Characteristics of Rapid Pyrolysis for Upgrading Heavy Oils in a

Article

Characteristics of rapid pyrolysis for upgrading heavy oils in circulating fluidized bed reactor Kang Seok Go, Myung Won Seo, Young-tae Guahk, Heyn Sung Chang, Nam-Sun Nho, Kwang-ho Kim, Yong Ku Kim, and Jae Goo Lee Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.7b00488 • Publication Date (Web): 12 May 2017 Downloaded from http://pubs.acs.org on May 26, 2017

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Energy & Fuels

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Characteristics of rapid pyrolysis for upgrading heavy oils in circulating fluidized bed

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reactor

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Kang Seok Go1, Myung Won Seo1, Young Tae Guahk2, Heyn Sung Chang1, Nam Sun Nho1*,

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Kwang Ho Kim1, Yong Ku Kim1, Jae Goo Lee1

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1

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Gajeongro, Yuseong-gu, Daejeon 305-343, Republic of Korea

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2

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Climate Change Research Division, Korea Institute of Energy Research (KIER), 152

Energy Efficiency and Materials Research Division, Korea Institute of Energy Research

(KIER), 152 Gajeongro, Yuseong-gu, Daejeon 305-343, Republic of Korea

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*Corresponding Author (Nam Sun Nho)

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Tel.: 82-42-860-3631, Fax: 82-42-860-3739, E-mail: [email protected]

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Keywords: heavy oil, rapid pyrolysis, hydrodynamic velocity, circulating fluidized bed

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1

ACS Paragon Plus Environment

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Abstract

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Hydrodynamic velocity and the effect of reaction temperature on the cracking

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performance were investigated for the production of light oil through the rapid

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pyrolysis of heavy oil in a circulating fluidized bed reactor. From this study, the gas

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velocity in the loop seal to prevent slugging, transport gas velocity to form a dense

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phase of solids on the bottom of the riser, and the minimum bubbling fluidization

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velocity for stable solid circulation in the reheater were determined. Under the steady

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supply of feed through the internal mixing with steam inside the feed nozzle, the

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reactivity for a temperature change from 527 to 574 °C was investigated. As results,

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residue (>535 °C) conversion and liquid yield of up to 71.1% and 78.2%, respectively,

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were found. The maximum impurity removal rates for sulfur, Conradson Carbon

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Residue, and metals (nickel and vanadium) were found to be 18.7%, 50.4%, and

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83.6%, respectively.

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1. Introduction

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The growing global energy consumption has prompted increased interest and

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exploitation of abundant heavy oil reserves in recent times. However, the application

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of well production, pipeline transportation, and conventional refining to such oils is

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very inefficient owing to the unfavorable physical properties such as high viscosity,

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specific gravity, and sulfur and metal contents1-2. Particularly, the high amounts of

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Conradson Carbon Residue (CCR), asphaltene, and other impurities in heavy oils

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rapidly deactivate the catalysts used for their hydrotreating, hydrocracking, and fluid 2


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catalytic cracking through the formation of cokes.

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To address these issues, slurry-phase hydrocracking using dispersed catalysts or

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dispersants has been introduced. Examples of such technologies are ESTTM, VCCTM,

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and UniflexTM developed by Eni, KBR/BP, and UOP, respectively3–5. It is reported

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that slurry-phase hydrocracking technology has the potential for high upgrading

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performance, namely, high residue conversion and liquid yield and over 90% metal

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content reduction6. Nevertheless, it requires huge investment in complex equipment

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and is characterized by a high operation cost, particularly owing to the high hydrogen

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consumption and high catalyst replacement cost3.

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Compared to hydrogen addition processes, the use of thermal cracking for carbon

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rejection is easier in oilfields like those in Canada and Venezuela that contain extra-

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heavy oils and oil sands. This is because of the low capital and operation cost

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afforded by the non-use of hydrogen7. This has led to the wide adoption of related

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technologies such as delayed coking and visbreaking8. However, the low liquid

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product yield due to the accompanying formation of large amounts of gas and coke

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remains a drawback of these technologies7.

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Rapid pyrolysis of heavy oils has been investigated for use in improving

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conventional coking. Representative examples of this method are HTLTM (heavy to

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light), IYQTM, and FlexicokingTM developed by Ivanhoe, ETX, and Exxon Mobil,

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respectively9. These technologies with a short contact time cracking have strong

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benefits for the economics because it produces a high conversion of feedstock and a

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high yield of liquid product by inhibiting a secondary cracking of product vapor. In 3


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additions, it can realize a high throughput processes. These technologies use a

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thermal medium such as a carbonaceous material or silica sand particles to thermally

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crack heavy feedstock, after which coke deposited by a cracking reaction is

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combusted to generate the heat required for the pyrolysis reaction3,10–12. The thermal

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medium is continuously circulated between the reactor and the reheater for heat

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transfer. In this respects, the dispersion of feed as a form of small droplets with an

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injection of steam and rapid mixing of feed with solids as a heating medium on the

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bottom of the riser is crucial to achieve a high conversion and liquid yield by

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controlling the contact efficiency. Otherwise, it would be rather occurred a large

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amount of coke in the area of near feed nozzle so that it causes a severe plugging

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issue. A stable circulation of solids between riser (reactor) and bubbling bed (reheater)

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through loop seal is also very important design factors to maintain a steady reaction

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performance through the constant heat transfer from reheater to riser.

