28Q820 7.7 293,607 1Q2 308291 127

8158E 81279 -Q4 80998 -Q7 807.2c -1.1 804.63 -1.4 80211 -1.7 AcUtiorol irfarnation \*por to Shads Ta-rp=ratLre(C) 514 513 -02 512 -Q3 512 -05 511 -Q6 510 -Q7 IpgaraSjBam TerparalrefC) 534 534 -Q1 533 •02 533 -02 533 -03 532 -Q3 Sheds SageEffidary Q53 Q53 QO Q53 QO Q53 0.0 Q53 QO Q53 QO KrfiGid Stage Efiidaxy Q75 Q75 QO Q75 QO Q75 QO Q75 QO Q75 QO SFLOda- rOfy(MvButi) 44.72 4565 21 4681 4.7 4812 7.6 4941 1Q5 5072 134

109


Table VIII-2 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Overhead properties

Gut Point [%] 0.4 2.5 17 27

TBP["C] TBP["C] 0 -253 -253 1 -237 -239 2 -207 -209

3.5 -167 -171 5 -136 -140

7.5 -102 -104 10 -85 -90

12.5 -51 -57 15 -34 -42

17.5 -3 -6 20 266 244 25 310 305 30 336 325 35 354 350 40 373 365 45 391 379 50 405 395 55 420 407 60 439 421 65 442 440 70 459 447 75 471 466 80 486 482 85 496 493 90 525 525

92.5 537 536 95 550 549

96.5 556 555 98 616 615 99 744 743

100 871 869

Pressure drop in the Grid, in. water Pressure in the Scrubber, psig

Temperature [ UC] 393 395 Pressure [psig] 16.99 26.91 Molecular Weight 70.77 62.93 Mass Density [kg/m 3] 2.83 3.27 Act. Volume Flow [m3/h] 274,787 207,839 Mass Enthalpy [kj/kg] -2924 -3179 Mass Entropy [kJ/kg-C] 5.30 5.49 Mass Heat Capacity [kJ/kg-C] 2.74 2.71 Vapour Phase Fraction ( Mass Basis) 0.93 0.91 Specific Heat [kJ/kgmole-C] 193.80 170.42 Std. Gas Flow [STD_m3/h] 260,141 255,492 Watson K 11.38 11.43 Liq. Mass Density (Std. Cond) [kg/m3] 930.01 921.04 Molar Volume [m3/kgmole] 24.98 19.23 Mass Heat of Vap. [kj/kg] 2825 2853

Fraction Distribution Data Volume fraction

C4-(<177 °C) 0.060 0.067 LGO (173-743 °C) 0.259 0.268 HGO (343-524 °C) 0.580 0.564 524+ (>524 °C) 0.101 0.100

0.4 17

2.5 27

100

90

80

70

60

» 50

5 40

30

20

10

0

•Pdrop 0.4 in; Abs. P=17 psig

Pdrop 2.5 in; Abs. P=27 psig

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

1 1 0


Table VIII-3 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom properties

Pre s su re d rop in the G r i d , in. water 0.4 2.5 Cut Point [%] 0.4 2.5 P r e s s u r e in the S c r u b b e r , p s i g 17 27

Cut Point [%] 17 27

Temperature [UC] 375 375 TBP [°C] TBP [UC] Pressure [psig] 17.00 27.00 0 412 389 Molecular Weight 637.04 610.37 1 441 423 Mass Density [kg/m3] 815.85 802.11 2 466 452 Act. Volume Flow [m3/h] 789 905 3.5 490 466 Mass Enthalpy [kj/kg] -1329 -1324 5 504 485 Mass Entropy [kJ/kg-C] 3.39 3.40 7.5 514 492 Mass Heat Capacity [kJ/kg-C] 2.89 2.91 10 517 505 Phase Fraction ( Mass Basis) 0.00 0.00 12.5 520 513 Specific Heat [kJ/kgmole-C] 1843.61 1777.47 15 525 515 Std. Gas Flow [STD_m3/h] 23,893 28,107 17.5 548 517 Watson K 11.41 11.44 20 556 520 Kinematic Viscosity [cSt] 0.69 0.69 25 564 527 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1028.41 30 592 556 Molar Volume [m3/kgmole] 0.78 0.76 35 605 584 Mass Heat of Vap. [kj/kg] 1309 1654 40 630 597 Surface Tension [dyne/cm] 15.47 14.91 45 635 626 Thermal Conductivity [W/m-K] 0.13 0.13 50 682 634 Viscosity [cP] 0.57 0.55 55 684 681

0.78 0.76 60 693 685 Fract ion D i s t r ibut ion Data 65 706 693


92.5 892 885 95 917 904

96.5 964 959 98 1031 1023 99 1047 1038

100 1055 1053 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


• 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>



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>


Figure VIII-8 Effect of pressure drop in the Grid and absolute pressure in the Scrubber on Scrubber Bottom composition

112


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


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:


-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


500 •

480

460

440 A

420 •

400

380

360

- Grid Top

- Grid Bottom

- Shed Top

- Shed Bottom

- Scrubber Pool

500

480

460

440

420 4

400

380

360

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

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


Figure IX-1 Effect of water instead of HGO

Underwash on temperatures along the

Scrubber





Underwash on temperature profile along the

Scrubber


-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/m3 (Table IX-1).

115


-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).


-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/m3 (Table IX-1).

-Average molecular weight changes from 637 to 654, 655, 664 (Table LX-3).


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


780

760

740

720 \

700

680 -I

660 -I 640

620

600 -I 580

560

540

520

500

480

460

440

420

- Overhead

- Scrubber Bottom

1

C a s e s : 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

900

850

800

750

700 -I 650

600 -

550 -

500

450

400

- Grid Liquid

- Shed Vapor

- Shed Liquid

C a s e s : 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


Underwash on mass flow rate of Scrubber

Overhead and Bottom

Figure IX-4 Effect of water instead of

HGO Underwash on mass flow rate of

other streams

Note: Lines that connect data points do not present trend lines. They are shown to help


117


Table IX-1 Effect of water instead of HGO Underwash on Scrubber parameters

Water Flow Rate Water flow rate (instead of HGO Underwash) (barrel/day) m3/h 0 3 12 12

Position ft kbarrel/day 0 (10 kbl/day HGO) 0.70 2.50 2.50 from the (Base Case) Uncontrolled Controlled by ATB Controlled by HGO

pool bottom Case: 1 2 % 3 % 4 % Koch Grid Top 43 Top Stage Temp Est (°C) 393 393 0.0 389 -1.1 390 -0.8 Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 395 396 0.3 390 -1.2 387 -2.1 Sheds Top 34 Top Stage Temp Est (°C) 405 406 0.3 401 -1.0 395 -2.4 Sheds Bot. 22 Bottom Stage Temp Est (°C) 473 483 2.1 484 2.3 483 2.0 Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 376 0.3


Scrubber Overhead Actual Volume Flow (rrfVh) 274,787 279,116 1.6 292,649 6.5 290,197 5.6 Mass Flow (kg/h) 778,651 751,648 -3.5 748,329 -3.9 697,520 -10.4 Mass Density (kg/m3) 2.83 2.69 -5.0 2.56 -9.8 2.40 -15.2

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 721 -8.6 580 -26.5 679 -13.9 Mass Flow (kg/h) 643,728 593,311 -7.8 476,648 -26.0 559,999 -13.0 Mass Density (kg/m3) 816 823 0.9 822 0.8 824 1.0

Overhead / ATB Mass Flow Ratio 2.58 2.49 -3.5 3.15 22.2 2.31 -10.4 Grid Liquid Actual Volume Flow (m3/h) 305 309 1.3 318 4.2 202 -33.7

Mass Flow (kg/h) 217,861 220,857 1.4 227,254 4.3 146,778 -32.6 Mass Density (kg/m3) 714.97 715.70 0.1 715.74 0.1 726.36 1.6

Shed Vapor Actual Volume Flow (m3/h) 277,329 281,729 1.6 295,180 6.4 290,634 4.8 Mass Flow (kg/h) 870,450 846,437 -2.8 849,522 -2.4 798,526 -8.3 Mass Density (kg/m3) 3.14 3.00 -4.3 2.88 -8.3 2.75 -12.5

Shed Liquid Actual Volume Flow (m3/h) 497 455 -8.4 384.31 -22.6 423 -14.9 Mass Flow (kg/h) 361,323 329,985 -8.7 277,213 -23.3 308,849 -14.5 Mass Density (kg/m3) 724.76 725.14 0.1 721.32 -0.5 730.15 0.7

To Coker Actual Volume Flow (m/h) 330 301 -8.6 242 -26.5 284 -13.9 Mass Flow (kg/h) 269,078 248,004 -7.8 199,239 -26.0 234,080 -13.0 Mass Density (kg/m3) 815.85 822.97 0.9 822.45 0.8 824.18 1.0


Vapour to Sheds Temperature (°C) 514 518 0.7 514 0.0 518 0.9 Upgoing Stream Temperature (°C) 534 535 0.1 534 0.0 535 0.1 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 1.00 33.3 1.00 33.3 SPL Coler Duty (MMBtu/h) 44.72 44.26 -1.0 42.74 -4.4 39.06 -12.7

118


Table IX-2 Effect of water instead of HGO Underwash on Scrubber Overhead properties

Water flow rate Basic 0.7 kbpd H20 2.5 kbpd H20 2.5 kbpd H20 Cut Point [%] Basic 0.7 kbpd HjO

(10 kbpd HGO Wash) Uncontrolled Controlled by ATB Controlled by HGO (10 kbpd HGO Wash) Uncontrolled Temperature fC] 393 393 389 390 TBP fC]

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

Mass Density [kg/m3] 2.83 2.70 2.56 2.40 2 •207 -209

Act. Volume Flow [m3/h] 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 •51 -56 Specific Heat [kJ/kgmole-C] 193.80 184.96 173.16 164.05 15 •34 •43 Std. Gas Flow [STD_m3/h] 260,141 263,719 278,271 274,537 17.5 -3 -5 Watson K 11.38 11.42 11.44 11.51 20 266 248 Liq. Mass Density (Std. Cond) [kg/m3] 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 2862 2899 2939 35 354 351

1.05 1.05 1.05 1.05 40 373 367

Fraction Distribution Data 45 391 391

Volume fraction 50 405 406

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)

524+(>524°C)

0.580 0.569 0.558 0.554 65 442 444 HGO(343-524°C)

524+(>524°C) 0.101 0.107 0.107 0.108 70 459 465

75 471 481 80 486 489

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

70

60

| 50

O > 40

30 \

20

• O kbarrel/day water, 10 kbarrel/day HGO

• 0.7 kbarrel/day w ater, uncontrolled

-2.5 kbarrel/day water, controlled by ATB

•2.5 kbarrel/day w ater, controlled by HGO wash

10 \

0 -I , , , , , • , • , 1 O 100 200 300 400 500 600 ' 700 800 900 1000

Temperature, °C

Figure IX-5 Effect of water instead of HGO Underwash on Scrubber Overhead TBP curve

119


Table IX-3 Effect of water instead of HGO Underwash on Scrubber Bottom properties

Water flow rate Basic 0.7 kbpd H 20 2.5 kbpd H20 2.5 kbpd HjO Cut Point Basic 0.7 kbpd H 20 2500 kbpd H20 2500 kbpd H20

(10kbpd Uncontrolled Controlled Controlled [%] (10 kbpd Uncontrolled Controlled Controlled HGO Wash) by ATB by HGO HGO Wash) by ATB by HGO

Temperature f C] 375 375 375 375 TBPfC] TBP[°C] TBPfC] TBPfC] Pressure [psig] 17.00 17.00 17.00 17.00 0 412 420 425 423 Molecular Weight 637.04 654.11 654.85 664.51 1 441 456 460 458 Mass Density [kg/m3] 815.85 822.97 822.45 824.18 2 466 485 487 486 Act. Volume Flow[m3/h] 812 741 597 699 3.5 490 509 512 512 Mass Enthalpy [kJ/kg] -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 Vapor Phase Fraction (Mass Basis) 0.00 0.00 0.00 0.00 12.5 520 555 551 556 Specific Heal [kJ/kgmole-C] 1843.61 1887.95 1890.50 1920.13 15 525 556 556 558 Std. Gas Flow [STD_m3/h] 23,893 21,417 17,210 19,926 17.5 548 559 558 563 Watson K 11.41 11.40 11.40 11.41 20 556 564 564 591 Kinematic Viscosity [cSt] 0.69 0.69 0.71 0.69 25 564 592 592 597 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1044.56 1043.80 1046.30 30 592 605 604 626 Molar Volume [m3/kgmole] 0.78 • 0.79 ,. 0.80 0.81 35 605 630 630 632 Mass Heat of Vap. [kJ/kg] 1309 1685 1703 1707 40 630 635 635 680 Surface Tension [dyne/cm] 15.47 15.72 15.72 1572 45 635 681 681 683 Thermal Conductivity [W/m-K] 0.13 0.14 0.14 0.14 50 682 684 684 685 Viscosity [cP] 0.57 0.57 0.58 0.57 55 684 691 691 693 Fraction Distribution Data 60 693 700 701 704

Volume fractior 65 706 740 740 742 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 1049

100 1055 1056 1060 1057

100

400 500 600 700 800 900 1000 1100 Temperature, °C

Figure IX-6 Effect of water instead of HGO Underwash on Scrubber Bottom TBP curve

120


200-300 300-400

• 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%

400-500 500-600 600-700 700-800 800-900 Components ' 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%

• r r r g ] m 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 1000>


Figure IX-8 Effect of water instead of HGO Underwash on Scrubber Bottom composition

121


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


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:


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


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

of HGO Underwash on temperatures along

the Scrubber




Figure X-2 Effect of saturated steam

instead of HGO Underwash on

temperature profile along the Scrubber


-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


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/m3 (Table X- l ) .

-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).


-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/m3 (Table X- l ) .

-Average molecular weight changes from 637 to 652, 664, 670 (Table X-3).