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Therefore, the determination of hydrodynamic velocities is a prerequisite to obtain

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an optimum pyrolysis performance. There have, however, been few published works

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on this subject. Hence, in the present study, the hydrodynamics of a circulating

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fluidized bed reactor was investigated, as well as the proper method for steady feed

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supply for the rapid pyrolysis of heavy oil. The observations were used to determine

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the effects of the reaction temperature on the product yield and properties. For this

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purpose, a mixture of vacuum residue and B-C oil was used as model bitumen feed.

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2. Methods

4


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2.1.Experimental apparatus

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Fig. 1 is a schematic diagram of the considered rapid pyrolysis system. The pyrolysis

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process was designed for a throughput of 1-barrel per day of heavy oil. The system is

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based on the concept of a circulating fluidized bed and is composed of four main

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sections, namely, the feeding (1–2), reaction and reheating (3–6), gas separation, and

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liquid products (7– 10) sections. The feed is supplied to the reactor through the feed

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pump, with the consumption rate monitored by the change in weigh of the feed vessel.

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Steam is added at the feeding nozzle to improve the feed injection and atomization.

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The riser reactor comprises a pipe of inner diameter 0.075 m and height 7.0 m. A

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cyclone at the top of the riser is used to separate the effluent and circulating solid.

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Two loop seals with an inner diameter of 0.075 m and height of 0.405 m are used to

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control the solid circulation rate and prevent gas leakage between the riser and the

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reheater. The products from the cyclone are transported through the scrubber and two

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condensers, which respectively extract the gas and liquid products.

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With regard to achieving a steady feed supply, it is generally known that steam can

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be used for feed atomization in a nozzle to facilitate dispersion, vaporization, and high

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solid contact efficiency in a reactor21. In the present study, the effects of internal and

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external mixing of the feed and steam (see Fig. 2) on the stability of the feed supply

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were respectively investigated. During the reaction between the steam and oil in the

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nozzle using a weight ratio of approximately 1.7, the variations of the feed vessel

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weight and nozzle pressure were monitored to evaluate the stability of the feed

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injection for two different types of mixing nozzles.

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2.2. Hydrodynamic test procedure

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The hydrodynamic test was conducted in a Plexiglas cold model reactor with the

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same configuration as the second section (3–6) of the system in Fig. 1. The physical

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properties of the employed bed material are given in Table 1. A solid inventory of 40

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kg was initially loaded to the circulating fluidized bed for an experimental run. At a

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given superficial gas velocity in the riser, air was injected into the bubbling fluidized

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bed and loop seal at the desired gas velocities. When a steady state was achieved, the

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bed height in the bubbling fluidized bed became constant13. The pressure drop in the

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bed was measured by transducers (DPLH series, Sensys, Korea) and recorded in a

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data acquisition system via an HMI system. Twelve pressure taps were mounted flush

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with the wall of the riser, and five transducers were attached to the walls of the

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bubbling fluidized bed. The axial solid holdup (εs) was determined from the measured

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differential pressure (∆P/∆L) profile as follows:

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∆P/∆L = ρsεsg

(1)

where ρs is the solid particle density and g is the gravitational acceleration.

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To determine the solid circulation rate, the hydraulic ball valve of the downcomer

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was closed and the ascending time (t) of the particles to a given distance in the

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transparent downcomer was measured. Based on the bulk density and the measured

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time, the solid circulation rate (Gs) was calculated. This procedure is referred to as the 6


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ascending time method14. The variables and operational ranges of the cold model test

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are presented in Table 2.

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2.3. Reaction test procedure

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Feed was filled into the feed vessel and mixed at 80 °C. Cooling water at room

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temperature was supplied to the condensers to collect the liquid products. Solid

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particles with the same properties as presented in Table 1 were fluidized by air in the

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reheater and heated to 650 °C. When the temperature in the reheater had reached the

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desired level, steam was supplied to the feed nozzle and distributor. The loop seal was

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then operated by using nitrogen to circulate the solid particles. The pressure profile

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was monitored to confirm the stability of the solid circulation under controlled gas

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flow into the recycle and supply chambers. When the pressure profile became stable,

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feed was injected into the riser and the feed rate and temperature and pressure in the

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riser and reheater were observed. The achievement of steady reaction could be

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confirmed by the trends of the O2 and CO2 concentrations of the flue gas.

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Under steady state conditions, the test was conducted using various reaction

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temperatures between 529 and 574 °C at a fixed vapor residence time of about 1.6

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second and a sand-to-oil ratio of about 53-61 without makeup silica sands.

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2.4.Physical property analyses of feedstock and products As indicated in Table 3, the feedstock used in this study was prepared by mixing 7


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vacuum residue and B-C oil. The mixture was used to simulate oil sand bitumen

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based on the API gravity and kinematic viscosity. The liquid products could be

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separated from the water by distillation (150 °C, 650 mmH2O). The physical

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properties of the liquid products were determined using the American Standard Test

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Methods (ASTM), namely, D5002-99 for API gravity, D445-15 for kinematic

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viscosity, D4294-03 for sulfur and metals, D4530-06 for CCR, IP469 for SARA

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(saturate, aromatic, resin, and asphaltene), and D7500 for simulated distillation

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distribution. After the completion of the rapid pyrolysis reaction in the riser, the solid

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sand particles, which included coke, were transferred into the reheater and

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combusted in air. Wherein, the full combustion of coke was confirmed by

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maintaining O2 content more than 1 vol. % of flue gas. The SO2, CO, and CO2

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contents of the flue gas were determined by the infrared absorption method, while

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the galvanic method was used to determine the O2 content. The coke yield was

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calculated from the CO, CO2, SO2, and H2O contents of the flue gases. That is,

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during the material balance the total mass of each component was measured and

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then the amount of carbon was obtained from the conversion of atomic ratio. The gas

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yield was determined from the material balance based on the liquid and coke yields.