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


2 3 4

C a s e s : 1: 0 kg/h steam, 10 kbarrel/day HGO; 2: 7000 kg/h steam, uncontrolled; 3:22000 kg/h

steam, controlled by ATB; 4:22000 kg/h steam, controlled by HGO wash; 5:4627 kg/h steam,

uncontrolled Pressure drop in Grid, in. of water

2 3 4

C a s e s : 1:0 kg/h s team, 10 kbarrel/day HGO; 2: 7000 kg/h steam, uncontrolled; 3: 22000 kg/h

steam, controlled by ATB; 4:22000 kg/h steam, controlled by HGO wash; 5: 4627 kg/h steam,

uncontrolled

Figure X-3 Effect of saturated steam instead

of HGO Underwash on mass flow rate of




Figure X-4 Effect of saturated steam

instead of HGO Underwash on mass

flow rate of other streams


126


Table X-l Effect of saturated steam instead of HGO Underwash on Scrubber parameters

Saturated Steam Flow Rate Saturated steam flow rate instead of HGO Underwash

Position ft kg/h 0 (10 kbl/day HGO) 7000 22000 22000 4627

from the Base Case Uncontrolled Controlled by ATB Controlled by HGO Same as water pool bottom Case: 1 2 % 3 % 4 % 5 %

Koch Grid Top 43 Top Stage Temp Est (°C) 393 392 -0.2 389 -1.0 390 -0.7 394 0.2

Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 395 395 0.0 391 -1.1 386 -2.4 397 0.5

Sheds Top 34 Top Stage Temp Est (°C) 405 405 0.1 401 -0.9 394 -2.7 407 0.5

Sheds Bot. 22 Bottom Stage Temp Est (°C) 473 482 1.7 497 5.0 486 2.7 486 2.6

Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 375 0.0 375 0.0

Flow Rates& Densities

Scrubber Overhead Actual Volume Flow (m3/h) 274,787 281,987 2.6 300,449 9.3 297,419 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

Mass Density (kg/m3) 2.83 2.67 -5.8 2.52 -11.2 2.34 -17.5 2.70 -4.7

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 727 -7.9 510 -35.3 661 -16.2 713 -9.6

Mass Flow (kg/h) 643,728 597,497 -7.2 422,224 -34.4 546,779 -15.1 587,558 -8.7

Mass Density (kg/m3) 816 822 0.8 827 1.4 827 1.4 824 1.0

Overhead / ATB Mass Flow Ratio 2.58 2.49 -3.3 3.48 35.0 2.30 -10.7 2.50 -3.1

Grid Liquid Actual Volume Flow (m3/h) 305 307 0.6 324 6.3 189 -38.1 310 1.8

Mass Flow (kg/h) 217,861 219,461 0.7 232,176 6.6 137,495 -36.9 221,960 1.9

Mass Density (kg/m3) 714.97 715.98 0.1 716.77 0.3 729.26 2.0 715.73 0.1

Shed Vapor Actual Volume Flow (m3/h) 277,329 284,611 2.6 303,004 9.3 297,520 7.3 282,240 1.8

Mass Flow (kg/h) 870,450 846,044 •2.8 861,981 -1.0 800,297 -8.1 850,652 -2.3

Mass Density (kg/m3) 3.14 2.97 -5.3 2.84 -9.4 2.69 -14.3 3.01 -4.0

Shed Liquid Actual Volume Flow (m3/h) 497 458 -7.8 349.92 -29.6 411 -17.3 450 -9.3

Mass Flow (kg/h) 361,323 332,365 -8.0 249,529 -30.9 300,237 -16.9 326,195 -9.7

Mass Density (kg/m3) 724.76 725.57 0.1 713.11 -1.6 730.43 0.8 724.23 -0.1

To Coker Actual Volume Flow (m3/h) 330 304 -7.9 213 -35.3 276 -16.2 298 -9.6

Mass Flow (kg/h) 269,078 249,754 -7.2 176,490 -34.4 228,554 -15.1 245,599 -8.7

Mass Density (kg/m3) 815.85 822.01 0.8 827.32 1.4 826.99 1.4 824.18 1.0


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


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 Basic 7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam

(10 kbpd Uncontrolled Controlled Controlled [%] (10 kbpd Uncontrolled Controlled Controlled HGO Wash) by ATB by HGO HGO Wash) by ATB by HGO

Temperature ft] 393 392 389 390 TBP [°C] TBP ft] Pressure [psig] 16.99 16.99 16.99 16.99 0 -253 -253 -253 •253 Molecular Weight 70.77 66.72 62.60 58.50 1 -237 -239 -239 •242 Mass Density [kg/m3] 2.83 2.67 2.52 2.34 2 -207 -209 -210 -215 Act. Volume Flow [m3/h] 282,860 290,271 309,276 306,156 3.5 •167 -171 -172 •179 Mass Enthalpy [kJ/kg] -2924 •3095 -3326 -3510 5 -136 •140 •141 -149 Mass Entropy [kJ/kg-C] 5.30 5.46 5.60 5.82 7.5 -102 •104 -105 •112 Mass Heat Capacity [kJ/kg-C] 2.74 2.74 2.72 2.73 10 -85 •90 •91 •96 Vapor Phase Fraction (Mass Basis) 0.93 0.93 0.92 0.95 12.5 -51 •57 -59 •73 Specific Heat [kJ/kgmole-CJ 182.60 170.32 159.56 164.05 15 •34 •42 •44 •48 Std. Gas Flow[STD_m3/h] 266,718 285,477 281,095 274,537 17.5 -3 -5 -6 •29 Watson K 11.38 11.42 11.44 11.53 20 266 246 237 •3 Liq. Mass Density (Std. Cond) [kg/m3] 930.01 922.64 922.97 906.82 25 310 308 303 289 Molar Volume [m3/kgmole] 24.98 25.00 24.88 25.02 30 336 334 325 322 Mass HeatofVap. [kJ/kg] 2825 2868 2950 2960 35 354 351 351 340

1.05 1.05 1.05 1.06 40 373 367 366 364 Fraction Distribution Data 45 391 390 390 380

C4-(<177°C) Volume fraction 50 405 406 406 399

C4-(<177°C) 0.060 0.059 0.075 0.093 55 420 420 420 420 LGO(177-343°C) 0.259 0.266 0.259 0.262 60 439 439 439 438 HGO (343-524°C) 0.580 0.568 0.557 0.529 65 442 443 444 444 524+(>524°C) 0.101 0.107 0.110 0.116 70 459 464 465 464

75 471 480 482 482 80 486 489 491 490 85 496 512 513 515 90 525 526 527 527

92.5 537 539 541 542 95 550 552 553 555

96.5 556 556 557 563 98 616 630 640 655 99 744 755 767 782

100 871 882 889 909

HGO

— — 7000 kg/h water,uncontrolled

22000 kg/h water, controlled by ATB

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


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 Basic 7000 kg/h Steam 22000 kg/h Steam 22000 kg/h Steam

(10 kbpd Uncontrolled Controlled Controlled [%] (10 kbpd Uncontrolled Controlled Controlled HGO Wash) by ATB by HGO HGO Wash) by ATB by HGO

Temperature fC] 375 375 375 375 TBPfC] TBP fC] TBPfC] TBPfC)

Pressure [psig] 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 Act. Volume Flow [m3/h] 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 Std. Gas Flow [STD_m3/h] 23,893 21,659 15,018 19,293 15 525 556 560 561 Watson K 11.41 11.40 11.39 11.41 17.5 548 557 569 590 Kinematic Viscosity [cSt] 0.69 0.69 0.72 0.70 20 556 562 591 592 Liq. Mass Density (Std. Cond) [kg/m3] 1038.96 1043.84 1047.30 1047.97 25 564 593 598 602 Molar Volume [m3/kg mole) 0.78 0.79 0.80 0.81 30 592 603 626 628 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 50 682 684 685 687

55 684 690 694 695 Fraction Distribution Data 60 693 699 706 706

Volume fraction 65 706 740 743 742 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

95 917 943 962 953 96.5 964 964 987 965

98 1031 1035 1043 1039 99 1047 1048 1053 1049

100 1055 1056 1062 1057


Figure X-6 Effect of sat. steam instead of HGO Underwash on Scrubber Bottom TBP curve

129


• 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 wash: Light Ends: 19%; Water: 70% ; 100> fraction: 11%

1 200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 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%

200-300 300-400 400-500 500-600 600-700 700-800 800-900 900-1000 Components' Boiling Temperatures Range, °C

1 , wrrm , mrm 1000>

Figure X-8 Effect of saturated steam instead of HGO Underwash on Scrubber Bottom

composition

130


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 middle fractions that 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


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"10) for high boiling components (524°C+) that