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167 168 169

3. Results and discussion 3.1. Hydrodynamics in rapid pyrolysis reactor 3.1.1. Hydrodynamic velocities for different flow regimes Hydrodynamic test was conducted in the cold model reactor using solid sand 8


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particles and air at room temperature and pressure to determine the hydrodynamic

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velocity for stable reaction. As can be seen from Fig. 3(a), the minimum fluidizing

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gas velocity was about 0.050 m/s, which is comparable to the 0.058 m/s

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theoretically determined by the Grace’s correlation17. Fig. 3(b) shows the effect of

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the gas velocity on the pressure fluctuation. As also observed by Yerushalmi and

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Cankurt18, who used silica alumina catalysts, the maximum fluctuation peak with

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increasing gas velocity Uc was approximately 1 m/s. The fluctuation eventually

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leveled off at 1.8 m/s with the onset of turbulent fluidization with gas velocity Uk.

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As shown in Fig. 3(c), the transport velocity Utr between the turbulent and fast

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fluidization regimes for pneumatic solid transportation was 2 m/s, as determined by

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the empty time method15,16. This result is similar to the 2.21 m/s predicted by the

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Goo’s correlation15.

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3.1.2. Solid circulation rate

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The solid circulation rate Gs was observed based on the variations of the gas

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velocity to the riser Ug,r, the supply velocity Ug,sc, and the velocity in the loop seal

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recycle chamber Ug,rc. As can be seen from Fig. 4(a), the solid circulation rate

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increased with increasing gas velocity to the supply chamber for a given gas

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velocity in the riser and recycle chamber. It was, however, observed that a slug was

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formed when the gas velocity to the supply chamber exceeded 0.1 m/s (2Umf), and

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this prevented further increase of the solid circulation rate. In other words, the

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circulation rate increased with increasing gas velocity to the recycle chamber, but

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became constant after reaching twice the minimum fluidization velocity (2Umf) 9


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under the same condition as that in the supply chamber. The maximum circulation

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rate was determined to be about 70 kg/m2s for Ug,r = 3.0 m/s. Fig. 4(b) shows the

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effect of the gas velocity to the recycle chamber on the solid circulation rate for a

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fixed gas velocity in the supply chamber of 0.10 m/s, with respect to the riser

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velocity. The solid circulation rate increased with increasing gas velocity to the

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recycle chamber to about 0.10 m/s (2Umf), beyond which point it tended to stabilize.

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However, with increasing gas velocity in the riser, the solid circulation rate can

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reach to about 80 kg/m2s. In particular, the change was only shown at values of Gs

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higher than 40 kg/m2s. This can be explained by the fact that a dense phase may be

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formed at the bottom of the riser at low gas velocities, thereby hampering solid

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circulation from the loop seal to the riser. Fig. 4(c) shows the effect of the gas

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velocity to the bubbling fluidized bed (reheater) on the solid circulation rate. As

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can be observed from the figure, Gs increases with increasing Ug,b up to

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approximately 0.1 m/s (2Umf), beyond which it saturates at 60 kg/m2s. This

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suggests that a certain minimum gas velocity in the bubbling fluidized bed is

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required to transport the desired amount of solids to the loop seal. This agrees with

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the observation of Seo et al.19.

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3.1.3. Solid holdup profile in riser

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The feed is injected with steam through the nozzle of the reactor, and the heat

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required for the cracking reaction of the feed is supplied through the solid sand

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particles, which are heated by the combustion of coke. The formation of a dense

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phase at the bottom of the riser is important for the enhancement of the reaction

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rate by improving the heat transfer. In a typical fast fluidization regime, the dense

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solid phase formed at the bottom of the riser has a volumetric fraction of

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approximately 0.1–0.220. The effects of the riser velocity and solid circulation rate

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on the formation of the dense phase and the solid holdup profile along the bed

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height were examined. The results are shown in Fig. 5. As can be observed from

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Fig. 5(a), for a fixed solid circulation rate of 43.7 kg/m2s, the dense phase only

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occurred for Ug,r = 3.0 m/s, while the solid holdup was limited to 0.05 for gas

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velocity in the riser in excess of 3.5 m/s. Based on these observations, the variation

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of the solid holdup with the solid circulation rate was investigated for Ug,r = 3.0

225

m/s. The results are shown in Fig. 5(b). The dense phase appeared when the solid

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circulation rate was increased to 34.5 kg/m2s, and was maintained at 1 m from the

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bottom of the riser until the circulation rate exceeded 48.9 kg/m2s.

228 229

3.1.4. Determination of operation range from hydrodynamic results

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Based on the above hydrodynamic results, the stable operation conditions were determined as follows:

232 233

-

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Gas velocity in the supply chamber of the loop seal for stable solid circulation without slugging: < 2Umf

235 236

-

Bubbling bed gas velocity for stable solid circulation: > 2Umf

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Gas velocity in the riser for stable solid circulation and formation of a dense

237 238

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phase: 1–1.5Utr

240 241

3.2.Effect of feed nozzle configuration on the feed supply

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Fig. 6 shows the effect of different types of mixing of the steam and feed on the

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operation stability. As can be seen from the figure, internal mixing in Fig. 6(a)

244

enabled the maintenance of stable feed supply at nearly 1.5 bar of nozzle pressure

245

over the entire duration of the experiment with constant temperature at the vicinity

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of nozzle in the bed, while external mixing caused an abrupt increase of the nozzle

247

pressure, resulting in observed temperature reduction. That is because the atomized

248

feed vapor stopped to contact to the temperature sensor near feed nozzle. When the

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feed is resupplied into the bed, the temperature is recovered to the formal state.

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However, the pressure in the nozzle continuously shows unstable pattern, which

251

shows the feed injection with external nozzle is not desirable. During some trials of

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external mixing, emergency shutdown was required owing to radical pressure

253

increase, and a large amount of coke was found around the nozzle when the reactor

254

was disassembled. These observations can be attributed to external mixing of the

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feed and steam producing inferior dispersion and atomization of the feed. The

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atomization efficiency of an external mixing nozzle particularly depends on the jet

257

angle of the steam and direct contact between the thermal medium and the

258

undispersed feed oil, with the latter factor increasing the possibility of coke

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formation. Conversely, internal mixing facilitates atomization of the feed oil into

260

small droplets through the turbulence generated by the steam inside the nozzle

261

before it contacts the heated thermal medium at the nozzle tip. 12


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3.3. Steady state operation

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Stable solid circulation was first confirmed by checking for constant temperature

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at the riser outlet and constant pressure at the bottom of the riser and reheater, as

266

shown in Figs. 7(a) and (b), respectively. Fig. 7 (c) shows the solid holdup along

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the height above the distributor. As can be seen from the figure, the dense phase at

268

the bottom of the riser was maintained during the reaction as that of the cold model

269

test. After the introduction of feed into the reactor, time was required for the

270

establishment of a steady state. This was because of the need to generate coke from

271

thermal cracking of the feedstock and transfer combustion heat to the reaction zone.

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The O2 and CO2 concentrations of the flue gas were used to monitor the reaction.

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The achievement of stable reaction can be confirmed from the stability of the

274

concentration profiles in Fig. 7(d).

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276 277

3.4. Effects of reaction temperature on conversion of heavy oil and product yields and properties

278 279

The steady-state residue conversion, product yields, and product properties were

280

investigated for different reaction temperatures. The conversion was calculated

281

using the following equation:

282 283

Percentage conversion = [1 – (fractional residue content of liquid product * liquid 13


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yield) / fractional residue content of feed] * 100, residue cut: >535 °C

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(2)

285 286

As shown in Fig. 8(a), the residue conversion increased with increasing riser

287

outlet temperature (ROT) up to 566 °C, attaining a maximum value of about 71%,

288

but it is shown a little bit reduction at 574 °C. The liquid product yield was 72–78

289

wt.% within 529–574 °C. The amount of coke present was not observed to change

290

abruptly, while the increase in gas yield was greater.

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The variation of the products of the fractional distillation by rapid pyrolysis is

292

shown in Fig. 9. The residue (>535 °C cut), which initially comprised about 70

293

wt.% of the feed, was reduced to 27.4 wt.%, while the distillate (IBP to 350 °C)

294

and vacuum gas oil (VGO , 350–535 °C) increased to about 43.8 and 28.8 wt.%,

295

respectively. Considering the conversion of residue fraction with respect to

296

increasing temperature, the unexpected data point out of trend at 574℃ is thought

297

to be an experimental error due to the small change of reaction temperature. Fig.

298

10 shows the results of the SARA analysis of the liquid product. As can be seen,

299

resin and asphaltene content, which was about 36% of the initial feedstock, was

300

reduced to about 17% after rapid pyrolysis, resulting in increased concentration of

301

saturates and aromatics in the liquid product. This indicates that the residue could

302

be converted into light oils by rapid pyrolysis.

303

With regard to pipeline transportation of the upgraded oil, the API gravity and

304

viscosity are the important parameters of the liquid product. From Fig. 11 (a), it

305

can be seen that the API gravity increased to as much as 12.8, and slightly

306

decreased with increasing ROT. This might be attributed to the increase of the 14


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Energy & Fuels

asphaltene content in the liquid product, as shown in Fig. 10.

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Fig. 11(b) indicates that the dynamic viscosity was within 166–387 cP.

309

Considering the typical requirement of an API gravity of 19 and viscosity of 350

310

cSt for upstream bitumen transportation through pipeline, there is a need for

311

further improvement of API gravity1.

312

Impurity contents of the liquid product and their removal rates are shown in Fig.