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


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


Table XI-1 ATB flow rate effect on Overhead TBP distillation curve

ATB Flow Rate (kbarrel/day)

Cut Point [%] 55 80 100

TBP fC] TBP ["C] TBP fC] 0 -253 -253 -253 1 -237 -240 -243 2 -207 -211 -216

3.5 -167 -174 -181 5 -136 -143 -151

7.5 -102 -107 -114 10 -85 -92 -97

12.5 -51 -62 -76 15 -34 -45 -50

17.5 -3 -8 -42 20 266 206 -5 25 310 298 282 30 336 324 310 35 354 341 335 40 373 363 349 45 391 377 364 50 405 391 377 55 420 405 391 60 439 419 403 65 442 436 407 70 459 441 420 75 471 448 440 80 486 466 445 85 496 489 480 90 525 523 517

92.5 537 532 531 5M 1 ri4(. '545

" 9 6 . 5 556 556 555 98 616 617 648 99 744 753 780

100 871 888 911

134


Table XI-2 Effect of ATB flow rate on Scrubber parameters

Position ft from the

pool bottom

ATB Flow Rate m3/h

kbarrel/day 273 55

(Base Case)

397 80

%

497 100

% Koch Grid Top 43 Koch Grid Bot. 38 Sheds Top 34 Sheds Bot. 22 Scrubber Pool 0

Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Top Stage Temp Est (°C) Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)

393 395 405 473 375

376 376 386 447 398

-4.3 -4.9 -4.8 -5.6 6.2

364 362 372 436 371

-7.4 -8.3 -8.1 -7.9 -0.9

Flow Rates& Densities Scrubber Overhead Actual Volume Flow (mJ/h)

Mass Flow (kg/h) Mass Density (kq/m3)

274,787 778,651

2.83

266,026 723,279

2.72

-3.2 -7.1 -4.1

257,200 663,909

2.58

-6.4 -14.7 -8.9

Scrubb.Pool Liquid Actual Volume Flow (m3/h) Mass Flow (kg/h) Mass Density (kq/m3)

789 643,728

816

1,385 1,103,082

796

75.5 71.4 -2.4

1,997 1,575,214

789

153.1 144.7 -3.3

Overhead / ATB Mass Flow Ratio 2.58 1.65 -36.1 1.15 -55.4 Grid Liquid Actual Volume Flow (m3/h)

Mass Flow (kg/h) Mass Density (kq/m3)

305 217,861

714.97

260 188,811

724.99

-14.5 -13.3

1.4

227 167,283

736.24

-25.4 -23.2 3.0

Shed Vapor Actual Volume Flow (mJ/h) Mass Flow (kg/h) Mass Density (kq/m3)

277,329 870,450

3.14

268,281 786,029

2.93

-3.3 -9.7 -6.7

258,944 705,130

2.72

-6.6 -19.0 -13.2

Shed Liquid Actual Volume Flow (m3/h) Mass Flow (kg/h) Mass Density (kq/m3)

497 361,323

724.76

819 593,614

724.56

64.9 64.3 0.0

1,115 810,640

727.01

124.4 124.4 0.3

To Coker Actual Volume Flow (mJ/h) Mass Flow (kg/h) Mass Density (kg/m3)

330 269,078

815.85

579 461,088

796.42

75.5 71.4 -2.4

835 658,440

788.78

153.1 144.7 -3.3

Vapour to Sheds Temperature (°C) 514 501 -2.6 492 -4.3 Upqoinq Stream Temperature (°C) 534 533 -0.3 532 -0.5 Sheds Staqe Efficiency 0.53 0.53 0.0 0.53 0.0 Koch Grid Staqe Efficiency 0.75 0.75 0.0 0.75 0.0 SPL Coler Duty (MMBtu/h) 44.32 75.43 70.2 102.31 130.9

1 3 5


Table XI-3 HGO Wash flow rate effect on Overhead TBP distillation curve

HGO Wash Flow Rate (kbarrel/day)

Cut Point [%] 24 40 60 80 100

TBP [°C] TBP ["C] TBP ["CJ TBP [°C] TBP ["C] 0 -253 -253 -253 -253 -253 1 -237 -237 -237 -237 -238 2 -207 -206 -206 -206 -208

3.5 -167 -166 -166 -167 -170 5 -136 -134 -134 -135 -138

7.5 -102 -101 -101 -102 -103 10 -85 -81 -81 -84 -89

12.5 -51 -49 -49 -51 -55 15 -34 -22 -22 -31 -43

17.5 -3 2 2 -1 -4 20 266 274 274 268 254 25 310 311 310 308 300 30 336 336 333 324 322 35 354 352 349 342 336 40 373 366 364 357 349 45 391 382 377 366 363 50 405 397 391 378 371 55 420 407 404 392 378 60 439 420 407 405 391 65 442 440 420 407 400 70 459 444 440 420 406 75 471 465 443 440 420 80 486 477 465 443 440 85 496 491 485 480 475 90 525 524 521 517 496

92.5 537 532 527 527 526 • 550 546 543 543 5|2

96*5 556 555 555 "554 553 98 616 600 601 596 598 99 744 733 731 730 736

100 871 863 861 864 874

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 Temperature, °C

700 800 900 1000

Figure XI-2 HGO Wash flow rate effect on Overhead TBP distillation curve

136


Table XI-4 Effect of HGO Wash flow rate on Scrubber parameters

HGO Wash Flow Rate Posit ion ft m3/h 119 199 298 397 497

from the pool bottom

kbarrel/day 24 (Base Case)

40 %

60 %

80 100 %

Koch Grid Top 43 Top Stage Temp Est (°C) 393 386 -1.9 376 -4.3 368.5 -6.3 359 -8.6 Koch Grid Bot. 38 Bottom Stage Temp Est (°C) 395 389 -1.6 380 -3.9 371.2 -6.0 361 -8.6 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 Scrubber Pool 0

Bottom Stage Temp Est (°C) Bulk Liquid Temperature (°C)

473 375

456 375

-3.7 0.0

441 375

-6.9 0.0

430.6 375.0

-9.0 0.0

420 375

-11.4 0.0


Scrubber Overhead Actual Volume Flow (m'Vh) 274,787 273,178 -0.6 270,319 -1.6 267410.3 -2.7 263,442 -4.1 Mass Flow (kg/h) 778,651 797,485 2.4 798,630 2.6 787391.8 1.1 762,774 -2.0 Mass Density (kg/m3) 2.83 2.92 3.0 2.95 4.3 2.9 3.9 2.90 2.2

Scrubb.Pool Liquid Actual Volume Flow (m3/h) 789 1,001 26.8 1,352 71.4 1723.4 118.4 2,251 185.3 Mass Flow (kg/h) 643,728 797,839 23.9 1,053,191 63.6 1323490.9 105.6 1,709,377 165.5 Mass Density (kg/m3) 816 797 -2.3 779 -4.5 768.0 -5.9 759 -6.9

Overhead / A I B Mass Flow Ratio 2.58 2.64 2.4 2.65 2.6 2.6 1.1 2.53 -2.0 Grid Liquid Actual Volume Flow (m3/h) 305 425 39.5 588 93.0 739.7 142.8 928 204.5

Mass Flow (kg/h) 217,861 303,893 39.5 425,029 95.1 541125.0 148.4 688,636 216.1 Mass Density (kg/m3) 714.97 715.03 0.0 722.87 1.1 731.5 2.3 742.07 3.8

Shed Vapor Actual Volume Flow (mJ/h) 277,329 277,899 0.2 277,562 0.1 276478.3 -0.3 273,949 -1.2 Mass Flow (kg/h) 870,450 892,151 2.5 906,554 4.1 910060.9 4.6 895,642 2.9 Mass Density (kg/m3) 3.14 3.21 2.3 3.27 4.1 3.3 4.9 3.27 4.2