313

12. As can be observed from Fig. 12(a), the sulfur in the liquid product was

314

greater than that in the feed and slightly increased with increasing ROT. This can

315

be explained by the generally low desulfurization of thermal cracking without the

316

use of a hydrogen atmosphere and the even distribution of the feed sulfur to all

317

the distillation fractions. The sulfur content of the liquid product was thus almost

318

retained or slightly increased by the formation of asphaltene or coke. Conversely,

319

the vanadium and nickel contents of the liquid product were reduced to as low as

320

27 and 28 wt.ppm, respectively. The CCR was also decreased to 10.8 wt.%. With

321

regard to the impurity removal rates, Fig. 12(d) indicates values of 18.7, 83.6, and

322

50.4 wt.% for sulfur, metals, and CCR, respectively. In contrast to the sulfur,

323

metals, and CCR being generally concentrated in the asphaltene fraction of the oil,

324

coke formation through thermal cracking could have reduced the metal and CCR

325

contents of the liquid product22.

326

The properties of initial feed and product in this study are compared with other

327

carbon rejection process10,24 in Table 4. The liquid yield in this study lies in 72.2 –

328

78.2%, while the liquid yield of HTL process is around 71.5 – 83%. The data in

329

this study was obtained from single stage processing, which is same as HTL 15


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

330

process. The API of the product oil is around 11.3 – 12.8, which is similar with

331

the HTL process (11.8-16.2). The impurity (metal, sulfur) removal conversion is

332

around 10-20 % higher than that of HTL process. Therefore, it can be concluded

333

that this process is comparable to other carbon rejection process (HTL) process.

334 335

4. Conclusions

336

The hydrodynamic velocity and reaction performance with respect to the reaction

337

temperature in a circulating fluidized bed reactor used for the production of light oils

338

by rapid pyrolysis were investigated. It was found that, to prevent slugging, the gas

339

velocity in the supply chamber of the loop seal should be maintained at lower than

340

twice the minimum fluidizing velocity (Umf), and that a gas velocity in the riser of 1–

341

1.5 times the transport velocity was required to produce a dense phase at the bottom

342

of the riser. In addition, the minimum bubbling fluidization velocity in the reheater

343

should be twice the value of Umf to achieve stable solid circulation.

344

The feed nozzle design that ensures steady supply and atomization of feed was

345

determined by comparing the results of internal and external mixing of the feed and

346

steam. The application of the findings to a once-through test conducted using a 1-

347

BPD pilot plant produced a residue (>535 °C) conversion and liquid yield of as

348

much as 71% and 78.2%, respectively. The maximum impurity removal rates were

349

also found to be 18.7%, 50.4%, and 83.6% for sulfur, CCR, and metals (nickel and

350

vanadium), respectively. Further, the upstream application of rapid pyrolysis

351

produced a liquid product viscosity that met the pipeline transportation requirement

352

of 350 cSt, although the API requires further improvement through the product 16


Page 16 of 40

Page 17 of 40 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

353

Energy & Fuels

recycling into the reactor.

354

355

17


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

356

Acknowledgement

357

This study was conducted within the framework of the Research and Development Program

358

of the Korea Institute of Energy Research (KIER) (B4-2433-02).

359

.

360

References

361

[1] Banerjee, D.K., Oil sands, Heavy oil & Bitumen, 2012. 11–33

362

[2] HUC, A.Y., Heavy crude oils: from geology to upgrading: an overview. Paris: Technip;

363

2011, 215-223

364

[3] Castañeda, L.C.; Muñoz, J. a. D.; Ancheyta, J., Current situation of emerging technologies

365

for upgrading of heavy oils, Catal. Today, 2014, 220–222: p. 248–273.

366

[4] Delbianco, A.; Eni slurry technology: a new process for heavy oil upgrading. 19th World

367

Petroleum Congress, Spain, 2008

368

[5] Daniel, G.; Mark, V.W.; Paul, Z., H. Ed, Upgrading Residues to Maximize Distillate

369

Yields with UOP Uniflex TM Process, 2010, 53: p. 33–41.

370

[6] Bellussi, G.; Rispoli, G.; Landoni, A.; Millini, R.; Molinari, D.; Montanari, E.; Moscotti,

371

D.; Pollesel, P., Hydroconversion of heavy residues in slurry reactors: Developments and

372

perspectives, J. Catal., 2013, 308: p. 189–200.

373

[7] Rana, M.S.; Sámano, V.; Ancheyta, J.; Diaz, J. a. I, A review of recent advances on

374

process technologies for upgrading of heavy oils and residua, Fuel, 2007, 86: p. 1216–1231.

375

[8] Sawarkar, A.N.; Pandit, A.B.; Samant, S.D.; Joshi, J.B., Petroleum residue upgrading via 18


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Energy & Fuels

376

delayed coking: a review. Can J Chem Eng., 2007, 85: p. 1–24.

377

[9] Belt, O.; East, M., Emerging technologies for the conversion of residues, Encyclopaedia

378

of hydrocarbons. Vol. III, 2005, 147

379

[10] Freel, B.; Graham, R.G., Rapid thermal processing of heavy hydrocarbon feedstocks.

380

US2012/0279825.

381

[11] ETX-Systems. IYQ upgrading – an industry altering technology for primary upgrading

382

of heavy oil. In: World Heavy Oil Congress, November 3–5, Puerto la Cruz, Venezuela, 2009.

383

[12] Furimsky, E., Characterization of cokes from fluid/flexi-coking of heavy feeds, Fuel

384

Process. Technol., 2000, 67: p. 205–230.

385

[13] Seo, M.W.; Suh, Y.H.; Kim, S.D.; Park, S.W.; Lee, D.H.; Song, B.H., Cluster and bed-to-

386

wall heat transfer characters in a dual circulating fluidized bed. Ind Eng Chem Res., 2012,

387

51(4): p. 2048–61.

388

[14] Kim, S.W.; Namkung, W.; Kim, S.D., Solid recycle characteristics of loop-seals in a

389

circulating fluidized bed. Chem Eng Technol., 2001, 24: p. 843–849.

390

[15] Goo, J.H.; Seo, M.W.; Kim, S.D.; Song, B.H., Effect of temperature and particle size on

391

minimum fluidization and transport velocities in a dual fluidized bed. Proceedings of the 20th

392

International Conference on Fluidized Bed Combustion, Xian, 2009: p. 305–309.

393

[16] Adànez, J.; de Diego, L.F.; Gayan, P., Transport velocities of coal and sand particles.

394

Powder Technol., 1993, 77: p. 61–68.

395

[17] Grace, J.R.; In: Hetsroni G, editors. Handbook of multiphase systems, Washington D.C.: 19


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396

Hemisphere, 1982, 8-1.

397

[18] Yerushalmi, J.; Cankurt, N.T., Chemtech, 8 (1978) 564. Powder Technol., 1979, 24: p.

398

187.

399

[19] Seo, M.W.; Nguyen, T.C.B.; Lim, Y.I.; Kim, S.D.; Park, S.W.; Song, B.H.; Kim, Y.J.,

400

Solid circulation and pressure drop characteristics of a dual circulating fluidized bed reactor:

401

experiments and CFD simulation. Chem Eng J., 2011, 168: p. 803−811.

402

[20] Kunii, D.; Levenspiel, O., Fluidization engineering. 2nd ed. Butterworth-Heinemann

403

Series in Chemical Engineering, 1991: p. 193–196.

404

[21] Chen, Y.M., Recent advances in FCC technology. Powder Technology, 2006, 163: p. 2–8.

405

[22] Yen, T.F.; Chilingarian, G.V., Asphaltenes and asphalts. In: Developments in petroleum

406

science. Amsterdam: Elsevier, 1994, 40: p. 281-304.

407

[23] Haider, A.; Levenspiel, O., Drag coefficient and terminal velocity of spherical and

408

nonspherical particles. Power Technol., 1989, 58: p. 63.

409

[24] Castaneda, L.C.; Munoz, J.A.D.; Ancheyta, J., Current situation of emerging

410

technologies for upgrading of heavy oils, 2014, 220-222: p.248-273.

20


Page 20 of 40

Page 21 of 40 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Fig. 1 Schematic diagram of rapid pyrolysis system 1) feed tank, 2) feed pump, 3) riser, 4) cyclones, 5) loop seals, 6) reheater, 7) scrubber, 8) condensers, 9) product tank, 10) blower, 11) steam generator

8 Gas meter 4 7

10

9

4

3 5 6

5 11 5 1

Recycle chamber

2


Supply chamber

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 22 of 40

Fig. 2 Conceptual geometry of feed nozzles with a) internal mixing b) external mixing between steam and feed

(a)

(b)

steam

feed

steam

steam


feed

steam

Page 23 of 40

Fig. 3 Hydrodynamic velocity a) of minimum fluidization Umf, b) at onset of turbulent fluidization Uk, c) at onset of transportation (b)

(a)

4000

1400

120

Upward Downward

800

Umf = 0.05 m/s

600

400

Pressure standard deviation [Pa]

1000

3000

90

2000

60

1000

30 Uk = 1.8 m/s

200

0

0 0

0.02 0.04 0.06 0.08 0.1 0.12 Gas velocity to the bubbling fluidized bed, Ug,b [m/s]


0 0

0.5 1 1.5 Gas velocity to the riser, Ug,r [m/s]

2

Suspension density, ρsus [kg/m3]

Uc = 1 m/s

1200

Pressure drop [Pa]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Energy & Fuels

Fig. 3 Hydrodynamic velocity a) of minimum fluidization Umf, b) at onset of turbulent fluidization Uk, c) at onset of transportation Utr

(c) 250

200

Emptying time [s]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 24 of 40

150

100

50 Utr = 2 m/s

0 0

1 2 3 Gas velocity to the riser, Ug,r [m/s]


4

Page 25 of 40

Fig. 4 Effect of gas velocity a) to the supply chamber with variation of recycle chamber velocity (Ug,rc), b) to the recycle chamber with variation of riser velocity (Ug,r) on solid circulation rate (Gs), c) Effect of gas velocity in the bubbling fluidized bed (Ug,b) on Gs

(a)

(b) 90

80 Ug,r 3.0 m/s

Ug,sc 0.10 m/s

80 Solid circulation rate, Gs [kg/m2s]

70 Solid circulation rate, Gs [kg/m2s]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

60 50 Slug formation in supply chamber

40 30

Ug,rc = 0.05 m/s (1Umf)

20

70 60 50 40 30 Ug,r = 3 m/s

20

Ug,rc = 0.1 m/s (2Umf) 10

Ug,r = 3.5 m/s 10

Ug,rc = 0.15 m/s (3Umf)

Ug,r = 4 m/s

0

0 0

0.05 0.1 Gas velocity to the supply chamber, Ug,sc [m/s]

0.15

0


0.05 0.1 0.15 Gas velocity to the recycle chamber, Ug,rc [m/s]