Shed Liquid Actual Volume Flow (m3/h) 497 639.68 28.8 860 73.1 1064.9 114.3 1,325.36 166.8 Mass Flow (kg/h) 361,323 459,545 27.2 613,054 69.7 758496.6 109.9 947,845 162.3 Mass Density (kg/m3) 724.76 718.40 -0.9 712.77 -1.7 712.3 -1.7 715.16 -1.3

To Coker Actual Volume Flow (mJ/h) 330 418 26.8 565 71.4 720.4 118.4 941 185.3 Mass Flow (kg/h) 269,078 333,497 23.9 440,234 63.6 553219.2 105.6 714,520 165.5 Mass Density (kg/m3) 815.85 797.16 -2.3 778.82 -4.5 768.0 -5.9 759.29 -6.9


Vapour to Sheds Temperature (°C) 514 505 -1.8 494 -4.0 485.4 -5.6 477 -7.2 Upgoing Stream Temperature (°C) 534 533 -0.2 532 -0.5 530.2 -0.7 528 -1.1 Sheds Stage Efficiency 0.53 0.53 0.0 0.53 0.0 0.5 0.0 0.53 0.0 Koch Grid Stage Efficiency 0.75 0.75 0.0 0.75 0.0 0.8 0.0 0.75 0.0 SPL Coler Duty (MMBtu/h) 44.72 48.25 7.9 56.03 25.3 62.2 39.0 67.02 49.9

137


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 R a t e ( k b a r r e l / d a y )

Cut Point [%] 0 10 20 50 100 Cut Point [%] TBP [°C] TBP [°C] TBP ["C] TBP fC] TBP fC]

0 -253 -253 -253 -253 -253 1 -238 -237 -237 -237 -239 2 -208 -207 -206 -206 -209

3.5 -170 -167 -166 -166 -171 5 -138 -136 -135 -134 -139

7.5 -103 -102 -101 -101 -104 10 -88 -85 -82 -82 -89

12.5 -55 -51 -50 -50 -56 15 -43 -34 -26 -24 -43

17.5 -4 -3 -1 0 -5 20 255 266 271 272 249 25 309 310 310 309 296 30 336 336 336 330 320 35 353 354 352 349 335 40 374 373 367 363 345 45 392 391 385 377 355 50 406 405 402 391 364 55 420 420 417 397 377 60 440 439 423 406 381 65 446 442 441 420 392 70 465 459 448 440 405 75 482 471 466 442 420 80 491 486 482 465 451 85 514 496 493 473 460 90 526 525 524 521 490

92.5 539 537 535 527 526 95 552 550 HIBi^g8 543 542

96.5 556 556 556 555 553 98 632 616 608 601 597 99 754 744 737 731 736

100 879 871 866 862 875

10

O-l , , , , , , , , , 1

0 100 200 300 400 500 600 700 800 900 1000 Temperature, °C

Figure XI-3 HGO Underwash flow rate effect on Overhead TBP distillation curve

138


Table XI-6 Effect of HGO Underwash flow rate on Scrubber parameters

HGO Uderw. Flow Rate HGO Underwash Flow Rate

Position ft m3/h 50 0 50 99 248 497

from the kbarrel/day 10 0 10 20 50 100

pool bottom (Base Case) % % % % %

Koch Grid Top 43 Top Stage Temp Est fC ) 393 397 0.9 393 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

Scrubber Pool 0 Bulk Liquid Temperature (°C) 375 375 0.0 375 0.0 375 0.0 375 0.0 375 0.0

Flow Rates& Densities

Scrubber Overhead Actual Volume Flow (m3/h) 274,787 274,718 0.0 274,787 0.0 273,909 -0.3 269,959 -1.8 262,182 -4.6

Mass Flow (kg/h) 778,651 757,412 -2.7 778,651 0.0 790,657 1.5 795,252 2.1 755,852 -2.9

Mass Density (kg/m3) 2.83 2.76 -2.7 2.83 0.0 2.89 1.9 2.95 4.0 2.88 1.7

Scrubb.Pool Liquid Actual Volume Flow (nvVh) 789 688 -12.9 789 0.0 921 16.8 1,422 80.2 2,555 223.8

Mass Flow (kg/h) 643,728 569,357 -11.6 643,728 0.0 739,959 14.9 1,103,587 71.4 1,931,997 200.1

Mass Density (kg/m3) 816 828 1.5 816 0.0 803 -1.6 776 -4.9 756 -7.3

Overhead/ATB Mass Flow Ratio 2.58 2.51 -2.7 2.58 0.0 2.62 1.5 2.63 2.1 2.50 -2.9

Grid Liquid Actual Volume Flow (m3/h) 305 314 3.2 305 0.0 296 -2.7 268 -12.0 232 -23.9

Mass Flow (kg/h) 217,861 224,963 3.3 217,861 0.0 212,704 -2.4 195,892 -10.1 174,124 -20.1

Mass Density (kg/m3) 714.97 715.50 0.1 714.97 0.0 717.71 0.4 730.60 2.2 750.76 5.0

Shed Vapor Actual Volume Flow (m3/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

Mass Density (kg/m3) 3.14 3.09 -1.6 3.14 0.0 3.17 1.1 3.18 1.3 3.05 -2.9

Shed Liquid Actual Volume Flow (m3/h) 497 440 -11.4 496.80 0.0 587 18.1 898.84 80.9 1,456 193.1

Mass Flow (kg/h) 361,323 316,491 -12.4 361,323 0.0 423,013 17.1 641,101 77.4 1,047,135 189.8

Mass Density (kq/m3) 724.76 718.94 -0.8 724.76 0.0 720.99 -0.5 713.25 -1.6 719.23 -0.8

To Coker Actual Volume Flow (m Ih) 330 287 -12.9 330 0.0 385 16.8 594 80.2 1,068 223.8

Mass Flow (kg/h) 269,078 237,991 -11.6 269,078 0.0 309,303 14.9 461,299 71.4 807,575 200.1

Mass Density (kg/m3) 815.85 828.00 1.5 815.85 0.0 803.04 -1.6 776.05 -4.9 756.30 -7.3


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


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 of Sheds over 6 trays,

and changing the number of Koch 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 Koch 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


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 ATB 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

ATB, 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 HGO 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 HGO

143


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


8.1. Conclusions

H Y S Y S process simulation of the Scrubber Section of the Fluid Coker gave insights into the

process behaviour and improved understanding of the whole process. Case studies showed trends

and quantitative outcomes of 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 of the Fluid

Coker. Results o f the simulation match the plant data very well (within 3.2% of the plant

data), once separation efficiencies near zero were assigned to the heaviest fractions.

• Calculations in Chapter 4 suggest that liquid entrainment may be present in the system.

Entrainment o f heavy species into the vapours decreases the efficiency o f the Shed section

radically.

• Consideration in Chapter 5 suggests that the K o c h Grid operates out of the designed

conditions. Too high gas loading and too low liquid flow rate result in 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 Koch Gr id is within the design range of operating

conditions could help improving the efficiency of 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


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


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


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


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 (343-

524°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

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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 th, Chapter 17-Gas-solids separation, 1997.


Edition 7th, Chapter 6- Fluid and particle dynamics, 1997.


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

http://www.koch-glitsch.com

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. 2 <,.„ V-b 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 c and Tc, and the acentric

factor co for pure substances [42].

For a pure component:

a = aca

ac = 0.45724^--^- (1-2)

b = 0.077480—— -

a represents temperature dependence of parameter a .