0.2

Energy & Fuels

Fig. 4 Effect of gas velocity a) to the supply chamber with variation of recycle chamber velocity (Ug,rc), b) to the recycle chamber with variation of riser velocity (Ug,r) on solid circulation rate (Gs), c) Effect of gas velocity in the bubbling fluidized bed (Ug,b) on Gs

(c) 80

Solid circulation rate, Gs [kg/m2s]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 26 of 40

60

40

20 Ug,rc = 0.15 m/s ( 3 Umf) Seo et al. (2011) 0 0

0.1 0.2 Gas velocity to the recycle chamber, Ug,rc [m/s]


0.3

Page 27 of 40

Fig. 5 Effect of a) riser gas velocity at Gs = 43.7 kg/m2s and b) solid circulation rate at Ug,r = 3.0 m/s on solid holdup εs in riser

(a)

(b) 7

7

Gs = 19.2 kg/m2s 6

Ug,r = 3.0 m/s

6

Height above gas distributor [m]

Ug,r = 3.5 m/s Height above gas distributor [m]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Ug,r = 4.0 m/s

5

Ug,r = 4.5 m/s 4

3

2

Gs = 34.5 kg/m2s Gs = 48.9 kg/m2s

5

Gs = 60 kg/m2s

4

3

2

1

1

0

0 0

0.1

0.2 Solid holdup, εs [-]

0.3

0.4

0


0.05

0.1

0.15 0.2 0.25 Solid holdup, εs [-]

0.3

0.35

Energy & Fuels

Fig. 6 Effect of different types of mixing between steam and feed on the feed supply with a) internal mixing, b) external mixing

(a)

(b) 700

10

stable feed supply

10

unstable feed supply

9

9

600

600

8

7 6

400

5 300

4 3

200

500 Temperature [℃]

500

Feed nozzle pressure, [bar]

8

7 6

400

5 300

4 3

200

2 100

2 100

1 0 0

200

400 600 Time [s]

800

0 1000

1 0


0

200

400 600 Time [s]

800

0 1000

Feed nozzle pressure, [bar]

700

Temperature [℃]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 28 of 40

Page 29 of 40

Fig. 7 Steady state reaction condition with a) riser outlet temperature, b) riser and reheater bottom pressure, c) solid holdup profile in riser, d) O2 and CO2 concentration of flue gas

(a)

(b) 15000

600

riser

12000

Pressure [Pa]

580

ROT [℃]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

9000

reheater 6000

560 3000

0

540 0

5

10

15

20

25

30

35

Time [min]


0

5

10

15

20

Time [min]

25

30

35

Energy & Fuels

Fig. 7 Steady state reaction condition with a) riser outlet temperature, b) riser and reheater bottom pressure, c) solid holdup profile in riser, d) O2 and CO2 concentration of flue gas (d)

(c) 6

20

5 16 4

CO2 O2, CO2 [vol.%]

Height above distributor [m]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 30 of 40

3

2

12

8

O2 4

1

0 0.00

0.05

0.10

0.15

0.20

0.25

0.30

0 0

Solid holdup, εs [-]

5

10

15

20

Time [min]


25

30

35

Page 31 of 40

Fig. 8 (a) Residue (>535℃) conversion, (b) product yields of liquid □, gas(C1 to C4) ◇, coke △ with change of riser outlet temperature (ROT) (a)

(b)

100.0

100.0

80.0

80.0

60.0

40.0 71.1

67.1

Product yields [wt.%]

Residue conversion [wt.%]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

78.2

75.1

73.7

72.2

16.4

17.6

9.8

10.1

60.0

40.0

57.3

20.0

20.0

36.9

0.0 529.4

543.8

566.0

574

0.0 520.0

14.8

12.1

10.1

9.6

535.0

550.0 ROT [℃]

ROT [℃]


565.0

580.0

Energy & Fuels

Fig. 9 Fractional distillation of liquid product with a variation of riser outlet temperature (ROT)

100%

6.0

90% 80% Yield in Liquid [wt.%]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 32 of 40

12.3

20.3

28.8

24.0

25.1

31.3

70% 60%

39.9

50%

43.8

43.0

40% 30%

70.0 56.4

20%

39.8 27.4

10%

31.9

0% Feed

529.4

543.8

566.0

574.0

ROT [℃] Residue(535'C+)

VGO(350~535'C)

Distillate(IBP to 350'C)


Page 33 of 40

Fig. 10 Effect of reaction temperature on SARA distribution of liquid product

100.0

5.5

9.0

9.7

9.1

63.5

62.2

25.0

26.4

0.9

1.8

2.3

2.7

529.4

543.8

566.0

574.0

10.8

90.0 SARA in liquid product [wt. %]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

80.0 70.0 60.0

58.6

56.9

73.8

50.0 40.0 30.0

19.0

20.0 10.0

31.3

14.5

16.9

0.0 Feed

ROT [℃] Asp., wt%

Resin, wt.%

Aro., wt%


Sat. ,wt%

Energy & Fuels

Fig. 11 Effect of reaction temperature on (a) API and (b) viscosity of liquid product

(a)

(b)

14

100000

12 10000 Viscosity @35'C [cP]

10 API Gravity [-]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 34 of 40

8 12.7

6 4

12.8 11.4

11.3

8.9

1000

100

24460

387.5

301.3

529.4

543.8

10

2 0

166.3

172.9

566.0

574.0

1 Feed

529.4

543.8

566.0

574.0

ROT [℃]

Feed

ROT [℃]


Page 35 of 40

Fig. 12 Effect of reaction temperature on removal of impurities, (a) sulfur, (b) metals, (c) CCR in liquid product, (d) removal percentage of impurities(sulfur △, metal □, MCR ○) (a)

(b) 200

5

4

3

2 3.44

3.75

3.82

3.97

3.87

1

Metal in liquid product [wt. ppm]

V, wt.ppm

Sulfur, in liquid product [wt.%]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Ni, wt.ppm

160

120 137.5

80

40 27

0

0 Feed

529.4

543.8

566.0

574.0

Feed

31.5

32

7.5

10

10

543.8

566.0

574.0

28

29.5

6.5

529.4

ROT [℃]

ROT [℃]


Energy & Fuels

Fig. 12 Effect of reaction temperature on removal of impurities, (a) sulfur △, (b) metals □, (c) CCR ○ in liquid product, (d) percentage of impurities’ removal from feed

(c)

(d)

20

100.0

16

12

8

17.04

10.81

11.79

13.22

13.23

4

0 Feed

529.4

543.8

566.0

574.0

Pct. of Impurities' removal [wt.%]

83.6

CCR in liquid product [wt.%]

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 36 of 40

83.1

81.4

81.6

42.8

43.9

80.0

60.0

50.4

48.0

40.0

20.0

0.0 520.0

14.7

535.0

16.6

14.9

550.0 ROT [℃]

ROT [℃]


565.0

18.7

580.0

Page 37 of 40 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Table 1. Physical properties of silica sand particle

physical properties

values

Mean diameter, dp [㎛]

249

Particle density, ρs [kg/m3]

1423

Bulk density ρb [kg/m3]

2353

Minimum fluidization velocity, Umf (Calculated) [m/s], [16]

0.058

Terminal velocity, Ut (Calculated) [m/s], [23]

1.82

Transport velocity, Utr (Calculated) [m/s], [15]

2.21


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 38 of 40

Table 2. Experimental variables and operational range for cold mode test

Experimental variables for cold mode test Gas velocity to the riser (Ug,r, m/s)

Operational range 3.0–4.5

Gas velocity to the bubbling fluidized bed (Ug,b, m/s)

0–0.25 (5 Umf)

Gas velocity to loop-seal supply chamber (Ug,sc, m/s)

0–0.15 (3 Umf)

Gas velocity to loop-seal recycle chamber (Ug,rc, m/s)

0–0.2 (4 Umf)


Page 39 of 40 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Energy & Fuels

Table 3. Physical properties of feedstock Feedstock

Typical Athabasca bitumen[1]

API gravity

8.5

8.0

Density@15℃

1.0109

Kinematic viscosity (35℃), cSt

24,460

20,000

Carbon residue, wt.%

18.4

13.5

Sulfur, wt.%

3.6

5.0

Nickel, wt.ppm

28

80

Vanadium, wt.ppm

138

200

IBP-350 ℃

6

15-20

350-535 ℃

24

35-40

>535 ℃

70

45-50

5.5 / 58.6 / 19.0 / 16.9

15-20 / 30-35 / 25-30 / 15-20

Fractional distillation, wt.%

Saturate / Aromatic / Resin / Asphaltene, wt.%


Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46

Page 40 of 40

Table 4. Comparison with previous study Process (Company)

This study (KIER)

Feed

529.4℃

543.8℃

566.0℃

574.0℃

Liquid

78.2

75.1

73.7

72.2

Gas

12.1

14.8

16.4

17.6

Coke

9.6

10.1

9.8

10.2

36.9

57.3

71.1

67.1

Properties Product yield (%)

Residue conversion (wt.%, >535℃)

Fractional distillation of liquid product with a variation of ROT (wt.%)

560℃

592℃

620℃

79.9-83

74.5

71.5

Bitumen

528-538℃

545℃

590℃

82.7

77.4

74.6

6

12.2

20.3

28.8

25.1

9.8

20.4 (538℃)

9.8

VGO (350~535℃)

24

31.3

39.9

438

43

38.7

41.7 (538℃)

38.7

Residue (535℃+)

70

56.4

39.8

27.4

31.9

51.5

37.9 (538℃)

51.5

8.9

12.7

12.8

11.4

11.3

11

15.4-16.2

14.4

13.3

8.6

12.6-12.9

11.8

12.4

24460

387.5

301.3

166.3

172.9

6343 (@40℃)

10.01 (@40℃)

15.2 (@40℃)

4.6 (@40℃)

30380 (@40℃)

278 (@40℃)

151 (@40℃)

25.6 (@40℃)

164.5 3.44

34.5 3.75

37 3.82

41.5 3.97

42 3.87

127 3.6

48 3.5

55 3.9

49 3.5

295

CCR (wt.%) Metal (V+Ni) Sulfur

17.04 83.6 14.7

10.81 83.1 16.6

11.79 81.4 14.9

13.22 81.6 18.7

62.2 2.8

56.7

61.4 2.8

CCR

50.4

48

42.8

43.9

Viscosity @35℃ (cP)

Impurity removal conversion (%)

Heavy oil

Distillate (IBP to 350℃)

API (-)

Impurity in product

HTL Process (Ivanhoe Energy)

Metal (V+Ni, ppm) Sulfur (wt.%)


-

Characteristics of Rapid Pyrolysis for Upgrading Heavy Oils in a

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