^ = l + K(\-Tr

05) (I_3)

K = 0.37464 +1.5422G7 - 0.26992ET2

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 = tltt(x,xJav) (1-4) 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 j ) a-7)

„ . - 0 - 4 5 7 2 * ' ^ (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-Aii+BvT + ^L ( I_42)

where A y , B y and C y are asymmetric binary interaction parameters. Values forau, a ~, b{, K „

A y , B y and C y can be calculated or found in tables. H Y S Y S has a library with binary

interaction parameters for more than 16000 binaries. If the system contains pseudo-

components, H Y S Y S provides a wide selection of the estimation methods for all needed

parameters - T c , P c, co, K , etc. Methods for estimating the interaction binaries between pseudo-

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 n c)

L ( X i , X 2 , X 3 , . . . , X n c )

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 n c , ( where nc is number of components) separates into L

moles of liquid with the composition xi, X2, z n c, and V moles of vapour with the

composition yu y 2 , y n c .

Overall mole balance on this system is:

F (Zi,Z 2,Z3,...,Znc)

F (Zi,Z 2,Z3,...,Znc)

t Q

159

Appendix II

L + V = 1 (II-l)

And the component mole balance is:

z, = x,L + ytV = XiL + y . (1 - L) = x. (l-V) + yy 0 = 1,2 ,...,nc) (II-2)

Energy balance:

FhF +Q = Vhv + LhL (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:

where O", and O', are the fugacity coefficients for the component i in the vapour and liquid

(II-4)

phases. Combining Equations (II-2) and (II-4), two sets of equations can be obtained.

First set contains X j :

(II-5) X: =

160

Appendix II

Since X j must sum to unity:

nc nc g

^ • - ^ L + Kfl-L)'1 ( " - 6 )

Another set contains y;

y' = o-£'lV (II"7)

nc nc J£ g

^ ' - ^ o - ^ V ' 1 (I,-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 X j 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 rp

R• T• lncp,. = f(vf -R—)-dp (II-9a) o P

dV v . = (ill.) . . (H-9b)

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 (II-

13a).

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:

O; ,„ „ , .„ _ A ln^. = - J - ( Z - l ) - l n ( Z - 5 ) -bK ' v ' 242B

2lLxjaji bt | l n [ Z + (1 + 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 3 - ( 1 - 5 ) Z 2 +Z(A-3B2 -2B)-(AB-B2 - 5 3 ) = 0 (II-13a)

162

Appendix II

aP

R2T2 (II-13b)

B = — (II-13c) 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 (I-

3) in Appendix I, and presented in Table AII.2:

163

Appendix II

Table AII.2 PR EOS parameters for pure substances

Substance Pc Tc CO a a; bi

Hydrogen (1) 2.05E+06 43.6 0 0.327 9.61 E+03 1.38E-02

Methane (2) 4.62E+06 190.63 0.01 0.981 2.44E+05 2.67E-02 Ethane (3) 4.88E+06 305.43 0.099 1.210 7.31 E+05 4.05E-02

Interaction parameters a-rare calculated in somewhat simpler way then presented in

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.

a g = ^ ( l - k u ) ( IM4 )

Table AII.3 Interaction parameters for Hydrogen-Methane-Ethane system

Substance

Hydrogen (1) a„= 9.61 E+03 a1 2= 4.95E+04 °13= 9.78E+04

Methane (2) a 2 J= 4.95E+04 a 2 2= 2.44E+05 Q23= 4.26E+05

Ethane (3) 031= 9.78E+04 a 3 2= 4.26E+05 Q33= 7.31 E+05

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; x2= 0.70; x 3 = 0.25

2. Mixing rules ( Eq. (1-4) and (1-5)) are applied:

a = W l l + W l 2 + W l 3 + W , + + W 2 3 + W 3 , + ( I M 5 )

"^3"^2^32 "^3"^'3^33

164

Appendix II

fl=3.2-105Pa m6/kmol2

b = + x2bz + x3b3 (11-16)

b=2.95-10"2 m3/mol

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 (II-

13a) 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 = 0.07

5. To estimate vapour composition Kj should be calculated from the following equation:

Kt=^- (H-17)

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:

Ky = 6.68 K2 = 0.61 K3 = 0.07

6. Vapour composition can be further estimated:

165

Appendix II

yx =K,x, =6.68-0.05 = 0.334 y2 =K2x2 =0.61-0.70 = 0.427 y3 =K3x3 =0.07-0.25 = 0.018

7. Since the sum of the fractions should 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:

0-334 n „„ yx = = 0.42

1 0.779

° - 4 2 7 n « y, = = 0.55 2 0.779

, , = * ° 1 £ - 0.023 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;x2= 0.55; x 3 = 0.023

2. Mixing rules (Eq. (1-4) and (1-5)) are applied:

0=1.11-10s Pa m6/kmol2

b=2.14-10"2 m3/mol

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 (II-

13a) 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

f2 = 0.62

<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 3 0.32

6. Liquid composition can be further estimated:

J C , =yxIK, =0.42/5.43 = 0.077 x2 =y2/K2 =0.55/0.98 = 0.56

X i = 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, +x2 +x, =0.077 + 0.56 + 0.10 = 0.738

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

° - 0 7 7 n i n x, = = 0.10 1 0.738

0 5 6 ft if, y, = = 0.76 2 0.738

3 0.738

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:

Kx = 6.367 K2 = 0.936 K, =0.158

Composition of the vapour phase:

jy, =0.312 y2 = 0.647 y, = 0.041

Composition of the liquid phase:

x, = 0.049 x2 =0.691 J C 3 = 0.260

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:

yx = 0.294 y2 = 0.659 y3 = 0.047

Composition of the liquid phase:

x, = 0.049 x2 = 0.701 x3 = 0.250

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

Table AIII.l Composition of hypothetical cyclone stream

Appendix III

Fraction wt% Water 19 Light Ends 10 CGO 54 OTSB 15

Table AIII.2 Composition of Light Ends fraction of cyclone stream

Light Ends components wt fraction (of Light Ends)

Hydrogen 0.01 H2S 0.06 Methane 0.21 Ethane 0.16 Ethylene 0.08 Propane 0.12 Propylene 0.13 Butadiene 0.02 Butenes 0.12 i-Butane 0.01 n-Butane 0.06

171

Appendix III

Table AIII.3 CGO assay Table AIII.4 CGO TBP data Method: ASTM 2887 with HTSB Method: TBP calculated by enhancement HYSYS.

ASTM D2887 Vol % TBP (°C) wt% NBP (UC) 0 229

0 221 1 237 5 266 2 248

10 287 3.5 263 15 304 5 275 20 319 7.5 288 25 333 10 297 30 345 12.5 306 35 357 15 314 40 368 17.5 322 45 380 20 329 50 391 25 342 55 403 30 354 60 414 35 366 65 426 40 378 70 438 45 390 75 450 50 402 80 464 55 414 85 479 60 426 90 496 65 438 95 521 70 450

100 572 75 462 80 474 85 487 90 501

92.5 511 95 523

96.5 531 98 541 99 541

100 541

172

Appendix

Table AIII.5 OTSB Assay Method: ASTM 2887 & SCFE-composite data ASTM SCFE COMPOSITE

y. Boil Boil Temp (C)

MT't SBotl

Boil Temp (C)

1 KBoil (

3oi Temp Q

0 315.7 0 523.3 0 3157

i 318.7 1 523.5 0.5 318.7

2 327.4 2 523.7 1 327.4

3.5 343.7 3.5 524.0 1.75 343.7

5 360.7 5 524.3 2.5 360.7

7.5 3838 7.5 524.8 3.75 383.8

10 4007 10 525 6 5 400.7 12.5 413 6 12:5 526.0 6.25 413.6

15 423 7 15 526.3 7.5 423.7

17.5 432 0 17.5 526.8 8.75 432.0 20 43S0 20 527.3 10 439.0

25 45C 5 25 528.4 12.5 450.5

30 45S6 30 533.6 15 459.6

35 468 7 35 539.2 17.5 468.7

40 477 1 40 549.6 20 477.1

45 484 7 45 566.3 22.5 484.7 50 492 1 50 583.4 25 492.1 55 499 3 55 597.6 27.5 499.3

60 506:3 60 639.1 30 506.3

65 512.9 65 695.7 32.5 5119

70 518.5 70 759.0 35 518.5

75 5215 75 825.8 37.5 522.5

80 526.0 80 892.9 40 526.0

85 529.2 85 957.1 42.5 529 2

90 532.4 90 1015.1 45 5324

92.5 534.0 92.5 1040.9 46.25 534 0

95 535.7 95 1063.9 47.5 535.7

96.5 536.8 96.5 1076.2 48.25 536.8

98 5379 98 1086.1 49 537.9

99 538 6 99 1089.1 49.5 538.6

100 539 4 100 1092.2 50 539.4 50.5 523.5 51 523 7

51.75 524 0 52.5 524.3 53.75 524.8

55 525 6 56.25 526 0 57.5 526 3 58.75 526 8

60 527.3 62.5 528 4 65 533.6

67.5 539 2 70 549 6

72.5 566.3 75 583 4

775 597 6 80 639.1

825 695 7 85 759.0

87 5 825 8 90 892.9

925 957.1 95 1015 1

96.25 1040S 97 5 1063 9 98 25 1076 2

99 1086 1 99.5 1089.1 100 1092 2

Table AIII.6 OTSB TBP data Method: TBP calculated by HYSYS

ASTM S C F E

vol % TBP (°C) TBP (°C)

0 341 523 1 345 524 2 355 524

3.5 374 524 5 394 524

7.5 419 525 10 436 526

12.5 448 526 15 458 526

17.5 466 527 20 472 527 25 483 528 30 492 534 35 501 539 40 509 550 45 517 566 50 525 583 55 532 598 60 540 639 65 544 696 70 547 759 75 ^52 826 80 553 893 85 555 957 90 558 1015

92.5 558 1041 95 560 1064

96.5 561 1076 98 562 1086 99 562 1089

100 563 1092

Appendix III

Table AIII.7 Cyclone Product TBP data Method: TBP calculated by HYSYS

Vol % TBP [UC] 0 -253 1 -247 2 -223

3.5 -192 5 -165

7.5 -128 10 -104

12.5 -93 15 -72

17.5 -50 20 -43 25 6 30 289 35 318 40 340 45 361 50 380 55 399 60 420 65 439 70 456 75 466 80 493 85 517 90 540

92.5 546 95 565

96.5 708 98 928 99 981

100 1028

-400 -200 0 200 400 600 800 1000 1200 Temperature, °C

Figure AIII.l Cyclone Product TBP curve

300 600 900 1200 Molecular weight

1500 1800

Figure A I I L 2 Cyclone Product molecular weight distribution curve

Table AIII.8 Cyclone Product composition

Boiling range, "C Mole fractions Mass fractions <=100 0.8921 0.3042

100-200 0.0000 0.0000 200-300 0.0162 0.0599 300-400 0.0395 0.2021 400-500 0.0353 0.2543 500-600 0.0144 0.1345 600-700 0.0005 0.0057 700-800 0.0005 0.0075 800-900 0.0004 0.0064 900-1000 0.0005 0.0107

1000> 0.0006 0.0148 200 400 600 800

Liquid Density, kg/m3

1000 1200

Figure AIII.3 Cyclone Product density distribution curve

174

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% NBP (°C)

0 262 1 277.5 2 296.5 3 308.5 4 318 5 327.5 6 336 7 343.5 8 350.5 9 357

10 363 11 369.5 12 375.5 13 381 14 387 15 392.5 16 397.5 17 403 18 407.5 19 412.5 20 417 21 421.5 22 426 23 430.5 24 435 25 439 26 444 27 448.5 28 453.5 29 458.5 30 463.5 31 468.5 32 473.5 33 478.5 34 483.5 35 488.5 36 494 37 499 38 504 39 509 40 514.5 41 519.5 42 524.5 43 530 44 535 45 540 46 545 47 550 48 555 49 560 50 564.5 51 569 52 573.5 53 578 54 583 55 587.5 56 591 57 595 58 599.5 59 604 61 613 62 617.5 63 622 64 627 65 631.5 66 636 67 640.5 68 645.5 69 651 70 655.5

Table AIII.10 ATB TBP data Method: TBP data calculated by HYSYS

Vol % TBP ("C) 0 316 1 326 2 347

3.5 368 5 384

7.5 405 10 422

12.5 437 15 451

17.5 464 20 475 25 497 30 519 35 544 40 569 45 594 50 618 55 640 60 662 65 683 70 702 75 722 80 744 85 776 90 804

92.5 838 95 884

96.5 922 98 962 99 987

100 1009

200 400 600 800 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

Table AIII.12 HGO assay Method: ASTM 2887 with HTSB enhanc.

Table AIII.12 cont.

ASTM D2887 wt% NBP (°C)

0 263.5 1 275 2 289.5 3 298.5 4 305.5 5 310.5 6 315.5 7 320 8 324.5 9 328.5

10 332 11 336 12 339.5 13 342.5 14 345.5 15 348.5 16 351.5 17 354 18 357 19 359.9 20 362 21 364.5 22 367 23 369.5 24 372 25 374.5 26 377 27 379 28 381.5 29 384 30 386 31 388.5 32 390.5 33 393 34 395 35 397 36 399.5 37 401.5 38 403.5 39 405.5 40 408 41 410 42 412 43 414 44 416 45 418 46 419.5 47 421.5 48 423.5 49 425.5 50 427.5 51 429.5 52 431 53 433 54 435 55 437 56 439 57 441 58 443 59 445 60 447 61 449 62 451 63 453.5 64 455.5 65 457.5 66 460 67 462 68 464.5 69 467

Appendix III Table AIIL13 HGO TBP data Method: TBP calculated by HYSYS

70 469 Vol % TBP (UC) 71 471.5 0 281 72 473.5 1 289 73 476 2 302 74 478.5 3.5 317 75 481 5 327 76 483.5 7.5 339 77 486 10 348 78 488.5 12.5 357 79 491.5 15 365 80 494.5 17.5 372 81 497 20 378 82 500 25 390 83 503 30 402 84 506 35 413 85 509.5 40 424 86 513 45 435 87 517 50 445 88 520.5 55 455 89 525 60 465 90 529.5 65 475 91 535 70 486 92 540.5 75 498 93 547 80 510 94 555 85 526 95 564.5 90 544 96 576 92.5 567 97 591.5 95 603 98 615.5 96.5 643 99 680.5 98 719

100 750 99 763 100 797

400 600 Temperature, °C

1000

Figure AIII.5 HGO TBP curve

Table AIII.14 HGO composition (HYSYS) Boiling range, °C Mole fractions

<=100 0.000 100-200 0.000 200-300 0.027 300-400 0.334 400-500 0.493 500-600 0.113 600-700 0.022 700-800 0.008 800-900 0.004 900-1000 0.000

1000> 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

Table AIII.15 Scrubber Overhead fractions

Appendix III

Table AIII.16 Sour Gas composition

Fraction w t % Sour Gas 12 Butane 5 Naphtha 22 CGO 61

Component Mole Fractions Hydrogen 0.113 H20 0.029 H2S 0.138 Methane 0.317 Ethane 0.133 Ethylene 0.062 Propane 0.086 Propene 0.061 1-Butene 0.001 Biacetylene 0.001 i-Butane 0.004 n-Butane 0.010 i-Pentane 0.000 n-Pentane 0.001 13-Butadiene 0.002 cis2-Butene 0.002 tr2-Butene 0.011 CO 0.007 C02 0.021

180

Table AIII.17 CGO Assay Table AIII.17 Cont Method: SIM Dist

SIM Dist wt % NBP [UC]

0 198 1 225 2 255 3 257 4 264 5 270 6 275 7 280 8 285 9 290

10 294 11 298 12 302 13 306 14 310 15 314 16 318 17 321 18 325 19 328 20 332 21 335 22 338 23 342 24 345 25 348 26 351 27 354 28 357 29 360 30 363 31 366 32 369 33 371 34 374 35 377 36 380 37 383 38 386 39 388 40 391 41 393 42 396 43 399 45 401 46 404 47 407 48 409 49 411 50 414

51 416 52 419 53 421 54 423 55 425 56 428 57 430 58 433 59 435 60 438 61 440 62 443 63 445 64 448 65 450 66 453 67 456 68 458 69 461 70 464 71 466 72 469 73 471 74 474 75 477 76 480 77 482 78 485 79 487 80 491 81 493 82 497 83 500 84 503 85 507 86 510 87 514 88 518 89 523 90 527 91 533 92 539 93 545 94 554 95 564 96 578 97 597

Appendix III

Table AIII.18 CGO TBP data Method: TBP calculated by HYSYS

Vol % TBP (°C) 0 220 1 241 2 266

3.5 276 5 286

7.5 298 10 310

12.5 321 15 330

17.5 339 20 347 25 362 30 377 35 391 40 405 45 417 50 429 55 441 60 453 65 466 70 478 75 492 80 502 85 520 90 542

92.5 564 95 594

96.5 628 98 683 99 716

100 735

181

Appendix III

Table AIII. 19 Naphtha Assay Method: SIM Dist

SIM Dist wt % NBP [°C]

1 -30 2 -7 3 -3 5 1

10 34 15 47 20 67 25 80 30 93 35 106 40 116 45 127 50 140 55 151 60 165 65 176 70 190 75 201 80 212 85 223 90 235 95 251 97 259 98 265 99 276

Table AIII.20 Naphtha TBP data Method: TBP calculated by HYSYS

Vol % TBP (°C) 0 -62 1 -42 2 -17

3.5 -8 5 0

7.5 15 10 30

12.5 40 15 47

17.5 53 20 62 25 77 30 88 35 97 40 107 45 118 50 131 55 144 60 158 65 171 70 184 75 196 80 207 85 217 90 229

92.5 235 95 243

96.5 250 98 258 99 264

100 264

Appendix III

Table AIII.21 "Plant" Scrubber Overhead TBP data Method: TBP calculated by HYSYS

Vol % TBP (UC) 0 -253 1 -216 2 -178

3.5 -170 5 -154

7.5 -116 10 -94

12.5 -84 15 -60

17.5 -47 20 -41 25 -1 30 11 35 84 40 130 45 188 50 237 55 298 60 339 65 371 70 400 75 426 80 450 85 476 90 500

92.5 518 95 540

96.5 568 98 614 99 660

100 691

Temperature, C

Figure AIII.6 "Plant" Scrubber Overhead TBP curve

100 200 300 400 Molecular Weight

500 600

Table AIII.22 "Plant" Scrubber Overhead composition and fraction distribution

Boiling range, °C Mole fractions Mass fractions <=100 0.681 0.239

100-200 0.087 0.100 200-300 0.052 0.088 300-400 0.082 0.204 400-500 0.069 0.235 500-600 0.025 0.106 600-700 0.004 0.020 700-800 0.001 0.008 800-900 0.000 0.000 900-1000 0.000 0.000

1000> 0.000 0.000

Vol. fraction C4-(<177°C) 0.070 LGO (177-343°C) 0.250 HGO(343-524°C) 0.579 524+(>524°C) 0.101

Figure AIII.7 "Plant" Scrubber Overhead molecular weight distribution curve

0 200 400 600 800 1000 1200 Liquid Density, kg/m

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 V Q ^ S O 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 g = 0.11423 lb/ft3 = 1.7145 kg/m3 is density of the Cyclone Product vapour phase*

u.g = 3.11 • 10"2 cP = 3.11 • 105 kg/ms is viscosity of the Cyclone Product vapour phase*

pi = 759.32 kg/m3 is density of the Cyclone Product liquid phase - liquid droplet*

ai = 6.357-10"4 lbf/ft is surface tension of the Cyclone Product liquid phase - liquid droplet*

D0=0.58 m is nozzle diameter

dp=1.081-10"5 m is droplet diameter; the whole calculation is done using this diameter, calculated

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:

Re = D°'V°'Pg = 24-105 (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, =vn-K— for 7< — <100 or 4.06<X<58m A, A>

(IV-2)

AT = 6.2 for v 0 = 10 to 50 m/s

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:

" , W o < , < , « = v „ - 2 . 1 4 ^ 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 for distance 0 < X < 4.06m (IV-5)

X(t) = 16.5 +547.3(^-0.0665) for distance 4.06 <X< 58m (IV-6)

Vertical distribution of velocity (vv) can be calculated from the vertical force balance (weight

of the droplet against the drag force) and the terminal velocity of the droplet:

mg-FD=m^ (IV-7) dt

In this equation g is gravitational acceleration, m is droplet mass, and F D is a drag force:

FD=CD{\pgvv

2){^dp) (IV-8)

Velocity in this equation is relative velocity between droplet (particle) and surrounding gas,

vv=vp - vg, dp is droplet (particle) diameter, and Co is drag coefficient. For the spherical particles

it can be calculated from:

24 24 u„ c - = i r = — ( I V - 9 )

R e pg-v

v-dp

dv Terminal velocity can be calculated from Equations (IV-7), (IV-8) and (IV-9) when—- = 0:

dt dv

mg-2>n/ugdpvv = m - ^ = 0 (IV-10)

Since m- p, -Vp - p, -(^mi3

p) is the mass of the spherical liquid droplet, Equation (IV-10) for

droplet terminal velocity becomes:

mg Prd2

p-g 57tpgdp 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-3xjUgdpvv =m dy^ dt

dvv

dt g — ^ - = g - ^ = g

m To

(IV-12)

(IV-13) V yT0 J

Integrating velocity in time from t=0 to t:

v v(0 = v-ro l-exp(—^-0 (IV-14)

dY Since vv = -f-, using Equation (IV-14) and integrating vertical distance Y in time from t=0 to t,

dt

the following Equation for vertical distance change in time can be obtained:

(IV-15) Y0't) = vT0-t + ^ exp(—^--0-1 g VT0 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 cross-

section, 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:

VT = VT0 ~ V g (IV-16)

where vT is the terminal velocity in the flowing surrounding gas, vg = 9.8 m/s is velocity of the

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 v(0 = 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 = V V(0-V^

and hence:

(IV-18)

vreal (t) = vT 8 l-exp(-—t)

dY

— v„ (IV-19)

Again, vvrea, - — , using Equation 19 and integrating vertical distance Y in time from t=0 to t, dt

the following Equation for vertical distance change in time can be obtained:

(IV-20) Y(t) = Y0(t)-v-t

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"5m), assuming that all others would be carried even

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

delayed coker in hysys 1

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